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{{#Wiki_filter:Millstone Power Station Unit 2 Safety Analysis Report Chapter 3 MPS2 UFSAR 3-i Rev. 35CHAPTER 3-REACTOR Table of ContentsSection Title Page3.1
{{#Wiki_filter:Millstone Power Station Unit 2 Safety Analysis Report Chapter 3
 
Table of Contents tion      Title                                                                                                           Page


==SUMMARY==
==SUMMARY==
DESCRIPTION..............................................................................3.1-13.1.1References...................................................................................................3.1-33.2DESIGN BASES.................................................................................................3.2-13.2.1Mechanical Design Bases...........................................................................
DESCRIPTION.............................................................................. 3.1-1 1        References................................................................................................... 3.1-3 DESIGN BASES ................................................................................................. 3.2-1 1        Mechanical Design Bases ........................................................................... 3.2-1 1.1       Fuel Assembly Design Bases...................................................................... 3.2-1 1.2       AREVA Fuel Rod Cladding Design Bases................................................. 3.2-2 1.3      Control Element Assembly Design Bases .................................................. 3.2-2 1.4      Reactor Internals Design Bases .................................................................. 3.2-3 1.5      CEDM/RVLMS (HJTC) Pressure Housing Design Bases ......................... 3.2-5 2         Nuclear Design Bases ................................................................................. 3.2-6 3         Thermal and Hydraulic Design Basis ......................................................... 3.2-8 4        References................................................................................................... 3.2-8 MECHANICAL DESIGN ................................................................................... 3.3-1 1        Core Mechanical Design............................................................................. 3.3-1 1.1      AREVA Fuel Rod ....................................................................................... 3.3-1 1.1.1    Fuel Rod Mechanical Criteria..................................................................... 3.3-1 1.1.2    Fuel Rod Design Analyses.......................................................................... 3.3-3 1.2      (Deleted) ..................................................................................................... 3.3-6 1.3      AREVA Fuel Assembly.............................................................................. 3.3-6 1.3.1    Design Summary......................................................................................... 3.3-6 1.3.2    Fuel Assembly Mechanical Criteria ........................................................... 3.3-9 1.4      Fuel Assembly Holddown Device ............................................................ 3.3-12 1.5      Control Element Assembly....................................................................... 3.3-12 1.6      Neutron Source Design ............................................................................. 3.3-13 1.7      In-Core Instruments .................................................................................. 3.3-13 1.8      Heated Junction Thermocouples............................................................... 3.3-14 2        Reactor Internal Structures ....................................................................... 3.3-14 2.1      Core Support Assembly ............................................................................ 3.3-15 2.2      Core Support Barrel .................................................................................. 3.3-15 2.3      Core Support Plate and Support Columns ................................................ 3.3-16 2.4      Core Shroud .............................................................................................. 3.3-16 2.5      Flow Skirt ................................................................................................. 3.3-16 2.6      Upper Guide Structure Assembly ............................................................. 3.3-16 3        Control Element Drive Mechanism .......................................................... 3.3-17 3.1      Design ....................................................................................................... 3.3-17 3-i                                                              Rev. 35
3.2-13.2.1.1Fuel Assembly Design Bases......................................................................3.2-13.2.1.2AREVA Fuel Rod Cladding Design Bases.................................................3.2-2 3.2.1.3Control Element Assembly Design Bases..................................................3.2-23.2.1.4Reactor Internals Design Bases..................................................................
 
3.2-33.2.1.5CEDM/RVLMS (HJTC) Pressure Housing Design Bases.........................
tion    Title                                                                                                          Page 3.2      Control Element Drive Mechanism Pressure Housing ............................. 3.3-18 3.2.1    Heated Junction Thermocouple Pressure Boundary ................................. 3.3-19 3.3      Magnetic Jack Assembly .......................................................................... 3.3-19 3.4      Position Indication .................................................................................... 3.3-19 3.5      Control Element Assembly Disconnect .................................................... 3.3-20 3.6      Test Program............................................................................................. 3.3-20 4        References................................................................................................. 3.3-20 NUCLEAR DESIGN AND EVALUATION ...................................................... 3.4-1 1        General Summary ....................................................................................... 3.4-1 2        Core Description ......................................................................................... 3.4-1 3        Nuclear Core Design................................................................................... 3.4-1 3.1      Analytical Methodology ............................................................................. 3.4-2 3.2      Physics Characteristics ............................................................................... 3.4-2 3.2.1    Power Distribution Considerations ............................................................. 3.4-2 3.2.2    Control Rod Reactivity Requirements ........................................................ 3.4-2 3.2.3    Moderator Temperature Coefficient Considerations .................................. 3.4-3 4        Post-Reload Startup Testing ....................................................................... 3.4-3 5        Reactor Stability ......................................................................................... 3.4-4 5.1      General........................................................................................................ 3.4-4 5.2      Detection of Oscillations ............................................................................ 3.4-4 5.3      Control of Oscillations................................................................................ 3.4-5 5.4      Operating Experience ................................................................................. 3.4-6 5.5      Method of Analysis..................................................................................... 3.4-6 5.5.1    Radial Xenon Oscillations .......................................................................... 3.4-7 5.5.2    Azimuthal Xenon Oscillations.................................................................... 3.4-7 5.5.3    Axial Xenon Oscillations............................................................................ 3.4-7 6        References................................................................................................... 3.4-8 THERMAL-HYDRAULIC DESIGN.................................................................. 3.5-1 1        Design Bases............................................................................................... 3.5-1 1.1      Thermal Design........................................................................................... 3.5-1 1.2      Hydraulic Stability ...................................................................................... 3.5-1 1.3      Coolant Flow Rate, Distribution and Void Fraction................................... 3.5-1 2        Thermal and Hydraulic Characteristics of the Design................................ 3.5-2 2.1      Fuel Temperatures ...................................................................................... 3.5-2 2.1.1    Fuel Cladding Temperatures....................................................................... 3.5-2 2.1.2    Fuel Pellet Temperatures ............................................................................ 3.5-2 2.1.3    UO2 Thermal Conductivity ........................................................................ 3.5-3 3-ii                                                                Rev. 35
3.2-53.2.2Nuclear Design Bases.................................................................................3.2-6 3.2.3Thermal and Hydraulic Design Basis.........................................................
 
3.2-83.2.4References...................................................................................................3.2-8
tion      Title                                                                                                          Page 2.1.4    Gap Conductance ........................................................................................ 3.5-3 2.2      Departure from Nucleate Boiling Ratio...................................................... 3.5-3 2.2.1    Departure from Nucleate Boiling ............................................................... 3.5-3 2.2.2    Hot Channel Factors ................................................................................... 3.5-3 2.2.3    Effects of Rod Bow on DNBR ................................................................... 3.5-5 2.3      Void Fraction and Distribution ................................................................... 3.5-5 2.4      Coolant Flow Distribution .......................................................................... 3.5-5 2.4.1    Coolant Flow Distribution and Bypass Flow.............................................. 3.5-5 2.4.2    Core Flow Distribution ............................................................................... 3.5-6 2.5      Pressure Losses and Hydraulic Loads ........................................................ 3.5-6 2.5.1    Pressure Losses ........................................................................................... 3.5-6 2.5.2    Hydraulic Loads.......................................................................................... 3.5-7 2.6      Correlation and Physical Data .................................................................... 3.5-7 2.7      Plant Parameters for Thermal-Hydraulic Design........................................ 3.5-7 2.8      Summary of Thermal and Hydraulic Parameters ....................................... 3.5-8 3        Thermal And Hydraulic Evaluation............................................................ 3.5-8 3.1      Analytical Techniques and Uncertainties ................................................... 3.5-8 3.1.1    XCOBRA-IIIC DNBR Analyses ................................................................ 3.5-8 3.1.2    Parameter Uncertainties .............................................................................. 3.5-8 3.2      Hydraulic Instability Analysis .................................................................... 3.5-8 3.3      Core Hydraulics ........................................................................................ 3.5-11 3.3.1    Fuel Assembly Pressure Drop Coefficients .............................................. 3.5-11 3.3.2    Guide Tube Bypass Flow and Heating Analysis ...................................... 3.5-12 3.3.3    Control Element Assembly Insertion Time Analysis ............................... 3.5-13 3.3.4    Fuel Assembly Liftoff............................................................................... 3.5-13 4        Tests And Inspections ............................................................................... 3.5-14 4.1      Reactor Testing ......................................................................................... 3.5-14 4.2      AREVA DNB and Hydraulic Testing ...................................................... 3.5-14 4.2.1    DNB Testing ............................................................................................. 3.5-14 4.2.2    Fuel Assembly Hydraulic Testing ............................................................ 3.5-14 5        References................................................................................................. 3.5-15 ANALYSIS OF REACTOR VESSEL INTERNALS ........................................ 3.A-1
.1        Seismic Analysis........................................................................................ 3.A-1
.1.1      Introduction................................................................................................ 3.A-1
.1.2      Method of Analysis.................................................................................... 3.A-1
.1.2.1    General....................................................................................................... 3.A-1
.1.2.2    Mathematical Models ................................................................................ 3.A-1
.1.2.3    Natural Frequencies and Normal Modes ................................................... 3.A-3
.1.2.4    Response Calculations .............................................................................. 3.A-4 3-iii                                                              Rev. 35
 
tion Title                                                                                                          Page
.1.3 Results........................................................................................................ 3.A-5
.1.4 Conclusion ................................................................................................. 3.A-5
.2  Normal Operating Analysis ....................................................................... 3.A-5
.3  Loss of Coolant Accident Analysis ........................................................... 3.A-7
.3.1 Discussion .................................................................................................. 3.A-7
.3.2 Analysis Codes ........................................................................................ 3.A-10
.4  Effects of Thermal Shield Removal......................................................... 3.A-11
.5  Leak-Before-Break Analysis ................................................................... 3.A-11
.6  References................................................................................................ 3.A-12 3-iv                                                                Rev. 35


===3.3 MECHANICAL===
List of Tables mber Title 1   Stress Limits for Reactor Vessel Internal Structures 1   Mechanical Design Parameters
DESIGN...................................................................................
* 2   Pressurized Water Reactor Primary Coolant Water Chemistry Recommended Specifications 1   Fuel Characteristics for a Representative Reload Core 2    Neutronics Characteristics for a Representative Reload Core 3   Representative Shutdown Margin Requirements 1   Nominal Reactor and Fuel Design Parameters 2   Design Operating Hydraulic Loads on Vessel Internals 3   Uncertainty Sources for DNBR Calculations (DELETED)
3.3-13.3.1Core Mechanical Design.............................................................................3.3-13.3.1.1AREVA Fuel Rod.......................................................................................3.3-13.3.1.1.1Fuel Rod Mechanical Criteria.....................................................................3.3-1 3.3.1.1.2Fuel Rod Design Analyses..........................................................................3.3-33.3.1.2(Deleted).....................................................................................................3.3-63.3.1.3AREVA Fuel Assembly..............................................................................3.3-6 3.3.1.3.1Design Summary.........................................................................................
-1  Natural Frequencies for Vertical Seismic Analysis Mathematical Model
3.3-63.3.1.3.2Fuel Assembly Mechanical Criteria...........................................................
-2   Seismic Stresses in Critical Reactor Internals Components for the Design Basis Earthquake 3-v                                    Rev. 35
3.3-93.3.1.4Fuel Assembly Holddown Device............................................................3.3-12 3.3.1.5Control Element Assembly.......................................................................3.3-123.3.1.6Neutron Source Design.............................................................................
3.3-133.3.1.7In-Core Instruments..................................................................................3.3-133.3.1.8Heated Junction Thermocouples...............................................................
3.3-143.3.2Reactor Internal Structures.......................................................................
3.3-143.3.2.1Core Support Assembly............................................................................
3.3-153.3.2.2Core Support Barrel..................................................................................
3.3-153.3.2.3Core Support Plate and Support Columns................................................
3.3-163.3.2.4Core Shroud..............................................................................................3.3-163.3.2.5Flow Skirt.................................................................................................3.3-163.3.2.6Upper Guide Structure Assembly.............................................................3.3-16 3.3.3Control Element Drive Mechanism..........................................................3.3-173.3.3.1Design.......................................................................................................
3.3-17 MPS2 UFSAR Table of Contents (Continued)
Section Title Page 3-ii Rev. 353.3.3.2Control Element Drive Mechanism Pressure Housing.............................3.3-183.3.3.2.1Heated Junction Thermocouple Pressure Boundary.................................3.3-193.3.3.3Magnetic Jack Assembly..........................................................................3.3-193.3.3.4Position Indication....................................................................................
3.3-193.3.3.5Control Element Assembly Disconnect....................................................3.3-20 3.3.3.6Test Program.............................................................................................3.3-20 3.3.4References.................................................................................................3.3-20


===3.4 NUCLEAR===
List of Figures ure Title 1  Reactor Vertical Arrangement 2  Reactor Core Cross Section 1  Fuel Rod Assembly 2A AREVA - Reload Fuel Assembly Batch S and Prior 2B AREVA - Reload Fuel Assembly Batch T and Later 3A AREVA - Reload Fuel Assembly Components Batch S and Prior 3B AREVA - Reload Fuel Assembly Components Batch T and Later 4A Bi-Metallic Fuel Spacer Assembly 4B HTP Fuel Space Assembly 5  Fuel Assembly Hold Down Device
DESIGN AND EVALUATION......................................................
-6  Control Element Assembly 7  Control Element Assembly Materials 8  Control Element Assemblies Group and Number Designation 9  Core Orientation 10 In-Core Instrumentation Assembly
3.4-13.4.1General Summary.......................................................................................3.4-13.4.2Core Description.........................................................................................3.4-1 3.4.3Nuclear Core Design...................................................................................3.4-13.4.3.1Analytical Methodology.............................................................................
-11 Reactor Internals Assembly 12 Pressure Vessel-Core Support Barrel Snubber Assembly 13 Core Shroud Assembly 14 Upper Guide Structure Assembly 15 Control Element Drive Mechanism (Magnetic Jack) 16 (Left Blank Intentionally) 17 Heated Junction Thermocouple Probe Pressure Boundary Installation 18 Typical Heated Junction Thermocouple Probe Assembly Installation 19 Placement of Natural Uranium Replacement Fuel Rods and Fuel Assembly Orientation Relative to the Core Baffle for Cycle 19 1  Representative Full Core Loading Pattern 2  Representative Quarter Core Loading Pattern 3-vi                              Rev. 35
3.4-23.4.3.2Physics Characteristics...............................................................................
 
3.4-23.4.3.2.1Power Distribution Considerations.............................................................
NOTE: REFER TO THE CONTROLLED PLANT DRAWING FOR THE LATEST REVISION.
3.4-23.4.3.2.2Control Rod Reactivity Requirements........................................................3.4-23.4.3.2.3Moderator Temperature Coefficient Considerations..................................
ure    Title 3      Representative BOC and EOC Exposure Distribution 4     Representative Boron Letdown, HFP, ARO 5      Representative Normalized Power Distributions, Hot Full Power, Equilibrium Xenon, 150 MWD/MTU 6      Representative Normalized Power Distribution, Hot Full Power, Equilibrium Xenon, 18,020 MWD/MTU
3.4-33.4.4Post-Reload Startup Testing.......................................................................
-1      Representative Node Locations - Horizontal Mathematical Model
3.4-33.4.5Reactor Stability.........................................................................................
-2      Mathematical Model - Horizontal Seismic Analysis
3.4-43.4.5.1General........................................................................................................3.4-4 3.4.5.2Detection of Oscillations............................................................................
-3      Mathematical Model - Vertical Seismic Analysis
3.4-43.4.5.3Control of Oscillations................................................................................
-4     Core Support Barrel Upper Flange - Finite Element Model
3.4-53.4.5.4Operating Experience.................................................................................
-5      Core Support Barrel Lower Flange - Finite Element Model
3.4-63.4.5.5Method of Analysis.....................................................................................
-6      Lateral Seismic Model - Mode 1, 3.065 CPS
3.4-63.4.5.5.1Radial Xenon Oscillations..........................................................................
-7      Lateral Seismic Model - Mode 2, 5.118 CPS
3.4-73.4.5.5.2Azimuthal Xenon Oscillations....................................................................
-8      Lateral Seismic Model - Mode 2, 5.118 CPS
3.4-73.4.5.5.3Axial Xenon Oscillations............................................................................
-9      Reactor Vessel Flange Vertical Response Spectrum (1% Damping)
3.4-73.4.6References...................................................................................................3.4-8 3.5THERMAL-HYDRAULIC DESIGN..................................................................3.5-13.5.1Design Bases...............................................................................................3.5-13.5.1.1Thermal Design...........................................................................................3.5-13.5.1.2Hydraulic Stability......................................................................................3.5-13.5.1.3Coolant Flow Rate, Distribution and Void Fraction...................................
-10    ASHSD Finite Element Model of the Core Support Barrel/Thermal Shield System
3.5-13.5.2Thermal and Hydraulic Characteristics of the Design................................
-11    Vertical Shock Model
3.5-23.5.2.1Fuel Temperatures......................................................................................3.5-2 3.5.2.1.1Fuel Cladding Temperatures.......................................................................
-12    Lateral Shock Mode
3.5-23.5.2.1.2Fuel Pellet Temperatures............................................................................
-13    SAMMSOR DYNASOR Finite Element Model of Core Support Barrel 3-vii                                    Rev. 35
3.5-23.5.2.1.3UO2 Thermal Conductivity........................................................................
3.5-3 MPS2 UFSAR Table of Contents (Continued)
Section Title Page 3-iii Rev. 353.5.2.1.4Gap Conductance........................................................................................3.5-33.5.2.2Departure from Nucleate Boiling Ratio......................................................3.5-33.5.2.2.1Departure from Nucleate Boiling...............................................................
3.5-33.5.2.2.2Hot Channel Factors...................................................................................
3.5-33.5.2.2.3Effects of Rod Bow on DNBR...................................................................
3.5-53.5.2.3Void Fraction and Distribution...................................................................
3.5-53.5.2.4Coolant Flow Distribution..........................................................................3.5-53.5.2.4.1Coolant Flow Distribution and Bypass Flow..............................................
3.5-53.5.2.4.2Core Flow Distribution...............................................................................
3.5-63.5.2.5Pressure Losses and Hydraulic Loads........................................................3.5-6 3.5.2.5.1Pressure Losses...........................................................................................
3.5-63.5.2.5.2Hydraulic Loads..........................................................................................3.5-73.5.2.6Correlation and Physical Data....................................................................
3.5-73.5.2.7Plant Parameters for Thermal-Hydraulic Design........................................
3.5-73.5.2.8Summary of Thermal and Hydraulic Parameters.......................................
3.5-83.5.3Thermal And Hydraulic Evaluation............................................................3.5-83.5.3.1Analytical Techniques and Uncertainties...................................................
3.5-83.5.3.1.1XCOBRA-IIIC DNBR Analyses................................................................
3.5-83.5.3.1.2Parameter Uncertainties..............................................................................
3.5-83.5.3.2Hydraulic Instability Analysis....................................................................3.5-83.5.3.3Core Hydraulics........................................................................................3.5-11 3.5.3.3.1Fuel Assembly Pressure Drop Coefficients..............................................
3.5-113.5.3.3.2Guide Tube Bypass Flow and Heating Analysis......................................
3.5-123.5.3.3.3Control Element Assembly Insertion Time Analysis...............................3.5-13 3.5.3.3.4Fuel Assembly Liftoff...............................................................................
3.5-133.5.4Tests And Inspections...............................................................................
3.5-143.5.4.1Reactor Testing.........................................................................................
3.5-143.5.4.2AREVA DNB and Hydraulic Testing......................................................3.5-143.5.4.2.1DNB Testing.............................................................................................
3.5-143.5.4.2.2Fuel Assembly Hydraulic Testing............................................................
3.5-143.5.5References.................................................................................................3.5-15 3.A ANALYSIS OF REACTOR VESSEL INTERNALS........................................
3.A-13.A.1Seismic Analysis........................................................................................3.A-13.A.1.1Introduction................................................................................................3.A-13.A.1.2Method of Analysis....................................................................................
3.A-13.A.1.2.1General.......................................................................................................3.A-13.A.1.2.2Mathematical Models................................................................................
3.A-13.A.1.2.3Natural Frequencies and Normal Modes...................................................
3.A-33.A.1.2.4 Response Calculations..............................................................................
3.A-4 MPS2 UFSAR Table of Contents (Continued)
Section Title Page 3-iv Rev. 353.A.1.3Results........................................................................................................3.A-53.A.1.4Conclusion.................................................................................................3.A-53.A.2Normal Operating Analysis.......................................................................3.A-5 3.A.3Loss of Coolant Accident Analysis...........................................................3.A-73.A.3.1Discussion..................................................................................................3.A-73.A.3.2Analysis Codes........................................................................................
3.A-103.A.4Effects of Thermal Shield Removal.........................................................3.A-113.A.5Leak-Before-Break Analysis...................................................................
3.A-113.A.6References................................................................................................
3.A-12 MPS2 UFSAR 3-v Rev. 35CHAPTER 3-REACTOR List of Tables Number Title3.2-1Stress Limits for Reacto r Vessel Internal Structures3.3-1Mechanical Design Parameters
*3.3-2Pressurized Water Reac tor Primary Coolant Water Chemistry Recommended Specifications3.4-1Fuel Characteristics for a Representative Reload Core3.4-2Neutronics Characteristics for a Representative Reload Core 3.4-3Representative Shutdown Margin Requirements 3.5-1Nominal Reactor and Fuel Design Parameters 3.5-2Design Operating Hydraulic Loads on Vessel Internals 3.5-3Uncertainty Sources for DNBR Calculations (DELETED)3.A-1Natural Frequencies for Vertical Seismic Analys is Mathematical Model3.A-2Seismic Stresses in Critical Reactor Internals Components for the Design Basis Earthquake MPS2 UFSARNOTE: REFER TO THE CONTROLLED PLANT DRAWING FOR THE LATEST REVISION.
3-vi Rev. 35CHAPTER 3 - REACTOR List of Figures Figure Title3.1-1Reactor Vertical Arrangement3.1-2Reactor Core Cross Section 3.3-1Fuel Rod Assembly 3.3-2AAREVA - Reload Fuel Assembly Batch "S" and Prior 3.3-2BAREVA - Reload Fuel Assembly Batch "T" and Later 3.3-3AAREVA - Reload Fuel Assembly Components Batch "S" and Prior3.3-3BAREVA - Reload Fuel Assembly Components Batch "T" and Later3.3-4ABi-Metallic Fuel Spacer Assembly 3.3-4BHTP Fuel Space Assembly 3.3-5Fuel Assembly Hold Down Device 3.3-6Control Element Assembly 3.3-7Control Element Assembly Materials 3.3-8Control Element Assemblies Group and Number Designation 3.3-9Core Orientation 3.3-10In-Core Instrumentation Assembly 3.3-11Reactor Internals Assembly 3.3-12Pressure Vessel-Core S upport Barrel Snubber Assembly3.3-13Core Shroud Assembly 3.3-14Upper Guide Structure Assembly 3.3-15Control Element Drive Mechanism (Magnetic Jack) 3.3-16(Left Blank Intentionally) 3.3-17Heated Junction Thermocouple Probe Pressure Boundary Installation3.3-18Typical Heated Junction Thermoc ouple Probe Assembly Installation3.3-19Placement of Natural Uranium Replacement Fuel Rods and Fuel Assembly Orientation Relative to the Core Baffle for Cycle 193.4-1Representative Full Core Loading Pattern 3.4-2Representative Quarter Core Loading Pattern MPS2 UFSAR List of Figures (Continued)NOTE: REFER TO THE CONTROLLED PLANT DRAWING FOR THE LATEST REVISION.
Figure Title 3-vii Rev. 353.4-3Representative BOC and EOC Exposure Distribution3.4-4Representative Boron Letdown, HFP, ARO3.4-5Representative Normalized Power Dist ributions, Hot Full Power, Equilibrium Xenon, 150 MWD/MTU3.4-6Representative Normalized Power Di stribution, Hot Full Power, Equilibrium Xenon, 18,020 MWD/MTU3.A-1Representative Node Locations - Horizontal Mathematical Model3.A-2Mathematical Model - Ho rizontal Seismic Analysis3.A-3Mathematical Model - Vertical Seismic Analysis3.A-4Core Support Barrel Upper Flange - Finite Element Model3.A-5Core Support Barrel Lower Flange - Finite Element Model3.A-6Lateral Seismic Model - Mode 1, 3.065 CPS 3.A-7Lateral Seismic Model - Mode 2, 5.118 CPS 3.A-8Lateral Seismic Model - Mode 2, 5.118 CPS 3.A-9Reactor Vessel Flange Verti cal Response Spectrum (1% Damping)3.A-10ASHSD Finite Element Model of the Co re Support Barrel/Thermal Shield System3.A-11Vertical Shock Model 3.A-12Lateral Shock Mode 3.A-13SAMMSOR DYNASOR Finite Elemen t Model of Core Support Barrel MPS2 UFSAR3.1-1Rev. 35CHAPTER 3 - REACTOR 3.1


==SUMMARY==
==SUMMARY==
DESCRIPTION The reactor is of the pressurized water type using two reactor coolant loops. A vertical cross section of the reactor is shown in Figure 3.1-1. The reactor core is comp osed of 217 fuel assemblies, 73 control element a ssemblies (CEA) and four neutr on source assemblies. The fuel assemblies are arranged to approximate a right circ ular cylinder with an e quivalent diameter of 136 inches and an active length of 136.7 inches. The fuel assemblies are co mprised of a structure and fuel and poison rods. The st ructure, which provides for 176 ro d positions, consists of five guide tubes attached to spacer grids and is encl osed at the top and botto m by end fittings. Each of the guide tubes replaces four fu el rod positions and provides a channel which guides the control element over its entire le ngth of travel. In selected fuel asse mblies the central guide tube houses in-core instrumentation. The reactor is currently fueled by assemblies produced by AREVA.The fuel is low enrichment UO 2 in the form of ceramic pellets an d encapsulated in zircaloy tubes. These tubes are seal welded as hermetic enclosures.Figure 3.1-2 shows a view of the reactor core cross section a nd some dimensional relations between fuel assemblies, fuel rods and CEA guide tubes.
DESCRIPTION reactor is of the pressurized water type using two reactor coolant loops. A vertical cross ion of the reactor is shown in Figure 3.1-1. The reactor core is composed of 217 fuel mblies, 73 control element assemblies (CEA) and four neutron source assemblies. The fuel mblies are arranged to approximate a right circular cylinder with an equivalent diameter of inches and an active length of 136.7 inches. The fuel assemblies are comprised of a structure fuel and poison rods. The structure, which provides for 176 rod positions, consists of five de tubes attached to spacer grids and is enclosed at the top and bottom by end fittings. Each of guide tubes replaces four fuel rod positions and provides a channel which guides the control ment over its entire length of travel. In selected fuel assemblies the central guide tube houses ore instrumentation. The reactor is currently fueled by assemblies produced by AREVA.
The reactor internals support and orient the fuel assemblies and CEAs, absorb the static and dynamic loads and transmit the loads to the reactor vessel flange, provide a passage way for the reactor coolant, and guide in-core instrumentation. The internals will safely perform their func tion during normal operating, upset and emergency conditions. The internals are designe d to safely withstand the forces due to dead weight, pressure differential, flow impingement, temperature differential, vibrat ions and seismic acceleration. All reactor components are consider ed category 1 for seismic desi gn. The reactor internals design limits deflection where required by function. Where necessary, components have been subjected to fatigue analysis. Where appropriate, the effect of neutron irradi ation on the mate rials concerned is included in the design evaluation. The effects of shock loadings on the in ternals is included in the design analysis.
fuel is low enrichment UO2 in the form of ceramic pellets and encapsulated in zircaloy tubes.
Reactivity control is provided by two independent systems: The control element drive system (CEDS) and the chemical and volume control system (CVCS). The CEDS controls short term reactivity changes and is used for rapid shutdown.
se tubes are seal welded as hermetic enclosures.
The CVCS is used to co mpensate for long term reactivity changes and can make the reactor subc ritical without the bene fit of the CEDS. The design of the core and the reactor protective system (RPS) prevents fuel damage limits from being exceeded for any single malfunction in either of the reactivity control systems.
ure 3.1-2 shows a view of the reactor core cross section and some dimensional relations ween fuel assemblies, fuel rods and CEA guide tubes.
The CEAs consist of five poison rods (control elements) assembled in a square array, with one rod in the center. The rods are connect ed to a spider casting which is coupled to the control element drive mechanism (CEDM) shaft.
reactor internals support and orient the fuel assemblies and CEAs, absorb the static and amic loads and transmit the loads to the reactor vessel flange, provide a passage way for the tor coolant, and guide in-core instrumentation.
There are a total of 73 CEAs.
internals will safely perform their function during normal operating, upset and emergency ditions. The internals are designed to safely withstand the forces due to dead weight, pressure erential, flow impingement, temperature differential, vibrations and seismic acceleration. All tor components are considered category 1 for seismic design. The reactor internals design ts deflection where required by function. Where necessary, components have been subjected atigue analysis. Where appropriate, the effect of neutron irradiation on the materials concerned cluded in the design evaluation. The effects of shock loadings on the internals is included in design analysis.
Some CEAs are mechanically connected in pairs and are known as dual CEAs.
ctivity control is provided by two independent systems: The control element drive system DS) and the chemical and volume control system (CVCS). The CEDS controls short term tivity changes and is used for rapid shutdown. The CVCS is used to compensate for long term tivity changes and can make the reactor subcritical without the benefit of the CEDS. The gn of the core and the reactor protective system (RPS) prevents fuel damage limits from being eeded for any single malfunction in either of the reactivity control systems.
MPS2 UFSAR3.1-2Rev. 35 Both dual and single CEAs are maneuvered by magnetic jack type CEDM's mounted on the reactor vessel head.The maximum reactivity worth of the CEAs and the associated reactivity a ddition rate are limited by core, CEA and CEDS design to prevent sudden large reactivity in creases. The design restraints are such that reactivity increases will not result in violation of th e fuel damage limits, rupture of the reactor coolant pressure boundary (RCPB), or disruption of the core or other internals sufficient to impair the effectiveness of emergency cooling.
CEAs consist of five poison rods (control elements) assembled in a square array, with one rod he center. The rods are connected to a spider casting which is coupled to the control element e mechanism (CEDM) shaft. There are a total of 73 CEAs. Some CEAs are mechanically nected in pairs and are known as dual CEAs.
The three-batch fuel management scheme is employed, where approximate ly 40 percent of the core is replaced at each refueling. Sufficient margin is provided to ensure that peak burnups of the individual fuel assemblies ar e within acceptable limits.
3.1-1                                    Rev. 35
The nuclear design of the core will ensure that th e combined response of all reactivity coefficients to an increase in reactor thermal power yields a net decrease in reactiv ity and that CEAs are moved in groups to satisfy the requirements of shutdown, power level changes and operational maneuvering. The control systems are designed to produce power dist ributions that are within the acceptable limits on overall nuc lear heat flux factor (F N Q) and departure from nucleate boiling ratio (DNBR). The RPS and administ rative controls ensure that these limits are not exceeded.
 
The reactor coolant enters the upper section of the reactor vessel through f our inlet nozzles, flows downward between the reactor vessel shell and the core barrel, and passes th rough the flow skirt and into the lower plenum where the flow dist ribution is equalized. The coolant then flows upward through the core removing h eat from the fuel rods, exits from the reactor vessel through two outlet nozzles and passes thr ough the tube side of the verti cal "U" tube steam generators where heat is transferred to the secondary system. The reactor coolant pumps (RCPs) return the coolant to the reactor vessel.
maximum reactivity worth of the CEAs and the associated reactivity addition rate are limited ore, CEA and CEDS design to prevent sudden large reactivity increases. The design restraints such that reactivity increases will not result in violation of the fuel damage limits, rupture of reactor coolant pressure boundary (RCPB), or disruption of the core or other internals icient to impair the effectiveness of emergency cooling.
The principal objective of the th ermal-hydraulic design is to a void fuel damage during normal operation and anticipated transients. It is recogni zed that there is a small probability of limited fuel damage in certain situations as discussed in Chapter 14.In order to meet the objective of the thermal-hydraulic design the following design limits are established, but violation of either is not necessarily equivalent to fuel damage:a.There is a high confidence level that departure from nucl ear boiling (DNB) is avoided during normal operation and anticipated transients. This is achieved by confirming the minimum DNBR calculated according to the HTP correlation (Reference 3.1-1) is greater than the 95/95 limit for the correlation;b.The melting point of the UO 2 fuel is not reached during normal operation or anticipated transients.
three-batch fuel management scheme is employed, where approximately 40 percent of the is replaced at each refueling. Sufficient margin is provided to ensure that peak burnups of the vidual fuel assemblies are within acceptable limits.
The RPS and the reactor c ontrol system (RCS) provide for automatic reactor trip or corrective actions before these design limits are exceeded.
nuclear design of the core will ensure that the combined response of all reactivity coefficients n increase in reactor thermal power yields a net decrease in reactivity and that CEAs are ved in groups to satisfy the requirements of shutdown, power level changes and operational euvering. The control systems are designed to produce power distributions that are within the eptable limits on overall nuclear heat flux factor (FNQ) and departure from nucleate boiling o (DNBR). The RPS and administrative controls ensure that these limits are not exceeded.
MPS2 UFSAR3.1-3Rev. 35The core design bases are presented in Section 3.2; the core mechanical design is discussed in Section 3.3; the nuclear design of the core is discussed in Section 3.4; and the thermal and hydraulic design is discussed in Section 3.5.
reactor coolant enters the upper section of the reactor vessel through four inlet nozzles, flows nward between the reactor vessel shell and the core barrel, and passes through the flow skirt into the lower plenum where the flow distribution is equalized. The coolant then flows ard through the core removing heat from the fuel rods, exits from the reactor vessel through outlet nozzles and passes through the tube side of the vertical U tube steam generators re heat is transferred to the secondary system. The reactor coolant pumps (RCPs) return the lant to the reactor vessel.
3.
principal objective of the thermal-hydraulic design is to avoid fuel damage during normal ration and anticipated transients. It is recognized that there is a small probability of limited damage in certain situations as discussed in Chapter 14.
rder to meet the objective of the thermal-hydraulic design the following design limits are blished, but violation of either is not necessarily equivalent to fuel damage:
: a. There is a high confidence level that departure from nuclear boiling (DNB) is avoided during normal operation and anticipated transients. This is achieved by confirming the minimum DNBR calculated according to the HTP correlation (Reference 3.1-1) is greater than the 95/95 limit for the correlation;
: b. The melting point of the UO2 fuel is not reached during normal operation or anticipated transients.
RPS and the reactor control system (RCS) provide for automatic reactor trip or corrective ons before these design limits are exceeded.
3.1-2                                    Rev. 35
 
raulic design is discussed in Section 3.5.
1  REFERENCES 1    EMF-92-153(P)(A) Rev. 1, HTP: Departure From Nucleate Boiling Correlation for High Thermal Performance Fuel, Siemens Power Corporation, January 2005.
3.1-3                                Rev. 35


==1.1 REFERENCES==
MPS-2 FSAR FIGURE 3.1-1 REACTOR VERTICAL ARRANGEMENT
3.1-1EMF-92-153(P)(A) Rev. 1, "H TP: Departure From Nucleate Boiling Correlation for High Thermal Performance Fuel," Siem ens Power Corporation, January 2005.
          'N'-CORE tNSTRUMENT ATI ON ASSEMBLY CONTROl ELEMENT ASSEMBLY EX'PANSION                              CFUllYWITHDRAWNl RI~
MPS-2 FSAR April 1998 Rev. 26.2FIGURE 3.1-1REACTOR VE RTICAL ARRANGEMENT MPS-2 FSAR April 1998 Rev. 26.2FIGURE 3.1-2REACTOR CORE CROSS SECTION MPS2 UFSAR3.2-1Rev. 35
ALIGNMENT PIN
                                                ~        UPPER 40' 11 7h" GUIDE
                                                        ~
                                                      .....--INLET NOZZLE Nt--CORE
              ..---==--!J-J~~.",MM                      SUPPORT BARREL*
n5-7i!OH                        ~...,..,.- FUEL A,CTIVE ASSEMBLY CORE UENGTH                            ""-CORE SHROUD CORE CORE: STOP                                        SUPPORT ASSEMBLY
.t FlCNY SKIRT April 1998                     Rev. 26.2


===3.2 DESIGN===
MPS-2 FSAR FIGURE 3.1-2 REACTOR CORE CROSS SECTION CORE                                              REACTOR
BASES The full power thermal rating of the core is 2,700 MWt. The physics a nd thermal and hydraulic information presented in this secti on is based on this core power level.
        ~~UIVALENT                                        VESSEl.
13~ETER.
CORE SUPPORT BARREl----.;;~                                          CORE SHROUD FUEl ROD                                                GUIDE TUBE 0.440' OD o!~..,
                ~TT          TT
                          "";~_'        "TT  r  T' '.'
0.S8O".........
      ~~~nOD
                      ~ .. 1-0.1;:11- ~ Oltr?P~E
                                ;..tHua    RODS J
0.200  11 WATER GAP April 1998                    Rev. 26.2


====3.2.1 MECHANICAL====
full power thermal rating of the core is 2,700 MWt. The physics and thermal and hydraulic rmation presented in this section is based on this core power level.
DESIGN BASES 3.2.1.1 Fuel Assembly Design Bases The design bases for evaluating the structural integrity of AREVA fuel assemblies are:
1   MECHANICAL DESIGN BASES 1.1   Fuel Assembly Design Bases design bases for evaluating the structural integrity of AREVA fuel assemblies are:
A.Fuel Assembly Handling The fuel assembly is evaluated for dynamic ax ial loads of approximately 2.5 times the fuel assembly weight.B.For All Applied Loads for Normal Oper ation and Anticipated Operational EventsFuel assembly component strength is evaluated against either prototype testing or elastic stress analysis. When the stress analysis method is used, the stress limits presented in the ASME Boiler and Pressure Vessel Code, Section III, Division 1, are used as a guide.
Fuel Assembly Handling The fuel assembly is evaluated for dynamic axial loads of approximately 2.5 times the fuel assembly weight.
The stress design limits for structural components are:
For All Applied Loads for Normal Operation and Anticipated Operational Events Fuel assembly component strength is evaluated against either prototype testing or elastic stress analysis. When the stress analysis method is used, the stress limits presented in the ASME Boiler and Pressure Vessel Code, Section III, Division 1, are used as a guide.
P m 1.0S m P m + P b 1.5S m P + Q 3.0S m where: P m is the primary membrane stress intensity P b is the primary bending stress intensityP is the primary stress intensity Q is the secondary stress intensity The design stress, S m is identified in the ASME Boiler and Pressure Vessel Code for austenitic stainless steel as a function of temperature. In the case of Zircaloy, which is not specifically identified in the ASME Boiler and Pressure Vess el Code, the design stress is identified as the lesser of two-thirds the yield stress, S y , or one-third the ultimate stress, S u.The ASME Boiler and Pressure Vessel Code de fines the stress intensity based on the maximum shear stress theory. The stress intensity is equal to one-half the largest algebraic difference between two principal stresses.
stress design limits for structural components are:
MPS2 UFSAR3.2-2Rev. 35 Primary stresses are deve loped by the imposed loading which is necessary to satisfy the laws of equilibrium between external and internal forces and moments. The basic characteristic 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 wall thickness, the prevention of fail ure is entirely dependent on the strain-hardening properties of the material.
Pm  1.0Sm Pm + Pb  1.5Sm P + Q 3.0Sm re:
Secondary stresses are developed by the self-constrai nt 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 yi elding and minor distor tions can satisfy the discontinuity conditions due to thermal expansions which cause the stress to occur.C.Loads during Postulated Accidents Deflection or failure of components shall not interfere with reactor shutdown or emergency cooling of the fuel rods.
Pm is the primary membrane stress intensity Pb is the primary bending stress intensity P is the primary stress intensity Q is the secondary stress intensity design stress, Sm is identified in the ASME Boiler and Pressure Vessel Code for austenitic nless steel as a function of temperature. In the case of Zircaloy, which is not specifically tified in the ASME Boiler and Pressure Vessel Code, the design stress is identified as the er of two-thirds the yield stress, Sy, or one-third the ultimate stress, Su.
ASME Boiler and Pressure Vessel Code defines the stress intensity based on the maximum ar stress theory. The stress intensity is equal to one-half the largest algebraic difference ween two principal stresses.
3.2-1                                    Rev. 35


The fuel assembly structural component stresses under faulted conditions are evaluated using primarily the methods outlined in Appendix F of the ASME Boiler and Pressure Vessel Code, Section III. The current methods utilize the limits provided fo r elastic system analysis.
ary stress is that it is not self-limiting. If a primary stress exceeds the yield strength of the erial through the entire wall thickness, the prevention of failure is entirely dependent on the in-hardening properties of the material.
The design stress intensity value (S m) is defined the same as fo r normal operating conditions.
ondary stresses are developed by the self-constraint of a structure. It must satisfy an imposed in pattern rather than being in equilibrium with an external load. The basic characteristic of a ondary stress is that it is self-limiting. Local yielding and minor distortions can satisfy the ontinuity conditions due to thermal expansions which cause the stress to occur.
Spacer grid crush load strength is based on the 95% confidence le vel on the true mean as taken from test measurements on unirradiated production grids at (or corrected to) operating temperature.
Loads during Postulated Accidents lection or failure of components shall not interfere with reactor shutdown or emergency ling of the fuel rods.
fuel assembly structural component stresses under faulted conditions are evaluated using arily the methods outlined in Appendix F of the ASME Boiler and Pressure Vessel Code, tion III. The current methods utilize the limits provided for elastic system analysis.
design stress intensity value (Sm) is defined the same as for normal operating conditions.
cer grid crush load strength is based on the 95% confidence level on the true mean as taken m test measurements on unirradiated production grids at (or corrected to) operating perature.
1.2      AREVA Fuel Rod Cladding Design Bases iscussion of the AREVA fuel rod cladding is given as part of the AREVA fuel rod discussion ection 3.3.1.1.
1.3 Control Element Assembly Design Bases CEA has been designed to ensure that the stress intensities in the individual structural ponents do not exceed the allowable limits for the appropriate material established in tion III of the ASME Boiler and Pressure Vessel Code. The exceptions to this criterion are that he Inconel 625 cladding is permitted to sustain plastic strain up to 3 percent due to irradiation uced expansion of the filler materials, and (b) because the ASME Code does not apply to ngs, the allowable stresses for the CEA springs are based on values which have been proven in tice.
CEA stress analyses consider the following load sources:
: a.      Internal pressure build up due to the effect of irradiation on B4C (production of helium).
3.2-2                                  Rev. 35


3.2.1.2 AREVA Fuel Rod Cladding Design BasesA discussion of the AREVA fuel ro d cladding is given as part of the AREVA fuel rod discussion in Section 3.3.1.1.
assumed).
3.2.1.3 Control Element Assembly Design Bases The CEA has been designed to ensure that the stress intensitie s in the individual structural components do not exceed the allowable limits for the appropriate material established in Section III of the ASME Boiler and Pressure Vessel Code. The exceptions to this criterion are that (a) the Inconel 625 cladding is permit ted to sustain plastic strain up to 3 percent due to irradiation induced expansion of the filler materials, and (b) because th e ASME Code does not apply to springs, the allowable stresses fo r the CEA springs are based on valu es which have been proven in practice.The CEA stress analyses consider the following load sources:a.Internal pressure build up due to the effect of irradiation on B 4 C (production of helium).
: c.     Dynamic stresses produced by seismic loading.
MPS2 UFSAR3.2-3Rev. 35b.External pressure of r eactor coolant (in the computation for determining the maximum stress in the cladding due to inte rnal pressure, no inte rnal pressure is assumed).c.Dynamic stresses produced by seismic loading.d.Dynamic loads produced by stepping motion of the magnetic jack.
: d.     Dynamic loads produced by stepping motion of the magnetic jack.
e.Mechanical and hydraulic loads produced during SCRAM.f.Cladding loads produced by differential expansion between clad and filler materials.In addition to the comparison of calculated stress levels with allowable stresses, the fatigue damage produced by significant cyclic stresses is also determined.
: e.     Mechanical and hydraulic loads produced during SCRAM.
It is a design requirement that the calculated cumulative damage factor for any location may not be equal to or greater than 1.0.
: f.     Cladding loads produced by differential expansion between clad and filler materials.
The fatigue usage factor calculations are based on the fatigue curves (str ess range vs. number of cycles) contained in Section III of the ASME Boiler and Pressure Vessel Code.
addition to the comparison of calculated stress levels with allowable stresses, the fatigue age produced by significant cyclic stresses is also determined. It is a design requirement that calculated cumulative damage factor for any location may not be equal to or greater than 1.0.
3.2.1.4 Reactor Internals Design Bases The reactor vessel internals are designed to meet the loading c onditions and the design limits specified below. The materials used in fabrication of the reactor internal structures are primarily Type 304 stainless steel. Th e flow skirt is fabricated from Inconel. Welded connections are used where feasible; however, in locations where mechanical connections are required, structural fasteners are used which are designed to remain captured in the event of a single failure.
fatigue usage factor calculations are based on the fatigue curves (stress range vs. number of les) contained in Section III of the ASME Boiler and Pressure Vessel Code.
Structural fastener material is typically a high strength austenitic stainless steel; however, in less critical applications, Type 316 stai nless steel is employed. Hardfacing, of Stellite material, is used at wear points. The effe ct of irradiation on the properties of the materials is considered in the design of the reactor internal structures.A.Categorization and Combination of Loadings1.Normal Operating and Upset Conditions The reactor vessel internals are designed to perform their functi ons safely without shutdown. The combination of design lo adings for these conditions are the following:Normal operating temperature differencesNormal operating pressure differences
1.4     Reactor Internals Design Bases reactor vessel internals are designed to meet the loading conditions and the design limits cified below. The materials used in fabrication of the reactor internal structures are primarily e 304 stainless steel. The flow skirt is fabricated from Inconel. Welded connections are used re feasible; however, in locations where mechanical connections are required, structural eners are used which are designed to remain captured in the event of a single failure.
ctural fastener material is typically a high strength austenitic stainless steel; however, in less cal applications, Type 316 stainless steel is employed. Hardfacing, of Stellite material, is used ear points. The effect of irradiation on the properties of the materials is considered in the gn of the reactor internal structures.
Categorization and Combination of Loadings
: 1.     Normal Operating and Upset Conditions The reactor vessel internals are designed to perform their functions safely without shutdown. The combination of design loadings for these conditions are the following:
Normal operating temperature differences Normal operating pressure differences Low impingement loads Weights, reactions and superimposed loads 3.2-3                                    Rev. 35


Low impingement loads Weights, reactions and superimposed loads MPS2 UFSAR3.2-4Rev. 35Vibration loads Shock loads (including OBE)Transient loadings of frequent occurrences not requiring shutdown Handling loads2.Emergency Conditions The internals are designed to permit an acceptable amount of local yielding while experiencing the loadings listed above with the SSE load replacing the OBE load.3.Faulted Conditions Permanent deformation of the reactor internal structur es is permitted. The loadings for these conditions include all the loadings listed for emerge ncy conditions plus the loadings resulting from the postulated LOCA.B.Design Limits Reactor internal compone nts are designed to ensure that th e stress levels and deflections are within an accept able range. The stress values for core support structures are not greater than those given in the May 1972 draft of Section III of the ASME Boiler and Pressure Vessel Code, Subsection NG, including Appendix F, "Rules for Evaluation of Faulted Conditions." St ress limits for the reactor ve ssel core support structures are presented in Table 3.2-1. In addition, to properly pe rform their functions, the r eactor internal structures will satisfy the deformation limits listed below
Shock loads (including OBE)
.1.Under design loadings plus operating ba sis earthquake forces or normal operating loadings plus SSE forces, de flections will be limited so that the CEAs can function and adequate core cooling is preserved.2.Under normal operating loadings plus SS E forces plus pipe rupture loadings resulting from a break of the largest line connect to the primary system piping, deflections will be limited so that the core will be held in place, adequate core cooling is preserved, and all CEAs can be inserted. Those deflections which would influence CEA movement will be limited to less than 80 percen t of the deflections required to prevent CEA insertion.3.Under normal operating loadings plus SSE forces plus the maximum pipe rupture loadings resulting from the full spectrum of pipe breaks, deflections will be limited so that the core will be held in place and adequate core cooling is preserved.
Transient loadings of frequent occurrences not requiring shutdown Handling loads
Although CEA insertion is not required for a safe and orderly shutdown for break sizes greater than the largest line connected to the primary system piping, calculations show that the CEAs will be insertable for larger breaks except for a MPS2 UFSAR3.2-5Rev. 35 few CEAs located near the vessel outlet nozzle which is feeding the postulated rupture.3.2.1.5 CEDM/RVLMS (HJTC) Pressu re Housing Design Bases The control element drive mechanism and Reactor Vessel Level Monitoring System (RVLMS) pressure housings form part of the reactor coolant boundary and are, therefore, designed to meet the stress requirements consistent with those of the reactor ve ssel closure head. The limiting stresses in the CEDM's and RV LMS pressure boundary components due to the design, Level A, Level B, Level C, Level D and Test conditions satisfy ASME Boiler Pressure Vessel Code, Section III, Subsection NB plus Appendix 1 and Section II, Pa rt D, 1998 Edition through 2000 Addenda, including Code Case N-4-12 for the CEDM motor housing material.The CEDMs and the RVLMs are designed to function normally during and after exposure to normal operating conditions plus the design basis earthquake (DBE). Under normal operating conditions, plus DBE, plus pipe rupture loadings, de flections of the CEDM will be limited so that the CEAs can be inserted afte r exposure to these conditions.
: 2.     Emergency Conditions The internals are designed to permit an acceptable amount of local yielding while experiencing the loadings listed above with the SSE load replacing the OBE load.
Those deflections, which could influence CEA movement, will be limited to less than 80 percent of the deflections required to prevent CEA movement. The RVLMS and the adja cent CEDMs do not cont act each other with maximum lateral displacement of the pressure housings.
: 3.     Faulted Conditions Permanent deformation of the reactor internal structures is permitted. The loadings for these conditions include all the loadings listed for emergency conditions plus the loadings resulting from the postulated LOCA.
Loading Combinations ASME Code SubsectionDesign ConditionP m  S m NB-3221 P 1  1.5S m P 1 + P b < 1.5S m Level A and Level BP 1 + P b + Q  3S m NB-3222 and NB3223(Normal and Upset)U  1Level C ConditionP m  greater of [1.2S m , S y]NB-3224(Emergency)P 1 + P b  greater of [1.8S m , 1.5S y]Level D ConditionP m  lesser of [2.4S m , 0.7S u]Paragraph F-1330 or F-1340, Appendix F(Faulted)P 1 + P b  lesser of [3.6S m , 1.05S u]Test ConditionsP m  0.9S y NB-3226 P m + P b  1.35S y when P m  0.67S y or  P m + P b  (2.15 S y - 1.2P m) when 0.67S y < P m  0.9S y Design ConditionP m  S m NB-3221 MPS2 UFSAR3.2-6Rev. 35Where P m = General primary membrane stress intensity P 1 = Primary local membrane stress intensity P 1 + P b = Primary membrane plus bending stress intensity P 1 + P b + Q = Primary plus sec ondary stress intensity S m = Design stress intensity S y = Yield strength S u = Tensile strength U = Cumulative fatigue usage factor
Design Limits Reactor internal components are designed to ensure that the stress levels and deflections are within an acceptable range. The stress values for core support structures are not greater than those given in the May 1972 draft of Section III of the ASME Boiler and Pressure Vessel Code, Subsection NG, including Appendix F, Rules for Evaluation of Faulted Conditions. Stress limits for the reactor vessel core support structures are presented in Table 3.2-1. In addition, to properly perform their functions, the reactor internal structures will satisfy the deformation limits listed below.
: 1.     Under design loadings plus operating basis earthquake forces or normal operating loadings plus SSE forces, deflections will be limited so that the CEAs can function and adequate core cooling is preserved.
: 2.     Under normal operating loadings plus SSE forces plus pipe rupture loadings resulting from a break of the largest line connect to the primary system piping, deflections will be limited so that the core will be held in place, adequate core cooling is preserved, and all CEAs can be inserted. Those deflections which would influence CEA movement will be limited to less than 80 percent of the deflections required to prevent CEA insertion.
: 3.     Under normal operating loadings plus SSE forces plus the maximum pipe rupture loadings resulting from the full spectrum of pipe breaks, deflections will be limited so that the core will be held in place and adequate core cooling is preserved.
Although CEA insertion is not required for a safe and orderly shutdown for break sizes greater than the largest line connected to the primary system piping, calculations show that the CEAs will be insertable for larger breaks except for a 3.2-4                                     Rev. 35


====3.2.2 NUCLEAR====
1.5    CEDM/RVLMS (HJTC) Pressure Housing Design Bases control element drive mechanism and Reactor Vessel Level Monitoring System (RVLMS) sure housings form part of the reactor coolant boundary and are, therefore, designed to meet stress requirements consistent with those of the reactor vessel closure head. The limiting sses in the CEDMs and RVLMS pressure boundary components due to the design, Level A, el B, Level C, Level D and Test conditions satisfy ASME Boiler Pressure Vessel Code, tion III, Subsection NB plus Appendix 1 and Section II, Part D, 1998 Edition through 2000 enda, including Code Case N-4-12 for the CEDM motor housing material.
DESIGN BASES The initial full power thermal rati ng of the core is 2700 MWt. It is upon this power level that the physics and thermal and hydraulic information presented in this section are based. The design basis for the nuclear design of the fuel and reactivity control systems are: a.Excess Reactivity and Fuel BurnupThe excess reactivity provided for each cycle is based on the depletion characteristics of the fuel and burnabl e poison and the desired burnup for each cycle. The desired burnup is based on the ec onomic analysis of both the fuel cost and the projected operating load demand cycle for the plant. The average burnup in
CEDMs and the RVLMs are designed to function normally during and after exposure to mal operating conditions plus the design basis earthquake (DBE). Under normal operating ditions, plus DBE, plus pipe rupture loadings, deflections of the CEDM will be limited so that CEAs can be inserted after exposure to these conditions. Those deflections, which could uence CEA movement, will be limited to less than 80 percent of the deflections required to vent CEA movement. The RVLMS and the adjacent CEDMs do not contact each other with imum lateral displacement of the pressure housings.
Loading Combinations                            ASME Code Subsection sign Condition      Pm  Sm                                    NB-3221 P1  1.5Sm P1 + Pb < 1.5Sm vel A and Level B P1 + Pb + Q  3Sm                              NB-3222 and NB3223 ormal and Upset)    U1 vel C Condition      Pm  greater of [1.2Sm, Sy]                NB-3224 mergency)            P1 + Pb  greater of [1.8Sm, 1.5Sy]
vel D Condition      Pm  lesser of [2.4Sm, 0.7Su]              Paragraph F-1330 or F-1340, Appendix F ulted)              P1 + Pb  lesser of [3.6Sm, 1.05Su]
t Conditions        Pm  0.9Sy                                  NB-3226 Pm + Pb  1.35Sy when Pm  0.67Sy or Pm + Pb  (2.15 Sy - 1.2Pm) when 0.67Sy
                    < Pm  0.9Sy sign Condition      Pm  Sm                                    NB-3221 3.2-5                                  Rev. 35


the core is chosen so as to insure that the peak assembly burnup is not greater than 56,000 MWD/MTU for Batch N, 52,500 MWD/MTU for Batch P, and 57,400  
ar Stress              0.6Sm                                      NB-3227.2 ere Pm = General primary membrane stress intensity P1 = Primary local membrane stress intensity P1 + Pb = Primary membrane plus bending stress intensity P1 + Pb + Q = Primary plus secondary stress intensity Sm = Design stress intensity Sy = Yield strength Su = Tensile strength U = Cumulative fatigue usage factor 2    NUCLEAR DESIGN BASES initial full power thermal rating of the core is 2700 MWt. It is upon this power level that the sics and thermal and hydraulic information presented in this section are based. The design s for the nuclear design of the fuel and reactivity control systems are:
: a.      Excess Reactivity and Fuel Burnup The excess reactivity provided for each cycle is based on the depletion characteristics of the fuel and burnable poison and the desired burnup for each cycle. The desired burnup is based on the economic analysis of both the fuel cost and the projected operating load demand cycle for the plant. The average burnup in the core is chosen so as to insure that the peak assembly burnup is not greater than 56,000 MWD/MTU for Batch N, 52,500 MWD/MTU for Batch P, and 57,400 MWD/MTU for Batch R and later.
: b.      Core Design Lifetime and Fuel Replacement Program The core design lifetime and fuel replacement program are based on a three region core with approximately 40 percent of the fuel assemblies replaced at each refueling.
: c.      Negative Reactivity Feedback and Reactivity Coefficients The negative reactivity feedback provided by the design is based on the requirement of General Design Criterion (GDC) 11. In the power operating range, the inherent combined response of the reactivity feedback characteristics (fuel temperature coefficient (FTC), moderator temperature coefficient (MTC),
moderator void coefficient (MVC), and moderator pressure coefficient (MPC)) to an increase in reactor thermal power will be a decrease in reactivity.
3.2-6                                    Rev. 35


MWD/MTU for Batch R and later
The burnable poison reactivity worth provided in the design will be sufficient to ensure that moderator coefficients of reactivity have magnitudes and algebraic signs consistent with the requirements for negative reactivity feedback and acceptable consequence in the event of postulated accidents or anticipated operational occurrences, viewed in conjunction with the supplied protective equipment.
.b.Core Design Lifetime and Fuel Replacement ProgramThe core design lifetime and fuel repl acement program are based on a th ree region core with approximately 40 percent of the fuel assemblies replaced at each refueling.c.Negative Reactivity Feedback and Reactivity Coefficients The negative reactivity feedback provided by the design is based on the requirement of General Design Criterion (GDC) 11. In the power operating range, the inherent combined response of the reactivity feedback characteristics (fuel temperature coefficient (FTC), moderator temperature coefficient (MTC), moderator void coefficient (MVC), and moderator pressure coefficient (MPC)) to
: e. Stability Criteria The design of the reactor and the instrumentation and control systems is based on meeting the requirements of GDC 12 with respect to spatial oscillations and stability. Sufficient CEA rod worth will be available to suppress xenon-induced power oscillations.
: f. Maximum Controlled Reactivity Insertion Rates The maximum reactivity addition rates are limited by core, CEA, and reactor regulating system (RRS) design based on preventing increases in reactivity which would result in the violation of specified acceptable fuel design limits, damage to the reactor pressure boundary, or disruption of the core or other internals sufficient to impair the effectiveness of emergency core cooling.
: g. Power Distribution Control Acceptable operation of the reactor in the absence of an accidental transient depends on maintaining a relationship among many parameters, some of which depend on the power distribution. In the absence of an accidental transient the power distribution is controlled such that in conjunction with other controlled parameters, limiting conditions of operation (LCO) are not violated. LCO are not less conservative than the initial conditions used in the accident analyses in Chapter 14. LCO and limiting safety system settings (LSSS) are determined such that specified acceptable fuel design limits are not violated as a result of anticipated operational occurrences and such that specified predicted acceptable consequence are not exceeded for other postulated accidents.
: h. Shutdown Margins and Stuck Rod Criteria The amount of reactivity available from insertion of withdrawn CEAs is required to be sufficient, under all power operating conditions, to ensure that the reactor can be brought to at least 3.6 percent  subcritical from the existing condition, including the effects of cooldown to an average coolant temperature of 532&deg;F, even when the highest worth CEA fails to insert. This criteria is exclusive of any safety allowance and is consistent with the most pessimistic analysis in Chapter 14.
3.2-7                                      Rev. 35


an increase in reactor thermal power will be a decrease in reactivity.Shear Stress 0.6S m NB-3227.2Loading Combinations ASME Code Subsection MPS2 UFSAR3.2-7Rev. 35d.Burnable Poison Requirements The burnable poison reactivity worth provided in the design will be suff icient to ensure that moderator coefficients of reactivity have magnitudes and algebraic signs consistent with the requirements for negative reactivity feedback and acceptable consequence in the event of postulated accidents or anticipated operational occurrences, viewed in conjunction with the supplied protective equipment.e.Stability Criteria The design of the reactor and the instrume ntation and control systems is based on meeting the requirements of GDC 12 with respect to spatial oscillations and stability. Sufficient CEA rod worth will be availabl e to suppress xenon-induced power oscillations.f.Maximum Controlled Reactivity Insertion Rates The maximum reactivity addition rates are limited by core, CEA, and reactor regulating system (RRS) design based on pr eventing increases in reactivity which would result in the violation of specified acceptable fuel design limits, damage to the reactor pressure boundary, or disruption of the core or other internals sufficient to impair the effectiveness of emergency core cooling.g.Power Distribution Control Acceptable operation of the reactor in the absence of an accidental transient depends on maintaining a relationship among many parameters, some of which depend on the power distribution. In the ab sence of an accidental transient the power distribution is controlled such th at in conjunction with other controlled parameters, limiting conditi ons of operation (LCO) are not violated. LCO are not less conservative than the initial conditions used in the accident analyses in Chapter 14. LCO and limiting safety system settings (LSSS) are determined such that specified acceptable fuel design limits are not violated as a result of anticipated operational occurrences and such that specified predicted acceptable consequence are not exceeded for other postulated accidents.h.Shutdown Margins and Stuck Rod Criteria The amount of reactivity available from insertion of withdr awn CEAs is required to be sufficient, under all pow er operating conditions, to en sure that the reactor can be brought to at least 3.6 percent  subcritical from th e existing condition, including the effects of cooldown to an average coolant temperature of 5 32&deg;F, even when the highest worth CEA fails to in sert. This criteria is exclusive of any safety allowance and is consistent with the most pessimistic analysis in Chapter 14.
The chemical and volume control system (CVCS) (Section 9.2) is used to adjust dissolved boron concentration in the moderator. After a reactor shutdown, this system is able to compensate for the reactivity changes associated with xenon decay and reactor coolant temperature decrease to ambient temperature. It also provides adequate shutdown margin during refueling. This system also has the capability of controlling long term reactivity changes due to fuel burnup, and reactivity changes during xenon transients resulting from changes in reactor load independently of the CEAs. In particular, any xenon transient may be accommodated at any time in the fuel cycle.
MPS2 UFSAR3.2-8Rev. 35i.Chemical Shim Control The chemical and volume control system (CVCS) (Section 9.2) is used to adjust dissolved boron concentration in the moderator. After a reactor shutdown, this system is able to compensate for the reactivity changes as sociated with xenon decay and reactor coolant te mperature decrease to ambi ent temperature. It also provides adequate shutdown mar gin during refueling. This system also has the capability of controlling long term reactivity changes due to fuel burnup, and reactivity changes during xenon transients resulting from changes in reactor load independently of the CEAs. In particular, any xenon transient may be accommodated at any time in the fuel cycle.
3    THERMAL AND HYDRAULIC DESIGN BASIS idance of thermally induced fuel damage during normal steady state and anticipated transient ration is the principal thermal and hydraulic design basis. It is recognized that there is a small bability of limited fuel damage in certain unlikely accident situations discussed in Chapter 14.
following corollary design basis are established, but violation of them is not necessarily ivalent to fuel damage.
: a.      A limit corresponding to 95% probability with 95% confidence (Reference 3.2-1) is set on the departure from nucleate boiling ratio (DNBR) during normal operation and any anticipated transients as calculated according to the HTP correlation.
: b.      The peak temperature of the fuel will be less than the melting point during normal operation and anticipated transients.
reactor control and protection system will provide for automatic reactor trip or other ective action before these design limits are exceeded.
core hydraulic resistance was considered in establishing the operational limits curves vided in Figures 4.5-4 and 4.5-5, and the Low Temperature Overpressure Protection (LTOP) tem described in Section 7.4.8. As fuel design changes, effects on the flow resistance will be luated to determine the impact.
4    REFERENCES 1      EMF-92-153(P)(A) Rev. 1, HTP: Departure From Nucleate Boiling Correlation for High Thermal Performance Fuel, Siemens Power Corporation, January 2005.
3.2-8                                      Rev. 35


====3.2.3 THERMAL====
Operating Conditions    Stress Categories and Limits of Stress Intensities Normal and Upset      Figure NG 3221.1 including notes Emergency              Figure NG 3224.1 including notes Faulted                Appendix F, Rules for Evaluating Faulted Conditions 3.2-9                                  Rev. 35
AND HYDRAU LIC DESIGN BASISAvoidance of thermally induced fuel damage during normal steady state and anticipated transient operation is the principal thermal and hydraulic design basis. It is recognized that there is a small probability of limited fuel damage in certain unl ikely accident situations discussed in Chapter 14.
The following corollary design ba sis are established, but violati on of them is not necessarily equivalent to fuel damage.a.A limit corresponding to 95% proba bility with 95% confidence (Reference 3.2-1) is set on the departure from nucleat e boiling ratio (DNBR) during normal operation and any anticipated transients as calculated according to the HTP correlation.b.The peak temperature of the fuel will be less than the melti ng point during normal operation and anticipated transients.
The reactor control and protecti on system will provide for automatic reactor trip or other corrective action before thes e design limits are exceeded.
The core hydraulic resistance was considered in establishing the ope rational limits curves provided in Figures 4.5-4 and 4.5-5, and the Low Temperature Overpressure Protection (LTOP) System described in Section 7.4.8. As fuel design changes, effects on the flow resistance will be


evaluated to determine the impact.
reactor core and internals are shown in Figure 3.3-1. A cross section of the reactor core and rnals is shown in Figure 3.1-2. Mechanical design features of the reactor internals, the control ment drive mechanisms (CEDM) and the core are described below. Mechanical design meters are listed in Table 3.3-1.
3.
1    CORE MECHANICAL DESIGN core approximates a right circular cylinder with an equivalent diameter of 136 inches and an ve height of 136.7 inches. It is made up of Zircaloy-4 clad fuel rods containing slightly ched uranium in the form of sintered UO2 pellets and UO2-Gd2O3 pellets. The fuel rods are uped into 217 assemblies.
rt term reactivity control is provided by 73 control element assemblies (CEA). The CEAs are ded within the core by the guide tubes which are integral parts of the fuel assemblies.
1.1    AREVA Fuel Rod detailed fuel rod design (see Figure 3.3-1) establishes such parameters as pellet diameter and th, density, cladding-pellet diametral gap, fission gas plenum size, and rod pre-pressurization
: l. The design also considers effects and physical properties of fuel rod components which y with burnup.
integrity of the fuel rods is ensured by designing to prevent excessive fuel temperatures, essive internal rod gas pressures, and excessive cladding stresses and strains. This end is ieved by designing the fuel rods to satisfy the design criteria during normal operation and cipated operational occurrences over the fuel lifetime. For each design criteria, the ormance of the most limiting fuel rod shall not exceed the specified limits.
l rods are designed to function throughout the design life of the fuel based upon the reactor rating conditions designated below without loss of mechanical integrity, significant ensional distortion, or release of fuel or fission products.
assemblies were evaluated for a peak assembly burnup of 56,000 MDW/MTU for Batch N, 00 MWD/MTU for Batch P, and 57,400 MWD/MTU for Batch R and later.
Millstone Unit 2 Cycle 19 reload core included four fuel assemblies with natural uranium acement fuel rods with an anti-rotation feature designed to prevent spinning of the rod during rations. The four assemblies containing replacement rods, and the conditions under which were evaluated for use, are discussed in Section 3.3.1.3.1, "Design Summary".
1.1.1    Fuel Rod Mechanical Criteria cladding primary and secondary stresses shall meet the 1977 ASME Boiler and Pressure sel Code Section III (Reference 3.3-1) requirements summarized below:
3.3-1                                  Rev. 35


==2.4 REFERENCES==
Stress Intensity Limits (Parameter)                          Yield Strength  Ultimate Tensile Strength mary Membrane (Pm)                                      < 2/3 Sy        < 1/3 Su mary Membrane Plus Primary Bending (Pm + Pb) < 1.0 Sy                    < 0.5 Su mary Plus Secondary (P + Q)                              < 2.0 Sy        < 1.0 Su mary stresses are developed by the imposed loading which is necessary to satisfy the laws of ilibrium between external and internal forces and moments. The basic characteristic of a ary stress is that it is not self-limiting. If a primary stress exceeds the yield strength of the erial through the entire thickness, the prevention of failure is entirely dependent on the strain-dening properties of the material.
3.2-1EMF-92-153(P)(A) Rev. 1, "H TP: Departure From Nucleate Boiling Correlation for High Thermal Performance Fuel," Siem ens Power Corporation, January 2005.
ondary stresses are developed by the self constraint of a structure. It must satisfy an imposed in pattern rather than being in equilibrium with an external load. The basic characteristic of a ondary stress is that it is self-limiting. Local yielding and minor distortions can satisfy the ontinuity conditions due to thermal expansions which cause the stress to occur.
MPS2 UFSAR3.2-9Rev. 35TABLE 3.2-1  STRESS LIMITS FOR REACTOR VESSEL INTERNAL STRUCTURESOperating ConditionsStress Categories and Limits of Stress Intensities1.Normal and UpsetFigure NG 3221.1 including notes2.EmergencyFigure NG 3224.1 including notes3.FaultedAppendix F, Rules for Evaluating Faulted Conditions MPS2 UFSAR3.3-1Rev. 35
dding circumferential strain shall not exceed the design limit through end-of-life (EOL).
total uniform strain, elastic and plastic shall not exceed the design limit during a transient.
strain analysis was performed with the RODEX2 (Reference 3.3-2) RAMPEX codes chmarked to available power ramp test data, i.e., INTERRAMP, OVERRAMP, and PERRAMP.
fuel rod shall be designed such that at a rod average burnup when substantial axial solidation has occurred, the total clad creep deformation shall not exceed the initial minimum metral fuel cladding gap. This will prevent pellet hangups allowing the plenum spring to close l gaps until densification is substantially complete, thus preventing the formation of pellet mn gaps of sufficient size for clad flattening.
fuel rod pressure at EOL shall not exceed the criteria approved by the NRC (Ref. 3.3-3). A ew of departure from nucleate boiling ratio (DNBR) limits for condition III or IV postulated dents events is required for fuel rods that exceed nominal system pressure. When fuel rod sure is predicted to exceed system pressure, the pellet-cladding gap shall not increase for dy or increasing power conditions. Analysis approved by the NRC has shown that the fuel rod pressure can safely exceed system pressure without causing any damage to the cladding.
al cladding wall thinning due to generalized external and internal corrosion shall not exceed a e which will impair mechanical performance over the projected fuel rod design lifetime under most adverse projected power conditions within coolant chemistry limits recommendations of le 3.3-2. It will also assure that the metal/oxide interface temperature will remain well below 3.3-2                                    Rev. 35


===3.3 MECHANICAL===
cumulative usage factor for cyclic stresses for all important cyclic loading conditions shall exceed the design limit.
DESIGN The reactor core and internals are shown in Figure 3.3-1. A cross section of the reactor core and internals is shown in Figure 3.1-2. Mechanical design features of the reactor internals, the control element drive mechanisms (CEDM) and the core are described below. Mechanical design parameters are listed in Table 3.3-1.
clearance between the upper and lower tie plate shall be able to accommodate the maximum erential fuel rod and fuel assembly growth to the designed burnup.
3.3.1 CORE MECHANICAL DESIGN The core approximates a right circular cylinder with an equivalent diameter of 136 inches and an active height of 136.7 in ches. It is made up of Zircaloy-4 clad fuel r ods containing slightly enriched uranium in the form of sintered UO 2 pellets and UO 2-Gd 2 O 3 pellets. The fuel rods are grouped into 217 assemblies.
centerline temperature of the hottest pellet shall be below the melting temperature. Fuel terline temperature is calculated at overpower conditions to verify that fuel pellet overheating s not occur during normal operation and anticipated operational occurrences.
1.1.2    Fuel Rod Design Analyses h design analysis was performed with AREVA methodology which involves a well defined ction of appropriate data and parameters, and the latest approved versions of computer codes.
s methodology, as required, has been submitted to the Nuclear Regulatory Commission (NRC) approved. The analysis is performed in accordance with the methods described in AREVAs alification of Exxon Nuclear Fuel For Extended Burnup (Reference 3.3-3).
cladding steady state stress analysis was performed by considering primary and secondary mbrane and bending stresses due to hydrostatic pressure, flow-induced vibration, spacer tact, pellet cladding interaction (PCI), thermal and mechanical bow and thermal gradients.
sses were calculated for the various combinations of the following conditions:
: a.      beginning of life (BOL) and EOL
: b.      cold and hot conditions
: c.     at mid-span and at spacer locations
: d.      at both the inner and outer surfaces of the cladding analysis was performed for the various sources of stress, including pressure, thermal, spacer tact, PCI, and rod bow. The applicable stresses at each orthogonal direction were combined to ulate the maximum stress intensities which are compared to the ASME design criteria. The lts of the analysis indicate that all stress values are within acceptable design limits for both L and EOL, hot and cold conditions. The EOL stresses have ample margin for both the hot and condition stresses.
cladding steady state strain is evaluated with the RODEX2 code, which has been approved by NRC (Reference 3.3-2). The code considers the thermal-hydraulic environment at the ding surface, the pressure inside the cladding, and the thermal, mechanical and compositional e of the fuel and cladding. Pellet density, swelling, densification, and fission gas release or orption models, and cladding and pellet diameters are input to RODEX2 to provide the most 3.3-3                                    Rev. 35
: a.      Radial Thermal Conduction and Gap Conductance
: b.      Fuel Swelling, Densification, Cracking, and Crack Healing
: c.      Gas Release and Absorption
: d.      Cladding Creep Deformation and Irradiation-Induced Growth
: e.     Cladding Corrosion
: f.     PCI
: g.     Free Rod Volume and Gas Pressure calculations are performed on a time incremental basis with conditions updated at each ulated increment so that the power history and path dependent processes can be modeled. The l dependence of the power and burnup distributions are handled by dividing the fuel rod into a ber of axial and radial regions. Power distributions can be changed at any desired time, and coolant and cladding temperatures are readjusted in all the regions. All the performance dels, e.g., giving the deformations of the fuel and cladding and gas release, are calculated at cessive times during each period of assumed constant power generation. The calculated ding strain is reviewed throughout the life of the fuel and both the maximum circumferential in and the maximum strain increment are compared with the design criteria. The calculated in did not exceed the strain limit. Both the maximum strain and the positive strain increment below the design limit strain.
ramping strain and the fatigue evaluation of the fuel rod were evaluated. The ramps are med to occur anytime during the irradiation and may reach the maximum peaking factor wed by the limits of operation. The ramps are analyzed either from cold shutdown or from a ety of hot powered starting conditions. The approach to rated power at the beginning of each tor cycle is performed to satisfy the AREVA maneuvering and conditioning mmendations. The clad response during ramping power changes is calculated with the MPEX code. This code calculates the PCI during a power ramp for one axial node at a time.
initial conditions are obtained from RODEX2 output. The RAMPEX code considers the mal condition of the rod in its flow channel, and the mechanical interactions that result from and cladding creep at any desired axial section in the rod during the power ramp. As pared to RODEX2, RAMPEX additionally models the pellet cladding axial stress interaction, mary creep with strain hardening, the effects of pellet chips, and localized stresses due to ing.
RAMPEX code provides the hoop stress and the stress intensity. The stress results of the ping analysis are used to evaluate the cladding fatigue damage through life due to the cyclic 3.3-4                                    Rev. 35


Short term reactivity control is provided by 73 control element assemblies (CEA). The CEAs are guided within the core by the guide tubes which are integral parts of the fuel assemblies.
MPEX over the power cycling range, are compared with this curve to determine the allowed les for each stress range. This result is combined with the projected number of duty cycles to rmine a fatigue usage factor. All of the reactor cycle (startup) ramp stresses were within the gn limit.
3.3.1.1 AREVA Fuel Rod The detailed fuel rod design (see Figure 3.3-1) establishes such parameters as pellet diameter and length, density, cladding-pellet di ametral gap, fission gas plenum size, and rod pre-pressurization level. The design also considers effects and physical properties of fuel rod components which vary with burnup.
ep collapse calculations are performed with RODEX2 and COLAPX codes. The RODEX2 e determines the cladding temperature and internal pressure history based on a model which ounts for changes in fuel rod volumes, fuel densification and swelling, and fill gas absorption.
reactor coolant, fuel rod internal temperature, and pressure histories generated by the DEX2 analysis are input to the COLAPX code along with a conservative statistical estimate of al cladding ovality and the fast flux history. The COLAPX code calculates, by large deflection ry, the ovality of the cladding as a function of time while the uniform cladding creepdown is ined by the RODEX2 analysis. The cladding ovality increase and creepdown are summed, at d average burnup when substantial axial consolidation has occurred, to show that they remain than the initial minimum pellet clad gap. Measurements of highly densifying irradiated fuel e demonstrated that pellet densification is essentially complete by the time the fuel has ined this burnup so that further creepdown after this phase will not result in significant pellet ellet gaps. The combined radial creepdown was shown to meet the design criteria. This will vent pellet hangups due to cladding creep, allowing the plenum spring to close axial gaps until sification is substantially complete, and thus assures that clad collapse will not occur. The h of the plenum spring is less than the spacing calculated for stiffening rings in a cylindrical l under external pressure which will prevent clad collapse in the plenum area.
culation of the gas pressure within a fuel rod is performed with the RODEX2 code. The initial gas is found by calculating the initial free volume and using the ideal gas law, along with input es for fill gas pressure and reference fill gas temperature. The free gaseous fission product d is calculated for each axial region and the total yield obtained by summing the axial region tributions. The power of each history used was multiplied for each cycle by a factor required the highest projected rod power to reach the Fr limit plus uncertainties. The calculations show for all power histories analyzed, the rod internal gas pressure will remain below the criteria roved by the NRC (Reference 3.3-3) for use in extended burnup gas pressure analysis.
waterside corrosion of fuel rods is evaluated with the MATPRO-11 (Reference 3.3-5) elation. The MATPRO-11 model is a two-stage corrosion rate model which is cubic in endence on oxide thickness until a transition to a subsequent linear dependence occurs. To ulate the rate changes as a function of both oxide thickness and the operating conditions of the rod, the MATPRO model is incorporated into AREVAs RODEX2 fuel performance code.
RODEX2 code determines the temperature increase of the water along the fuel rod assuming t balance within a channel for the prescribed mass flow and inlet temperature. The radial perature drops are evaluated successively between the water, the oxide surface, the metal/
de interface, and the inside of the cladding using RODEX2 correlations and methods. To ount for the change in corrosion rate due to the changing oxide layer and thermal conditions, code includes an update in cladding temperature at every calculation step. This is an iterative cess due to the continuously changing oxide thickness. Conditions are also revised at times 3.3-5                                  Rev. 35


The integrity of the fuel rods is ensured by designing to prevent excessive fuel temperatures, excessive internal rod gas pressures, and excessi ve cladding stresses and strains. This end is achieved by designing the fuel r ods to satisfy the design crit eria during norma l operation and anticipated operational occurrences over the fuel lifetime. Fo r each design criteria, the performance of the most lim iting fuel rod shall not exceed the specified limits.
mblies in seven separate reactors. Each data point represents the maximum thickness sured along a rod length. The enhancement factor is based on a best fit regression analysis of data. A final multiplier is also applied which envelopes the data. The waterside corrosion in cladding was evaluated with RODEX2 for the steady state strain analysis. A best-fit corrosion lification factor was applied to the MATPRO model along with a final multiplier to bound the sured data on AREVA standard cladding. The maximum calculated oxide thickness was w the design limit.
Fuel rods are designed to function throughout the design life of the fuel based upon the reactor operating conditions designated below without loss of mechanical integrity, significant dimensional distortion, or releas e of fuel or fission products.
l rod and fuel assembly growth projected to occur during irradiation was based on servative design curves established from measured irradiation growth data. The rod growth us the assembly growth plus tolerances was compared with the clearance within the assembly fuel rod growth. Differential thermal expansion between the fuel rods and guide tubes was considered. There is space between the upper and lower tie plates to accommodate the imum differential growth out to a rod burnup of 62,000 MWd/MTU.
The assemblies were evaluated for a peak assembly burnup of 56,000 MDW/MTU for Batch N, 52,500 MWD/MTU for Batch P, and 57,400 MWD/MTU for Batch R and later.The Millstone Unit 2 Cycle 19 reload core included four fuel assemblies with natural uranium replacement fuel rods with an anti-rotation feature designed to pr event spinning of the rod during operations. The four assemblies containing replacement rods, and the conditions under which they were evaluated for use, are discussed in Section 3.3.1.3.1, "Design Summary".
pellet centerline temperature calculation was performed with the RODEX2 code. Fuel pellet terline temperatures were calculated at overpower conditions. The high power cycle of each er history was modified to include a spike in each cycle. This spike increased the maximum er of a pellet in the rod up to FTQ. Pellet melting temperature is a function of burnup.
3.3.1.1.1 Fuel Rod Mechanical Criteria The cladding primary and secondary stresses sh all meet the 1977 ASME Boiler and Pressure Vessel Code Section III (Reference 3
sidering a conservative peak pellet burnup to determine the minimum pellet melting perature at EOL, the maximum pellet centerline temperature is well below both BOL and L limits.
.3-1) requirements summarized below:
1.2    (Deleted) 1.3    AREVA Fuel Assembly 1.3.1    Design Summary AREVA fuel assemblies are 14 by 14 arrays containing 176 fuel rods in a cage structure of 5 de tubes and 9 spacer grids. Both the guide tubes and the fuel rod cladding are made of aloy-4 for low neutron absorption and high corrosion resistance. The fuel assembly upper tie es are stainless steel castings with Inconel holddown springs. The fuel assembly upper tie e is mechanically locked to the guide tubes and may be easily removed to allow inspection of diated fuel rods. For Reload T (Cycle 15) and beyond, lower tie plates are the ELGUARD' debris resistant design.
MPS2 UFSAR3.3-2Rev. 35 Primary stresses are deve loped by the imposed loading which is necessary to satisfy the laws of equilibrium between external and internal forces and moments. The basic characteristic 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 prevention of failure is entirely dependent on the strain-hardening properties of the material.
eloads M, N, and P (Cycles 10-12), eight of the nine spacers in each fuel assembly are made Zircaloy-4 structure with Inconel-718 springs (i.e., bi-metallic spacer). The ninth spacer, ted just above the lower tie plate, is made of Inconel-718 and, using features of the AREVA h Thermal Performance (HTP) spacer design, has been adapted to provide fuel assembly ris resistance.
Secondary stresses are developed by th e self constraint of a structur
3.3-6                                    Rev. 35
: e. 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 yi elding and minor distor tions can satisfy the discontinuity conditions due to thermal expansions which cause the stress to occur.Cladding circumferential strain shall not exceed the design limit through end-of-life (EOL).The total uniform strain, elastic and plastic shal l not exceed the design limit during a transient.
The strain analysis was performed with th e RODEX2 (Reference 3.3-2) RAMPEX codes benchmarked to available power ramp test data, i.e., INTERRAMP, OVERRAMP, and SUPERRAMP.


The fuel rod shall be designed such that at a rod average burnup when substantial axial consolidation has occurred, the total clad creep deformation shall not exceed the initial minimum diametral fuel cladding gap. This will prevent pellet hangups allowi ng the plenum spring to close axial gaps until densification is substantially co mplete, thus preventing the formation of pellet column gaps of sufficient size for clad flattening.
ger end cap serves to raise the fuel rod cladding above the debris trapping region of the ninth tom) spacer.
Reloads T through X (Cycles 15-18), the High Thermal Performance (HTP) fuel assembly gn was implemented in which all nine spacers are of the Zircaloy-4 HTP design. This design ined the longer, solid fuel rod lower end cap.
fuel assembly design for Reload Y (Cycle 19) and later utilized eight Zircaloy-4 HTP spacers replaces the ninth, bottom spacer with an Inconel High Mechanical Performance (HMP) cer. The HMP spacer is similar to the HTP spacer, except that it is constructed of Inconel-718 the flow channels are parallel to the fuel.
wings of the AREVA fuel assemblies are given in Figure 3.3-2A and Figure 3.3-3A. Fuel mbly drawings for Reload T (Cycle 15) and beyond are included in Figures 3.3-2B and 3.3-analysis has shown that the AREVA reload fuel assemblies will meet the design criteria:
: a. The maximum steady state cladding strain is well below the design limit.
: b.      The maximum steady state cladding stress meets the ASME Boiler and Pressure Vessel Code requirements.
: c.      The transient strain is within the circumferential limit.
: d.      Cladding creep collapse is precluded.
: e.      The fuel rod internal pressure at the EOL remains below the criteria approved by the NRC (Ref. 3.3-3).
: f.      The maximum clad oxidation is below the design limit.
: g.      The cladding fatigue usage factor is well below the design limit.
: h.     There is space between the upper and lower tie plate to accommodate fuel rod growth.
: i.      Pellet centerline temperatures remain below the design criteria.
: j.      The fuel assembly growth is within the space available between the upper and lower core plates in the reactor core.
: k. The assembly holddown springs will prevent bundle lift-off.
3.3-7                                  Rev. 35


The fuel rod pressure at EOL shall not exceed the criteria appr oved by the NRC (Ref. 3.3-3). A review of departure from nucleate boiling ratio (DNBR) limits fo r condition III or IV postulated accidents events is required for fuel rods that exceed nominal system pressure. When fuel rod pressure is predicted to exceed system pressu re, the pellet-cladding gap shall not increase for steady or increasing power conditi ons. Analysis approved by the NRC has shown that the fuel rod gas pressure can safely exceed system pressu re without causing any da mage to the cladding.Total cladding wall thinning due to generalized external and internal corrosion shall not exceed a value which will impair mechanic al performance over the projected fuel rod design lifetime under the most adverse projected power conditions within coolant ch emistry limits recommendations of Table 3.3-2. It will also assure that the metal/oxide interface temperature will re main well below Stress Intensity Limits (Parameter)Yield StrengthUltimate Tensile Strength Primary Membrane (P m)< 2/3 S y< 1/3 S u Primary Membrane Plus Primary Bending (P m + P b)< 1.0 S y< 0.5 S u Primary Plus Secondary (P + Q)
aloy-4 tubular cladding. Zircaloy-4 end caps are welded to each end to give a hermetic seal.
< 2.0 S y< 1.0 S u MPS2 UFSAR3.3-3Rev. 35the level where large increases in corrosion, due to the insulating eff ect of the oxide, would adversely affect the mechanic al behavior of the cladding.The cumulative usage factor for cyclic stresses for all important cyclic loading conditions shall not exceed the design limit.
fuel rod upper plenum contains a high strength alloy compression spring to prevent fuel mn separation during fabrication and shipping, and during incore operation. The rods are surized with helium to improve heat transfer and reduce clad creep ovality.
The clearance between the upper and lower tie plate shall be able to accommodate the maximum differential fuel rod and fuel assembly growth to the designed burnup.
fuel assembly structure consists of an upper tie plate assembly, lower tie plate, guide tubes spacer grids, which together provide the support for the fuel rods.
The centerline temperature of th e hottest pellet shall be below the melting temperature. Fuel centerline temperature is calculated at overpower conditions to verify that fuel pellet overheating does not occur during normal operation and anticipated operational occurrences.
lower tie plate is a machined stainless steel casting which provides the lower end support for guide tubes. The Zircaloy guide tubes are attached to the lower tie plate by means of Inconel screws. The FUELGUARDTM lower tie plate, included in Reload T and beyond provides ection to the fuel from debris in the primary coolant.
3.3.1.1.2 Fuel Rod Design Analyses Each design analysis was performed with AREVA methodology which invol ves a well defined selection of appropria te data and parameters, a nd the latest approved versio ns of computer codes.
upper tie plate assembly latches to and provides the upper end support for the guide tubes.
This methodology , as required, ha s been submitted to the Nucl ear Regulatory Commission (NRC) and approved. The analysis is performed in accordance with the methods described in AREVA's "Qualification of Exxon Nuclear Fuel Fo r Extended Burnup" (Reference 3.3-3).The cladding steady state stress analysis was performed by cons idering primary and secondary membrane and bending stresses due to hydrostatic pressure, flow-induced vibration, spacer contact, pellet cladding interaction (PCI), thermal and mechanic al bow and thermal gradients. Stresses were calculated for the various combinations of the following conditions:a.beginning of life (BOL) and EOLb.cold and hot conditionsc.at mid-span and at spacer locationsd.at both the inner and outer surfaces of the claddingThe analysis was performed for the various sources of stress, in cluding pressure, thermal, s pacer contact, PCI, and rod bow. The app licable stresses at each orthogonal direction were combined to calculate the maximum stress intensities which are compared to the ASME design criteria. The results of the analysis indicate that all stress values are within acceptable design limits for both BOL and EOL, hot and cold condi tions. The EOL stresses have ample margin for both the hot and cold condition stresses.
upper tie plate assembly consists of a stainless steel grid structure and reaction plate taining five Inconel X-750 holddown springs. The springs are located around Inconel locking and sleeves which mechanically attach to the guide tubes and pilot into the reactor alignment
The cladding steady state strain is evaluated w ith the RODEX2 code, which has been approved by the NRC (Reference 3.3-2). The code consider s the thermal-hydraulic environment at the cladding surface, the pressure insi de the cladding, and the thermal, mechanical and compositional state of the fuel and cladding. Pellet density, sw elling, densification, and fission gas release or absorption models, and cladding a nd pellet diameters ar e input to RODEX2 to provide the most MPS2 UFSAR3.3-4Rev. 35conservative strain calculation or subsequent ramping or collapse calculations for the reference fuel rod design. The major fuel rod performanc e characteristics modeled by the RODEX2 code are:a.Radial Thermal Conduc tion and Gap Conductanceb.Fuel Swelling, Densification, Cracking, and Crack Healingc.Gas Release and Absorption d.Cladding Creep Deformation and Irradiation-Induced Growthe.Cladding Corrosionf.PCI g.Free Rod Volume and Gas Pressure The calculations are performed on a time incremental basis with conditions updated at each calculated increment so that th e power history and path dependent processes can be modeled. The axial dependence of the power a nd burnup distributions are handled by dividing the fuel rod into a number of axial and radial regions. Power distributions can be ch anged at any desired time, and the coolant and cladding temperatures are readjusted in all the region
: e. The springs are partially shrouded on the outside diameter by stainless steel cups to prevent induced spring vibration.
: s. All the performance models, e.g., giving the defo rmations of the fuel and cladding a nd gas release, ar e calculated at successive times during each period of assume d constant power generation. The calculated cladding strain is reviewed throughout the life of the fuel and both the maximum circumferential strain and the maximum strain increment are comp ared with the design criteria. The calculated strain did not exceed the strain limit. Both the maximum strain and the positive strain increment are below the design limit strain.
guide tubes, in conjunction with the spacers and tie plates, form the structural skeleton of the assembly and provide channels for insertion of the control rods. The guide tubes are icated from Zircaloy-4 tubing and are fully annealed. The center tube is of uniform diameter reas the outer four guide tubes have a reduced diameter section at the bottom which produces shpot action to decelerate the dropped CEAs.
The ramping strain and the fati gue evaluation of the fuel rod were evaluated. The ramps are assumed to occur anytime duri ng the irradiation and may reac h the maximum peaking factor allowed by the limits of operation. The ramps ar e analyzed either from cold shutdown or from a variety of hot powered st arting conditions. The approach to ra ted power at the beginning of each reactor cycle is performed to satisfy the AREVA maneuvering and conditioning recommendations. The clad response during ramp ing power changes is calculated with the RAMPEX code. This code calculates the PCI duri ng a power ramp for one axial node at a time.
end plug is welded to the lower end of the guide tube and is drilled and threaded to accept the er cap screws. At the upper end, the guide tube is crimped into an external stainless steel ing sleeve which engages the upper tie plate assembly. The upper tie plate assembly is locked he guide tube end fittings and can be unlocked for reconstitution or for fuel examination using cial tools.
The initial conditions are obtai ned from RODEX2 output. The RAMPEX code considers the thermal condition of the rod in its flow channel, and the mechanical interactions that result from fuel and cladding creep at any desired axial section in the rod during the power ramp. As compared to RODEX2, RAMPEX a dditionally models the pellet cl adding axial stress interaction, primary creep with strain hardening, the effects of pellet chips, and localized stresses due to ridging.The RAMPEX code provides the hoop stress and the stress intensity. The stress results of the ramping analysis are used to ev aluate the cladding fatigue damage through life due to the cyclic MPS2 UFSAR3.3-5Rev. 35 power variations. The fatigue an alysis is based on the O'Donnel and Langer (Reference 3.3-4) design curve. The cyclic amplitudes of the maximum local stress intensity, as determined by RAMPEX over the power cycling range, are compared with this curve to determine the allowed cycles for each stress range. This result is combined with the projected number of duty cycles to determine a fatigue usage factor.
ainless steel sleeve assembly with a chrome plated inside diameter is inserted in the top end of guide tube assembly. This sleeve protects the guide tube from control rod fretting and wear n the rod is in the withdrawn/ready position. The sleeve is mechanically captured by the upper late.
All of the reactor cycle (startup) ramp stresses were within the design limit.
l rod pitch and position is maintained by nine spacer grids. The spacers are axially positioned hat the assemblies will be compatible with existing fuel assemblies.
Creep collapse calculations are performed with RODEX2 a nd COLAPX codes. The RODEX2 code determines the cladding temperature and in ternal pressure history based on a model which accounts for changes in fuel r od volumes, fuel densification a nd swelling, and fill gas absorption.
bi-metallic spacers used in Reloads M through S (Cycles 10-14) are formed by an rlocking rectangular grid of Zircaloy-4 structural strips (see Figure 3.3-4A). Inconel-718 ng strips are mechanically secured within these strips. The Zircaloy-4 structural strips are ded at all intersections and to the side plates. Dimples formed in the structural strips center the 3.3-8                                    Rev. 35
The reactor coolant, fuel rod internal temperature, and pressure histories generated by the RODEX2 analysis are input to th e COLAPX code along with a cons ervative statisti cal estimate of initial cladding ovality and the fast flux history. The COLAPX code calculates, by large deflection theory, the ovality of the claddi ng as a function of time while the uniform cladding creepdown is obtained by the RODEX2 an alysis. The cladding ovality increa se and creepdown are summed, at a rod average burnup when substantia l axial consolidation has occurre d, to show that they remain less than the initial minimum pell et clad gap. Measurem ents of highly densif ying irradiated fuel have demonstrated that pellet densification is essentially complete by the time the fuel has attained this burnup so that furthe r creepdown after this phase will not result in significant pellet to pellet gaps. The combined radial creepdown wa s shown to meet the desi gn criteria. This will prevent pellet hangups due to cla dding creep, allowing the plenum sp ring to close axial gaps until densification is substantially complete, and thus assures that clad collapse will not occur. The pitch of the plenum spring is less than the spacing calculated for stiffening rings in a cylindrical shell under external pressure which will pr event clad collapse in the plenum area.
Calculation of the gas pressure within a fuel r od is performed with the RODEX2 code. The initial fill gas is found by calcul ating the initial free volume and using the ideal gas law, along with input values for fill gas pressure a nd reference fill gas temperatur
: e. The free gaseous fission product yield is calculated for each ax ial region and the total yield obta ined by summing the axial region contributions. The power of each history used was multiplied fo r each cycle by a factor required for the highest projected rod power to reach the F r limit plus uncertainties. The calculations show that for all power histories analy zed, the rod internal gas pressure will remain below the criteria approved by the NRC (Reference 3.3-3) for us e in extended burnup gas pressure analysis.The waterside corrosion of fuel rods is evaluated with the MATPRO-11 (Reference 3.3-5) correlation. The MATPRO-11 mode l is a two-stage corrosion ra te model which is cubic in dependence on oxide thickness until a transition to a subsequent linear dependence occurs. To calculate the rate changes as a function of both oxide thickness a nd the operating conditions of the fuel rod, the MATPRO model is incorporated into AREVA's RODEX2 fuel performance code.
The RODEX2 code determines the temperature incr ease of the water along the fuel rod assuming heat balance within a channel for the prescribed mass flow and inlet temperature. The radial temperature drops are evaluated successively between the water, the oxide surface, the metal/
oxide interface, and the inside of the cladding using RODEX2 correlations and methods. To account for the change in corros ion rate due to the changing oxi de layer and thermal conditions, the code includes an update in cl adding temperature at ev ery calculation step. This is an iterative process due to the continuously changing oxide thickness. Conditio ns are also revised at times MPS2 UFSAR3.3-6Rev. 35where new power or flow conditions are prescribed. The MATPRO model incorporated in RODEX2 is benchmarked via an overall enhanc ement factor to oxide thickness data from assemblies in seven separate reactors. Each data point represents the maximum thickness measured along a rod length. The enha ncement factor is based on a be st fit regression analysis of the data. A final multiplier is also applied which envelopes the data. The waterside corrosion in the cladding was evaluated with RODEX2 for the steady state strain analysis. A best-fit corrosion amplification factor was applied to the MATPRO model along with a final mu ltiplier to bound the measured data on AREVA standard cladding.
The maximum calculated oxide thickness was below the design limit.
Fuel rod and fuel assembly growth projected to occur during irradiation was based on conservative design curves establ ished from measured irradiati on growth data. The rod growth minus the assembly growth plus tolerances was compared with the clearance within the assembly for fuel rod growth. Differential thermal expansion between the fuel rods and guide tubes was also considered. There is space between the uppe r and lower tie plates to accommodate the maximum differential growth out to a rod burnup of 62,000 MWd/MTU.The pellet centerline temperatur e calculation was performed with the RODEX2 code. Fuel pellet centerline temperatures were calculated at overpower conditions. The high power cycle of each power history was modified to include a spike in each cycle. This spike increased the maximum power of a pellet in the rod up to F T Q. Pellet melting temperature is a function of burnup.
Considering a conservative peak pellet bur nup to determine the minimum pellet melting temperature at EOL, the maximu m pellet centerline temperatur e is well below both BOL and EOL limits.


3.3.1.2 (Deleted)3.3.1.3 AREVA Fuel Assembly 3.3.1.3.1 Design SummaryThe AREVA fuel assemblies are 14 by 14 arrays containing 176 fuel rods in a cage structure of 5 guide tubes and 9 spacer grids. Both the guide tubes and the fuel r od cladding are made of Zircaloy-4 for low neutron absorp tion and high corrosion resistance.
eloads M, N, and P (Cycles 10-12), the debris resistant Inconel HTP spacer grid in the ninth, om location is located just above the lower tie plate. It is formed by an interlocking angular grid of Inconel-718 strips. The strips are welded at all intersections and to the side es. The spacer is positioned on top of the lower tie plate with the strip intersections directly ve the tie plate flow holes. This reduces the size of debris that may pass through the flow holes eby reducing the possibility of fretting against the cladding. Reloads R and S (Cycles 13 and use a similar debris resistant concept with the Inconel HTP spacer replaced by a bimetallic cer coupled with a longer lower end cap on the fuel rods.
The fuel assembly upper tie plates are stainless steel cast ings with Inconel holddown spri ngs. The fuel assembly upper tie plate is mechanically locked to the guide tubes and may be easil y removed to allo w inspection of irradiated fuel rods. For Reload T (Cycle
HTP spacers for Reloads T through X (Cycles 15-18) are all Zircaloy-4 (Figure 3.3-4B). The ps are welded at the intersections and side plates. The structure of the Zircaloy-4 strips vides the rod support.
: 15) and beyond, lower tie plates are the FUEL GUARD debris resistant design.
Reload Y (Cycle 19) and later, all Zircaloy-4 HTP spacers are used in eight locations. The onel-718 HMP spacer is used in the ninth, bottom location. The Inconel-718 HMP bottom cer is similar in design to the HTP spacers except for the flow channels, which are not canted.
In Reloads M, N, and P (Cycles 10-12), eight of the nine spacers in each fuel assembly are made of a Zircaloy-4 structure with Inconel-718 springs (i.e., bi-metallic spacer). The ninth spacer, located just above the lower tie plate, is made of Inconel-718 and, using features of the AREVA High Thermal Performance (HTP) spacer design, has been adapted to provide fuel assembly debris resistance.
Millstone Unit 2 Cycle 19 reload core included four fuel assemblies with natural uranium acement fuel rods with an anti-rotation feature designed to prevent spinning of the rod during rations. The four assemblies containing replacement rods were installed in symmetric, pheral core locations against the baffle as shown in Figure 3.3-19 (Reference 3.3-9). The core tions into which the assemblies were placed where P-1, A-8, H-21, and Y-14 (see ure 3.4-1). The replacement rods installed under these conditions were evaluated against blished mechanical, nuclear, and thermal/hydraulic design criteria for Millstone Unit 2 fuel, were determined to be compliant with their design and licensing bases (Reference 3.3-10).
MPS2 UFSAR3.3-7Rev. 35The fuel assembly design for Re loads R and S (Cycles 13 and 14) has all nine spacers of the bimetallic design. Additionally, in this design a longer solid fu el rod lower end cap is used. The longer end cap serves to raise th e fuel rod cladding above the debr is trapping regi on of the ninth (bottom) spacer.
1.3.2    Fuel Assembly Mechanical Criteria structural integrity of the fuel assemblies is assured by setting design limits on stresses and ormations due to various handling operational and accident loads. These limits are applied to design and evaluation of upper and lower tie plates, grid spacers, guide tubes, holddown ngs, and locking hardware.
In Reloads T through X (Cycles 15-18), the Hi gh Thermal Performance (HTP) fuel assembly design was implemented in which all nine spacers are of the Zi rcaloy-4 HTP desi gn. This design retained the longer, solid fuel rod lower end cap. The fuel assembly design for Relo ad Y (Cycle 19) and later utili zed eight Zircaloy-4 HTP spacers and replaces the ninth, bottom spacer with an Inconel High Mechanical Performance (HMP) spacer. The HMP spacer is similar to the HTP spacer, except that it is constructed of Inconel-718 and the flow channels are parallel to the fuel. Drawings of the AREVA fuel assemblies are given in Figure 3.3-2A and Figure 3.3-3A. Fuel assembly drawings for Reload T (Cycle 15) and beyond are in cluded in Figur es 3.3-2B and 3.3-3B.The analysis has shown that the AREVA reload fuel assemblies will meet the design criteria:a.The maximum steady state cladding strain is well below the design limit.b.The maximum steady state cladding stress meets the ASME Boiler and Pressure Vessel Code requirements.c.The transient strain is within the circumferential limit.
design bases for evaluating the structural integrity of the fuel assemblies are:
d.Cladding creep collapse is precluded.e.The fuel rod internal pressure at the EOL remains below the criteria approved by the NRC (Ref. 3.3-3).f.The maximum clad oxidation is below the design limit.
: a.     Fuel Assembly Handling - Dynamic axial loads approximately 2.5 times assembly weight.
g.The cladding fatigue usage factor is well below the design limit.h.There is space between the upper and lower tie plate to accommodate fuel rod growth.i.Pellet centerline temperatures remain below the design criteria.j.The fuel assembly growth is within the space available between the upper and lower core plates in the reactor core.
: b.     For All Applied Loads for Normal Operation and Anticipated Operational Events -
k.The assembly holddown springs will prevent bundle lift-off.
The fuel assembly component structural design criteria are established for the two primary material categories, austenitic stainless steels (tie plates), and Zircaloy (guide tubes, grids, spacer sleeves). The stress categories and strength theory for 3.3-9                                        Rev. 35
MPS2 UFSAR3.3-8Rev. 35The fuel rods consist of short cylindrical UO 2 pellets or UO 2-Gd 2 O 3 pellets contained in Zircaloy-4 tubular cladding. Zircaloy-4 end caps are welded to each end to give a hermetic seal.
The fuel rod upper plenum contains a high stre ngth alloy compression spring to prevent fuel column separation during fabric ation and shipping, and during in core operation.
The rods are pressurized with helium to improve heat transfer and reduce clad creep ovality.
The fuel assembly structure cons ists of an upper tie plate assembly, lower tie plate, guide tubes and spacer grids, which together pr ovide the support for the fuel rods.
The lower tie plate is a machined stainless steel castin g which provides the lo wer end support for the guide tubes. The Zircaloy guide tubes are attached to the lowe r tie plate by means of Inconel cap screws. The FUELGUARD TM lower tie plate, in cluded in Reload T and beyond provides protection to the fuel from de bris in the primary coolant.
The upper tie plate assembly latc hes to and provides the upper e nd support for the guide tubes. The upper tie plate assembly consists of a st ainless steel grid structure and reaction plate containing five Inconel X-750 ho lddown springs. The springs are located around Inconel locking nuts and sleeves which mechanically attach to the guide tubes and pilot into the reactor alignment plate. The springs are partially shrouded on the outside di ameter by stainless st eel cups to prevent flow induced spring vibration.


The guide tubes, in conjunction wi th the spacers and tie plates, form the structural skeleton of the fuel assembly and provide cha nnels for insertion of the cont rol rods. The guide tubes are fabricated from Zircaloy-4 tubing and are fully annealed. The center tube is of uniform diameter whereas the outer four gui de tubes have a reduced diameter section at the bot tom which produces a dashpot action to decelerate the dropped CEAs.
Steady state stress limits are given in FSAR Section 3.3.1.1.1. Stress nomenclature is per the ASME Boiler and Pressure Vessel Code, Section III.
: c.      Loads During Postulated Accidents - Deflection or failure of components shall not interfere with reactor shutdown or emergency cooling of the fuel rods during postulated seismic and loss of coolant accident (LOCA) occurrences.
The assembly structural component stresses under faulted conditions are evaluated using primarily the methods outlined in Appendix F of the ASME Boiler and Pressure Vessel Code, Section III.
design basis for the guide tube wear sleeves is that the fuel assembly shall not be damaged by A induced fretting-wear. Flow tests at reactor conditions of prototypic fuel and guide tube r sleeve assemblies have been used in establishing the performance of the CEA wear sleeve bination.
holddown springs, as compressed by the upper core plate during reactor operation, shall vide a net positive downward force during steady state operation, based on the most adverse bination of component dimensional and material property tolerances. In addition, the ddown springs are designed to accommodate the additional load associated with a pump rspeed transient (resulting in possible temporary liftoff of the fuel assemblies), and to continue nsure fuel assembly holddown following such an occurrence.
fuel assembly growth plus BOL length shall not exceed the minimum space between the er and lower core plates in the reactor cold condition (70&deg;F). The reactor cold condition is ting since the expansion coefficient of the stainless steel core barrel is greater than the fficient of expansion of the Zircaloy guide tubes.
spacer assembly is designed to withstand the thermal and irradiation induced differential ansion between the fuel rods and guide tubes and to withstand the design handling and dent loads discussed above. The debris resistant Inconel-718 HTP spacer used in the ninth, om location for reloads M, N and P (Cycles 10-12) was positioned such that the internal strip rsections are directly above the lower tie plate flow holes, thus reducing the size of debris ch could pass through the lower tie plate.
Reloads R and S (Cycles 13 and 14), the Inconel-718 HTP spacer grid at the ninth, bottom tion was replaced with a bimetallic spacer which is raised off the upper surface of the lower plate. The gap between the upper surface of the lower tie plate and the lower surface of the etallic spacer is spanned by a long fuel rod end cap of solid Zircaloy-4.
Zircaloy-4 HTP spacer grid is used in all nine locations in Reloads T through X (Cycles 15-This design is typically referred to as the HTP Fuel Assembly. This spacer grid design 3.3-10                                    Rev. 35


An end plug is welded to the lowe r end of the guide tube and is drilled and threaded to accept the lower cap screws. At the upper end, the guide tube is crimped into an external stainless steel locking sleeve which engages the upper tie plate assembly. The upper tie plate assembly is locked to the guide tube end fittings a nd can be unlocked for re constitution or for fu el examination using special tools.A stainless steel sleeve assembly with a chrome plated inside diameter is inse rted in the top end of the guide tube assembly. This sleeve protects th e guide tube from control rod fretting and wear when the rod is in the withdrawn/ready position. The sleeve is mechanically captured by the upper tie plate.
eload Y (Cycle 19) and later, the Zircaloy-4 HTP spacer grid is used in eight locations and an onel HMP spacer grid is used in the ninth, bottom location. This design retains the long fuel lower end cap and is typically referred to as the HTP+HMP Fuel Assembly. The HTP+HMP gn has improved structural strength, and fretting resistance compared to the HTP design.
Fuel rod pitch and position is maintained by nine spacer grids. The spacers are axially positioned so that the assemblies will be compatible with existing fuel assemblies.
design analysis is based upon reactor operating conditions. Typically, these conditions are:
The bi-metallic spacers used in Reloads M through S (Cyc les 10-14) are formed by an interlocking rectangular grid of Zircaloy-4 structural strips (see Figure 3.3-4A). Inconel-718 spring strips are mechanically secured within th ese strips. The Zircaloy-4 structural strips are welded at all intersections and to the side plates. Dimples formed in the structural strips center the MPS2 UFSAR3.3-9Rev. 35 fuel rod within the cell and along with the springs provide a positive but compliant support for each rod, sufficient to prevent fretting vibration.
Nominal Core Thermal Power = 2700 MW Nominal Coolant Pressure = 2250 psia Maximum Flow for Fuel Assembly Liftoff = 422,466 gallons per minute (at 480&deg;F)
In Reloads M, N, and P (Cycles 10-12), the debris resistant Incone l HTP spacer grid in the ninth, bottom location is located just above the lower tie plate. It is formed by an interlocking rectangular grid of Inconel-718 stri ps. The strips are welded at all intersections and to the side plates. The spacer is positioned on top of the lower tie plate with the strip intersections directly above the tie plate flow holes. This reduces the size of debris that may pass through the flow holes thereby reducing the possibility of fretting against the cladding.
Maximum Core Coolant Inlet Temperature at Nominal Power = 549&deg;F Total Average Linear Power = 6.206 kW/ft power histories used in the design analysis are designed to achieve a peak assembly burnup 6,000 MWD/MTU for Batch N, 52,500 MDW/MTU for Batch P, and 57,400 MDW/MTU for ch R and later.
Reloads R and S (Cycles 13 and
servative rod local peaking factors are used which result in a peak rod burnup of 62,000 d/MTU. Each of the rod design histories follows the single hottest rod in the first cycle ration, the hottest rod in second cycle operation, etc.
: 14) use a similar debris resistant concept with the Inconel HTP spacer replaced by a bimetallic spacer coupled with a longer lo wer end cap on the fuel rods. The HTP spacers for Reloads T through X (Cycles 15-18) are all Zircaloy-4 (Figure 3.3-4B). The strips are welded at the inters ections and side plates. The stru cture of the Zircaloy-4 strips provides the rod support.  
l assembly components must be able to withstand anticipated seismic and LOCA forces.
se may result from bundle vibration and impact due to a seismic or LOCA event. An analysis performed for the previous reloads to determine the maximum bundle displacements and the imum spacer grid forces expected during postulated accidents for Millstone 2. The loads and lacements analysis, which was performed by CE (Reference 3.3-6), considered the safe tdown earthquake (SSE) and limiting Branch Line LOCA events, and the dynamic properties he AREVA reload fuel assemblies. The resulting fuel assembly displacements and the bined seismic and LOCA grid spacer impact forces were provided to AREVA.
loads and displacements were conservatively adjusted for the Batch R design due to the mization of the fuel rod. The fuel weight was increased and the assembly stiffness was reased. The spacer impact loads and the fuel assembly maximum deflections were servatively recalculated from the reference analysis values. The spacer strength margin, the de tube stresses, and the fuel rod stresses were calculated for the adjusted loads.
culated stresses at the appropriate deflections were combined with the steady state stresses and pared with the ASME Design Criteria for faulted conditions. This limit is 0.7 times the mate strength for the primary stress combinations as compared to 0.5 times ultimate for steady e loadings. This criteria was met for both the fuel rods and the guide tubes.
3.3-11                                  Rev. 35


In Reload Y (Cycle 19) and later, all Zircaloy-4 HTP spacers ar e used in eight locations. The Inconel-718 HMP spacer is used in the ninth, bottom location. The Inconel-718 HMP bottom spacer is similar in design to the HTP spacers except for the flow channels, which are not canted.
imum projected one-sided impact load and the maximum through grid load. The maximum wable crushing load is the 95 percent confidence lower limit of the true mean of the ribution of crush test measurements. The allowable through grid strength is well above the imum through grid load. It is also above the maximum one-sided impact load. For Reload R beyond, the seismic/LOCA calculations were reviewed and determined to be bounding.
The Millstone Unit 2 Cycle 19 reload core included four fuel assemblies with natural uranium replacement fuel rods with an anti-rotation feature designed to pr event spinning of the rod during operations. The four assemblies containing replacement rods were installed in symmetric, peripheral core locations against the baffle as shown in Figure 3.3-19 (Reference 3.3-9). The core locations into which the assemblies were placed where P-1, A-8, H-21, and Y-14 (see Figure 3.4-1). The replacement rods installed unde r these conditions were evaluated against established mechanical, nuclear, and thermal/hydraulic design cr iteria for Millstone Unit 2 fuel, and were determined to be compliant with their design and licensing bases (Reference 3.3-10).
1.4    Fuel Assembly Holddown Device uel assembly holddown device has been incorporated to prevent the possibility of lifting the assembly by hydraulic forces under all normal flow conditions with temperature greater than
3.3.1.3.2 Fuel Assembly Mechanical Criteria The structural integrity of the fu el assemblies is assured by set ting design limits on stresses and deformations due to various handl ing operational and accident loads
&deg;F. The holddown device consists of a spring-loaded plate which is integral to the fuel mbly. The springs are compressed as the upper guide structure is lowered into place. The ed spring load, together with the weight of the fuel assembly, prevents possible axial motion he fuel assembly during operating conditions.
. These limits are applied to the design and evaluation of upper and lower tie plates, grid spacers, guide tubes, holddown springs, and locking hardware.
holddown device is incorporated into the upper end fitting and features a movable holddown e which acts on the underside of the fuel alignment plate (refer to Figure 3.3-5). The movable e is loaded by coil springs which are located around the upper end fitting posts. The springs positioned at the upper end of the assembly so that the spring load combines with the mbly weight in counteracting the upward hydraulic forces. The springs are sized and the ng preload selected, such that a net downward force will be maintained for all normal and cipated transient flow and temperature conditions. It should be noted that the movable plate serves as the lifting surface during handling of the fuel assembly.
The design bases for evaluating the structural integrity of the fuel assemblies are:
embly holddown was previously analyzed in Section 3.6.1 of Reference 3.3-8. The analysis been reperformed for Batch T and beyond fuel and is conservative.
a.Fuel Assembly Handling - Dynamic axial loads appr oximately 2.5 times assembly weight.b.For All Applied Loads for Normal Operat ion and Anticipated Operational Events -The fuel assembly component structural design criteria are established for the two primary material categories, austenitic stainless steels (tie plates), and Zircaloy (guide tubes, grids, spacer sleeves). Th e stress categories and strength theory for MPS2 UFSAR3.3-10Rev. 35austenitic stainless steel presented in the ASME Boiler and Pressure Ve ssel Code, Section III (Reference 3.3-1) are used as a general guide.Steady state stress limi ts are given in FSAR Section 3.3.1.1.1. Stress nomenclature is per the ASME Boiler and Pressure Vessel Code, Section III.c.Loads During Postulated Accidents - Deflection or failure of components shall not interfere with reactor shutdown or emer gency cooling of the fuel rods during postulated seismic and loss of cool ant accident (LOCA) occurrences.
1.5    Control Element Assembly As are provided by Combustion Engineering (CE) and AREVA. The CEA (shown in ure 3.3-6) is comprised of five Inconel tubes 0.948 inch in diameter. All tubes contain neutron on materials with the distribution of the poison materials as depicted in Figure 3.3-7. Each is sealed by welded end caps. A gas expansion space is provided to limit maximum tube ss due to internal pressure developed by the release of gas and moisture from the boron ide. The overall length of the CEA is provided in Table 3.3-1. Four tubes are assembled in a are array around the centrally located fifth tube. The tubes are welded to an upper end fitting.
The assembly structural component stress es under faulted condi tions are evaluated using primarily the methods outlined in Appendix F of the ASME Boiler and Pressure Ve ssel Code, Section III.The design basis for the guide tube wear sleeves is that the fuel assembly shall not be damaged by CEA induced fretting-wear. Flow tests at reactor conditions of prototypic fuel and guide tube wear sleeve assemblies have been used in establ ishing the performance of the CEA wear sleeve combination.
upper end fittings are attached to a spider hub which couples the CEA to the drive mechanism ugh the extension shaft.
chanical reactivity control is achieved by operational maneuvering of groups of single CEAs.
dual CEA is made up of two single CEAs connected to separate grippers attached to single nsion shaft. The arrangement of the CEAs in the core is shown in Figures 3.3-8 and 3.3-9.
3.3-12                                  Rev. 35


The holddown springs, as compressed by the upper core plate during re actor operation, shall provide a net positive downward force during st eady state operation, based on the most adverse combination of compone nt dimensional and ma terial property tolerances. In addition, the holddown springs are designed to accommodate the additional lo ad associated with a pump overspeed transient (re sulting in possible temporary liftoff of the fuel asse mblies), and to continue to ensure fuel assembly holddown following such an occurrence.
uffer (deceleration dashpot) system is used for slowing down the CEAs at the end of a reactor
The fuel assembly growth pl us BOL length shall not exceed the minimum space between the upper and lower core plates in the reactor cold condition (70
. The buffering action is accomplished by guide tubes which have a reduced diameter in the er section. When the tip of a CEA falls into the buffer region, the pressure buildup in the lower de tube supplies the force to slow down the CEA. The velocity is decreased to a level which minimize impact. The final impact is further cushioned by a coil spring arrangement mounted und the center CEA finger.
&deg;F). The reactor cold condition is limiting since the expansion coefficient of the stainless steel core barrel is greater than the coefficient of expansion of the Zircaloy guide tubes.
four outer guide tubes have the reduced diameter lower section (dashpot). There is no dashpot he center guide tube. There are four bleed holes above the dashpot region for the four outer de tubes. For the center guide tube, these four bleed holes are at a lower elevation. For all de tubes, there is a small drain hole at the bottom. The CEA tip is filled with a Silver-Indium-mium alloy. This replaces the B4C to avoid the change of buffer characteristics that B4C ation-induced swelling might bring about.
The spacer assembly is designed to withstand the thermal and irradiation induced differential expansion between the fuel r ods and guide tubes and to wi thstand the design handling and accident loads discussed above. Th e debris resistant Inconel-718 HTP spacer used in the ninth, bottom location for reloads M, N a nd P (Cycles 10-12) was positioned such that the internal strip intersections are directly above the lower tie plate flow holes, t hus reducing the size of debris which could pass through the lower tie plate.  
design parameters have been optimized to establish the best combination of buffer stroke and er annulus. A significant analytical effort has shown that the pressure buildup and the impact s are not damaging to the system. In addition, a test program has confirmed the feasibility of system. It has demonstrated that the buffer will work under the worst expected tolerance dition.
1.6    Neutron Source Design Cycle 18 and beyond, the reactor core will not utilize neutron sources. It has been determined during startups without neutron sources, there will continue to be a sufficient neutron count at each of the four Wide Range (WR) Excore fission detectors due to the high burnup fuel mblies that will be positioned on the core periphery.
Cycle 17 and earlier, four neutron sources were installed in the reactor core. They were held acant CEA guide tubes by means of an externally loaded spring reacting between the upper alignment plate and the top of the fuel assembly. The cladding of the neutron source rods is of ee standing design. The internal pressure is always less than reactor operating pressure.
rnal gaps and clearances are provided to allow for differential expansion between the source erial and cladding.
1.7    In-Core Instruments in-core instruments (refer to Section 7.5.4) are located in the in-core instrumentation mbly (Figure 3.3-10). The in-core instrumentated thimble support frame and guide tubes are ported by the upper guide structure (UGS) assembly. The tubes are conduits which protect the ore instruments and guide them during removal and insertion operations. The thimble support e supports the 43 in-core thimble assemblies and acts as an elevator to lift the thimbles from core into the UGS during the refueling operation.
3.3-13                                    Rev. 35


In Reloads R and S (Cycles 13 and 14), the In conel-718 HTP spacer grid at the ninth, bottom location was replaced with a bimetallic spacer which is raised off the upper surface of the lower tie plate. The gap between the upper surface of the lower tie plate and the lower surface of the bimetallic spacer is spanned by a long fuel rod end cap of solid Zircaloy-4.
heated junction thermocouple (HJTC) system is composed of two channels of HJTC ruments. Each HJTC instrument channel is manufactured into a probe assembly consisting of t HJTC sensors, a seal plug, and electrical connectors (Figure 7.5-6). The eight HJTC sensors physically independent and located at eight levels from the reactor vessel head to the fuel nment plate.
The Zircaloy-4 HTP spacer grid is used in all nine locations in Reloads T through X (Cycles 15-18). This design is typically referred to as the 'HTP Fuel Assembly'. This spacer grid design MPS2 UFSAR3.3-11Rev. 35 provides improved DNB performanc e, structural strength, and fret ting resistance. The long fuel rod end cap is maintained in the HTP Fuel Assembly.In Reload Y (Cycle 19) and later, the Zircaloy-4 HTP spacer grid is used in eight locations and an Inconel HMP spacer grid is used in the ninth, bottom location. This design retains the long fuel rod lower end cap and is typically referred to as the 'HTP
probe assembly is housed in a stainless steel support tube structure that protects the sensors m flow loads and serves as the guide path for the sensors. Figure 3.3-18 describes the locations he HJTC probe assemblies.
+HMP Fuel Assembly'. The HTP+HMP design has improved structural strength, and fretting resistance compared to the HTP design.
C Probes and Support Tubes in Upper Guide Structure HJTC probes and support tubes are installed inside two-part length CEA shrouds which ect the support tubes from normal operating cross-flow loads as well as blowdown loads. The port tubes are latched to the bottom of the CEA shroud and permanently tensioned by means threaded spanner nut at the top. Operating loads are far less than the preload developed by the ioning operation. Therefore, the support tubes will not be affected by thermal or flow loads.
The design analysis is based upon reactor operating conditions. Typically, these conditions are:
support tubes are designed to account for all tolerance conditions so that proper clearances be assured. Physically, the support tubes are similar in mass and size to a typical control ment assembly drive shaft, which would reside in the same area of the upper guide structure.
Nominal Core Thermal Power = 2700 MW Nominal Coolant Pressure = 2250 psia
presence or absence of the HJTC probes within the support tubes will in no way affect the grity of the support tubes, the UGS, the pressure boundary, and will have no significant effect n the hydraulic conditions within the reactor vessel head.
2    REACTOR INTERNAL STRUCTURES reactor internals are designed to support and orient the reactor core fuel assemblies and As, absorb the CEA dynamic loads and transmit these and other loads to the reactor vessel ge, provide flow paths for the reactor coolant, and guide in-core instrumentation.
internals are designed to safely perform their function during all steady state conditions and ng normal operating transients. The internals are designed to safely withstand the forces due eadweight, handling, system pressure, flow impingement, temperature differential, vibration seismic acceleration. All reactor components are considered Class 1 for seismic design. The tor internals design limits deflection where required by function. In most cases the design of tor internals components is limited by stress, not deflection. For the CEA shroud which is the t limiting internal component for deflection, the allowable design deflection limit is 0.5 inch.
s limit is two-thirds of the conservatively established loss-of-function deformation limit, 0.75 and applies to a break whose equivalent diameter is no larger than the largest line connected he primary coolant line. The structural components satisfy stress values given in Section III of ASME Pressure Vessel Code. Certain components have been subjected to a fatigue analysis.
ere appropriate, the effect of neutron irradiation on the materials concerned is included in the gn evaluation.
3.3-14                                  Rev. 35


Maximum Flow for Fuel Assembly Liftoff = 422,466 gallons per minute (at 480
CEA shrouds, the in-core instrumentation guide tubes and the HJTC support tubes). The flow t, although functioning as an integral part of the coolant flow path is separate from the rnals and is affixed to the bottom head of the pressure vessel. These components are shown in ure 3.1-1 and 3.3-11. The in-core instrumentation is described in Section 7.5.4.
&deg;F) Maximum Core Coolant Inlet Temp erature at Nominal Power = 549
amic system analysis methods and procedures which have been used to determine dynamic onses of reactor internals have been provided in CE, Report CENPD-42, Topical Report of amic Analysis of Reactor Vessel Internals under Loss-of-Coolant Accident Conditions with lication of Analysis to CE 800 MWe Class Reactors.
&deg;F Total Average Linear Power = 6.206 kW/ft The power histories used in the design analysis are designed to achiev e a peak assembly burnup of 56,000 MWD/MTU for Batch N, 52,500 MDW/MTU for Batch P, and 57,400 MDW/MTU for Batch R and later.
2.1     Core Support Assembly major support member of the reactor internals is the core support assembly. This assembled cture consists of the core support barrel, the lower support structure, and the core shroud. The or materials for the assembly is Type 304 stainless steel.
Conservative rod local peaking factors are used which result in a peak rod burnup of 62,000 MWd/MTU. Each of the rod design histories foll ows the single hottest r od in the first cycle operation, the hottest rod in second cycle operation, etc.
core support assembly is supported at its upper end by the upper flange of the core support el which rests on a ledge in the reactor vessel flange.
Fuel assembly components must be able to wi thstand anticipated seis mic and LOCA forces.
lower flange of the core support barrel supports and positions the lower support structure.
These may result from bundle vibrati on and impact due to a seismic or LOCA event. An analysis was performed for the previous reloads to determine the maxi mum bundle displacements and the maximum spacer grid forces expected during postulated accidents for Mill stone 2. The loads and displacements analysis, which was performed by CE (Reference 3.3-6), considered the safe shutdown earthquake (SSE) and limiting Branch Line LOCA events, a nd the dynamic properties of the AREVA reload fuel assemblies. The re sulting fuel assembly displacements and the combined seismic and LOCA grid spacer impact forces were provided to AREVA.The loads and displacements were conservatively adjusted for the Batch R design due to the optimization of the fuel rod. The fuel weight was increased and the assembly stiffness was decreased. The spacer impact loads and the fuel assembly maximu m deflections were conservatively recalculated from the reference analysis values. The spacer strength margin, the guide tube stresses, and the fuel rod stresses were calculated for the adjusted loads.
lower support structure provides support for the core by means of a core support plate ported by columns resting on beam assemblies. The core support plate provides support and ntation for the fuel assemblies. The core shroud which provides lateral support for the fuel mblies is also supported by the core support plate. The lower end attaches the core barrel to pressure vessel.
Calculated stresses at the appropria te deflections were combined wi th the steady state stresses and compared with the ASME Design Criteria for faulted conditions. This limit is 0.7 times the ultimate strength for the primary stress combinati ons as compared to 0.5 times ultimate for steady state loadings. This criteria was met fo r both the fuel rods and the guide tubes.
2.2    Core Support Barrel core support barrel is a right circular cylinder with a nominal inside diameter of 148 inches a minimum wall thickness of 1.75 inch. It is suspended by a 4 inch thick flange from a ledge he pressure vessel. The core support barrel, in turn, supports the lower support structure upon ch the fuel assemblies rest. Press fitted into the flange of the core support barrel are four nment keys located 90 degrees apart. The reactor vessel, closure head and upper guide cture assembly flanges are slotted in locations corresponding to the alignment key locations to vide proper alignment between these components in the vessel flange region.
MPS2 UFSAR3.3-12Rev. 35 The calculated grid spacer load s during each accident and the combined loads were compared with the allowable grid spacer strength at operating temperature. The loads evaluated were the maximum projected one-sided impact load and the maximum through grid load. The maximum allowable crushing load is the 95 percent confidence lower limi t of the true mean of the distribution of crush test measur ements. The allowable through gr id strength is well above the maximum through grid load. It is also above the maximum one-sid ed impact load. For Reload R and beyond, the seismic/LOCA calculations were reviewed and determined to be bounding.
ce the core support barrel is over 27 feet long and is supported only at its upper end, it is sible that coolant flow could induce vibrations in the structure. Therefore, amplitude limiting ices, or snubbers are installed on the outside of the core support barrel near the bottom end.
3.3.1.4 Fuel Assembly Holddown Device A fuel assembly holddown device ha s been incorporated to preven t the possibility of lifting the fuel assembly by hydraulic forces under all normal flow conditions with temperature greater than 500&deg;F. The holddown device consists of a spring-load ed plate which is integral to the fuel assembly. The springs are compre ssed as the upper guide structure is lowered into place. The added spring load, together with the weight of the fuel assembly, prevents possible axial motion of the fuel assembly during operating conditions.
snubbers consist of six equally spaced double lugs around the circumference and are the oves of a tongue-and groove assembly; the pressure vessel lugs are the tongues. Minimizing clearance between the two mating pieces limits the amplitude of any vibration. During mbly, as the internals are lowered into the vessel, the pressure vessel tongues engage the core port grooves in an axial direction. With this design, the internals may be viewed as a beam 3.3-15                                  Rev. 35
The holddown device is incorporat ed into the upper end fitting a nd features a movable holddown plate which acts on the underside of the fuel alignment plate (ref er to Figure 3.3-5). The movable plate is loaded by coil springs which are loca ted around the upper end fitting posts. The springs are positioned at the upper end of the assembly so that the spring load combines with the assembly weight in counteracting the upward hydr aulic forces. The spri ngs are sized and the spring preload selected, such th at a net downward force will be maintained for all normal and anticipated transient flow and te mperature conditions. It should be noted that the movable plate also serves as the lifting surface during handling of the fuel assembly.
Assembly holddown was previously analyzed in Section 3.6.1 of Reference 3.3-8. The analysis has been reperformed for Batch T a nd beyond fuel and is conservative.
3.3.1.5 Control Element Assembly CEAs are provided by Combustion Engineering (CE) and AREVA. The CEA (shown in Figure 3.3-6) is comprised of five Inconel tubes 0.948 inch in diameter. All tube s contain neutron poison materials with the distri bution of the poison materials as depicted in Figure 3.3-7. Each tube is sealed by welded end caps. A gas expa nsion space is provided to limit maximum tube stress due to internal pressure developed by th e release of gas and mo isture from the boron carbide. The overall length of the CEA is provided in Table 3.3-1. Four tubes are assembled in a square array around the centrally lo cated fifth tube. The tubes are we lded to an upper end fitting.
The upper end fittings are attached to a spider hub wh ich couples the CEA to the drive mechanism through the extension shaft.
Mechanical reactivity control is achieved by operational maneuvering of groups of single CEAs.
The dual CEA is made up of two si ngle CEAs connected to separate grippers attached to single extension shaft. The arrangement of the CEAs in the core is shown in Figures 3.3-8 and 3.3-9.
MPS2 UFSAR3.3-13Rev. 35 There are 49 single CEAs and 12 dual CEAs all operated by a tota l of 61 CEDMs. Considering the 12 dual CEAs as 24 single CEAs gives an overall number of 73 CEAs in the core.A buffer (deceleration dashpot) system is used for slowing down the CEAs at the end of a reactor trip. The buffering action is accomplished by guide tubes which have a reduc ed diameter in the lower section. When the tip of a CEA falls into the buffer regi on, the pressure buil dup in the lower guide tube supplies the force to slow down the CE A. The velocity is decreased to a level which will minimize impact. The final impact is furthe r cushioned by a coil spring arrangement mounted around the center CEA finger.
The four outer guide tube s have the reduced diamet er lower section (dashpot
). There is no dashpot in the center guide tube. There ar e four bleed holes above the da shpot region for the four outer guide tubes. For the center guide tube, these four bleed holes ar e at a lower elevation. For all guide tubes, there is a small drain hole at the bottom. The CEA tip is filled with a Silver-Indium-Cadmium alloy. This replaces the B 4C to avoid the change of buffer characteristics that B 4 C radiation-induced swel ling might bring about.
The design parameters have been optimized to establish the best combination of buffer stroke and buffer annulus. A significant analytical effort has shown that the pressure buildup and the impact loads are not damaging to the syst em. In addition, a test program has confirme d the feasibility of the system. It has demonstrated that the buffer will work under the worst expected tolerance condition.


3.3.1.6 Neutron Source Design For Cycle 18 and beyond, the reactor core will not utilize neutron s ources. It has been determined that during startups without neutron sources, there will continue to be a suff icient ne utron count rate at each of the four Wide Range (WR) Excore fission dete ctors due to the high burnup fuel assemblies that will be positioned on the core periphery. For Cycle 17 and earlier, four neut ron sources were installed in the reactor core. They were held in vacant CEA guide tubes by mean s of an externally loaded sp ring reacting between the upper fuel alignment plate and the top of the fuel assembly. The cladding of the neutron source rods is of a free standing design. The internal pressure is always less th an reactor operating pressure.
sure vessel tongues have bolted, lock welded Inconel X shims and the core support barrel oves are hardfaced with Stellite to minimize wear. The snubber assembly is shown in ure 3.3-12.
Internal gaps and clearances are provided to allow for differenti al expansion between the source material and cladding.
2.3    Core Support Plate and Support Columns core support plate is a 147 inch diameter, 2 inch thick, Type 304 stainless steel plate into ch the necessary flow distributor holes for the fuel assemblies have been machined. Fuel mbly locating pins (four for each assembly) are shrunk-fit into this plate. Columns and port beams are located between this plate and the bottom of the core support barrel in order to vide support for this plate and transmit the core load to the bottom flange of the core support el.
3.3.1.7 In-Core InstrumentsThe in-core instruments (refer to Section 7.5.4) are located in the in-core instrumentation assembly (Figure 3.3-10). The in-c ore instrumentated thimble su pport frame and gui de tubes are supported by the upper guide structure (UGS) assembly. The tubes are condui ts which protect the in-core instruments and guide th em during removal and inserti on operations. The thimble support frame supports the 43 in-core thimble assemblies and acts as an elevat or to lift the thimbles from the core into the UGS during the refueling operation.
2.4    Core Shroud core shroud provides an envelope for the core and limits the amount of coolant bypass flow.
MPS2 UFSAR3.3-14Rev. 35 3.3.1.8 Heated Junction Thermocouples The heated junction thermocoupl e (HJTC) system is composed of two channels of HJTC instruments. Each HJ TC instrument channel is manufactured into a probe assembly consisting of eight HJTC sensors, a seal plug, and electrical connectors (Figure 7.5-6). The eight HJTC sensors are physically independent and loca ted at eight levels from the reactor vessel head to the fuel alignment plate.
shroud (Figure 3.3-13) consists of two Type 304 stainless steel ring sections, aligned by ns of radial shear pins and attached to the core support plate by Type 348 stainless steel tie
The probe assembly is housed in a stainless steel suppor t tube structure that protects the sensors from flow loads and serves as the guide path for the sensors. Figure 3.3-18 describes the locations of the HJTC probe assemblies.HJTC Probes and Support Tubes in Upper Guide Structure The HJTC probes and support t ubes are installed inside two-part length CEA shrouds which protect the support tubes from norma l operating cross-flow loads as well as blowdown loads. The support tubes are latched to the bottom of th e CEA shroud and permanently tensioned by means of a threaded spanner nut at the top. Operating lo ads are far less than the preload developed by the tensioning operation. Therefore, the support tubes will not be affect ed by thermal or flow loads.
: s. A gap is maintained between the core shroud outer perimeter and the core support barrel in er to provide some coolant flow upward between the core shroud and core support barrel, eby minimizing thermal stresses in the core shroud and eliminating stagnant pockets.
The support tubes are designed to account for all tolerance conditi ons so that proper clearances will be assured. Physically, the support tubes are similar in mass and size to a typical control element assembly drive shaft, wh ich would reside in the same ar ea of the upper guide structure.
2.5    Flow Skirt Inconel flow skirt is a right circular cylinder, perforated with 2-11/16 inch diameter holes, reinforced at the top and bottom with stiffening rings. The flow skirt is used to reduce ualities in core inlet flow distributions and to prevent formation of large vortices in the lower um. The skirt provides a nearly equalized pressure distribution across the bottom of the core port barrel. The skirt is supported by nine equally spaced machined sections which are welded he bottom of the pressure vessel.
The presence or absence of the HJTC probes within the support tubes will in no way affect the integrity of the support tubes, the UGS, the pressure boundary, and will have no significant effect upon the hydraulic conditions within the reactor vessel head.
2.6    Upper Guide Structure Assembly s assembly (Figure 3.3-14) consists of the upper support structure, 69 CEA shrouds, a fuel mbly alignment plate and an expansion compensating ring. The UGS assembly aligns and rally supports the upper end of the fuel assemblies, maintains the CEA spacing, prevents fuel mblies from being lifted out of position during a severe accident condition and protects the As from the effect of coolant crossflow in the upper plenum. The UGS is handled as one unit ng installation and refueling.
upper end of the assembly is a structure consisting of a support plate welded to a grid array of nch deep beams and a 24 inch deep cylinder which encloses and is welded to the ends of the ms. The periphery of the plate contains four accurately machined and located alignment ways, equally spaced at 90 degree intervals, which engage the core barrel alignment keys. The 3.3-16                                    Rev. 35


====3.3.2 REACTOR====
ure head. The grid aligns and supports the upper end of CEA shrouds.
INTERNAL STRUCTURES The reactor internals are designe d to support and orient the reac tor core fuel assemblies and CEAs, absorb the CEA dynamic loads and transmit these and other loads to the reactor vessel flange, provide flow paths for the reactor coolant, and guide in-core instrumentation.
CEA shrouds extend from the fuel assembly alignment plate to an elevation about three feet ve the UGS support plate. There are 57 single-type shrouds. These consist of cylindrical upper ions welded to integral bottom sections, which are shaped to provide flow passages for the lant passing through the alignment plate while shrouding the CEAs from cross-flow. There are 12 dual-type shrouds which in configuration consist of two single-type shrouds connected by ctangular section shaped to accommodate the dual CEAs. The shrouds are bolted to the fuel mbly alignment plate. At the UGS support plate, the single shrouds are connected to the plate spanner nuts which permit axial adjustment. The spanner nuts are tightened to proper torque lockwelded. The dual shrouds are attached to the upper plate by welding.
fuel assembly alignment plate is designed to align the upper ends of the fuel assemblies and upport and align the lower ends of the CEA shrouds.
cision machined and located holes in the fuel assembly alignment plate align the fuel mblies. The fuel assembly alignment plate also has four equally spaced slots on its outer edge ch engage with Stellite hardfaced pins protruding from the core shroud to limit lateral motion he UGS assembly during operation. The fuel alignment plate bears the upward force of the assembly holddown devices. This force is transmitted from the alignment plate through the A shrouds to the UGS support plate and hence to the expansion compensating ring.
expansion compensating ring bears on the flange at the top of the assembly to accommodate l differential thermal expansion between the core barrel flange, UGS flange and pressure sel flange support edge and head flange recess.
UGS assembly also supports the in-core instrumentation thimble support frame, guide tubes, HJTC support tubes.
integral connections in the reactor internals are designed within the stress intensity limits d in Tables N-422 and N-416.1 of Section III of the ASME code for normal and upset ditions. For emergency and faulted conditions, the design limits are as given in Table 3.2-1.
3    CONTROL ELEMENT DRIVE MECHANISM 3.1    Design CEDM is of the magnetic jack type drive. Each CEDM is capable of withdrawing, inserting, ding or tripping the CEA from any point within its 137-inch stroke. The design of the CEDM hown in Figure 3.3-15 and is identical to that for Maine Yankee (AEC Docket Number 50-
) and Calvert Cliffs Units 1 and 2 (AEC Docket Numbers. 50-317 and 50-318).
CEDM drives the CEA within the reactor core and indicates the position of the CEA with ect to the core. The speed at which the CEA is inserted or withdrawn from the core is 3.3-17                                    Rev. 35


The internals are designed to sa fely perform their f unction during all steady state conditions and during normal operating transients.
nergized, allowing the CEA and the supporting CEDM components to drop into the core by vity. The CEA drop time is 2.75 seconds, where drop time is defined as the interval between time power is removed from the CEDM coils and the time the CEA has reached 90 percent of fully inserted position. The reactivity is reduced during such a drop at a rate sufficient to trol the core under any operating transient or accident condition. The CEA accelerates to about t/sec and is decelerated at the end of the drop by the buffer section of the CEA guide tubes.
The internals are designed to sa fely withstand the forces due to deadweight, handling, system pressure, flow impingement, temperature diff erential, vibration and seismic acceleration. All reac tor components are considered Class 1 for seismic design. The reactor internals design limits de flection where required by function. In most cases the design of reactor internals components is limited by stress, not deflection.
ve down capability following a reactor trip is not required for safety purposes. The safety lyses of Chapter 14 assume the CEA of highest reactivity worth sticks in the fully withdrawn ition. A drive down feature would introduce the possibility of a failure which would prevent er from being removed from the CEDMs during a trip, which would lead to a reduction in t safety.
For the CEA shroud which is the most limiting internal component for deflection, the allowable de sign deflection limit is 0.5 inch.
re are 69 CEDM nozzles on top of the reactor vessel closure head. Eight of the 69 nozzles e used for the part length CEAs in Cycle 1, six of which are no longer used, and two of which used for HJTC/RVLMS instrumentation. There are 61 CEDMs in current use. The six spare zles are capped with adapters. Each CEDM is connected to a CEA by a locked coupling. The ght of the CEAs and CEDMs is carried by the vessel head.
This limit is two-thirds of the conservatively established loss-of-function deformation limit, 0.75 inch and applies to a break whose equivalent diameter is no larger than the largest line connected to the primary coolan t line. The structural com ponents satisfy stress values given in Section III of the ASME Pressure Vessel Code. Certain component s have been subjected to a fatigue analysis. Where appropriate, the effect of neutron irradiation on the materials concerned is included in the design evaluation.
CEDM is designed to handle dual, single or part length CEAs. The maximum operating ed capability of the CEDMs is 40 inches per minute for single CEAs and 20 inches per minute dual CEAs.
MPS2 UFSAR3.3-15Rev. 35The components of the reactor internals are divided into four major parts consisting of the core support barrel, the lower core support structure (including the core shroud), the UGS (including the CEA shrouds, the in-core instrumentation guide tubes and the HJTC s upport tubes). The flow skirt, although functioning as an integral part of the coolant flow path is separate from the internals and is affixed to the bottom head of th e pressure vessel. These components are shown in Figure 3.1-1 and 3.3-11. The in-core instrume ntation is describe d in Section 7.5.4.
3.2    Control Element Drive Mechanism Pressure Housing CEDM housing is attached to the reactor vessel head nozzle by means of a threaded joint and welded. The CEDM nozzles are made of Inconel Alloy 690 to minimize Primary Water Stress rosion Cracking. The CEDM pressure housings including the magnetic coil jack assemblies e replaced as part of the replacement reactor vessel closure head project.
Dynamic system analysis methods and procedures which have be en used to determine dynamic responses of reactor internals have been pr ovided in CE, Report CENPD-42, "Topical Report of Dynamic Analysis of Reactor Vessel Internals under Loss-of-Coolant Acci dent Conditions with Application of Analysis to CE 800 MWe Class Reactors".
CEDM upper housing design and fabrication conform to the requirements of the ASME ler and Pressure Vessel Code, Section III, 1998 Edition through 2000 Addenda. The housing is gned for steady state conditions as well as all anticipated pressure and thermal transients.
3.3.2.1 Core Support Assembly The major support member of the reactor internals is the core support assembly. This assembled structure consists of the core support barrel, the lowe r support structure, and the core shroud. The major materials for the assembly is T ype 304 stainless steel.
ce the CEDM housing is seal welded to the head nozzle, it need not be removed since all icing of the CEDM is performed from the top of the CEDM housing. This opening is closed means of an upper housing and an omega seal weld. The CEDM pressure housing is capable of g vented after major coolant refills of the reactor coolant system (RCS), such as after a eling and after reactor coolant pump (RCP) maintenance. However, venting of the CEDM sure housing is no longer necessary after major refills of the Reactor Coolant System (RCS),
The core support assembly is s upported at its upper end by the uppe r flange of the core support barrel which rests on a ledge in the reactor vessel flange.
e a vacuum refill method is used. The vacuum refill process involves a partial vacuum in the S while at mid-loop level and then slowly refilling the RCS.
The lower flange of the core support barrel supports and positi ons the lower support structure.
3.3-18                                    Rev. 35
The lower support structure pr ovides support for the core by means of a core support plate supported by columns resting on beam assemblies. The core s upport plate provides support and orientation for the fuel assemblies. The core shroud which provides lateral support for the fuel assemblies is also supported by th e core support plate. The lower end attaches the core barrel to the pressure vessel.
3.3.2.2 Core Support Barrel The core support barrel is a right circular cylinder with a nominal inside diameter of 148 inches and a minimum wall thickness of 1.75 inch. It is suspended by a 4 in ch thick flange from a ledge on the pressure vessel. The core support barrel, in turn, supports the lower support structure upon which the fuel assemblies rest. Pr ess fitted into the flange of the core support barrel are four alignment keys located 90 degrees apart. The reactor vessel, closure head and upper guide structure assembly flanges are sl otted in locations corresponding to the alignment key locations to provide proper alignment between these components in the vessel flange region.
Since the core support barrel is over 27 feet long and is supporte d only at its upper end, it is possible that coolant flow could induce vibrations in th e structure. Therefor e, amplitude limiting devices, or snubbers are installe d on the outside of the core s upport barrel near the bottom end.
The snubbers consist of six equally spaced double lugs ar ound the circumference and are the grooves of a "tongue-and groove" asse mbly; the pressure vessel l ugs are the tongues. Minimizing the clearance between the two mating pieces limits the amplitude of a ny vibration. During assembly, as the internals are lowered into the ve ssel, the pressure vessel tongues engage the core support grooves in an axial direction. With this design, the intern als may be viewed as a beam MPS2 UFSAR3.3-16Rev. 35 with supports at the furthest extr emities. Radial and ax ial expansion of the co re support barrel are accommodated, but lateral movement of the core support barrel is restri cted by this design. The pressure vessel tongues have bolted, lock welded Inconel X shims and th e core support barrel grooves are hardfaced with Stellite to minimize wear. The snubber as sembly is shown in Figure 3.3-12.
3.3.2.3 Core Support Plate and Support Columns The core support plate is a 147 inch diameter, 2 inch thick, T ype 304 stainless steel plate into which the necessary flow distributor holes for the fuel assemblies have been machined. Fuel assembly locating pins (four fo r each assembly) are shrunk-fit into this plate. Columns and support beams are located between this plate and the bottom of the core support barrel in order to provide support for this plate and transmit the core load to the bottom flange of the core support barrel.3.3.2.4 Core Shroud The core shroud provides an envel ope for the core and limits the amount of coolant bypass flow.
The shroud (Figure 3.3-13) consists of two Type 304 stainless st eel ring sections, aligned by means of radial shear pins and attached to the core support plate by Type 348 stainless steel tie rods. A gap is maintained between the core shr oud outer perimeter and the core support barrel in order to provide some coolant flow upward between the core sh roud and core support barrel, thereby minimizing thermal stresses in the core shroud and eliminating stagnant pockets.
3.3.2.5 Flow Skirt The Inconel flow skirt is a right circular cylinder, perforated wi th 2-1 1/16 inch diameter holes, and reinforced at the top and bottom with stif fening rings. The flow sk irt is used to reduce inequalities in core inlet flow distributions and to prevent formation of large vortices in the lower plenum. The skirt provides a nearly equalized pressure distri bution across the bottom of the core support barrel. The skirt is suppor ted by nine equally spaced machin ed sections which are welded to the bottom of the pressure vessel.


3.3.2.6 Upper Guide Structure AssemblyThis assembly (Figure 3.3-14) consists of the upper support structure, 69 CEA shrouds, a fuel assembly alignment plate and an expansion compensating ring. The UGS assembly aligns and laterally supports the upper end of the fuel assemb lies, maintains the CEA spacing, prevents fuel assemblies from being lifted out of position during a severe acci dent condition and protects the CEAs from the effect of coolan t crossflow in the upper plenum.
HJTC probe assemblies are located at the two original locations (CEDMs 11 and 13) on the acement reactor vessel closure head. The HJTC pressure boundary also known as the Reactor sel Level Monitoring System (RVLMS) pressure housing assembly consists of upper pressure sing tube, upper flange type Grayloc connection and lower housing. The lower housing is ed to the reactor vessel head nozzle by means of a threaded joint and an omega seal weld. The sure boundary at the top of the RVLMS pressure housing is maintained by a quick disconnect yloc type flange (See Figure 3.3-17). The components are designed to ASME Section III, PV Code 1998 Edition through 2000 Addenda.
The UGS is handled as one unit during installation and refueling.
pressure and thermal loads associated with normal operation and transient conditions have n included in stress analyses performed in accordance with ASME BPVC criteria. All stresses within allowable limits.
The upper end of the assembly is a structure consis ting of a support plate welded to a grid array of 24 inch deep beams and a 24 inch deep cylinder which encl oses and is welded to the ends of the beams. The periphery of the pl ate contains four accurately ma chined and located alignment keyways, equally spaced at 90 degree intervals, which engage the core barrel alignment keys. The MPS2 UFSAR3.3-17Rev. 35 reactor vessel closure head flange is slotted to engage the upper ends of the alignment keys in the core barrel. This system of keys and slots provides an accurate me ans of aligning the core with the closure head. The grid aligns and supports the upper end of CEA shrouds.
3.3    Magnetic Jack Assembly magnetic jack motor assembly is an integral unit which fits into the CEDM housing through opening in the top of the housing. This unit carries the motor tube, lift and hold pawls and nets. The drive power is supplied by electrical coils positioned around the CEDM housing.
The CEA shrouds extend from the fuel assembly al ignment plate to an elev ation about three feet above the UGS support plate. There are 57 single-type shrouds. These cons ist of cylindrical upper sections welded to integral bot tom sections, which are shaped to provide flow passages for the coolant passing through the alignment plate while shrouding the CEAs from cross-flow. There are also 12 dual-type shrouds which in configuration consist of two single-type shrouds connected by a rectangular secti on shaped to accommodate the dual CEAs. The shrouds are bolted to the fuel assembly alignment plate. At the UGS support plat e, the single shrouds are connected to the plate by spanner nuts which permit axia l adjustment. The spanner nuts ar e tightened to proper torque and lockwelded. The dual shrouds are at tached to the upper plate by welding.
CEDMs are cooled by air supplied at 900 CFM at 95&deg;F (maximum) to each CEDM. The gn of the control element drive mechanism is such that loss of cooling air will not prevent the DM from releasing the CEA. The ability of the CEDM to release the rods is not dependent on cooling flow provided by the CEDM Cooling System. Cooling function is only to ensure ability of the CEDM coil stack. Following insertion of the CEDM motor assembly, the upper sure housing is threaded into the CEDM motor housing and seal welded. This upper pressure sing encloses the CEDM extension shaft and supports the shroud assembly. The reed switch mbly is supported by the shroud assembly.
The fuel assembly alignment plat e is designed to align the upper ends of the fuel assemblies and to support and align the lower ends of the CEA shrouds.Precision machined and located holes in the fuel assembly alignment plate align the fuel assemblies. The fuel assembly ali gnment plate also has four equally spaced slots on its outer edge which engage with Stellite hard faced pins protruding from the co re shroud to limit lateral motion of the UGS assembly during operation. The fuel alignment plate bears th e upward force of the fuel assembly holddown devices. Th is force is transmitted from the alignment plate through the CEA shrouds to the UGS support plate and hence to the expansion compensating ring.
lifting operation consists of magnetically operated step movements. Two sets of mechanical hes (one holding, one lifting) are utilized engaging a notched drive shaft. To prevent excessive h wear, a means has been provided to unload the lifting latches during the engaging and ngaging operations.
The expansion compensating ring bear s on the flange at the top of the assembly to accommodate axial differential thermal expans ion between the core barrel flange, UGS flange and pressure vessel flange support edge and head flange recess.
magnetic force is obtained from large DC magnet coils mounted on the outside of the motor er for the electromagnets is obtained from one of two separate supplies. A control grammer actuates the stepping cycle and obtains the CEA location by a forward or reverse ping sequence. CEDM hold for shutdown and regulating CEAs is obtained by energizing a d coil at a reduced current while all other coils are deenergized. The full length CEAs are ped upon interruption of electrical power to all coils.
The UGS assembly also supports the in-core inst rumentation thimble suppor t frame, guide tubes, and HJTC support tubes.
3.4    Position Indication ee separate means are provided for transmitting CEA position indication.
3.3-19                                    Rev. 35


All integral connections in the reactor internals are designed within the stress intensity limits listed in Tables N-422 and N-416.1 of Section III of the ASME code for normal and upset conditions. For emergency and faulted conditions, the design limits are as given in Table 3.2-1.
vide an output voltage proportional to CEA position. The third method utilizes three pairs of switches spaced at discrete locations within a position transmitter assembly. A permanent net built into the drive shaft actuates the reed switches one at a time as it passes by them. CEA ition instrumentation is discussed in detail in Section 7.5.3.
3.5    Control Element Assembly Disconnect CEA is connected to the drive shaft extension with an internal collet-type coupling at its er end. (Coupling is performed before the vessel head is installed). In order to disengage the A from the drive shaft extension, a tool is attached to the top end of the drive shaft when the tor vessel head has been removed.
pulling up on the spring-loaded operating rod in the center of the drive shaft, a tapered plunger ithdrawn from the center of the collet-type gripper causing it to collapse due to axial pressure m the CEA, thus permitting removal of the coupler from the CEA. Releasing the operating rod nger after the coupler has been withdrawn from the CEA expands the coupler to a diameter prevents recoupling to the CEA.
3.6 Test Program est program has been conducted to verify the adequacy of the magnetic jack CEDM. The gram is described in Section 1.5.4.
4    REFERENCES 1     ASME Boiler and Pressure Vessel Code, Section III, 1977 Edition, ASME New York, NY.
2    K. R. Merckx, RODEX2 - Fuel Rod Thermal-Mechanical Response Evaluation Model, XN-NF-81-58 (NP)(A), Revision 2, March 1985 and Supplements.
3    Qualification of Exxon Nuclear Fuel for Extended Burnup (PWR), XN-NF-82-06 (NP)(A), Revision 1, Supplements 2, 4, 5, October 1985.
4    W. J. O'Donnel and B. F. Langer, Fatigue Design Bases for Zircaloy Components, Nuclear Science and Engineering, Volume 20, January 1964.
5    MATPRO Version, A Handbook of Material Properties for Use in the Analysis of Light Water Reactor Fuel Rod Behavior, TREE-NUREG 1008, December 1976.
6    J. C. Winslow (CE) to T. J. Honan (NU), CE Letter, Seismic and Branch Line LOCA Analysis of SPC Reload Fuel for Millstone 2, NU-88-043 (March 31, 1988).
3.3-20                                    Rev. 35


====3.3.3 CONTROL====
8  ANF-88-88(P), Rev. 1, Design Report for Millstone Point Unit 2 Reload ANF-1, August 29, 1988.
ELEMENT DRIVE MECHANISM 3.3.3.1 Design The CEDM is of the magnetic jack type drive. Each CEDM is cap able of withdrawing, inserting, holding or tripping the CEA from any point within its 137-inch st roke. The design of the CEDM is shown in Figure 3.3-15 and is identical to th at for Maine Y ankee (A EC Docket Number 50-309) and Calvert Cliffs Units 1 and 2 (A EC Docket Numbers. 50-317 and 50-318).
9  AREVA Contract Requirements Document Number 89-9070921-001-AREVA Contract No. J37MIL219B, January 28, 2008.
The CEDM drives the CEA within the reactor co re and indicates the position of the CEA with respect to the core. The speed at which the CEA is inserted or withdrawn from the core is MPS2 UFSAR3.3-18Rev. 35 consistent with the reactivity change requireme nts during reactor operati on. For conditions that require a rapid shutdown of the reactor, the CEDM coils of the shutdown and regulating CEAs are deenergized, allowing the CEA and the supporting CEDM components to drop into the core by gravity. The CEA drop time is 2.75 seconds, where drop time is de fined as the interval between the time power is removed from the CEDM coils and the time th e CEA has reached 90 percent of its fully inserted position. The reactivity is reduced during such a drop at a rate sufficient to control the core under any opera ting transient or accident conditi on. The CEA accelerates to about 11 ft/sec and is decelerated at the end of the drop by the buffer se ction of the CEA guide tubes. Drive down capability following a reactor trip is not required for safe ty purposes. The safety analyses of Chapter 14 assume th e CEA of highest reacti vity worth sticks in the fully withdrawn position. A drive down feature would introduce the pos sibility of a failure which would prevent power from being remove d from the CEDMs during a trip, wh ich would lead to a reduction in plant safety.There are 69 CEDM nozzles on top of the reactor vessel closure head. Eight of the 69 nozzles were used for the part length CEAs in Cycle 1, six of which are no longer used, and two of which are used for HJTC/RVLMS instru mentation. There are 61 CEDMs in current use. The six spare nozzles are capped with adapters. Each CEDM is connected to a CEA by a locked coupling. The weight of the CEAs and CEDMs is carried by the vessel head.
10 AREVA Document 51-9074000-000, Compliance Document - Replacement Fuel Rod -
Millstone 2 Fuel Failure Mitigation, March 5, 2008.
3.3-21                              Rev. 35


The CEDM is designed to handl e dual, single or part lengt h CEAs. The maximum operating speed capability of the CEDMs is 40 inches per minut e for single CEAs and 20 inches per minute for dual CEAs.
TABLE 3.3-1 MECHANICAL DESIGN PARAMETERS
3.3.3.2 Control Element Drive Mechanism Pressure HousingThe CEDM housing is attached to the reactor vessel head nozzle by means of a threaded joint and seal welded. The CEDM nozzles are made of Inconel Alloy 690 to minimize Primary Water Stress Corrosion Cracking. The CEDM pressure housings including the ma gnetic coil ja ck assemblies were replaced as part of the replacement reactor vessel closure head project.
* l Assembly Geometry                                    14 by 14 Assembly Pitch, inches                      8.180 Assembly Envelope, inches                    8.160 Rod Pitch, inches                            0.580 Number of Grids per Assembly                9 Approximate Assembly Weight, lb.             1280/1313
The CEDM upper housing design and fabrication conform to the requirements of the ASME Boiler and Pressure Vessel Code, Section III, 1998 Edition th rough 2000 Addenda. The housing is designed for steady state conditions as well as all anticipated pr essure and thermal transients. Once the CEDM housing is seal welded to th e head nozzle, it need no t be removed since all servicing of the CEDM is perf ormed from the top of the CEDM housing. This opening is closed by means of an upper housing and an omega seal weld. The CEDM pr essure housing is capable of being vented after major coolant refills of the reactor coolant system (RCS), such as after a refueling and after react or coolant pump (RCP) maintenance. However, venting of the CEDM pressure housing is no longer necessary after majo r refills of the Reactor Coolant System (RCS), since a vacuum refill me thod is used. The vacuum refill process involves a partial vacuum in the RCS while at mid-loop level and then slowly refilling the RCS.
* Fuel Rod to Fuel Rod Outside Dimension, inches 7.980 l Rod and Pellet Clad OD, inches                              0.440 Clad thickness, inches                      0.031/0.028
MPS2 UFSAR3.3-19Rev. 35 3.3.3.2.1 Heated Junction Thermoc ouple Pressure BoundaryThe HJTC probe assemblies are located at the two original locations (CEDMs 11 and 13) on the replacement reactor vess el closure head. The HJTC pressure boundary also known as the Reactor Vessel Level Monitoring System (RVLMS) pressure housing assembly consists of upper pressure housing tube, upper flange type Grayloc connection and lowe r housing. The lower housing is joined to the reactor vessel head nozzle by means of a threaded joint and an omega seal weld. The pressure boundary at the top of the RVLMS pressu re housing is maintained by a quick disconnect Grayloc type flange (See Figure 3.3-17). The components are designed to ASME Section III, B&PV Code 1998 Edition through 2000 Addenda.
* Pellet Diameter, inches                      0.3700/0.3770
The pressure and thermal loads associated with normal operation and tr ansient conditions have been included in stress analyses performed in accordance with ASME BPVC criteri
* Pellet Length, inches                        0.425/0.435
: a. All stresses are within allowable limits.
* Pellet Density (% Theoretical)               94.0/95.0/95.35 **
3.3.3.3 Magnetic Jack Assembly The magnetic jack motor assembly is an integr al unit which fi ts into the CEDM housing through an opening in the top of the hous ing. This unit carries the motor tube, lift and hold pawls and magnets. The drive power is supplied by electri cal coils positioned around the CEDM housing.
Active Stack Length, Cold, inches            136.7 trol Rod Guide Tube Number per assembly                         4 Tube ID, above dashpot, inches              1.035 Wall Thickness, inches                      0.040 rumentation Tube Number per Assembly                          1 Tube ID, inches                              1.035 Wall Thickness, inches                      0.040 cer Grid Material                                    Zircaloy-4 / Inconel-718 3.3-22                          Rev. 35
The CEDMs are cooled by ai r supplied at 900 CFM at 95
&deg;F (maximum) to each CEDM. The design of the control element drive mechanism is such that loss of cooling air will not prevent the CEDM from releasing the CEA. Th e ability of the CEDM to release the rods is not dependent on the cooling flow provided by th e CEDM Cooling System. Cooling function is only to ensure reliability of the CEDM coil stack. Following insertion of the CEDM motor assembly, the upper pressure housing is threaded into the CEDM motor housing and seal welded. This upper pressure housing encloses the CEDM extension shaft and supports the shroud assembly. The reed switch assembly is supported by the shroud assembly.
The lifting operation consis ts of magnetically operated step movements. Two sets of mechanical latches (one holding, one lifting) are utilized engaging a notched drive shaft. To prevent excessive latch wear, a means has been provided to unload th e lifting latches during the engaging and disengaging operations.


The magnetic force is obtained from large DC magnet coils m ounted on the outside of the motor tube.
9/0 for Batch R, S 9/0 for Batch T - X 8/1 Batch Y and beyond ves (Wear)
Power for the electromagnets is obtained from one of two separate supplies. A control programmer actuates the stepping cycle and obtains the CEA lo cation by a forward or reverse stepping sequence. CEDM hold for shutdown and regulating CEAs is obtained by energizing a hold coil at a reduced current while all other coils are deenergized. The full length CEAs are tripped upon interruption of elec trical power to all coils.
Material                                      SS/Chrome Plate nable Poison Rod Active Length, inches                        124.7 + UO2 blankets Material                                      Gd2O3 / U02 Pellet Diameter, inches                      0.3700/0.3770 Clad Material                                Zircaloy-4 Clad ID, inches                              0.378/0.384 Clad OD, inches                              0.440 Clad Thickness, (nominal) inches              0.031/0.028 Diametral Gap, (cold, nominal), inches        0.008/0.007 Pellet Length, inches                        0.545 trol Element Assembly Number                                        73 Number of Absorber Elements per Assembly      5 Type                                          Cylindrical Rods Clad Material                                Inconel 625 Clad Thickness, inches                        0.036 Clad OD, inches                              0.948 Poison Material                              B4C & Ag-In-CD Corner Element Pitch, inches                  4.64 Total CEA Length, inches                      161.31- CE / 161.25 - AREVA Poison Length, inches                        132 -CE / 133.5 - AREVA CEA Dry Weight, lb.                           95 - CE / 85 - AREVA 3.3-23                            Rev. 35
3.3.3.4 Position IndicationThree separate means are provided for tr ansmitting CEA position indication.
MPS2 UFSAR3.3-20Rev. 35 The first method utilizes the el ectrical pulses from the magnetic coil power programmer. The second method utilizes reed swit ches and a voltage divider ne twork mounted on the CEDM to provide an output voltage proportional to CEA position. The thir d method utilizes three pairs of reed switches spaced at discre te locations within a position transmitter assembly. A permanent magnet built into the drive shaft actuates the reed switches one at a time as it passes by them. CEA position instrumentation is discu ssed in detail in Section 7.5.3.
3.3.3.5 Control Element Assembly Disconnect The CEA is connected to the drive shaft extensio n with an internal colle t-type coupling at its lower end. (Coupling is performed be fore the vessel head is installe d). In order to disengage the CEA from the drive shaft extension, a tool is att ached to the top end of the drive shaft when the reactor vessel head has been removed.
By pulling up on the spring-loaded operating rod in th e center of the drive shaft, a tapered plunger is withdrawn from the center of the collet-type gripper causing it to collapse due to axial pressure from the CEA, thus permitting removal of the c oupler from the CEA. Releasing the operating rod plunger after the coupler has been withdrawn from the CEA expands the coupler to a diameter that prevents recoupling to the CEA.


3.3.3.6 Test Program A test program has been conducte d to verify the adequacy of the magnetic jack CEDM. The program is described in Section 1.5.4.
Single                                    210 - CE / 200 - AREVA Dual                                      334 - CE / 314 AREVA e Arrangement Number of Fuel Assemblies in Core Total          217 Number of Single CEAs                            49 Number of Dual CEAs                              12 CEA Pitch, minimum, inches                      11.57 Spacing Between Fuel Assemblies, Fuel Rod Surface to Surface, inches              0.200 Spacing, Outer Fuel Rod Surface to Core Shroud, inches                              0.18 Hydraulic Diameter, Nominal Channel, feet        0.04445 Total Flow Area (Excluding Guide Tubes), square feet 53.5 Total Core Area, square feet                    101.1 Core Equivalent Diameter, inches                136 Core Circumscribed Diameter, inches              143.1 Core Volume, liters                              32,526 Total Fuel Loading, MTU (Typical)                83.65 Total Heat Transfer Area, square feet            50,117 Applicable to Batches N, P/applicable to Batch R and subsequent Batches.
Applicable to Batches N, P/applicable to Batches R, S/applicable to Batch T and subsequent Batches.
3.3-24                                  Rev. 35


3.
ABLE 3.3-2 PRESSURIZED WATER REACTOR PRIMARY COOLANT WATER CHEMISTRY RECOMMENDED SPECIFICATIONS ductivity (S/cm at 25&deg;C)        Relative to Lithium and Boron concentration.
at 25&deg;C                            Determined by the concentration of boric acid and lithium present. Consistent with the Primary Chemistry Control Program.(4) solved Oxygen, at power            < 0.1 ppm (1) (2) (3) oride                              < 0.15 ppm oride                              < 0.10 ppm rogen                            25-50 cc (STP)/KgH2O pended Solids                      0.35 ppm prior to reactor startup Consistent with the Primary Chemistry Control Program.(4) on, as boric acid                  0-2620 ppm (5)
TES:
)  The temperature at which the Oxygen limit applies is > 250&deg;F.
)  The at power operation residual Oxygen concentration control value is  0.005 ppm.
)  During plant startup, Hydrazine may be used to control dissolved Oxygen concentration at 0.1 ppm.
)  During power operation lithium is coordinated with boron to maintain a pH(t) of  7.0, but 7.4, consistent with the Primary Chemistry Control Program. Lithium is added to the RCS during plant startup, but prior to reactor criticality, and is in specification per the Primary Chemistry Control Program within 72 hours after criticality. Lithium may be removed from the reactor coolant immediately before, or during, shutdown periods to aid in the cleanup of corrosion products. By evaluation, a maximum lithium concentration of 4.5 ppm is permissible with a target lithium concentration of 4.3 ppm for 100% power operations.
)  RCS boron concentration is maintained as necessary to ensure core reactivity or shutdown margin requirements are met. Although the RCS and related auxiliary systems containing reactor coolant are designed for a maximum concentration of 2620 ppm boron, it should be noted the design basis for the TSP baskets in the containment sump assumes the RCS, SITs, and RWST are at a maximum boron concentration of 2400 ppm.
3.3-25                                      Rev. 35


==3.4 REFERENCES==
MPS-2 FSAR FIGURE 3.3-1 FUEL ROD ASSEMBLY UPPER END CAP PLENUM SPRING DISHED PELLETS FUEL CLADDING 136.70 ACTIVE FUEL LENGTH 146.25
3.3-1ASME Boiler and Pressure Vessel Code, Section III, 1977 Edition, ASME New York, NY.3.3-2K. R. Merckx, "RODEX2 - Fuel Rod Th ermal-Mechanical Response Evaluation Model," XN-NF-81-58 (NP)(A), Revi sion 2, March 1985 and Supplements. 3.3-3"Qualification of Exxon Nuclear Fuel for Extended Burnup (PWR)," XN-NF-82-06 (NP)(A), Revision 1, Suppl ements 2, 4, 5, October 1985.3.3-4W. J. O'Donnel and B. F. Langer, "Fat igue Design Bases for Zircaloy Components,"
                                                        .440 CLADDING OD
Nuclear Science and Engineering, Volume 20, January 1964.
                                                        .3770 PELLET DIAMETER
3.3-5MATPRO Version, "A Handbook of Material Properties for Use in the Analysis of Light Water Reactor Fuel Rod Behavior," TREE-NUREG 1008, December 1976.3.3-6J. C. Winslow (CE) to T. J. Honan (NU), CE Letter, "Seismic and Branch Line LOCA Analysis of SPC Reload Fuel for Millstone 2," NU-88-043 (March 31, 1988).
                                                            .028 CLADDING WALL 136.70 ACTIVE FUEL LENGTH April 1998                          Rev. 24.8
MPS2 UFSAR3.3-21Rev. 353.3-7"PWR Primary Water Chemistry Guidelines
," Revision 2, Electric Power Research Institute (EPRI) Final Report, EPRI NP7077, dated November 1990.3.3-8ANF-88-88(P), Rev. 1, "Design Report for Millstone Point Unit 2 Reload ANF-1,"
August 29, 1988.3.3-9AREVA Contract Requirements Document Number 89-9070921-001-AREVA Contract No. J37MIL219B, January 28, 2008.
3.3-10AREVA Document 51-9074000-000, "Complianc e Document - Replacement Fuel Rod -
Millstone 2 Fuel Failure Mitigation," March 5, 2008.
MPS2 UFSAR3.3-22Rev. 35TABLE 3.3-1  MECHANICAL DESIGN PARAMETERS
*Fuel Assembly Geometry 14 by 14 Assembly Pitch, inches 8.180 Assembly Envelope, inches 8.160 Rod Pitch, inches 0.580Number of Grids per Assembly 9 Approximate Assembly Weight, lb.
1280/1313
* Fuel Rod to Fuel Rod Outside Dimension, inches 7.980 Fuel Rod and Pellet Clad OD, inches 0.440Clad thickness, inches 0.031/0.028
* Pellet Diameter, inches 0.3700/0.3770
* Pellet Length, inches 0.425/0.435 *Pellet Density (% Theoretical) 94.0/95.0/95.35
**Active Stack Length, Cold, inches 136.7Control Rod Guide Tube Number per assembly 4Tube ID, above dashpot, inches 1.035 Wall Thickness, inches 0.040Instrumentation Tube Number per Assembly 1 Tube ID, inches 1.035 Wall Thickness, inches 0.040 Spacer Grid Material Zircaloy-4 / Inconel-718 MPS2 UFSAR3.3-23Rev. 35Number per Assembly 8/1 for Batches N, P 9/0 for Batch R, S9/0 for Batch T - X


8/1 Batch Y and beyond Sleeves (Wear)
MPS-2 FSAR FIGURE 3.3-2A AREVA - RELOAD FUEL ASSEMBLY BATCH "S" AND PRIOR 9X TYPICAL SPACER UPPER TIE PLATE ASSEMBLY              LOWER TIE PLATE Rev. 24.8
Material SS/Chrome Plate Burnable Poison Rod Active Length, inches 124.7 + UO 2 blankets Material Gd 2 O 3 / U0 2 Pellet Diameter, inches 0.3700/0.3770 Clad Material Zircaloy-4 Clad ID, inches 0.378/0.384 Clad OD, inches 0.440 Clad Thickness, (nominal) inches 0.031/0.028Diametral Gap, (cold, nominal), inches 0.008/0.007


Pellet Length, inches 0.545Control Element Assembly Number 73 Number of Absorber Elements per Assembly 5 Type Cylindrical Rods Clad Material Inconel 625 Clad Thickness, inches 0.036 Clad OD, inches 0.948 Poison Material B 4C & Ag-In-CD Corner Element Pitch, inches 4.64Total CEA Length, inches 161.31- CE / 161.25 - AREVA Poison Length, inches 132 -CE / 133.5 - AREVACEA Dry Weight, lb. 95 - CE / 85 - AREVA MPS2 UFSAR3.3-24Rev. 35Total Operating Assembly Dry Weight, lb. Single 210 - CE / 200 - AREVADual 334 - CE / 314 AREVACore Arrangement Number of Fuel Assemblies in Core Total 217Number of Single CEAs 49 Number of Dual CEAs 12 CEA Pitch, minimum, inches 11.57 Spacing Between Fuel Assemblies,  Fuel Rod Surface to Surface, inches0.200
MPS-2 FSAR FIGURE 3.3-2B AREVA - RELOAD FUEL ASSEMBLY BATCH "T" AND LATER UPPER TIE PLATE                                            LOWER TIE PLATE Rev. 24.8


Spacing, Outer Fuel Rod Surface to Core Shroud, inches0.18 Hydraulic Diameter, Nominal Channel, feet 0.04445Total Flow Area (Excluding Guide Tubes), square feet 53.5Total Core Area, square feet 101.1 Core Equivalent Diameter, inches 136 Core Circumscribed Diameter, inches 143.1 Core Volume, liters 32,526 Total Fuel Loading, MTU (Typical) 83.65 Total Heat Transfer Area, square feet 50,117*Applicable to Batches N, P/applicable to Batch R and subsequent Batches.**Applicable to Batches N, P/applicable to Batc hes R, S/applicable to Batch T and subsequent Batches.
MPS-2 FSAR FIGURE 3.3-3A AREVA - RELOAD FUEL ASSEMBLY COMPONENTS BATCH "S" AND PRIOR 9X TYPICAL SPACER UPPER TIE PLATE ASSEMBLY             LOWER TIE PLATE Rev. 24.8
MPS2 UFSAR3.3-25Rev. 35TABLE 3.3-2  PRESSURIZED WATER REACTOR PRIMARY COOLANT WATER CHEMISTRY RECOMMENDED SPECIFICATIONS Conductivity (&#xb5;S/cm at 25
&deg;C) Relative to Lithium and Boron concentration.
pH at 25&deg;C Determined by the concentrati on of boric acid and lithium present. Consistent with th e Primary Chemistry Control Program.(4) Dissolved Oxygen, at power < 0.1 ppm (1) (2) (3)
Chloride < 0.15 ppm Fluoride < 0.10 ppm Hydrogen 25-50 cc (STP)/KgH 2 O Suspended Solids 0.35 ppm prior to reactor startup Li Consistent with the Prim ary Chemistry Control Program.
(4) Boron, as boric acid 0-2620 ppm (5) NOTES:(1)The temperature at which th e Oxygen limit applies is > 250
&deg;F.(2)The at power operation residual Oxyg en concentration control value is  0.005 ppm
.(3)During plant startup, Hydrazine may be used to control disso lved Oxygen concentration at  0.1 ppm.(4)During power operation lithium is c oordinated with boron to maintain a pH (t) of  7.0, but  7.4, consistent with the Primary Chemistry Control Program.
Lithium is added to the RCS during plant startup, but prior to reactor criticality, and is in specification per the Primary Chemistry Control Program within 72 hours after criticality. Lithium may be removed from the reactor coolant immediately before, or dur ing, shutdown periods to aid in the cleanup of corrosion products. By eval uation, a maximum lithium concentration of 4.5 ppm is permissible with a target lithium concentration of 4.3 ppm for 100% power operations.(5)RCS boron concentration is maintained as necessary to ensu re core reactivity or shutdown margin requirements are met. Although the RC S and related auxiliary systems containing reactor coolant are designed for a maximum concentration of 2620 ppm boron, it should be noted the design basis for the TSP baskets in the containment sump assumes the RCS, SITs, and RWST are at a maximu m boron concentration of 2400 ppm.
MPS-2 FSAR April 1998 Rev. 24.8 FIGURE 3.3-1 FUEL ROD ASSEMBLY UPPER END CAP PLENUM SPRING DISHED PELLETS FUEL CLADDING
.440 CLADDING OD
.3770 PELLET DIAMETER
.028 CLADDING WALL 136.70 ACTIVE FUEL LENGTH 136.70 ACTIVE FUEL LENGTH 146.25 MPS-2 FSAR Rev. 24.8FIGURE 3.3-2AAREVA -
RELOAD FUEL ASSEMBLY BATCH "S" AND PRIOR 9X TYPICAL SPACER UPPER TIE PLATE ASSEMBLY LOWER TIE PLATE MPS-2 FSAR Rev. 24.8FIGURE 3.3-2BAREVA - RELOAD FUEL ASSEMBLY BATCH "T" AND LATER UPPER TIE PLATE LOWER TIE PLATE MPS-2 FSAR Rev. 24.8FIGURE 3.3-3AAREVA - RELOAD FUEL ASSEM BLY COMPONENTS BATCH "S" AND PRIOR 9X TYPICAL SPACER UPPER TIE PLATE ASSEMBLY LOWER TIE PLATE MPS-2 FSAR Rev. 24.8FIGURE 3.3-3BAREVA - RELOAD FUEL ASSEM BLY COMPONENTS BATCH "T" AND LATER UPPER TIE PLATE SPACER 136.70 ACTIVE FUEL LENGTH FUEL ROD LOWER TIE PLATE MPS-2 FSAR Rev. 21FIGURE 3.3-4ABI-METALLIC FUEL SPACER ASSEMBLY GUIDE TUBE L OC ATI O N FUEL ROD SPACER SIDEPLATE SPRING STRIP MPS-2 FSAR Rev. 21FIGURE 3.3-4BHTP FUEL SPACER ASSEMBLY GUIDE TUBE FUEL ROD MPS-2 FSAR Rev. 26.2FIGURE 3.3-5FUEL ASSEMBLY HOLD DOWN DEVICE LOCKING NUT UPPER REACTION PLATE FUEL ALIGNMENT


PLATE SPRING UPPER TIE  
MPS-2 FSAR FIGURE 3.3-3B AREVA - RELOAD FUEL ASSEMBLY COMPONENTS BATCH "T" AND LATER FUEL ROD 136.70 ACTIVE FUEL LENGTH UPPER TIE PLATE SPACER                                    LOWER TIE PLATE Rev. 24.8


PLATE MPS-2 FSAR Rev. 30.2FIGURE 3.3-6CONTROL ELEMENT ASSEMBLY MPS-2 FSAR Rev. 30.2FIGURE 3.3-7CONTROL ELEMEN T ASSEMBLY MATERIALS MPS-2 FSAR April 1998 Rev. 26.2FIGURE 3.3-8CONTROL ELEMENT ASSEMBLIES GROUP AND NUMBER DESIGNATION MPS-2 FSARApril 1998Rev. 26.2FIGURE 3.3-9CORE ORIENTATION Outlet Nozzle Alignment Key 4 Equally Spaced Inlet Nozzle See Figure 3.3-8 for Identification of Core
MPS-2 FSAR FIGURE 3.3-4A BI-METALLIC FUEL SPACER ASSEMBLY FUEL ROD GUIDE TUBE LOCATION SPACER SIDEPLATE SPRING STRIP Rev. 21


Arrangement and CEA
MPS-2 FSAR FIGURE 3.3-4B HTP FUEL SPACER ASSEMBLY GUIDE TUBE FUEL ROD Rev. 21


Groups Fuel Assembly CEDM CEA Building North Reactor Vessel Core Support Barrel Elevation View MPS-2 FSARApril 1998Rev. 26.2FIGURE 3.3-10IN-CORE INSTRUMENTATION ASSEMBLY 90 180 270 0 MPS-2 FSARApril 1990Rev. 26.2FIGURE 3.3-11REACTOR INTERNALS ASSEMBLY Upper Guide Structure Support Plate CEA Shroud In-Core Instrumentation
MPS-2 FSAR FIGURE 3.3-5 FUEL ASSEMBLY HOLD DOWN DEVICE FUEL ALIGNMENT PLATE LOCKING NUT UPPER                                                    SPRING REACTION PLATE UPPER TIE PLATE Rev. 26.2


Guide Tube Core Support
MPS-2 FSAR FIGURE 3.3-6 CONTROL ELEMENT ASSEMBLY SPIDER IDENTIFICATION CEA Serial Number SPRING POISON ROD ASSEMBLY POISON MATERIAL 132 Inches Total - CE 133.5 Inches Total - AREVA Rev. 30.2


Barrel Core Support
MPS-2 FSAR FIGURE 3.3-7 CONTROL ELEMENT ASSEMBLY MATERIALS
                                  ~ 8" - CE 0"          ~ 12.5" - AREVA            ZONE A ZONE A                    ZONE B MATERIALS NUMBER                        ZONE A                ZONE B Ag              Ag    B        B 73 Ag                    B Ag              Ag    B        B B  B4 C      Ag  Ag In  Cd Rev. 30.2


Assembly Snubber Core Support Plate Core Shroud Fuel Aignment
MPS-2 FSAR FIGURE 3.3-8 CONTROL ELEMENT ASSEMBLIES GROUP AND NUMBER DESIGNATION NORTH ABC                D E            F    GHJKLMNPR                      S T          V      W X      Y
                                                          ! II I I II II 1                                            'I I          t    t      1            I 2                                                  a-S3      3-64 3                            A~5                        7-40          4-64        A4
                                              ,      * "-53                                  ,"
f' 4                      7.Q            A~          ,.~        1-30        ~...wi        '-IS A..c4                        2*21                    12-22 S                                                                                                y, A"
                                "I'                                                                      ~
6                    I~              &-15                                  ~1a          M7
                            ...52        2-20            J.7          5-4          1*1        2*23          4-55      1            7--                                                                  ~
8-                                                                *fJ
                ~    3081                                                8-8                                    U6  I---
                                                            "I'8-1 9-                    '-28                                                                "'1 10-                                                                                                          ~
i---      7*:tg                        5-3        ,.,          $-5                        7~'            270 0
          -ll -- ~
J.~1 ,..--
13 -        ~*SO          '*21                      [1-6        1*1                        '*32 14-                                                                " 1\
15 --            '-51        2-'11          B~  "      5-2            8-9        2*24          4-51      ""--
16                    lA-Q            &1-4                                              11.,,",
                                ,V                                                      ... "
                                                                                                            ~
17              A~                          2*"                      2*25 "lA-48 18                    t7-S9          If..c4        ",25      1-33        iA~          t7.u
                                              .-                                              I' 19                            ~              ~50        7*38          4.07        A-G 20                                                  ~-5B lUi 21
                                                          -1      I    1 I l              e      SOURCE: LOCATION April 1998                                                                  Rev. 26.2


Plate Aignment Pins Outlet Nozzle Alignment Key Expansion
MPS-2 FSAR FIGURE 3.3-9 CORE ORIENTATION Outlet Nozzle Alignment Key 4 Equally Spaced Inlet Nozzle See Figure 3.3-8 for Identification of Core Arrangement and CEA Groups Fuel Assembly CEDM Core Support Reactor Barrel Vessel CEA Building North                    Elevation View April 1998                                                  Rev. 26.2


Compensating Ring MPS-2 FSARApril 1990Rev. 24.8FIGURE 3.3-12PRESSURE VESSEL-CORE SUPPORT BARREL SNUBBER ASSEMBLY CENTER SUPPORT BARREL HARD-FACED SURFACE BOLT (12 REQ'D PER ASSEMBLY)CORE STABILIZING
MPS-2 FSAR FIGURE 3.3-10 IN-CORE INSTRUMENTATION ASSEMBLY 90 0
180 270 April 1998                Rev. 26.2


LUG PRESSURE VESSEL SNUBBER SPACER BLOCK SHIM (2 REQ'D PER ASSEMBLY)
MPS-2 FSAR FIGURE 3.3-11 REACTOR INTERNALS ASSEMBLY Expansion                                                    Upper Guide Structure Compensating Ring                                            Support Plate Alignment Key                                                  CEA Shroud In-Core Outlet Nozzle Instrumentation Guide Tube Aignment Pins Core Support Barrel Fuel Aignment Plate Core Shroud Core Support Plate Core Support Assembly Snubber April 1990                        Rev. 26.2
 
MPS-2 FSAR FIGURE 3.3-12 PRESSURE VESSEL-CORE SUPPORT BARREL SNUBBER ASSEMBLY CENTER SUPPORT BARREL HARD-FACED SURFACE CORE STABILIZING LUG BOLT (12 REQ'D PER                    SNUBBER SPACER BLOCK ASSEMBLY)
SHIM (2 REQ'D PER ASSEMBLY)
PIN (4 REQ'D PER ASSEMBLY)
PIN (4 REQ'D PER ASSEMBLY)
BOLT (4 REQ'D PER ASSEMBLY)
BOLT (4 REQ'D PER ASSEMBLY)
MPS-2 FSARApril 1990Rev. 24.8FIGURE 3.3-13CORE SHROUD ASSEMBLY MPS-2 FSARApril 1990 Rev. 26.2FIGURE 3.3-14UPPER GUIDE STRUCTURE ASSEMBLY MPS-2 FSARApril 1990 Rev. 26.2 FIGURE 3.3-15 CONTROL ELEMENT DRIVE MECHANISM (MAGNETIC JACK)
PRESSURE VESSEL April 1990                                  Rev. 24.8
MPS-2 FSARApril 1998 Rev. 26.2 FIGURE 3.3-16 (LEFT BLANK INTENTIONALLY)
 
MPS-2 FSAR Rev. 23.3FIGURE 3.3-17HEATED JUNCTION THERMOCOUPLE PROBE PRESSURE BO UNDARY INSTALLATION MPS-2 FSARApril 1990 Rev. 26.2 FIGURE 3.3-18 T YPICAL HEATED JUNCTION THERMOCOUPLE PR OBE ASSEMBLY INSTALLATION MPS-2 FSAR Rev. 27.4 FIGURE 3.3-19 PLACEMENT OF NATURAL URANIUM REPLACEMENT FUEL RODS AND FUEL ASSEMBLY ORIENTATION RELATIVE TO THE COR E BAFFLE FOR CYCLE 19 MPS2 UFSAR3.4-1Rev. 35
MPS-2 FSAR FIGURE 3.3-13 CORE SHROUD ASSEMBLY Upper Segment
.J Lower Segment April 1990                 Rev. 24.8
 
MPS-2 FSAR FIGURE 3.3-14 UPPER GUIDE STRUCTURE ASSEMBLY EXPANSION COMPENSAtING RING
* CEA SHROUD GRID ASSEMBLY CEA SHROUDS "FUEL ASSEMBLY AUGNMENT PLATE April 1990                 Rev. 26.2
 
MPS-2 FSAR FIGURE 3.3-15 CONTROL ELEMENT DRIVE MECHANISM (MAGNETIC JACK)
U~PPER 1iT~--ES:XTENS ION ItW,PrltTCH                              HAFT MAGNfT NG                              PIRESSURE HOUSING UPPER I .--~---- O'PERATI NG ROD SHROUD--~
ELECTRICAl CO~DUIT LIFT COil DRJVING LATCH '
COIL DRIVING LATCHES PULlDOWN COil MOTOR TUBE LOAD
                                        ~ANSFER OIL HOLDING U~TCH HOLDING                            COIL LATCHES PRESSURE HOUSING LOWER '
April 1990                  Rev. 26.2
 
MPS-2 FSAR FIGURE 3.3-16 (LEFT BLANK INTENTIONALLY)
April 1998            Rev. 26.2
 
MPS-2 FSAR FIGURE 3.3-17 HEATED JUNCTION THERMOCOUPLE PROBE PRESSURE BOUNDARY INSTALLATION Pressure Housing          Shroud Assembly Grayloc Coupling Detail W Rev. 23.3


===3.4 NUCLEAR===
MPS-2 FSAR FIGURE 3.3-18 TYPICAL HEATED JUNCTION THERMOCOUPLE PROBE ASSEMBLY INSTALLATION P\.CEOM NOZ2~:
DESIGN AND EVALUATION
L =~ CS EX rr' ::~.~ w
                                              - - - S~l.. ~~WCS iN I.C.!,
NOZZl.E (C*~ i't.ANiSi l.C.I. Pt.Ai !
L.OWe~!O PCStT10N
-U.G.5. SUPPC RT  ~\.Ai'a
                                            ~ROBe ASS!M8LY VACANT PART l!NG1'H--.I C!A SHROUt) ASS!MSI. Y c&#xa3;. --------..i-HOT LSG HJ"iC ScNSCR .
t.OC*~*nCN CIA P1.UG FUEL ALIGNMENT PUTi m
April 1990                              Rev. 26.2


====3.4.1 GENERAL====
MPS-2 FSAR FIGURE 3.3-19 PLACEMENT OF NATURAL URANIUM REPLACEMENT FUEL RODS AND FUEL ASSEMBLY ORIENTATION RELATIVE TO THE CORE BAFFLE FOR CYCLE 19 CORE BAFFLE              NATURAL URANIUM REPLACEMENT FILLER ROD ENRICHED FUEL ROD          CEA GUIDE TUBE Rev. 27.4
 
1   GENERAL  


==SUMMARY==
==SUMMARY==
This section summarizes the nuclear characteristics of the core a nd discusses the design parameters which are of significan ce to the performance of the core in normal transient and steady state operational conditions. A discussion of the nuclear design methods employed and comparisons with experiment s which support the use of these methods is included.
The numerical values presented ar e based on a representative core design. Sufficient analyses are completed each cycle to ensure that actual reload batches k eep operating parameters within design limits, accommodate essential reactivity require ments with the cont rol system provided, and meet other requirements for safe operation.
3.4.2 CORE DESCRIPTIONThe Millstone Unit 2 reactor consists of 217 assemblies, each having a 14 by 14 fuel rod array. The assemblies are composed of up to 176 fuel rods, four control rod guide tubes, and one center control rod guide tube/instrum ent tube. The fuel rods consist of slightly enriched UO 2 or UO 2-Gd 2 O 3 pellets inserted into Zircaloy tubes. The control rod guide tubes and instrument tubes are also made of Zircaloy. Each AREVA assembly contains nine spacers. A description of the AREVA supplied fuel design and design methods is contained in References 3.4-1, 3.4-2 and 3.4-3.
A representative loading pattern is shown in Figure 3.4-1 and is expressed in terms of previous cycle core locations and fuel assembly identifi ers. A summary of fuel characteristics for a representative core design is presented in Table 3.4-1. Figure 3.4-2 pres ents representative quarter core assembly movements. Representati ve beginning of cycle (BOC) and end of cycle (EOC) assembly exposures are shown in a quarter core representation in Figure 3.4-3.A representative low radial leakag e fuel management plan results in scatter loading of the fresh fuel throughout the core. Some fresh assemblies loaded in the co re interior contain gadolinia-bearing fuel in order to control power peaki ng and reduce the initial boron concentration to maintain the moderator temperature coefficient (MTC) within its Technical Specification limit.
The exposed fuel is also scatter loaded in th e center in a manner to control the power peaking.


====3.4.3 NUCLEAR====
s section summarizes the nuclear characteristics of the core and discusses the design ameters which are of significance to the performance of the core in normal transient and steady e operational conditions. A discussion of the nuclear design methods employed and parisons with experiments which support the use of these methods is included.
CORE DESIGN The nuclear design bases for core design are as follows:a.The design shall permit operation within the Te chnical Specificat ions for Millstone Unit 2 Nuclear Plant.b.The design Cycle length (EFP D) shall be determined on the basis of an estimated Cycle energy and previous Cycle energy window.
numerical values presented are based on a representative core design. Sufficient analyses are pleted each cycle to ensure that actual reload batches keep operating parameters within gn limits, accommodate essential reactivity requirements with the control system provided, meet other requirements for safe operation.
MPS2 UFSAR3.4-2Rev. 35c.The loading pattern shall be designed to achieve power distributions and control rod reactivity worths according to the following constraints:1.The peak linear heat rate (LHR) and the peaking factor Fr shall not exceed Technical Specifications limits in a ny single fuel rod throughout the cycle under nominal full power operating conditions.2.The SCRAM worth of all rods minus the most reactive rod s hall exceed the shutdown requirement.
2    CORE DESCRIPTION Millstone Unit 2 reactor consists of 217 assemblies, each having a 14 by 14 fuel rod array.
The neutronic design methods used to ensure the above requireme nts are consistent with those described in Reference 3.4-4.
assemblies are composed of up to 176 fuel rods, four control rod guide tubes, and one center trol rod guide tube/instrument tube. The fuel rods consist of slightly enriched UO2 or 2-Gd2O3 pellets inserted into Zircaloy tubes. The control rod guide tubes and instrument tubes also made of Zircaloy. Each AREVA assembly contains nine spacers. A description of the EVA supplied fuel design and design methods is contained in References 3.4-1, 3.4-2 and 3.4-epresentative loading pattern is shown in Figure 3.4-1 and is expressed in terms of previous le core locations and fuel assembly identifiers. A summary of fuel characteristics for a esentative core design is presented in Table 3.4-1. Figure 3.4-2 presents representative rter core assembly movements. Representative beginning of cycle (BOC) and end of cycle C) assembly exposures are shown in a quarter core representation in Figure 3.4-3.
3.4.3.1 Analytical Methodology The neutronics methods us ed in the core analysis are descri bed in Reference 3.4-4. The neutronic design analysis for each reload core is performed using the PRISM reactor simulator code. Full-core depletion calculations perf ormed with PRISM are used to determine the core wide power distribution in three dimensions and to r econstruct the individual rod power and burnup distributions. Thermal-hydraulic f eedback and axial exposure distribution effects are explicitly accounted for in the PRISM cal culations. The CASMO/MICBURN a ssembly depletion model is used to generate the microscopic cr oss section input to the PRISM code.
epresentative low radial leakage fuel management plan results in scatter loading of the fresh throughout the core. Some fresh assemblies loaded in the core interior contain gadolinia-ring fuel in order to control power peaking and reduce the initial boron concentration to ntain the moderator temperature coefficient (MTC) within its Technical Specification limit.
3.4.3.2 Physics Characteristics The neutronics characteristics of a representative reload core are presented in Table 3.4-2. The safety analysis for each cycl e is applicable for a specified previous cycle energy window. A representative HFP letdown curve is shown in Figure 3.4-4.
exposed fuel is also scatter loaded in the center in a manner to control the power peaking.
3    NUCLEAR CORE DESIGN nuclear design bases for core design are as follows:
: a. The design shall permit operation within the Technical Specifications for Millstone Unit 2 Nuclear Plant.
: b. The design Cycle length (EFPD) shall be determined on the basis of an estimated Cycle energy and previous Cycle energy window.
3.4-1                                    Rev. 35
: 1. The peak linear heat rate (LHR) and the peaking factor Fr shall not exceed Technical Specifications limits in any single fuel rod throughout the cycle under nominal full power operating conditions.
: 2. The SCRAM worth of all rods minus the most reactive rod shall exceed the shutdown requirement.
neutronic design methods used to ensure the above requirements are consistent with those cribed in Reference 3.4-4.
3.1   Analytical Methodology neutronics methods used in the core analysis are described in Reference 3.4-4. The neutronic gn analysis for each reload core is performed using the PRISM reactor simulator code. Full-depletion calculations performed with PRISM are used to determine the core wide power ribution in three dimensions and to reconstruct the individual rod power and burnup ributions. Thermal-hydraulic feedback and axial exposure distribution effects are explicitly ounted for in the PRISM calculations. The CASMO/MICBURN assembly depletion model is d to generate the microscopic cross section input to the PRISM code.
3.2   Physics Characteristics neutronics characteristics of a representative reload core are presented in Table 3.4-2. The ty analysis for each cycle is applicable for a specified previous cycle energy window. A esentative HFP letdown curve is shown in Figure 3.4-4.
3.2.1    Power Distribution Considerations resentative calculated power maps are shown in Figures 3.4-5 and 3.4-6 for BOC uilibrium xenon), and EOC conditions, respectively. The power distributions were obtained m a three-dimensional neutronics model with moderator density and Doppler feedback effects rporated. The Technical Specification limits on Fr and LHR are 1.69 and 15.1 kW/ft, ectively.
3.2.2    Control Rod Reactivity Requirements epresentative shutdown margin evaluation is given in Table 3.4-3. The Millstone Unit 2 hnical Specifications require a minimum shutdown margin of 3,600 pcm.
3.4-2                                    Rev. 35
 
Technical Specifications require that the MTC be less than +7 pcm/&deg;F at or below 70 percent ated thermal power, less than +4 pcm/&deg;F above 70 percent power and greater than -32 pcm/&deg;F 00 percent of rated thermal power. Representative MTC calculation results are presented in le 3.4-2.
4    POST-RELOAD STARTUP TESTING tup tests will be performed at the beginning of each reload cycle to obtain the as-built core racteristics and to verify Technical Specification and core physics design parameters. The ad startup physics test program is based on ANSI-19.6-1 (Reference 3.4-9). The Startup Test ivity Reduction (STAR) Program (Reference 3.4-10) provides an alternative to the ANSI-6-1 test program provided that specific criteria for the reload core design and construction are sfied. The STAR Program criteria are established in station procedures and include additional licability requirements for core design, fuel and control element assembly (CEA) fabrication, A lifetime monitoring, refueling and startup testing.
reload startup physics test program shall consist of the following:
: a.      Critical Boron Concentration - HZP, Control Rods Withdrawn.
: b.      Critical Boron Concentration - HZP, Control Rod Group(s) of at least 1%
reactivity are fully inserted in the core. 1
: c.      Control Rod Group Worths - HZP, two or more control rod groups shall be measured which are well distributed radially and represent a predicted total worth of at least 3% reactivity. 1
: d.      Isothermal Temperature Coefficient - HZP.
: e.      Flux Symmetry - between 0 and 30% of full power.
: f.      Power Distribution - between 40 and 75% of full power.
: g.      Isothermal Temperature Coefficient - greater than 70% of full power.
: h.      Power Distribution - greater than 90% of full power.
: i.      Critical Boron Concentration - greater than 90% of full power.
: j.      HZP to full power reactivity difference.
his test may be eliminated if performing the STAR Program per Reference 3.4-10.
3.4-3                                  Rev. 35
 
5.1    General on induced spatial oscillations on the Millstone Unit 2 core fall into three classes or modes.
se are referred to as axial oscillations, azimuthal oscillations, and radial oscillations. An axial llation is one in which the axial power distribution periodically shifts to the top and bottom of core. An azimuthal oscillation is one in which the X-Y power distribution periodically shifts m one side of the core to the other. A radial oscillation is one in which the X-Y power ribution periodically shifts inward and outward from the center of the core to the periphery.
on stability analyses indicate that a number of general statements can be made:
: a.      The time scale on which the oscillations occur is long, and any induced oscillations typically exhibit a period of 25 to 30 hours.
: b.      As long as the initial power peaking associated with the perturbation initiating the oscillation is within the limiting conditions for operation, specified acceptable fuel design limits will not be approached for a period of hours allowing an operator time to decide upon and take appropriate remedial action prior to the time when allowable peaking factors would be exceeded.
: c.      The core will be stable to radial mode oscillations at all times in the burnup cycle.
: d.      The core will be stable to azimuthal mode oscillations at all times in the burnup cycle.
: e.      All possible modes of undamped oscillations can be detected by both exactor and in-core instrumentation as discussed below.
5.2    Detection of Oscillations mary reliance for the detection of any xenon oscillations is placed on the exactor flux nitoring instrumentation. The power range excore neutron detectors (one axial pair per drant) are used to monitor the symmetry of power distributions and are located at distinct muthal and axial positions. These detectors are sensitive primarily to the power density ations produced by peripheral fuel assemblies in the vicinity of the detectors. All possible on induced spatial oscillations will affect the power densities of the peripheral fuel assemblies he core.
ddition, the in-core instrumentation provides information which will be used in the early es of cycle operation to confirm predicted correlations between indications from the excore ctors and the space-dependent flux distribution within the core. Later on, during normal ration, the in-core detector system provides information which may be used to supplement that ilable from the excore detectors.
3.4-4                                      Rev. 35


3.4.3.2.1 Power Distribution Considerations Representative calculated power maps are shown in Figures 3.4-5 and 3.4-6 for BOC (equilibrium xenon), and EOC conditions, respectively. The power distributions were obtained from a three-dimensional neutronics model with moderator density and Doppler feedback effects incorporated. The Technical Specification limi ts on Fr and LHR are 1.69 and 15.1 kW/ft, respectively.
ce the reactor will not be operated under conditions that imply instability with respect to muthal xenon oscillation, no special protective system features are needed to accommodate muthal mode oscillations. Regardless, a maximum azimuthal power tilt is prescribed in the hnical Specifications along with prescribed operating restrictions in the event that the muthal power tilt limit is exceeded.
described earlier, the power range excore neutron detectors are used to monitor the azimuthal metry of the power distributions since they are located at distinct locations in the X-Y plane.
uld the excore detectors indicate different readings in the azimuthal direction, a tilt in the core er distribution would be indicated. When the tilt exceeds a preset magnitude an alarm will ur. In the event of an alarm, the orientation of the tilt will be determined and, on the basis of ntation, the proper CEAs will be manually adjusted to reduce the magnitude of the tilt.
features provided for azimuthal xenon oscillation control are:
: a.     instrumentation for monitoring azimuthal power tilt.
: b.      administrative limits on azimuthal power tilt.
excore detectors are used to monitor the axial power distribution and to detect deviations m the equilibrium distribution such as those which would occur during an axial xenon llation. This is done by monitoring variations in the external axial shape index, a parameter ved from the excore detector readings which is related to the axial power distribution. Control xial xenon oscillation is accomplished utilizing Regulating Bank 7. When it is determined that axial shape index may exceed the boundaries of a specified control band about the equilibrium e, this bank is slowly inserted and eventually withdrawn over a period of several hours. The is then stabilized until a new oscillation develops.
features provided for axial xenon control and protection are:
: a.     equipment for monitoring axial shape index.
: b.     administrative limits on axial power distribution, external axial shape index.
: c.      an axial shape index reactor trip (local power density - high).
: d.      use of Regulating Bank 7 for control of axial power distribution.
3.4-5                                      Rev. 35


3.4.3.2.2 Control Rod Reactivity Requirements A representative shutdown margin evaluation is given in T able 3.4-3. The Millstone Unit 2 Technical Specifications require a minimum shutdown margin of 3,600 pcm.
ent core designs for Millstone Unit 2 (Cycles 10 and beyond) have been developed to include ger fuel cycles along with low radial leakage fuel management. These current designs scatter fresh fuel assemblies throughout the interior of the core with the highest burnup fuel mblies being loaded along the core periphery. Core designs prior to Cycle 10 operation were of a low radial leakage design due to the loading of fresh fuel assemblies along the core phery.
MPS2 UFSAR3.4-3Rev. 35 3.4.3.2.3 Moderator Temperature Coefficient ConsiderationsThe Technical Specifications require that the MTC be less than +7 pcm/
h respect to xenon oscillations in the radial and azimuthal directions, studies indicate that core gns of a low radial leakage design (i.e., highest burnup assemblies loaded on the core phery with fresh fuel assemblies scatter loaded about the core interior) are more stable than e designs which load fresh fuel assemblies along the core periphery. Therefore, the clusions regarding xenon oscillations in the radial and azimuthal directions, which are ented in Section 3.4.5.5, remain applicable to current plant operations.
&deg;F at or below 70 percent of rated thermal power, less than +4 pcm/
h regard to axial xenon oscillations, the core near end-of-cycle may be naturally unstable in absence of any control rod action even if low leakage core designs are utilized. But axial on oscillations are sufficiently slow (the period of oscillation being 25 to 30 hours) so that e would be sufficient time to control the oscillations. In addition, automatic protection is vided if operator action is not taken to remedy the situation. Regulating Bank 7 CEAs are zed for controlling axial xenon oscillations.
&deg;F above 70 percent power a nd greater than -32 pcm/
5.5    Method of Analysis classic method for assessing spatial xenon oscillations is that developed by Randall and St.
&deg;F at 100 percent of rated thermal power. Representa tive MTC calculation resu lts are presented in Table 3.4-2.
n (Reference 3.4-5) which consists of expanding small perturbations of the flux and xenon centrations about equilibrium values in eigenfunctions of the system with equilibrium xenon ent. However, it is necessary to extend this simple linear analysis to treat cores which are uniform because of fuel zoning, depletion, and CEA patterns, for example. Such extensions e been worked out and are reported in References 3.4-6 and 3.4-8. In this extension, the nvalue separations between the excited state of interest and the fundamental are computed erically for symmetrical flux shapes. For nonsymmetrical flux shapes, the eigenvalue aration can usually be obtained indirectly from the dominance ratio 1/0, computed during iteration cycle of the spatial calculation.
3.4.4 POST-RELOAD STARTUP TESTINGStartup tests will be performed at the beginning of each reload cy cle to obtain the as-built core characteristics and to verify Technical Specification and core physics design parameters. The reload startup physics test progr am is based on ANSI-19.6-1 (Ref erence 3.4-9). The Startup Test Activity Reduction (STAR) Program (Reference 3.4-10) provides an alternative to the ANSI-19.6-1 test program provided that sp ecific criteria for the reload core design and construction are satisfied. The STAR Program criteria are established in station procedures and include additional applicability require ments for core design, fuel and contro l element assembly (CEA) fabrication, CEA lifetime monitoring, refu eling and startup testing.
merical space time calculations are performed in the required number of spatial dimensions for various modes as checkpoints for the predictions for the extended Randall-St. John treatment cribed above.
The reload startup physics test progr am shall consist of the following:a.Critical Boron Concentration - HZP, Control Rods Withdrawn.b.Critical Boron Concentration - HZP, Control Rod Group(s) of at least 1% reactivity are fully inserted in the core.
3.4-6                                    Rev. 35
1c.Control Rod Group Worths - HZP, two or more control rod groups shall be measured which are well dist ributed radially and repres ent a predicted total worth of at least 3% reactivity.
1 d.Isothermal Temperature Coefficient - HZP.e.Flux Symmetry - between 0 and 30% of full power.f.Power Distribution - between 40 and 75% of full power.
g.Isothermal Temperature Coefficient - greater than 70% of full power.
h.Power Distribution - greater than 90% of full power.
i.Critical Boron Concentration - greater than 90% of full power.
j.HZP to full power reactivity difference.
1.This test may be eliminated if performing the STAR Program per Reference 3.4-10.
MPS2 UFSAR3.4-4Rev. 35


====3.4.5 REACTOR====
confirm that the radial oscillation mode is extremely stable, a space-time calculation was run a reflected, zoned core 11 feet in diameter without including the damping effects of the ative power coefficient. The initial perturbation was a poison worth of 0.4 percent in reactivity ed in the central 20 percent in the core for 1 hour. Following removal of the perturbation, the lting oscillation was followed in 4-hour time steps for a period of 80 hours. Results show that oscillation died out very rapidly with a damping factor of about minus 0.06 per hour. When damping coefficient is corrected for a finite time mesh by the formula in Reference 3.4-7, it is e strongly convergent. On this basis, it is concluded that radial oscillation instability will not ur.
STABILITY 3.4.5.1 General Xenon induced spatial oscillations on the Millstone Unit 2 core fall into three classes or modes. These are referred to as axial os cillations, azimuthal oscillations, and radial oscillations. An axial oscillation is one in which the axial power distri bution periodically shifts to the top and bottom of the core. An azimuthal oscillation is one in which the X-Y power distribution periodically shifts from one side of the core to the other. A ra dial oscillation is one in which the X-Y power distribution periodically shifts inward and outward from the center of the core to the periphery.
s conclusion is of particular significance because it means that there is no type of oscillation re the inner portions of the core act independently of the peripheral portions of the core whose avior is most closely followed by the excore flux detectors. Radial mode oscillations, even ugh highly damped, would be manifested as periodic variation in the excore flux power signal le the delta-T power signals remained constant. Primary reliance is placed on the excore flux ctors for the detection of any xenon oscillations.
Xenon stability analyses indicate that a numb er of general statem ents can be made:a.The time scale on which the oscillations occur is long, and a ny induced oscillations typically exhibit a pe riod of 25 to 30 hours.b.As long as the initial power peaking asso ciated with the pert urbation initiating the oscillation is within the limiting conditions for operation, specified acceptable fuel design limits will not be approached for a period of hours allowing an operator time to decide upon and take appropriate remedial action prior to the time when allowable peaking factors would be exceeded.c.The core will be stable to radial mode oscillations at all tim es in the burnup cycle.d.The core will be stable to azimuthal m ode oscillations at all times in the burnup cycle.e.All possible modes of undamped oscillations can be detected by both exactor and in-core instrumentation as discussed below.
5.5.2    Azimuthal Xenon Oscillations lyses indicate that the eigenvalue separation between the first asimuthal harmonic and the damental is about 0.86 percent in . The calculated damping coefficient for the first azimuthal de is minus 0.016 per hour, and the higher modes will be even more strongly damped.
3.4.5.2 Detection of Oscillations Primary reliance for the det ection of any xenon oscillations is placed on the exactor flux monitoring instrumentation. The power range exco re neutron detectors (one axial pair per quadrant) are us ed to monitor the symmetry of power distributions and are located at distinct azimuthal and axial positions. These detectors are sensitive primarily to the power density variations produced by peripheral fu el assemblies in the vicinity of the detectors. All possible xenon induced spatial oscillations will affect the power densitie s of the peripheral fuel assemblies in the core.
thermore, the Doppler coefficient applicable to the Millstone Unit 2 reactor is calculated to be roximately minus 1.36 x 10-3 /(kW/ft) which is sufficiently negative to ensure stability of he azimuthal modes.
5.5.3    Axial Xenon Oscillations checkpoints for the predictions for the modified Randall-St. John approach, numerical spatial e calculations have been performed for the axial case at both beginning and end-of-cycle. The and poison burnup distributions were obtained by depletion with soluble boron control so that power distribution was strongly flattened. Spatial Doppler feedback was included in these ulations. The initial perturbation used to excite the oscillations was a 50 percent insertion into top of the core of a 1.5 percent reactivity CEA bank for 1 hour. The damping factor for this was calculated to be about +0.02 per hour; however, when corrected for finite time mesh rvals by the methods of Reference 3.4-7, the damping factor is increased to approximately
: 04. When this damping factor is plotted at the appropriate eigenvalue separation for this mode nd-of-cycle, it is apparent that good agreement is obtained with the modified Randall-St. John diction.
3.4-7                                      Rev. 35


In addition, the in-core instrumentation provides information which will be used in the early stages of cycle operation to confirm predicted correlations between indications from the excore detectors and the space-depende nt flux distribution within th e core. Later on, during normal operation, the in-core detector syst em provides information which ma y be used to supplement that available from the excore detectors.
s result suggests that the constant power condition which applies to the axial oscillations lts in a very weak moderator feedback since the moderator density distribution is fixed at the and bottom of the core and only the density distribution in between can change.
MPS2 UFSAR3.4-5Rev. 35 3.4.5.3 Control of Oscillations Since the reactor will not be ope rated under conditions that imply instability with respect to azimuthal xenon oscillation, no sp ecial protective system features are needed to accommodate asimuthal mode oscillations. Regardless, a maximu m azimuthal power tilt is prescribed in the Te chnical Specifications along with prescribed operating re strictions in the event that the azimuthal power tilt limit is exceeded.As described earlier, the power range excore neutron detectors ar e used to monitor the azimuthal symmetry of the power distributions since they are located at distinct locations in the X-Y plane. Should the excore detectors indicate different readings in the azimuthal direct ion, a tilt in the core power distribution would be indicated. When the tilt exceeds a preset magnitude an alarm will occur. In the event of an alarm, the orientation of the tilt will be determined and, on the basis of orientation, the proper CEA's will be manually adjusted to redu ce the magnitude of the tilt.
the calculated Doppler coefficient of minus 1.36 x 10-3 /(kW/ft), the damping factor toward end of the burnup cycle is positive. Thus, within the uncertainties in predicting power fficients and uncertainties in the analyses, there is a prediction of unstable axial xenon llations in the absence of any control action. These oscillations are sufficiently slow (the od of oscillation being 25 to 30 hours) so that there would be sufficient time to control the llations. In addition, automatic protection is provided if operator action is not taken to remedy situation. Regulating Bank 7 CEAs are utilized for controlling axial xenon oscillations.
The features provided for azimuthal xenon osci llation control are:a.instrumentation for monitoring azimuthal power tilt.
6    REFERENCES 1    Generic Mechanical Design Report Exxon Nuclear 14 x 14 Fuel Assemblies for Combustion Engineering Reactors, XN-NF-82-09(A), Exxon Nuclear Company, Richland, WA 99352, November 1982.
b.administrative limits on azimuthal power tilt.The excore detectors are used to monitor the ax ial power distribution and to detect deviations from the equilibrium distribut ion such as those which woul d occur during an axial xenon oscillation. This is done by monito ring variations in the external axial shape index, a parameter derived from the excore detector readings which is related to the axial power distribution. Control of axial xenon oscillation is accomp lished utilizing Regulating Bank 7. When it is determined that the axial shape index may exceed the boundaries of a sp ecified control band about the equilibrium value, this bank is slowly in serted and eventually withdrawn over a period of several hours. The core is then stabilized until a new oscillation develops.
2    Design Report for Millstone Point Unit 2 Reload ANF-1, ANF-88-088(P), Rev. 1, Advanced Nuclear Fuels Corporation, Richland, WA 99352, August 1988.
The features provided for axial xenon control and protection are:a.equipment for monitoring axial shape index.
3    Millstone Unit 2 Mechanical Design Report for Increased Peaking EMF-91-245(P),
b.administrative limits on axial power distribution, external axial shape index.c.an axial shape index reactor trip (local power density - high).d.use of Regulating Bank 7 for cont rol of axial power distribution.
Siemens Nuclear Power Corporation, January 1992.
MPS2 UFSAR3.4-6Rev. 35 3.4.5.4 Operating Experience Recent core designs for Millstone Unit 2 (Cycles 10 and beyond) ha ve been developed to include longer fuel cycles along with low radial leakage fuel management. These current designs scatter load fresh fuel assemblies throughout the interi or of the core with the highest burnup fuel assemblies being loaded along the core periphery. Core designs prior to Cycle 10 operation were not of a low radial leakage desi gn due to the loading of fresh fu el assemblies along the core periphery.With respect to xenon oscillations in the radial and azimuthal direct ions, studies indicate that core designs of a low radial leak age design (i.e., highest burnup a ssemblies loaded on the core periphery with fresh fuel assemblies scatter loaded about the core interior) are more stable than those designs which load fres h fuel assemblies along the core periphery. Therefore, the conclusions regarding xenon oscillations in th e radial and azimuthal directions, which are presented in Section 3.4.5.5, remain app licable to current plant operations.With regard to axial xenon oscillations, the core near end-of-cycle may be naturally unstable in the absence of any control rod action even if low leakage core designs are utilized. But axial xenon oscillations are sufficiently slow (the period of osc illation being 25 to 30 hours) so that there would be sufficient time to control the oscillations. In addition, automatic protection is provided if operator action is not taken to remedy the situati on. Regulating Bank 7 CEA's are utilized for controlling axial xenon oscillations.
4     EMF-96-029(P)(A) Volumes 1 and 2, Reactor Analysis System for PWRs Volume 1 -
3.4.5.5 Method of Analysis The classic method for assessing spatial xenon oscillations is that developed by Randall and St. John (Reference 3.4-5) which c onsists of expanding small pert urbations of the flux and xenon concentrations about equilibrium values in eigenfunctions of th e system with equilibrium xenon present. However, it is necessary to extend this simple linear analysis to treat cores which are nonuniform because of fuel zoning, depletion, and CEA patterns, fo r example. Such extensions have been worked out and are reported in Re ferences 3.4-6 and 3.4-8. In this extension, the eigenvalue separations between th e excited state of interest a nd the fundamental are computed numerically for symmetrical fl ux shapes. For nonsymmetrical flux shapes, the eigenvalue separation can usually be obtained indirectly from the dominance ratio 1/0 , computed during the iteration cycle of the spatial calculation.
Mehodology Description, Volume 2 - Benchmarking Results, Siemens Power Corporation, January 1997.
Numerical space time calcu lations are performed in the required number of spatial dimensions for the various modes as checkpoints for the predictions for the extended Randall-St.
5    Randall, D., Xenon Spatial Oscillations, Nucleonics, 16, 3, pages 82-86 (1958).
John treatment described above.
6    Stacey, Jr., W. M., Linear Analysis of Xenon Spatial Oscillations, Nuclear Sci. Eng.,
MPS2 UFSAR3.4-7Rev. 35 3.4.5.5.1 Radial Xenon OscillationsTo confirm that the radial oscillation mode is extremely stable, a space-time calculation was run for a reflected, zoned core 1 1 feet in diameter without including the damping effects of the negative power coefficient. The initial perturbati on was a poison worth of 0.4 percent in reactivity placed in the central 20 percent in the core for 1 hour. Following removal of the perturbation, the resulting oscillation was fo llowed in 4-hour time steps for a pe riod of 80 hours. Results show that the oscillation died out very rapidly with a da mping factor of about minus 0.06 per hour. When this damping coefficient is correct ed for a finite time mesh by the formula in Reference 3.4-7, it is more strongly convergent. On this ba sis, it is concluded that radial oscillation instability will not occur.This conclusion is of particular significance because it means that there is no type of oscillation where the inner portions of the core act independently of the periphe ral portions of the core whose behavior is most closely followe d by the excore flux detectors. Ra dial mode oscillations, even though highly damped, would be mani fested as periodic variation in the excore flux power signal while the delta-T power signals re mained constant. Primary reliance is placed on the excore flux detectors for the detection of any xenon oscillations.
30, pages 453-455 (1967).
3.4.5.5.2 Azimuthal Xenon Oscillations Analyses indicate that the eige nvalue separation between the first asimut hal harmonic and the fundamental is a bout 0.86 percent in . The calculated damping coefficient for the first azimuthal mode is minus 0.016 per hour, and the higher modes will be even more strongly damped. Furthermore, the Doppler coefficien t applicable to the Mi llstone Unit 2 reactor is calculated to be approximately minus 1.36 x 10
7    Poncelet, C. G., The Effect of a Finite Time Step Length on Calculated Spatial Xenon Stability Characteristics in Large PWR's Trans. ANS, 10, 2, page 571 (1967).
-3 /(kW/ft) which is sufficiently negative to ensure stability of all the azimuthal modes.
8    CEND-TP-26., Diatch, P.B.
3.4.5.5.3 Axial Xenon Oscillations As checkpoints for the predictions for the modified Randall-St. John appr oach, numerical spatial time calculations have been pe rformed for the axial case at bot h beginning and end-of-cycle. The fuel and poison burnup distributions were obtained by depletion with soluble boron control so that the power distribution was strongly flattened. Spatial Doppler fee dback was included in these calculations. The initial perturbati on used to excite the oscillations was a 50 percent insertion into the top of the core of a 1.5 percent reactivity CEA bank for 1 hour. The damping factor for this case was calculated to be about +0.02 per hour; however, when co rrected for fin ite time mesh intervals by the methods of Reference 3.4-7, the da mping factor is increased to approximately
9    ANSI/ANS-19.6-1 Reload Startup Physics Tests for Pressurized Water Reactors, 2005.
+0.04. When this damping factor is plotted at the appropriate eige nvalue separation for this mode at end-of-cycle, it is apparent that good agreement is obtained with the modified Randall-St. John prediction.
10   WCAP-16011-P-A, Revision 0, Startup Test Activity Reduction Program, February 2005.
MPS2 UFSAR3.4-8Rev. 35Calculations performed with both Doppler and moderator reactivity feedback have resulted in damping factors which are essentially the same as those obtained with Doppler feedback alone.
3.4-8                                      Rev. 35
This result suggests that the constant power condition which a pplies to the axial oscillations results in a very weak moderator feedback since the moderator density dist ribution is fixed at the top and bottom of the core and only the de nsity distribution in between can change.
For the calculated Doppler coefficient of minus 1.36 x 10
-3 /(kW/ft), the damping factor toward the end of the burnup cycle is positive. Thus, within the uncertainties in predicting power coefficients and uncertainties in the analyses, there is a pr ediction of unstable axial xenon oscillations in the absence of any control action. These oscillations are sufficiently slow (the period of oscillation being 25 to 30 hours) so that there would be sufficient time to control the oscillations. In addition, automati c protection is provided if operato r action is not taken to remedy the situation. Regulating Bank 7 CEA's are utili zed for controlling axial xenon oscillations.
3.


==4.6 REFERENCES==
Fuel Types        N1    N2    N3    N4    P1    P2    P3    P4    P5    R1    R2    R3    R4    R5    R Central Zone Assem- 3.94  3.90  3.87  3.82   3.87  3.86  3.84  3.81  3.76  4.49  4.49  4.47  4.39  4.33  4.42 bly Average Enrich-ment (w/o)
3.4-1"Generic Mechanical Design Report E xxon Nuclear 14 x 14 Fuel Assemblies for Combustion Engineering Reactors," XN-NF-82-09(A), Exxon Nuclear Company, Richland, WA 99352, November 1982.3.4-2"Design Report for Millstone Point Unit 2 Reload ANF-1," ANF-88-088(P), Rev. 1, Advanced Nuclear Fuels Corpor ation, Richland, WA 99352, August 1988.3.4-3"Millstone Unit 2 Mechanical Design Report for Increased Peaking" EMF-91-245(P), Siemens Nuclear Power Corporation, January 1992.3.4-4EMF-96-029(P)(A) Volumes 1 and 2, "Reactor Analysis System for PWRs Volume 1 -
Number Gadolinia    0      6      12    16    0      4      8      12    16    0      4      8      12    16    12 Bearing Rods Nominal Density (%  94    94    94    94    94    94    94    94    94    95    95    95    95    95    95 TD)
Mehodology Description, V olume 2 - Benc hmarking Results", Siemens Power Corporation, January 1997.3.4-5Randall, D., "Xenon Spatial Oscillations," Nucleonics, 16, 3, pages 82-86 (1958).3.4-6Stacey, Jr., W. M., "Linear Analysis of Xenon Spatial Oscillations
Pellet OD (inches)  0.370  0.370  0.370  0.370  0.370  0.370  0.370  0.370  0.370  0.377  0.377  0.377  0.377  0.377  0.37 Clad OD (inches)   0.440  0.440  0.440  0.440  0.440  0.440  0.440  0.440  0.440  0.440  0.440  0.440  0.440  0.440  0.44 Diametral Gap      0.0080 0.0080 0.0080 0.0080 0.0080 0.0080 0.0080 0.0080 0.0080 0.007  0.007  0.007  0.007  0.007  0.00 (inches)
," Nuclear Sci. Eng., 30, pages 453-455 (1967).3.4-7Poncelet, C. G., "The Effect of a Finite Time Step Length on Calculated Spatial Xenon Stability Characteristics in Large PWR's" Trans. ANS, 10, 2, page 571 (1967).3.4-8CEND-TP-26., Diatch, P.B.3.4-9ANSI/ANS-19.6-1 "Reload Startup Physics Tests for Pressurized Water Reactors,"
Clad Thickness      0.031  0.031  0.031  0.031  0.031  0.031  0.031  0.031  0.031  0.028  0.028  0.028  0.028  0.028  0.02 (inches)
2005.3.4-10WCAP-16011-P-A, Revision 0, "Startup Te st Activity Reduction Program," February 2005.
Rod Pitch (inches) 0.580  0.580  0.580  0.580  0.580  0.580  0.580  0.580  0.580  0.580  0.580  0.580  0.580  0.580  0.58 Spacer Material    Bime-  Bime-  Bime-  Bime-  Bime-  Bime-  Bime-  Bime- Bime- Bime- Bime- Bime- Bime- Bime- Bim tallic tallic tallic tallic tallic tallic tallic tallic tallic tallic tallic tallic tallic tallic talli Fuel Supplier      AREVA AREVA AREVA AREVA AREVA AREVA AREVA AREVA AREVA AREVA AREVA AREVA AREVA AREVA AR Fuel Stack Height  136.7  136.7  136.7  136.7  136.7  136.7  136.7  136.7  136.7  136.7  136.7  136.7  136.7  136.7 136.
MPS2 UFSARMPS2 UFSAR3.4-9Rev. 35TABLE 3.4-1 FUEL CHARACTERISTICS FOR A REPRESENTATIVE RELOAD COREFuel TypesN1N2N3N4P1P2P3P4P5R1R2R3R4R5R6Central Zone Assem-bly Average Enrich-ment (w/o)3.943.903.873.823.873.863.843.813.764.494.494.474.394.334.42Number Gadolinia
Nominal (inches)
Number of Assem-   8      20    8      25    8      8      12    8      36    8      8      8      8      48    4 blies Regionwise Loading  3.04  7.60  3.03  9.43  3.04  3.04  4.55  3.03  13.58  3.19  3.19  3.19  3.17  18.98  1.59 (MTU) 3.4-9                                                  Rev


Bearing Rods0612160481216048121612Nominal Density (%
RELOAD CORE
              <characteristic>                      BOC      EOC tical Boron (ppm): HZP, ARO, No Xenon            1453    ---
tical Boron (ppm): HFP, ARO, Equilibrium        1024    0 non derator Temperature Coefficient (pcm/&deg;F):        +2.0    -10.4 P
derator Temperature Coefficient (pcm/&deg;F):        -6.0    -23.3 P
ppler Coefficient (pcm/&deg;F)                      -1.17  -1.33 ron Worth (pcm/ppm): HZP                          -8.8    -10.8 ron Worth (pcm/ppm): HFP                          -8.4    -10.4 R (kW/ft) HFP (a)                                12.8    11.6 layed Neutron Fraction                            0.0064  0.0054 P, PDIL Worth (pcm)                              157    241 1 Rod Worth, HZP (pcm)                            6271    7696 cess Shutdown Margin (pcm): HFP                  124    323 cess Shutdown Margin (pcm): HZP                  140    751 Including uncertainties.
3.4-10                  Rev. 35


TD)949494949494949494959595959595Pellet OD (inches)0.3700.3700.3700.3700.3700.3700.3700.3700.3700.3770.3770.3770.3770.3770.377Clad OD (inches)0.4400.4400.4400.4400.4400.4400.4400.4400.4400.4400.4400.4400.4400.4400.440 Diametral Gap (inches)0.00800.00800.00800.00800.00800.00800.00800.00800.00800.0070.0070.0070.0070.0070.007Clad Thickness (inches)0.0310.0310.0310.0310.0310.0310.0310.0310.0310.0280.0280.0280.0280.0280.028Rod Pitch (inches) 0.5800.5800.5800.5800.5800.5800.5800.5800.5800.5800.5800.5800.5800.5800.580Spacer MaterialBime-tallic Bime-tallic Bime-tallic Bime-tallic Bime-tallic Bime-tallic Bime-tallic Bime-tallic Bime-tallic Bime-tallic Bime-tallic Bime-tallic Bime-tallic Bime-tallic Bime-tallicFuel SupplierAREVAAREVAAREVAAREVAAREVAAREVAAREVAAREVAAREVAAREVAAREVAAREVAAREVAAREVAAREVAFuel Stack Height Nominal (inches)136.7136.7136.7136.7136.7136.7136.7136.7136.7136.7136.7136.7136.7136.7136.7Number of Assem-blies82082588128368888484Regionwise Loading (MTU)3.047.603.039.433.043.044.553.0313.583.193.193.193.1718.981.59 MPS2 UFSAR3.4-10Rev. 35 (a) Including uncertainties.TABLE 3.4-2  NEUTRONICS CHARACTERISTICS FOR A REPRESENTATIVE RELOAD CORE
TABLE 3.4-3 REPRESENTATIVE SHUTDOWN MARGIN REQUIREMENTS trol Rod Worth (pcm)
<characteristic
BOC: BOC:   EOC:   EOC:
> B OC E O C Critical Boron (ppm): HZP, ARO, No Xenon1453
                <parameter>        HZP    HFP  HZP    HFP ARI                              9315  9315  10450  10450 N-1                              6271  6271  7696  7696 PDIL                            2116  157  2862  241
---Critical Boron (ppm): HFP, ARO, Equilibrium Xenon 1024 0Moderator Temperature Coefficient (pcm/
[(N-1) - PDIL]
&deg;F): HZP+2.0-10.4Moderator Temperature Coefficient (pcm/
* 0.9            3740  5503  4351  6710 ctivity Insertion (pcm)
&deg;F): HFP-6.0-23.3Doppler Coefficient (pcm/
BOC: BOC:   EOC:   EOC:
&deg;F)-1.17-1.33Boron Worth (pcm/ppm): HZP-8.8-10.8Boron Worth (pcm/ppm): HFP-8.4-10.4LHR (kW/ft) HFP (a)12.811.6Delayed Neutron Fraction0.00640.0054HFP, PDIL Worth (pcm)157241 N-1 Rod Worth, HZP (pcm)62717696Excess Shutdown Margin (pcm): HFP124323Excess Shutdown Margin (pcm): HZP140751 MPS2 UFSAR3.4-11Rev. 35TABLE 3.4-3 REPRESENTATIVE SHUTDOWN MARGIN REQUIREMENTSControl Rod Worth (pcm)
                <parameter>        HZP    HFP  HZP    HFP Power Defect                    0      1507  0      2515 Void                            0      50    0      50 Flux Redistribution              0      222  0      222 Total Requirements              0      1779  0      2787 tdown Margin (pcm)
<parameter>BOC: HZP BOC: HFPEOC: HZP EOC: HFPARI931593151045010450N-16271627176967696PDIL21161572862241[(N-1) - PDIL]
BOC: BOC:   EOC:   EOC:
* 0.93740550343516710Reactivity Insertion (pcm)
                <parameter>        HZP    HFP   HZP    HFP
<parameter>BOC: HZP BOC: HFPEOC: HZP EOC: HFPPower Defect0150702515 Void050050Flux Redistribution02220222Total Requirements0177902787Shutdown Margin (pcm)
[(N-1)
<parameter>BOC: HZP BOC: HFPEOC: HZP EOC: HFP[(N-1)
* PDIL]
* PDIL]
* 0.9 - Total3740372443513923Required Shutdown 3600360036003600 Excess Shutdown Margin140124751323 MPS-2 FSAR June 2000 FIGURE 3.4-1 REPR ESENTATIVE FULL COR E LOADING PATTERN MPS-2 FSAR June 2000FIGURE 3.4-2REPRESENTATIVE QUARTER CORE LOADING PATTERN MPS-2 FSAR June 2000FIGURE 3.4-3REPRESENTATIVE BOC AND EOC EXPOSURE DISTRIBUTION MPS-2 FSAR June 2000FIGURE 3.4-4REPRESENTATIVE BORON LETDOWN, HFP, ARO MPS-2 FSAR June 2000FIGURE 3.4-5 REPRESENTATIVE NORMALIZ ED POWER DISTRIBUTIONS, HOT FULL POWER, EQUILIBRIUM XENON, 150 MWD/MTU MPS-2 FSAR Rev. 32FIGURE 3.4-6REPRESENTATIVE NORMALI ZED POWER DISTRIBUTION, HOT FULL POWER, EQUILIBRIUM XENON, 18,020 MWD/MTU
* 0.9 - Total    3740  3724  4351  3923 Required Shutdown               3600  3600  3600  3600 Excess Shutdown Margin          140    124  751    323 3.4-11                    Rev. 35
 
MNPS-2 FSAR MPS-2 FSAR FIGURE 3.4-1 REPRESENTATIVE FULL CORE LOADING PATTERN 1      2    3      4      s        I      7
* 11    13      15    16      17    18      18    20      21 y
I=J N2B1          N191  N32/
0.16 e-os R-13 x                            H2O      R01    P64    Rl0 N11 R15 P45              ROO    N27 S..()5        F-04              L*'S        F-,S            S.17 w                      N55 R2D        P05 R39        P33 Ra1 P36 R80 P04                    R23 N57
                        $-07          X-16          C-07          0.15          X-06          8-15 v              N58    P09 P21 R65          N65 R49 N52 R55 NIt                    R70i  P26 P1. N54 R.oEi W-OS J.02            V*17              l-20        V..os          J.2O T*19 Re16 T        N22    R20i P1I R70 P63            R59 P52 R31 P.... A62 P40                      R69 P17 "R19 N25 T~          X*13          "N-04        E-07            E-15          N-1S          X-09          T*1' s        R07 P03      R75  P65      P19  P57 R43 N43 R4Ei PSl P27                      P47 R68 P02 R04 F.(J2        V.()9 V.04    8-13              T-09        S.Q9 V*18      V-l3            F*20 R        P66 R77 N62 R63 P58                P42 N03 R27 N06 P38 P71                        ASS N68 R38 P54
  ~
N29 V*1S          E..oc          J.06 -'-13 N-03              N*19 J.09 J.16                E*l8        V.()6
                                                                                                                      -N36    p N  N-U1 R1S P29 RS6 P37 R47                N07 R35 Pl0 R34 N02 R42                      P68    R52    P32    R09    N*'~
f---                                                                                                              I'"'-
N1.        R-19        R-17          W-09              T.03        W*13          R..05          R43          N13    M L  F*'9 N26    R82    N48 R32        N44  R28 P12 ~ P13 R26 N42                        R30    N51    R84    N1E F.oJ f---
N23 R*l1          X*11            J.05          C-05 0.11 W*17                                B-11          ~11 H-"
                                                                                                                      -N24    K J  S-19 Rl1    P31    RSO  P48      R44    N04 R36 P11 R33 N05 R45                      P51    R54    P30    R14    S..()J f---                                                                                                              10-N34        6-19          G-1,          e-09            ~.19          0.13          (;.OS        G-03          NJ,    H G  .un  P62 R40      N66  AGO      P72    P39 NOB R25            N01    P43 P56        AS1    N64 R79      P67    J.1S D-1E        T-G4          N-06  N-13    J.Q3          J.19    N..()9 N*1S            T-18          o-ee f        R02  P08 R66 PSO P22              P60    R48 N41 R41 P59 P20 P53                      R73    P06    R05 S-02          D-09 [).04    F-13            E*13          F.09 [).18 0-13              S-2O E        H15  R17 P24 R71 P70              R64    P49 R29 P41 R57 P55 R72                      P23    R22    N12 E..06        8-13          J.04          T-07          T*15          J.18          8.09          E-16 D              N56 P16 P2S R76              N63      R53 N50 ASl N67 R67 P28                      P15    N60 G-OE E~ N~                    0.17            L.Q2          [).()5          N*2(l 0-17 G-1E c                      N59 R21 P01          R78      P35 RaJ P34 R37 P07 R18                      N53 F-UT          8-16          W-UT          W-15          B.o&            F-1S 8                            N17 ROB        P46    R13 N21 R12 P69 R03 N10 Assembly 1.0.
F-OS          S..()4          l.07        $-18            F-17 Previous Cycle L ocatlon A
N18 N3S/
G-09 W-1S w~ ~13 I~Itail il
* 10    12    14 Figure 3.4-1 Representative Full Core Loading Pattern June 2000 June 2000
 
MPS-2 FSAR FIGURE 3.4-2 REPRESENTATIVE QUARTER CORE LOADING PATTERN MNPS-2 FSAR 11        13        15        16      17        18        19        20      21 1    N4  2    P2 3      R4 4    N4 5      R4 6      N4 7    R6 8    HZ L      C-l1      [-19                E-13              L-20              L-15  9    N2 90                  90              270                270 F-19    K 10    P2 11    R5 12      HI 13    R5 14    P5 IS    RS 16    P4 17    R2  180 J      E-19                C.. 13              [-15              e-15            18    H3 270                270 J-1S    H 19    R4 20    Nl 21      P5 22    PS 23    RS 24    N4 2S    R5 26    P5 G                J-19    J-13        F-13              D-17              F-18 180        90                90              270 27    "N4 28    R5 29      P5 30    P3 31    P5 32    R5 33    PI 34    RI F      E-13              J-16        D-18    D-13              B*16 270                                      90 35    R4 36    P5 37      RS 38    PS 39    R5 40    P3 41    R3 42    H2
[                G-17                J-18              J*20              [-16 90                                    270 43    N4 44    RS 45      N4 46    R5 47    P3 48    P2 49    H4 D      B-l1                [-18                B*13    e-17    G*16 270                270                  90 SO    R6 51    P4 52      RS 53    PI 54    R3 55    N4 t                G-19                F-20              F-15 90                270 56    HZ 57    R2 58      P5 S9    RI 60    H2 Region No./Subbatch Type B      G-11                0-16                F-17 Previous Cycle 1/4 Core Loca tion 270                  90                        Rotation (Deg. CCW) 61    H2 62 H3 A            e*16    G*13 180 12      14 FIGURE 3.4-2 REPRESENTATIVE QUARTER CORE LOADING PATTERN                                           00-.)3 June 2000 June 2000
 
MPS-2 FSAR MNPS*2    FSAR FIGURE 3.4-3 REPRESENTATIVE BOC AND EOC EXPOSURE DISTRIBUTION 11          13          15        16        17      18      19        20      21 1  "4      2  P2      3    a4  4  1'4    5  R4  6  114  7  1t6  8  112 34.036      16.585      0.000      35.125    0.000    30.580  0.000    32.683  9  112 50.652      37.623      23.655    52.112    24.328  48.977  23.811    46.522 32.000 10  '2  11    ItS    12    .1  13  ItS  14  '5  15  as  16  P4  17  R2  38.656 J            16.585      0.000      31.361    0.000      20.938  0.000    20.147    0.000 18 .3 37.623      23.096      48.812    22.853    Cl.311  23.C95  39.970    20.063 34.801    B 19  1t4  20    .1    21    PS  22  .5    23  as  24  .4  25  as  26 '5    39.877 c            0.000      31.375      18.933    20.637    0.000    32.250  0.000    20.199 23.655      48.835      37.652    39.981    23.397  49.790  22.002    34.127 27  .4  28    ItS    29    P5  30  P3    31  '5  32  1t5 33  PI  34  Itl 35.125      0.000      20.590    16.156    20.581  0.000    12.918    0.000 52.112      22.843      39.941    36.614    40.659  23.238  33.445    16.106 35  1t4  36    '5    37    itS 38  PS    39  itS 40  .3  41  R3  42  .2
* 0.000 24.328 20.946 41.288 0.000 23.383 20.562 40.643 0.000 23.301 16.'42 36.317 0.000 19.103 33.040 40.028 43  N4  .. 4  JlS    45    114 46  RS    47  '3  48  P2  49  .. 4 D            30.580      0.000      32.250    0.000      16.938  16.589  34.067 48.977      23.419      49.774    23.235    36.317  31.056  40.643 50  1t6  51    P4    52    :R5 53  PI    54  R3  55  Mol    Region No. I .atch t ype e            0.000      20.156      0.000      12.924    0.000    33.971 BOC Exposure {GWdl MTU) 23.811      39.964      21.995    33.449    19.107  40.559                                      00-.;13 EOC Exposure (Gwdl MTU) 56  "2  57    R2    58 . '5    59  1t1  60  112
* 32.683 46.522 0.000 20.109 20.166 3 ....107 0.000 16.108 33.070 40.056 61  112    62  113 31.943      34.813 38.624      39.900 12          14 FIGURE 3.4-3 REPRESENTATIVE BOC AND EOC EXPOSURE DISTRIBUTION
-- ~
June 2000 June 2000
 
MPS-2 FSAR FIGURE 3.4-4 REPRESENTATIVEMNPS-2 FSAR        BORON LETDOWN, HFP, ARO 1~
1100
-    1(XX)  '" I'..
E                  -,
a..
-Q.
      !IX)
                      "~
(
0    fa)
.J
.J                                    ~,
0                                                                  -
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.J
(                                            -,
e 0
c 600 0    SOO 0
                                                                "~
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0 400                                                          '" "\
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:m                                                                        ~
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                                                                                              " \.
o                                                                                        r\
o            2 3  t    5 6 1        8 9 10 II 12 13 Ii IS 16 11 18 19 C9cle Exposure lGWd/MTU)
Figure 3.4-4 Representative Boron Letdown, HFP, ARO Millstone Unit 2 June 2000 June 2000
 
MPS-2 FSAR FIGURE 3.4-5 REPRESENTATIVE NORMALIZED POWER DISTRIBUTIONS, HOT FULL POWER, EQUILIBRIUM XENON, 150 MWD/MTU MNPS-2 FSAR 11          13        15      16        17        18          19          20      21 i    M4      2    P2    3  R4  4  1f4    5  R4    6    ..4    7  R6      8  112 L      0.932        1.197      1.307    0.880      1.317      0.971      1.327      0.765    ,  .2
        . 0 . 981      1.313      1.600    0.933      1.5'4      1.047      1.577      0.'45 0.347 10      P2      11  JtS  12  51  13  ItS  14    P5  15    ItS  16    Pot  17    R2    0.626 1.197        1.261      0.950    1.214      1.085    1.241      1.108      1.212    18  53 1.313        1.570      1.048    1.517      1.155    1.543      1.179      1.573 0.263    B l'      R4    20    .1  21    75 22    P5  23    ItS  24    114  25    as  26    P5    0.545 o      1.307        0.950      1.058    1.089      1.259    0.926      1.211      0.781 1.600        1.049      1.173    1.186      1.57"    0.981      1.5"2      1.045 27      .... 28    ItS  29    P5 30    P3  31    P5  32    ItS  33    PI  34    JU p    0.880          1.214      1.089    1.215    1.130      1.278      1.212      1.027 0.933          1.516      1.187    1.376    1.222      1.596      1.31"      1.574 35      R4    36    PS  37  ItS 38    P5  39    ItS  ..0    P3  41  R3    .. 2  112
* 1.317 1.564 1.083 1.151 1.258 1.573 1.130 1.222 1.302 1.587 1.115 1.218 1.120 1.574 0.385 0.846 43      1f4    44  as    45
* N4  4'  Jl5  4'    P3  48    P2    49  114 D    0.971          1.236      0.924    1.277    1.115      0.823      0.346 1.047          1.536      0.geO    1.595    1.218      1.068      0.684 50      R6      51  P4    52  R5  53  PI  54    R3  55    114 e    1.327        1.107      1.210  1.212      1.120      0.347              Re9ioft 110. I ruel 1.577        1.1'7      1.541    1.313      1.574    0.686 A**-=bly Average Power 56      K2
* 57      R2    58  P5  59  JU  60    .2                        A** embly Peak Pin Power
* 0.765 0.945 1.216 1.578 0.782 1.046 1.027 1.57" 0.385 0.845        Pr
* 1.600 (e 11) 61    If2  62  1f3                                PEAK LRR (kW/ft)
* 12.8  (1 6 B 16) 0.349        0.264 0.630        0.547 12        ~.
FIG~RE 3.4-5 REPRESENTATIVE NORMALIZED POWER DISTRIBUTIONS 00*
                                                                                                                        .13 HFP, EQUILIBRIUM XENON,                150 -'IVi-WD/MTli- -- - ---
June 2000  June 2000
 
MPS-2 FSAR FIGURE 3.4-6 REPRESENTATIVE NORMALIZED            MNPS..2 FSAR    POWER DISTRIBUTION, HOT FULL POWER, EQUILIBRIUM XENON, 18,020 MWD/MTU
        ~1          13        15          16          17        18      19        20      21 1f4    2  P2    3  a4    4    )14    5    1t4  6  .4    7  1t6    8  .2
      ~
L    0.905      1.131      1.401    0.969        1.425      1.005    1.313      0.736        52 0.955      1.228      1.495    1.008        1.510      1.066    1.449      0.883 0.383  Jt 10    P2    11    as  12  Jfl  13      as    14    P5  15    as  16  Pol  17  R2    0.619 J  1.131      1.386      0.982    1.388        1.112      1.402    1.038      1.062    18  .3 1.228      1.503      1.017
* 1 . 514      1.161      1.51"    1.139      1.300 0.293  B 19 It"  20  Xl    21  P5  22      P5  23    ItS  24    .4  25  R5    26  PS    0.546 c  1.401      0.982      0.993      1.027        1.395    0.t81    1.300      0.740 1.495      1.017      1.053      1.129        1.517    1.032    1.459      0.984 27  .4    28  ItS  29  P5  30      P3    31    P5  32    as  33  P1    34  1t1
  .,  0.969      1.388      1.027      1.051        1.060    1.363    1.057      0.859 1.008      1.5106    1.130      1.183        1.125    1.478    1.140      1.261 35  1t4    36  P5    37  RS  38      PS    39    itS  40    P3  41  R3    42  .2 1.425      1.111      1.396    1.061        1.349    1.011    1.036      0.386 1.510      1.161      1.517    1.126        1.0663    1.104    1.356      0.759 43  Mol    44  itS  45  114 46      ItS  4'    P3  48    P2  49  N4 D  1.005      1.400      0.981    1.363        1.011    0.771    0.371 1.066      1.513      1.031    1.478        1.104    0.976    0.677 50  Jl6    51  Pol  S2  as  53      PI    54    Jl3  55  1f4 c  1.313      1.038      1.300    1.057        1.036      0.371            Region No. I Fuel 1.449      1.138      1.461    1.141        1.357      0.679 Aaaembly Average Power 56  .2    57    R2    58  P5  S9    III    60  52                      A** embly P.ak Pin Power
* 0.736 0.883 1.063 1.302 0.'40 0.985 0.859 1.262 0.386 0.758          "r
* 1.517    (E 15) 61    112  62  .3                                      PEAX LHR (kW/ft)
* 11.6 (21 E 15) 0.384      0.293 0.621      0.547 12        14 FIGURE 3.4-6 REPRESENT:ATIV~_~O~~~_I..I:l:~D~9WER DISTRIBUTION HFP, EQUILIBRIUM XENON, EOC                                            Rev. 32 June 2000


MPS2 UFSAR3.5-1Rev. 35 3.5 THERMAL-HYDRAULIC DESIGN This section presents thermal and hydraulic analysis of the reactor core, analytical methods utilized, and experiment al work supporting the analytical t echniques. The prime objective of the thermal and hydraulic design of the reactor is the assurance that th e core can meet normal steady state and anticipated transient performance requirements without exceeding the design bases. A summary of the significant reactor and fuel parameters used in the thermal and hydraulic design and analysis is presented in Table 3.5-1.
s section presents thermal and hydraulic analysis of the reactor core, analytical methods zed, and experimental work supporting the analytical techniques. The prime objective of the mal and hydraulic design of the reactor is the assurance that the core can meet normal steady e and anticipated transient performance requirements without exceeding the design bases. A mary of the significant reactor and fuel parameters used in the thermal and hydraulic design analysis is presented in Table 3.5-1.
1    DESIGN BASES 1.1    Thermal Design idance of thermally induced fuel damage during any normal steady state and anticipated sient operation is the principal thermal and hydraulic design basis. The following limits are blished, but violation of them will not necessarily result in fuel damage. The Reactor tection System will provide for automatic reactor trip or other corrective action before these gn limits are exceeded.
: a.      Avoidance of departure from nucleate boiling (DNB) for the limiting rod in the core with 95 percent probability at a 95 percent confidence level.
: b.      Limitation of the peak temperature of the fuel to less than the melting point during normal operation and anticipated transients.
ce the departure from nucleate boiling ratio (DNBR) criterion ensures that the cladding perature remains close to the coolant temperature, no additional criteria for cladding perature are required for normal operation and anticipated transients. For design basis dent conditions (loss of coolant accidents (LOCA)), under which the DNBR criterion does not ly, cladding temperatures are calculated to ensure that they remain below 2200&deg;F, which is the k clad temperature criterion of 10 CFR 50 Appendix K. For other postulated accidents, fuel ure is assumed to occur if the calculated DNBR is below the DNB correlation 95/95 limit.
1.2    Hydraulic Stability rating conditions shall not lead to flow instability during normal steady state and anticipated sient operation.
1.3    Coolant Flow Rate, Distribution and Void Fraction ower limit on the total primary coolant flow rate, called design flow, is set to assure that the is adequately cooled when uncertainties in system resistance, pump head, and core bypass are taken in the adverse direction. By design of the reactor internal flow passages, this flow istributed to the core such that the core is adequately cooled with all permissible core power ributions. The hydraulic loads for the design of the internals are based on the upper limit of the 3.5-1                                      Rev. 35


====3.5.1 DESIGN====
nsure that sufficient coolant flow reaches the fuel, the amount of coolant flow which bypasses core through the guide tubes must not excessively reduce the active core flow. The guide tube lant flow must, however, be sufficient to ensure that coolant in the guide tubes will not boil ensure adequate cooling of the CEA fingers. The CEA drop time in the guide tubes must also t the criterion of 90 percent insertion within 2.75 seconds to ensure that scram performance is ccordance with plant Technical Specifications.
BASES 3.5.1.1 Thermal DesignAvoidance of thermally induced fuel damage during any norma l steady state and anticipated transient operation is the principal thermal and hydraulic design basis. Th e following limits are established, but violati on of them will not necessarily result in fuel damage. The Reactor Protection System will provide for automatic reactor trip or other corrective action before these design limits are exceeded.a.Avoidance of departure from nucleate boiling (DNB) for the limiting rod in the core with 95 percent probability at a 95 percent confidence level.b.Limitation of the peak temperature of the fuel to less than the melting point during normal operation and anticipated transients.
hough the coolant velocity, its distribution, and the coolant voids affect the thermal margin, gn limits need not be applied to these parameters because they are not themselves limiting h respect to thermal margin. These parameters are included in the thermal margin analyses and affect the thermal margin to the design limits.
Since the departure from nucleate boiling ratio (DNBR) criter ion ensures that the cladding temperature remains close to the coolant te mperature, no additiona l criteria for cladding temperature are required for normal operation and anticipated transients. For design basis accident conditions (loss of coolant accidents (LOCA)), under wh ich the DNBR criterion does not apply, cladding temperatures are calculated to ensure that they remain below 2200
2    THERMAL AND HYDRAULIC CHARACTERISTICS OF THE DESIGN 2.1    Fuel Temperatures RODEX2 code (Reference 3.5-1) incorporates models to describe the thermal and hanical behavior of the fuel rod in a flow channel including the gas release, swelling, sification, and cracking in the pellet; the gap conductance; the radial thermal conduction; the volume and gas pressure internal to the fuel rod; the fuel and cladding deformations; and the ding corrosion as a function of burnup. The calculations are performed on a time-incremental s with conditions being updated at each calculated increment.
&deg;F, which is the peak clad temperature criterion of 10 CFR 50 Appendix K. For other postu lated accidents, fuel failure is assumed to occur if the calculate d DNBR is below the DNB correlation 95/95 limit.
2.1.1    Fuel Cladding Temperatures RODEX2 thermal-hydraulic model (Reference 3.5-1) calculates the lowest cladding surface perature based on one of two heat transfer regimes; i.e., forced convection and fully developed leate boiling. The forced convection and fully developed nucleate boiling heat transfer elations in RODEX2 were developed by Kays and Thom et al., respectively.
3.5.1.2 Hydraulic StabilityOperating conditions shall not lead to flow inst ability during normal stea dy state and anticipated transient operation.
2.1.2    Fuel Pellet Temperatures RODEX2 radial temperature distribution model begins with the standard differential equation eat conduction (Poisson Equation) for an isotropic solid with internal heat generation. The ation is written in cylindrical coordinates assuming that the thermal conductivity of the fuel is nction of fuel temperature, but is independent of position. With additional assumptions of l symmetry, negligible heat conduction in the axial direction, and steady state conditions, a
3.5.1.3 Coolant Flow Rate, Distribution and Void Fraction A lower limit on the total primary coolant flow rate, called "design" flow, is set to assure that the core is adequately cooled when uncertainties in system resistance, pum p head, and core bypass flow are taken in the adverse direction. By design of the reactor inte rnal flow passages, this flow is distributed to the core such that the core is adequately cooled with al l permissible core power distributions. The hydraulic loads fo r the design of the internals ar e based on the upper limit of the MPS2 UFSAR3.5-2Rev. 35flow. The upper limit is obtained in a similar manner as the design flow but with the uncertainties taken in the opposite direction.To ensure that sufficient coolant flow reaches the fuel, the amount of c oolant flow which bypasses the core through the guide tubes mu st not excessively reduce the active core flow. The guide tube coolant flow must, however, be sufficient to ensu re that coolant in the guide tubes will not boil and ensure adequate cooling of the CEA fingers. The CEA drop time in the guide tubes must also meet the criterion of 90 percent insertion within 2.75 seconds to ensure that scram performance is in accordance with plant Technical Specifications.
-dimensional (i.e., radial) steady state form of the equation is derived and employed.
Although the coolant velocity, its distribution, and the coolant voids affect the thermal margin, design limits need not be applied to these parameters because they are not themselves limiting with respect to thermal margin. These parameters are included in the thermal margin analyses and thus affect the thermal margin to the design limits.
minimum power level required to produce centerline melt in Zircaloy clad uranium fuel rods efined as the Fuel Centerline Melt Linear Heat Rate (FCMLHR) limit and is expressed in kW/
This FCMLHR is determined using the methodology of Reference 3.5-22. A conservative le specific FCMLHR limit is used for Millstone Unit 2. The maximum LHR for normal 3.5-2                                    Rev. 35


====3.5.2 THERMAL====
temperature than an all-uranium-bearing fuel rod. Gadolinia rods are specifically analyzed to terline melt criteria.
AND HYDRAULIC CHAR ACTERISTICS OF THE DESIGN 3.5.2.1 Fuel TemperaturesThe RODEX2 code (Reference 3.5-1) incorpor ates models to desc ribe the thermal and mechanical behavior of the fuel rod in a fl ow channel including the gas release, swelling, densification, and cracking in the pellet; the gap conductance; the radial thermal conduction; the free volume and gas pressu re internal to the fuel rod; the fuel and claddi ng deformations; and the cladding corrosion as a function of burnup. The calculations are performed on a time-incremental basis with conditions being updated at each calculated increment.
2.1.3   UO2 Thermal Conductivity ns expression for thermal conductivity of the fuel is used in RODEX2. Two corrections are lied: one for density and one to account for the gadolinia content in the fuel.
3.5.2.1.1 Fuel Cladding TemperaturesThe RODEX2 thermal-hydraulic model (Reference 3
2.1.4   Gap Conductance RODEX2 gap conductance model is based on that proposed by Kjaerheim and Rolstad. The l gap conductance has three components: (1) gas conductance, (2) radiation, and (3) fuel/
.5-1) calculates the lowest cladding surface temperature based on one of two h eat transfer regimes; i.e., forc ed convection and fully developed nucleate boiling. The forced conv ection and fully developed nucleate boiling heat transfer correlations in RODEX2 were developed by Kays and Thom et al., respectively.
ding solid-to-solid contact.
3.5.2.1.2 Fuel Pellet Temperatures The RODEX2 radial temperature distribution model begins with the standard diff er ential equation of heat conduction (Poisson Equation) for an isotropic solid with internal heat generation. The equation is written in cylindrical coordinates assuming that the thermal conduc tivity of the fuel is a function of fuel temperature, but is independent of position. With additional assu mptions of axial symmetry, negligible heat conduction in the axial directi on, and steady state conditions, a one-dimensional (i.e., radial) steady state form of the equa tion is derived and employed.
2.2   Departure from Nucleate Boiling Ratio BRs are calculated using approved correlations. An approved core thermal-hydraulic puter code is used to determine the flow and enthalpy distribution in the core and the local ditions in the hot channel for use in the DNB correlation.
The minimum power level required to produce center line melt in Zircaloy clad uranium fuel rods is defined as the Fuel Centerline Melt Linear Heat Rate (FCMLHR) limit and is expressed in kW/
2.2.1    Departure from Nucleate Boiling XCOBRA-IIIC (Reference 3.5-2) computer code is employed to evaluate the thermal-raulic conditions in the various assemblies and in the subchannels of the limiting assembly.
ft. This FCMLHR is determined using the methodology of Reference 3.5-22. A conservative cycle specific FCMLHR limit is used for Millstone Unit 2. The maximum LHR for normal MPS2 UFSAR3.5-3Rev. 35 operation and anticipated transients is typically well below the c onservative FCMLHR limit. It should be noted that a gadolinia-bearing fuel rod will, for a given LHGR , operate with a higher fuel temperature than an all-uranium-bearing fuel rod. Gadolinia rods are specifically analyzed to centerline melt criteria.
t, mass, and momentum fluxes between the inter-rod flow channels are explicitly calculated.
3.5.2.1.3 UO 2 Thermal ConductivityLyon's expression for thermal conduc tivity of the fuel is used in RODEX2. Tw o corrections are applied: one for density and one to account for the gadol inia content in the fuel.
l and reactor design conditions employed in these calculations are given in Table 3.5-1.
3.5.2.1.4 Gap Conductance The RODEX2 gap conductance mode l is based on that proposed by Kjaerheim and Rolstad. The total gap conductance has three components: (1) gas conductance, (2) radiation, and (3) fuel/
calculations include a statistically determined engineering factor to account for ufacturing tolerances, thermal expansion and densification effects. The engineering factor is lied to the local heat flux in the calculation of DNBR.
cladding solid-to-solid contact.
eactor densification results in a shortening of the fuel column. At power levels typical of BR-limiting rods, thermal expansion tends to offset the densification effect. The XCOBRA-model does not specifically model changes in stack length due to thermal expansion and sification.
3.5.2.2 Departure from Nucleate Boiling Ratio DNBRs are calculated using approved correlat ions. An approved core thermal-hydraulic computer code is used to determine the flow and enthalpy distribution in the core and the local conditions in the hot channel for use in the DNB correlation.
HTP DNB correlation, demonstrated to be applicable to the AREVA 14 by 14 reload fuel mblies for CE reactors, is described in Reference 3.5-3. A minimum allowable limit esponding to 95% probability with 95% confidence is set on the DNBR during normal ration and any anticipated transients.
2.2.2    Hot Channel Factors channel factors for heat flux and enthalpy rise, Fq and Fr:
3.5-3                                  Rev. 35


3.5.2.2.1 Departure from Nucleate BoilingThe XCOBRA-IIIC (Reference 3.5-2) computer c ode is employed to evaluate the thermal-hydraulic conditions in the various assemblies and in the subchannels of the limiting assembly.
rage ratios of these quantities. The heat flux hot channel factor (Fq) considers the local imum linear heat generation rate at a point (the hot spot), and the enthalpy rise hot channel or (Fr) involves the maximum integrated linear heat generation rate along a channel (the hot nnel).
Heat, mass, and momentum fluxes between the inter-rod flow cha nnels are explic itly calculated.
ineering hot channel factor, FE:
Fuel and reactor design conditions employed in these calculations are given in Table 3.5-1.
engineering hot channel factor is used to evaluate the maximum linear heat generation rate in core. This subfactor is determined by statistically combining the fabrication uncertainties for pellet diameter, density, and enrichment, as well as the effect of densification. A conservative e of 1.03 is used. The effect of variations in fabrication tolerances is considered in the lysis. To account for manufacturing uncertainties and densification, the peak rod heat flux is eased by 3% in the calculation of DNBR.
The calculations include a stat istically determined engineer ing factor to account for manufacturing tolerances, thermal expansion and densification effect
2.2.2.1     Nuclear Peaking Factors embly and rod peaking factors and axial power distributions are input into the XCOBRA-IIIC
: s. The engineering factor is applied to the local heat fl ux in the calculation of DNBR.In-reactor densification results in a shortening of the fuel column. At power levels typical of DNBR-limiting rods, thermal expansion tends to offset the densification effect. The XCOBRA-IIIC model does not specifically model changes in stack length due to thermal expansion and densification.
: e. Departure from nucleate boiling is dependent on the local rod heat flux and the local fluid ditions within the channel.
The HTP DNB correlation, demonstr ated to be applicable to the AREVA 14 by 14 reload fuel assemblies for CE reactors, is described in Reference 3.5-3. A minimum allowable limit corresponding to 95% probability with 95%
effect of asymmetries in core power distribution (specifically azimuthal power tilt) is not ctly taken into account in the XCOBRA-IIIC thermal-hydraulic calculations. The effects of muthal power tilt are accounted for in the generation (verification) of the TM/LP trip and LPD monitoring setpoints through the measurement of radial peaking factors.
confidence is set on the DNBR during normal operation and any anticipated transients.
2.2.2.2     Rod Bowing Factor the fuel assembly burnup increases, the gaps between fuel rods change. Decreased rod-to-rod s can occur, which can reduce the DNB ratio. Penalties are calculated as a function of burnup applied to the DNBR or peak linear power as appropriate.
3.5.2.2.2 Hot Channel Factors Hot channel factors for heat flux and enthalpy rise, F q and F r:
2.2.2.3     Inlet Flow Distribution Factor t flow maldistribution is treated in the XCOBRA-IIIC model by applying a generic inlet flow alty to the limiting assembly and its crossflow neighbors.
MPS2 UFSAR3.5-4Rev. 35The total hot channel factors for heat flux and enthalpy rise are defined as the maximum-to-core average ratios of these quantities. The heat fl ux hot channel factor (F q) considers the local maximum linear heat generation ra te at a point (the ho t spot), and the enthalpy rise hot channel factor (F r) involves the maximum integrat ed linear heat generation ra te along a channel (the hot channel).Engineering hot channel factor, F E: The engineering hot channel factor is used to evaluate the maximum linear heat generation rate in the core. This subfactor is determined by statistically combining the fabrication uncertainties for fuel pellet diameter, density, and enrichment, as well as the ef fect of densification. A conservative value of 1.03 is used. The effect of variations in fabrication tolerances is considered in the analysis. To account for manufacturing uncertainties and densification, the peak rod heat flux is increased by 3% in the calculation of DNBR.
2.2.2.4     Flow Mixing Factor effects of both pressure-driven and turbulent flow mixing between channels on the hot nnel enthalpy rise are calculated by the XCOBRA-IIIC computer code. The turbulent flow ing is modeled empirically and is based on the reduction of the data from hot mixing tests g XCOBRA-IIIC.
3.5.2.2.2.1 Nuclear Peaking Factors Assembly and rod peaking factors and axial power distributions are input into the XCOBRA-IIIC code. Departure from nucleate boili ng is dependent on the local rod heat flux and the local fluid conditions within the channel.The effect of asymmetries in core power distri bution (specifically azimuth al power tilt) is not directly taken into account in the XCOBRA-IIIC thermal-hydraulic calculations. The effects of azimuthal power tilt are accounted for in the generation (verifi cation) of the TM/LP trip and LPD trip monitoring setpoints through the m easurement of radial peaking factors.
3.5-4                                       Rev. 35
3.5.2.2.2.2 Rod Bowing Factor As the fuel assembly burnup increases, the gaps between fuel rods change. Decreased rod-to-rod gaps can occur, which can reduce the DNB ratio.
Penalties are calculated as a function of burnup and applied to the DNBR or peak linear power as appropriate.
3.5.2.2.2.3 Inlet Flow Distribution FactorInlet flow maldistribution is treated in the XC OBRA-IIIC model by applyi ng a generic inlet flow penalty to the limiting assemb ly and its crossflow neighbors.
3.5.2.2.2.4 Flow Mixing FactorThe effects of both pressure-driven and turbul ent flow mixing between channels on the hot channel enthalpy rise are calculated by the XCOBRA-IIIC comput er code. The turbulent flow mixing is modeled empiri cally and is based on the reduction of the data from hot mixing tests using XCOBRA-IIIC.
MPS2 UFSAR3.5-5Rev. 35 The geometry of the channels su rrounding the hot channel and the radial power distribution affect the lateral enthalpy transport for both the pressure-driven and tu rbulent flow mixing.
3.5.2.2.3 Effects of Rod Bow on DNBRIn accordance with AREVA rod bow methodology (Reference 3.5-4), the magnitude of rod bow for assemblies of the type used in Millstone Un it 2 has been estimated.
Significant impact on the DNBR due to rod bow does not occur until the gap closures exceed 50 percent. The maximum design exposure for AREVA reload fuel in Millstone Unit 2 is signifi cantly less than that at which 50 percent closure occurs; theref ore, rod bow does not significa ntly impact the minimum DNBR (MDNBR). A further consequence of the small amount of rod bow for AREVA fuel is that total power peaking is not si gnificantly impacted.
3.5.2.3 Void Fraction and Distribution The XCOBRA-IIIC model calculates the local thermal and hydraulic conditions for input to the DNB correlation. While local conditions of enthalpy, quality, flow rate and pressure are associated with a code-calcula ted local void fraction, the void fr action is not input to the DNB correlation. The DNB correlation is approved over a local quality range, but it is not a direct function of void fraction. Th erefore, there is no expl icit limit set on averag e or local void fraction beyond that implied in the test conditions used to develop the DNB correlation.
3.5.2.4 Coolant Flow Distribution 3.5.2.4.1 Coolant Flow Distribution and Bypass Flow The minimum primary coolant flow rate at fu ll power conditions is given in Ta ble 3.5-1.Tracing the coolant flow path in Figure 3.1-1, the c oolant enters the four inlet nozzles and flows into the annular plenum between the reactor vessel and core support barrel. It then flows down the annulus between the reactor vessel and core barrel and up through the flow skirt to the plenum below the core lower support structure. The skirt and lower support structure help to even out the inlet flow distribution to the core. The coolant passes through the openings in the lower core plate and flows axially through the fuel assemblies.
A portion of the coolant passes through the lower core plate and into the guide tubes in the fuel assemblies. The fuel assembly alignm ent plate is not drilled through in guide tube locations without CE As; therefore, core bypass flow is limited in these guide tubes. After passing th rough the core, the coolant flow s into the region outside the control element assembly shrouds.
From this region, the coolant fl ows across the control element assembly shrouds and passes out th rough the outlet sleeves on the core barrel to the outlet nozzles.
The coolant which does not contact any fuel rods is termed core bypass cool ant. The following are the principal core bypass routes:a.Direct inlet to outlet coolant flow at the joint between the core support barrel sleeve and reactor vess el nozzle.
MPS2 UFSAR3.5-6Rev. 35b.Coolant flow into the guide tubes in the fuel assemblies.c.Coolant flow in the region between the core support barrel and core shroud.d.Coolant flow from the inlet nozzle re gion through the alignment keyways to the vessel head region.Table 3.5-1 gives the "best estimat e" value for the core bypass flow rate as a fraction of the total primary flow rate. Ta king into a ccount the core bypass flow rate, th e core flow rate, which is the effective flow rate for heat transfer, can be calculated from the total primary coolant flow rate.
3.5.2.4.2 Core Flow Distribution The core flow distribution (CFD) analysis is performed to assess cross flow between assemblies in the core for use in subsequent MDNBR subcha nnel analyses. A full core model provides cross-flow boundary conditions to a full assembly m odel at the assembly boundaries. MDNBRs are computed from a full assembly simulation.
In the analysis, each fuel assemb ly in the Millstone Unit 2 core is modeled as a hydraulic channel.
The calculations are performed wi th the XCOBRA-IIIC computer code (Reference 3.5-2). Cross flow between adjacent assemblies in the open lattice core is di rectly modeled. The single-phase loss coefficients are used in the CFD analyses to hydraulically characterize the assemblies in the core.This computational procedure is designed to evaluate thermal-hydraulic conditions during boiling and non-boiling conditions. One-dimens ional, two phase se parated, slip flow is assumed in the XCOBRA-IIIC calculati on. These assumptions are valid onl y if the cross flow between connecting channels is small compared to the axial velocities in the individual channels. Because small cross flow does exist, mathematical models have to be postulated for both turbulent and diversion cross-flow mixing. Mode ls of the two-phase state are also defined in terms of void fraction, which is a function of enthalpy, flow rate, heat flux, pr essure, and axial position. This computational procedure is not applicable when large blockages exist in the fuel bundles since this leads to considerable cross flow wh ich cannot be adequately represented by the one-dimensional analysis.Table 3.5-1 summarizes the reactor and fuel desi gn parameters used in these CFD calculations and subsequent MDNBR analyses.


3.5.2.5 Pressure Losses and Hydraulic Loads 3.5.2.5.1 Pressure LossesThe fuel assembly irrecoverable pressure los ses have been calculated using standard loss coefficient methods and results from model tests. The pressure loss across the AREVA fuel assembly was determined based on the re sults of Reference 3.5-5 and analyses.
2.2.3    Effects of Rod Bow on DNBR ccordance with AREVA rod bow methodology (Reference 3.5-4), the magnitude of rod bow assemblies of the type used in Millstone Unit 2 has been estimated. Significant impact on the BR due to rod bow does not occur until the gap closures exceed 50 percent. The maximum gn exposure for AREVA reload fuel in Millstone Unit 2 is significantly less than that at which percent closure occurs; therefore, rod bow does not significantly impact the minimum DNBR DNBR). A further consequence of the small amount of rod bow for AREVA fuel is that total er peaking is not significantly impacted.
MPS2 UFSAR3.5-7Rev. 35 3.5.2.5.2 Hydraulic Loads 3.5.2.5.2.1 Hydraulic Loads on Vessel Internal Components The design hydraulic loads for the internal com ponents for steady state operating conditions are listed in Table 3.5-2. These loads were derived fr om analysis and from reactor flow model and component test results. All hydraulic loads in Table 3.5-2 are ba sed on the maximum expected system flow rate and a coolant temperature of 500
2.3     Void Fraction and Distribution XCOBRA-IIIC model calculates the local thermal and hydraulic conditions for input to the B correlation. While local conditions of enthalpy, quality, flow rate and pressure are ciated with a code-calculated local void fraction, the void fraction is not input to the DNB elation. The DNB correlation is approved over a local quality range, but it is not a direct ction of void fraction. Therefore, there is no explicit limit set on average or local void fraction ond that implied in the test conditions used to develop the DNB correlation.
&deg;F. When these hydraulic loads are used in the structural analysis, they are adjusted for coolant temperatur
2.4      Coolant Flow Distribution 2.4.1     Coolant Flow Distribution and Bypass Flow minimum primary coolant flow rate at full power conditions is given in Table 3.5-1.
: e. The worst condition (i.e., coolant temperature) is not necessarily the same for each internal component; therefore, the loads are adjusted to reflect the difference in coolant temperature. This is done to ensure the design hydraulic stresses are acceptable duri ng start-up and during power operation.The types of loads considered in the analysis ar e: (1) steady-state drag and impingement loads, and (2) fluctuating loads induced by pump pressure pu lsations, turbulence, and vortex shedding.
cing the coolant flow path in Figure 3.1-1, the coolant enters the four inlet nozzles and flows the annular plenum between the reactor vessel and core support barrel. It then flows down the ulus between the reactor vessel and core barrel and up through the flow skirt to the plenum w the core lower support structure. The skirt and lower support structure help to even out the t flow distribution to the core. The coolant passes through the openings in the lower core plate flows axially through the fuel assemblies. A portion of the coolant passes through the lower plate and into the guide tubes in the fuel assemblies. The fuel assembly alignment plate is not led through in guide tube locations without CEAs; therefore, core bypass flow is limited in e guide tubes. After passing through the core, the coolant flows into the region outside the trol element assembly shrouds. From this region, the coolant flows across the control element mbly shrouds and passes out through the outlet sleeves on the core barrel to the outlet nozzles.
All of these loads are not exerte d on each internal component, but each component sees at least one of the loads. Table 3.5-2 lists the components and type of loads that are exerted on them.
coolant which does not contact any fuel rods is termed core bypass coolant. The following are principal core bypass routes:
3.5.2.5.2.2 Core Hydraulic Loads/Fuel Assembly Liftoff The holddown spring force and the a ssembly weight force prevent th e fuel assembly from lifting off the core support plate duri ng reactor steady-state operation, based on the most adverse combination of compone nt dimensional and ma terial property tolerances. In addition, the holddown springs are designed to accommodate the additional lo ad associated with a pump overspeed transient (re sulting in possible temporary liftoff of the fuel asse mblies), and to continue to ensure fuel assembly holddow n following such occurrences.
: a.     Direct inlet to outlet coolant flow at the joint between the core support barrel sleeve and reactor vessel nozzle.
The limiting reactor steady-state conditions are the 4 th pump startup conditions. These corres pond to the minimum temperature and maximum pressure and coolant fl ow for reactor startup. Thermal expansion of the reactor vessel and fuel assembly is also considered.
3.5-5                                    Rev. 35
3.5.2.6 Correlation and Physical DataReference 3.5-1 describes the correlations and physical data employed in heat transfer calculations performed by RODEX2. Reference 3.5-7 describes the co rrelations and physical data employed in the hydraulic calc ulations performed by XCOBRA-I IIC. Reference 3.5-3 describes the correlations and physical data employed in the DNB correlation.
: c.     Coolant flow in the region between the core support barrel and core shroud.
3.5.2.7 Plant Parameters for Thermal-Hydraulic Design The plant parameters considered include total primary coolant flow rate, vessel inlet temperature, primary pressure, and core thermal power. Two se ts of thermal-hydraulic conditions are defined:
: d.      Coolant flow from the inlet nozzle region through the alignment keyways to the vessel head region.
nominal conditions and design co nditions. Nominal plant conditions represent the best estimate for the primary coolant flow ra te, pressure, and vessel inlet temperature and do not include allowances for instrument errors. Design plant conditions represent the lower limit on primary flow rate when uncertainties in system resistan ce and pump head are included, and represent the upper limit on vessel inlet temperature when design margins on st eam generator performance are MPS2 UFSAR3.5-8Rev. 35 included. Furthermore, the varia tions which occur during steady state operation in the power, pressure, and inlet temperature due to controlle r deadband and instrument error are considered with the design plant parameters. During steady state operation, the possible variations in these parameters define an operatin g envelope. One combination of these parameters gives the MDNBR, and this combination is utilized in Chapter 14 as the initial condi tions in transient and accident analysis. Table 3.5-1 lists the nominal plant parameters.
le 3.5-1 gives the best estimate value for the core bypass flow rate as a fraction of the total ary flow rate. Taking into account the core bypass flow rate, the core flow rate, which is the ctive flow rate for heat transfer, can be calculated from the total primary coolant flow rate.
3.5.2.8 Summary of Thermal a nd Hydraulic ParametersThe thermal and hydraulic parameters for the reactor are listed in Ta ble 3.5-1.
2.4.2      Core Flow Distribution core flow distribution (CFD) analysis is performed to assess cross flow between assemblies he core for use in subsequent MDNBR subchannel analyses. A full core model provides cross-boundary conditions to a full assembly model at the assembly boundaries. MDNBRs are puted from a full assembly simulation.
he analysis, each fuel assembly in the Millstone Unit 2 core is modeled as a hydraulic channel.
calculations are performed with the XCOBRA-IIIC computer code (Reference 3.5-2). Cross between adjacent assemblies in the open lattice core is directly modeled. The single-phase coefficients are used in the CFD analyses to hydraulically characterize the assemblies in the s computational procedure is designed to evaluate thermal-hydraulic conditions during boiling non-boiling conditions. One-dimensional, two phase separated, slip flow is assumed in the OBRA-IIIC calculation. These assumptions are valid only if the cross flow between necting channels is small compared to the axial velocities in the individual channels. Because ll cross flow does exist, mathematical models have to be postulated for both turbulent and ersion cross-flow mixing. Models of the two-phase state are also defined in terms of void tion, which is a function of enthalpy, flow rate, heat flux, pressure, and axial position. This putational procedure is not applicable when large blockages exist in the fuel bundles since leads to considerable cross flow which cannot be adequately represented by the
-dimensional analysis.
le 3.5-1 summarizes the reactor and fuel design parameters used in these CFD calculations subsequent MDNBR analyses.
2.5      Pressure Losses and Hydraulic Loads 2.5.1      Pressure Losses fuel assembly irrecoverable pressure losses have been calculated using standard loss fficient methods and results from model tests. The pressure loss across the AREVA fuel mbly was determined based on the results of Reference 3.5-5 and analyses.
3.5-6                                    Rev. 35


====3.5.3 THERMAL====
2.5.2.1      Hydraulic Loads on Vessel Internal Components design hydraulic loads for the internal components for steady state operating conditions are d in Table 3.5-2. These loads were derived from analysis and from reactor flow model and ponent test results. All hydraulic loads in Table 3.5-2 are based on the maximum expected em flow rate and a coolant temperature of 500&deg;F. When these hydraulic loads are used in the ctural analysis, they are adjusted for coolant temperature. The worst condition (i.e., coolant perature) is not necessarily the same for each internal component; therefore, the loads are sted to reflect the difference in coolant temperature. This is done to ensure the design raulic stresses are acceptable during start-up and during power operation.
AND HYDRAULIC EVALUATION 3.5.3.1 Analytical Techniques and Uncertainties 3.5.3.1.1 XCOBRA-IIIC DNBR Analyses The thermal-hydraulic simulations employed to evaluate the MDNBR were performed in accordance with AREVA's Nuclear Regulatory Commission (NRC) approved thermal-hydraulic methodology for mixed co res (Reference 3.5-8).
types of loads considered in the analysis are: (1) steady-state drag and impingement loads, (2) fluctuating loads induced by pump pressure pulsations, turbulence, and vortex shedding.
The MDNBR performanc e of the core during anticipated transi ents will be demonstrated to meet the thermal-hydraulic design crite rion on DNBR through th e performance of tran sient analysis of the limiting events. The results of this analysis are included in Chapter 14.
of these loads are not exerted on each internal component, but each component sees at least of the loads. Table 3.5-2 lists the components and type of loads that are exerted on them.
2.5.2.2      Core Hydraulic Loads/Fuel Assembly Liftoff holddown spring force and the assembly weight force prevent the fuel assembly from lifting the core support plate during reactor steady-state operation, based on the most adverse bination of component dimensional and material property tolerances. In addition, the ddown springs are designed to accommodate the additional load associated with a pump rspeed transient (resulting in possible temporary liftoff of the fuel assemblies), and to continue nsure fuel assembly holddown following such occurrences. The limiting reactor steady-state ditions are the 4th pump startup conditions. These correspond to the minimum temperature and imum pressure and coolant flow for reactor startup. Thermal expansion of the reactor vessel fuel assembly is also considered.
2.6    Correlation and Physical Data erence 3.5-1 describes the correlations and physical data employed in heat transfer ulations performed by RODEX2. Reference 3.5-7 describes the correlations and physical data loyed in the hydraulic calculations performed by XCOBRA-IIIC. Reference 3.5-3 describes correlations and physical data employed in the DNB correlation.
2.7    Plant Parameters for Thermal-Hydraulic Design plant parameters considered include total primary coolant flow rate, vessel inlet temperature, mary pressure, and core thermal power. Two sets of thermal-hydraulic conditions are defined:
inal conditions and design conditions. Nominal plant conditions represent the best estimate the primary coolant flow rate, pressure, and vessel inlet temperature and do not include wances for instrument errors. Design plant conditions represent the lower limit on primary rate when uncertainties in system resistance and pump head are included, and represent the er limit on vessel inlet temperature when design margins on steam generator performance are 3.5-7                                      Rev. 35


3.5.3.1.2 Parameter UncertaintiesTables 14.0.7-2 through 14.0.7-5 identif y parameter uncertainties included in the AREVA thermal and hydraulic and DNB methodology. Plant instrument calibrat ion procedures and related specification requirements are designed so that these uncertainties do not increase.
h the design plant parameters. During steady state operation, the possible variations in these meters define an operating envelope. One combination of these parameters gives the NBR, and this combination is utilized in Chapter 14 as the initial conditions in transient and dent analysis. Table 3.5-1 lists the nominal plant parameters.
3.5.3.2 Hydraulic Instability Analysis Boiling flows may be suscepti ble to thermohydrodynamic instabili ties. These instabilities are undesirable in reactors since th ey may cause a change in ther mohydraulic conditions that may lead to a reduction in the DNB h eat flux or to undesired forced vi brations of core components. However, unlike in Boiling Water Reactors (BWRs), hydraulic stability is not a concern in PWR cores. This statement, which is discussed below, is supported by the literature and the state of the art on instabilities occurring in two-phase flow systems.Instabilities in vertical up-flow of a two-phase mixture in a heated channel can be broadly classified into several categories. Of these, th e following relevant instabilities are discussed.1.Flow Excursion MPS2 UFSAR3.5-9Rev. 35Also called Ledinegg Instability, this is well described in Ref. 3.5-10. This instability occurs when the slope of th e boiling channel pressure drop-flow rate curve (internal characteristic) becomes sm aller than the slope of the loop supply pressure drop-flow rate curve (external characteristic), i.e., where P is the pressure drop and G is the mass flow rate.In this manner, a negative flow perturba tion will be amplified as the internal pressure drop becomes larger than the exte rnal at the perturbed flow and the flow decelerates further until a stable point is reached.
2.8      Summary of Thermal and Hydraulic Parameters thermal and hydraulic parameters for the reactor are listed in Table 3.5-1.
If the core is considered as a single averag e channel, the external pressure and flow characteristics as s een by the core exhibit due to the pump characteristics. Th is negative slope is stabilizing.
3    THERMAL AND HYDRAULIC EVALUATION 3.1      Analytical Techniques and Uncertainties 3.1.1    XCOBRA-IIIC DNBR Analyses thermal-hydraulic simulations employed to evaluate the MDNBR were performed in ordance with AREVAs Nuclear Regulatory Commission (NRC) approved thermal-hydraulic hodology for mixed cores (Reference 3.5-8).
MDNBR performance of the core during anticipated transients will be demonstrated to meet thermal-hydraulic design criterion on DNBR through the performance of transient analysis of limiting events. The results of this analysis are included in Chapter 14.
3.1.2     Parameter Uncertainties les 14.0.7-2 through 14.0.7-5 identify parameter uncertainties included in the AREVA thermal hydraulic and DNB methodology. Plant instrument calibration procedures and related cification requirements are designed so that these uncertainties do not increase.
3.2 Hydraulic Instability Analysis ling flows may be susceptible to thermohydrodynamic instabilities. These instabilities are esirable in reactors since they may cause a change in thermohydraulic conditions that may to a reduction in the DNB heat flux or to undesired forced vibrations of core components.
wever, unlike in Boiling Water Reactors (BWRs), hydraulic stability is not a concern in PWR
: s. This statement, which is discussed below, is supported by the literature and the state of the on instabilities occurring in two-phase flow systems.
abilities in vertical up-flow of a two-phase mixture in a heated channel can be broadly sified into several categories. Of these, the following relevant instabilities are discussed.
: 1.     Flow Excursion 3.5-8                                      Rev. 35
 
curve (internal characteristic) becomes smaller than the slope of the loop supply pressure drop-flow rate curve (external characteristic), i.e.,
( P -) internal < d--------------
d--------------                            ( P -) ----- external dG                                    dG where P is the pressure drop and G is the mass flow rate.
In this manner, a negative flow perturbation will be amplified as the internal pressure drop becomes larger than the external at the perturbed flow and the flow decelerates further until a stable point is reached.
If the core is considered as a single average channel, the external pressure and flow characteristics as seen by the core exhibit
( P -) ----- external < 0 d--------------
dG due to the pump characteristics. This negative slope is stabilizing.
On the other hand, considering flow in a single limiting bundle, the other parallel flow paths impose a flat pressure drop versus flow relation where d(P)/dG = 0.
On the other hand, considering flow in a single limiting bundle, the other parallel flow paths impose a flat pressure drop versus flow relation where d(P)/dG = 0.
While this situation is less stable than th e average core assumption, it is mitigated by the cross flow and mixing between th is limiting bundle and the neighboring bundles. Ref. 3.5-11 shows experimentally a definite stabilizing influence of cross flow mixing.
While this situation is less stable than the average core assumption, it is mitigated by the cross flow and mixing between this limiting bundle and the neighboring bundles. Ref. 3.5-11 shows experimentally a definite stabilizing influence of cross flow mixing.
The internal pressure drop versus flow characteristics were shown to satisfy the Ledinegg stability criterion
( P -) ----- internal > 0 d--------------
dG for a wide range of conditions in the LOFT reactor (Ref. 3.5-12) which closely approximates a PWR core during nominal and worst case operating conditions.
Therefore, in conclusion, Ledinegg Instability is not a concern in PWR cores.
: 2. Density Wave Instability Dynamic instabilities may occur even when the static stability criterion is satisfied (pressure drop increases when flow increases). For a density wave dynamic 3.5-9                    Rev. 35


The internal pressure drop ve rsus flow characteristics were shown to satisfy the Ledinegg stability criterion for a wide range of conditions in the LOFT reactor (Ref. 3.5-12) which closely approximates a PWR core during nominal and worst case operating conditions.
increase is delayed. In the case of a sinusoidal inlet flow perturbation of particular frequency, the lagging pressure drop response is such that its instantaneous value supports the growth of the initial perturbation (Ref. 3.5-13). Such unstable behavior requires the delayed portion of the total pressure drop (in the two-phase region) to be large compared with the single-phase pressure drop. The onset of this instability depends on the operating conditions and the distribution of pressure drop along the channel, as well as the external loop characteristics. A vast body of literature and several computer programs for the analysis of density waves exists mainly for BWR concerns (see for example the collection of papers in Ref.
Therefore, in conclusion, Ledinegg Inst ability is not a concern in PWR cores.2.Density Wave InstabilityDynamic instabilities may occur even when the static stab ility criterion is satisfied (pressure drop increases wh en flow increases). For a density wave dynamic dP ()dG---------------
3.5-14). Inferences from BWR experience are drawn to dismiss the possibility of density wave instabilities in a PWR core:
-----internal dP ()dG--------------------external<dP ()dG--------------------external0
* Unlike a BWR, there is no riser section contributing significantly to the 2-phase pressure drop.
<dP ()dG--------------------internal0
* For a single limiting channel with a constant pressure drop boundary condition, the cross flow in a PWR core has a stabilizing effect.
>
* Density wave oscillations are known to be stabilized with increasing pressure (decreasing enthalpy and density difference between the two phases). No unstable density wave oscillations could be obtained for pressures higher than 1200 psia (Ref. 3.5-15).
MPS2 UFSAR3.5-10Rev. 35instability, consider an inlet flow increase perturbing the initial value. The rate of enthalpy rise and density eff ects will travel up the channel, and the pressure drop increase is delayed. In the case of a sinusoi dal inlet flow pertur bation of particular frequency, the lagging pressure drop respons e is such that its instantaneous value supports the growth of the initial perturbation (Ref. 3.5-13). Such unstable behavior requires the delayed portion of the total pressure drop (in the two-phase region) to be large compared with the si ngle-phase pressure dr op. The onset of this instability depends on the operating conditi ons and the distribution of pressure drop along the channel, as well as the external loop char acteristics. A vast body of literature and several computer programs for the analysis of density waves exists mainly for BWR concerns (see for exampl e the collection of papers in Ref.
* BWR oscillations occur when the saturated boiling boundary is low (elevation <<4 feet). For a PWR, such boiling boundary can be achieved at nominal flow rates by more than doubling the power, which leaves a considerable stability margin even for the worst case transient.
3.5-14). Inferences from BWR experience ar e drawn to dismiss the possibility of density wave instabilities in a PWR core:*Unlike a BWR, there is no riser sect ion contributing significantly to the 2-phase pressure drop.*For a single limiting channel with a constant pressure drop boundary condition, the cross flow in a PWR core has a stabilizing effect.
* Considering the nuclear coupling, the void-reactivity coefficient in a PWR is reduced when the coolant is borated. Such reduction in the void-reactivity coefficient is stabilizing to this mode of oscillation.
*Density wave oscillations are known to be stabilized with increasing pressure (decreasing enthalpy and density diff erence between the two phases). No unstable density wave os cillations could be obtained for pressures higher than 1200 psia (Ref. 3.5-15).*BWR oscillations occur when the saturated boiling boundary is low (elevation <<4 feet). For a PWR, such boiling boundary can be achieved at nominal flow rates by more than doubling the power, which leaves a  
* For a density wave coupled with an out-of-phase neutron flux oscillation mode, the large subcritical reactivity of the first flux harmonic stabilizes this mode of hydraulic-neutronic oscillation. This is due to the PWR core being small compared with typical BWR cores.
The LOFT reactor stability study also addressed the density wave oscillations and concluded that these are not likely (Ref. 3.5-12).
In conclusion, Density Wave Instability is not a concern in PWR cores.
: 3. Flow Pattern Transition Instability 3.5-10                                      Rev. 35


considerable stability margin even for the worst case transient.*Considering the nuclear coupling, the void-reactivity coefficient in a PWR is reduced when the coolant is borated. Such reduction in the void-reactivity coefficient is stabiliz ing to this mode of oscillation.*For a density wave coupled with an out-of-phase neut ron flux oscillation mode, the larg e subcritical reactivity of the first flux harmonic stabilizes this mode of hydraulic-neu tronic oscillation. This is due to the PWR core being small compared with typical BWR cores.
channel experiences a succession of high void and low void flows as a vapor slug passes through (Ref. 3.5-12). As a vapor slug clears the channel exit, the average void content in the channel is temporarily reduced and vice versa resulting in pressure drop and flow rate oscillations. In a worst case condition in a PWR, slug flow may occur in a small number of channels near the exit. No significant oscillatory response is expected, particularly since the slug formation is limited to a short segment near the exit of the hot channels.
The LOFT reactor stability study also addr essed the density wave oscillations and concluded that these ar e not likely (Ref. 3.5-12).In conclusion, Density Wave Instability is not a concern in PWR cores.3.Flow Pattern Transition Instability MPS2 UFSAR3.5-11Rev. 35 The term "Flow Pattern Instability" is us ed in the literature in two connotations.
The more common meaning of the Flow Pattern Transition Instability refers to unstable transitions between bubbly and annular flow (Ref. 3.5-10). A flow rate perturbation decreasing the flow rate and increasing the void fraction will result in flow transition from bubbly-slug to annular pattern. The annular flow is characterized with lower pressure drop, which results in accelerating the flow. The increase in flow rate brings the void fraction back below the value required to support annular flow. Thus the transition back to bubbly-slug regime takes place.
The first refers to the slug flow pattern where a particular elevation in a heated channel experiences a succes sion of high void and low void flows as a vapor slug passes through (Ref. 3.5-12). As a vapor slug clears the channel exit, the average void content in the channel is temporaril y reduced and vice versa resulting in pressure drop and flow rate oscillations.
Extensive work has been done on flow pattern transition (see for example Ref.
In a worst case condition in a PWR, slug flow may occur in a small number of ch annels near the exit. No significant oscillatory response is expect ed, particularly since the slug formation is limited to a short segment near the ex it of the hot channels.The more common meaning of the "Flow Pattern Transition Instab ility" refers to unstable transitions between bubbly and a nnular flow (Ref. 3.5-10). A flow rate perturbation decreasing the flow rate and increasing the voi d fraction will result in flow transition from bubbly-slug to annul ar pattern. The annular flow is characterized with lower pressure drop, which results in accelerating the flow. The increase in flow rate brings the void fraction back below the value required to support annular flow. Thus the transiti on back to bubbly-slug regime takes place.
3.5-16). Most work was limited to pressures of 1000 psia and below where these transitions are more distinct. At higher pressures, Hosler (Ref. 3.5-17) notes for 1400 and 2000 psia, that the flow appears more homogeneous with no reliable observation of pattern transition.
Extensive work has been done on flow pa ttern transition (see for example Ref.
Weisman et. al. (Ref. 3.5-18) observed no premature DNB due to bubbly-to-slug flow transition which they expected as the range of tested void fractions covers the transition range. Hosler (Ref. 3.5-17), on the other hand, noted that CHF occurred via a film dryout mechanism in established annular flow, which is far from the transition boundary to bubbly-slug pattern.
3.5-16). Most work was limited to pressu res of 1000 psia and below where these transitions are more distin ct. At higher pressures, Ho sler (Ref. 3.5-17) notes for 1400 and 2000 psia, that the flow appear s more homogeneous with no reliable observation of pattern transition.Weisman et. al. (Ref. 3.5-18) observed no premature DNB due to bubbly-to-slug flow transition which they expected as the range of tested void fractions covers the transition range. Hosler (Ref. 3.5-17), on the other hand, noted that CHF occurred via a film dryout mechanism in established annular flow, which is far from the transition boundary to bubbly-slug pattern.
In conclusion, Flow Pattern Transition Instability is not a concern in PWR cores.
In conclusion, Flow Pattern Transition Instability is not a concern in PWR cores.
3.5.3.3 Core Hydraulics 3.5.3.3.1 Fuel Assembly Pressure Drop CoefficientsPressure drop coefficients for the AREVA reload fuel presented are deri ved from pressure drop tests performed in AREVA's portable hydraulic test facility (Reference 3.5-5). The pressure drop coefficients are for the liquid phase and ar e referenced to the bare rod flow area.For reload Batches M (Cycle 10), N (Cycle 11), a nd P (Cycle 12) the pressure drop coefficient for the lower tie plate/spacer combination includes the effects of a debris resistant spacer. The reload Batch R (Cycle 13) and S (Cycle 14) fuel assemblies implemented an alternative debris resistant design which has a slightly lowe r pressure drop across the lower tie plate/spacer combination compared to the Batch P and prio r fuel assemblies. As a result, for Cycles 13 and 14, this results MPS2 UFSAR3.5-12Rev. 35 in higher inlet flows to the Ba tch R and S assemblies and a re duction in flow to the surrounding Batch M, N, and P reinsert assemblies.Due to crossflow effects, the decr eased flow will equilibrate with adjacent assemblies within the next one or two spacers.
3.3   Core Hydraulics 3.3.1   Fuel Assembly Pressure Drop Coefficients ssure drop coefficients for the AREVA reload fuel presented are derived from pressure drop s performed in AREVAs portable hydraulic test facility (Reference 3.5-5). The pressure drop fficients are for the liquid phase and are referenced to the bare rod flow area.
Limiting MDNBRs occur toward the top of the core. Therefore, the slight redistribution in the inlet flows, due to the new lower tie plate and adjacent spacer, will not affect calculated MDNBRs.
reload Batches M (Cycle 10), N (Cycle 11), and P (Cycle 12) the pressure drop coefficient for lower tie plate/spacer combination includes the effects of a debris resistant spacer. The reload ch R (Cycle 13) and S (Cycle 14) fuel assemblies implemented an alternative debris resistant gn which has a slightly lower pressure drop across the lower tie plate/spacer combination pared to the Batch P and prior fuel assemblies. As a result, for Cycles 13 and 14, this results 3.5-11                                    Rev. 35
 
to crossflow effects, the decreased flow will equilibrate with adjacent assemblies within the t one or two spacers. Limiting MDNBRs occur toward the top of the core. Therefore, the slight stribution in the inlet flows, due to the new lower tie plate and adjacent spacer, will not affect ulated MDNBRs.
Batch T and later HTP fuel assemblies have a lower total pressure drop than the previous etallic fuel assemblies (i.e., Batch S and prior). A thermal hydraulic compatibility analysis performed in Reference 3.5-23 for HTP fuel assemblies co-resident with bimetallic fuel mblies in the Millstone Unit 2 core. This analysis demonstrates that the two fuel assembly es are compatible. Of note is that the core pressure drop would decrease by approximately
% from the all bimetallic core (Cycle 14) to an all HTP core. The core pressure drop decrease m Cycle 14 to Cycle 15 will be approximately 0.69% since the Cycle 15 core has 80 HTP tch T) fuel assemblies and 137 bimetallic (Batch N, P, and R) fuel assemblies. Use of the HMP cer in the lowermost position (Reload Y and later) has a negligible effect on core differential sure.
3.3.2    Guide Tube Bypass Flow and Heating Analysis guide tube thermal-hydraulic design calculations are performed to demonstrate adequate ling of the CEA fingers and to ensure that bypass flow through the guide tubes does not uly reduce core flow.
w enters the guide tube through the weep holes and cap screw and exits through the top of the de tube. In the Millstone Unit 2 core, there are 81 assemblies under CEA positions. Of these, assemblies are under active CEA positions. The CEA fingers extend a short distance into the de tube in these 73 assemblies at the all-rods-out (ARO) position which provides a substantial uction in the guide tube bypass flow. The remaining eight assemblies were originally under the length CEAs which have been removed. In these eight assemblies, the flow is unimpeded, e the last flow plugging devices were removed in Cycle 12. The assembly guide tubes of 91 mblies project a short distance into close fitting sockets in the upper alignment plate. The lting flow annulus represents a significant resistance to guide tube bypass flow in these mblies. The remaining 45 core locations are instrument tube locations. In these locations, the pheral guide tubes also project a short distance into close fitting sockets in the upper nment plate. The center guide tube contains instrumentation which produces a flow annulus ch in turn reduces the flow in the center guide tubes.
guide tube model employed in the flow and heating calculations uses loss coefficients to rmine the guide tube flow path hydraulic losses. The core pressure drop at rated power and is employed as the driving force for flow through the guide tube. The model permits ulation of the guide tube configurations described above. The guide tube thermal model udes the effects of coolant heating by gamma deposition and neutron deceleration. The effects heating due to neutron absorption and gamma deposition in the inserted control rod are 3.5-12                                    Rev. 35
 
culations were performed to assess the maximum expected guide tube bypass flow ference 3.5-6). At hot full power (HFP), ARO configuration was selected as that resulting in greatest bypass flow. The total core bypass flow, including flow through the guide tubes in this ance, was determined to be less than 4.0 percent of vessel flow. The result confirms that guide bypass flow does not unduly reduce core flow.
assess the adequacy of guide tube cooling, a simulation was also performed for a single mbly with the CEA fully inserted at HFP conditions. The fully inserted CEA fingers stantially increase the hydraulic resistance in the guide tube, and also represent a significant t source. The exit coolant temperature is well below saturation. Heat transfer through the guide wall provides a significant part of the cooling.
ed on the results described above, it is concluded that ample guide tube cooling is afforded by current design, and that bypass flow remains within acceptable limits.
3.3.3    Control Element Assembly Insertion Time Analysis rge data base of CEA insertion time measurements has been obtained at a CE plant similar to lstone Unit 2, with fuel identical in pertinent guide tube design characteristics to the Millstone t 2 AREVA reload fuel. The measurements span a time period during which reload quantities REVA fuel resided in the core. Statistical analysis (Reference 3.5-6) of this data indicates that CEA 90 percent insertion time is equal to or less than 2.5 seconds, which is well below the imum acceptable 90 percent insertion time of 2.75 seconds specified in the Technical cifications.
r 500 CEA insertion time measurements from nine different tests were analyzed. The surements reflect the time required to reach 90 percent insertion from interruption of power to CEA drive mechanism. Approximately six standard deviations separate the mean of the sured CEA insertion time data from the 2.75 second maximum allowable for Millstone Unit h over 500 data points, higher order statistics may also be applied to the data to conclude that rod drop time will be equal to or less than the greatest time measured in the tests with a bability of 99 percent at a 99 percent confidence level. The greatest rod drop time in the tests, oted above, was 2.50 seconds. The AREVA assemblies are, therefore, expected to conform to maximum CEA 90 percent insertion time of 2.75 seconds with a substantial margin.
3.3.4    Fuel Assembly Liftoff hydraulic lift force on the fuel assembly was calculated (Reference 3.5-6) using the drag fficient for a 14 by 14 fuel assembly with bimetallic spacer grids. This value differed slightly Reload Batches M, N, and P (Cycles 10, 11, and 12). The replacement of a bimetallic spacer h a debris resistant Inconel HTP spacer increased the drag while the thermal rounding of the 3.5-13                                    Rev. 35
 
1194 pounds. The assembly weight and spring force totals 1801 lbs, thus providing a 607 nd holddown margin. This margin, which is more than half of the worst case steady state lift e, will envelope any minor variation due to the spacer modifications. It will also provide ddown during and after a 20% pump overspeed resulting in a 44% lift force increase. For oad Batch R (Cycle 13) and Batch S, the fuel assembly weight increased by approximately 40 nds and a bimetallic spacer replaced the Inconel HTP spacer, increasing the margin to liftoff.
imilar analysis was performed for the Reload T design. The use of HMP spacers beginning h Reload Y has a negligible effect on lift.
maximum shear stress of 84,062 psi in the holddown springs occurs in the cold reactor dition. This is below the design criterion of 100,000 psi. The stress at reactor operating ditions is 74,188 psi, which is below the criterion of 90,000 psi at operating temperature.
diation may cause some stress relaxation of the Inconel X-750 holddown springs while sing irradiation induced growth of the fuel assemblies. The assembly growth results in higher ng deflection which offsets any radiation induced relaxation of the springs. The springs are ially shrouded in spring cups, which minimize flow-induced vibration of the springs and vent potential fretting wear.
4    TESTS AND INSPECTIONS 4.1    Reactor Testing rmal-hydraulic design criteria are verified during plant startup testing. This is accomplished measuring the primary intrinsic parameters (e.g., levels, pressures, temperatures, flows, tron fluence and differential pressures) and calculating the non-measurable and extrinsic meters (e.g., power level, core peaking factors). During the operating cycle, various mal-hydraulic parameters are periodically monitored to ensure compliance with the Technical cifications.
4.2    AREVA DNB and Hydraulic Testing 4.2.1    DNB Testing ails of the testing supporting the HTP DNB correlation are contained in Reference 3.5-3.
4.2.2    Fuel Assembly Hydraulic Testing gle-phase hydraulic characteristics of the AREVA Millstone Unit 2 fuel assembly were erimentally determined by hydraulic tests (Reference 3.5-5) performed in AREVAs Portable raulic Test Facility (PHTF).
pressure drop testing characterized the component loss/flow coefficients of the lower tie plate luding the inlet hardware), spacers, and the upper tie plate (including the exit hardware).
3.5-14                                    Rev. 35
 
were used to drive empirical relationships, which describe the single-phase pressure drops of Millstone Unit 2 fuel assembly and its components.
se test data from Reference 3.5-5 were used to calculate the Batch M, N, and P lower tie plate, cer, and upper tie plate pressure drop coefficients, and the bare rod friction factor. Additional data and analyses were used to determine the Batch R lower tie plate pressure drop coefficient elations. The loss/flow coefficients derived from these tests and calculations are all referenced he bare rod Reynolds Number.
5    REFERENCES 1    XN-NF-81-58(P)(A), Revision 2, and Supplements 1 and 2, RODEX2 Fuel Rod Thermal-Mechanical Response Evaluation Model, March 1984.
2    XN-NF-75-21(P)(A), Revision 2, XCOBRA-IIIC: A Computer Code to Determine the Distribution of Coolant During Steady-State and Transient Core Operation, January 1986.
3    EMF-92-153(P)(A) Rev. 1, HTP: Departure From Nucleate Boiling Correlation for High Thermal Performance Fuel, Siemens Power Corporation, January 2005.
4    XN-75-32(P)(A), Supplements 1, 2, 3, and 4, Computational Procedure for Evaluating Fuel Rod Bowing, October 1983.
5    ANF-89-018(P), Single-Phase Hydraulic Flow Test of ANF Millstone-2 Fuel Assembly, January 1989.
6    ANF-88-088(P), Revision 1, Design Report for Millstone Point Unit 2, Reload ANF-1, August 1988.
7    BNWL-1695, COBRA-IIIC: A Digital Computer Program for Steady-State and Transient Thermal-Hydraulic Analysis of Rod Bundle Nuclear Fuel Elements, March 1973.
8    XN-NF-82-21(P)(A), Revision 1, Application of Exxon Nuclear Company PWR Thermal Margin Methodology to Mixed Core Configurations, September 1983.
9    EMF-2135, Revision 0, Millstone Unit 2 Cycle 13 Extended Shutdown Safety Analysis Report, January 1999.
10    J. A. Boure, A. E. Bergles, and L. S. Tong, Review of Two-Phase Flow Instability, ASME Paper 71-HT-42, August 1971.
3.5-15                                    Rev. 35
 
Transfer Conf., pp. 235-239, Tokyo, Japan (September 1974).
12 S. A. Eide, Instability Study for LOFT for L2-1, L2-2 and L2-3 Pretest Steady State Operating Conditions, RE-A-78-096, Idaho National Engineering Laboratory, November 1978.
13 J. March-Leuba, Density-Wave Instabilities in Boiling Water Reactors, Oak Ridge National Laboratory Report ORNL/TM-12130 (September 1992).
14 Proceedings of the International Workshop on Boiling Water Reactor Stability, Committee on the Safety of Nuclear Reactors Installations, OECD Nuclear Energy Agency, Holtsville, NY (October 1990).
15 H. S. Kao, C. D. Morgan, and W. B. Parker, Prediction of Flow Oscillation in Reactor Core Channel, Trans. ANS Vol. 16, pp. 212-213 (1973).
16 A. E. Bergles and M. Suo, Investigation of Boiling Water Flow Regimes at High Pressure, Dynatech Corp. NYO-3304-8 (February 1966).
17 E. R. Hosler, Flow Patterns in High Pressure Two-Phase (Steam-Water) Flow with Heat Addition, 9th National Heat Transfer Conferrence, Chemical Engineering Progress Symposium Series, Number 82, Vol. 64, pp. 54-66 (August 1967).
18 Weisman et. al., Experimental Determination of the Departure from Nucleate Boiling in Large Rod Bundles at High Pressure, 9th National Heat Transfer Conference, Chemical Engineering Progress Symposium Series, Number 82, Vol. 64, pp. 114-125 (August 1967).
19 Reference Deleted 20 Letter, R. I. Wescott (SPC) to C. H. Wu (NU), Transmittal of Bases for New Uncertainties in the Setpoint Analysis for Millstone Unit 2, RIW:97:049, February 27, 1998.
21 Reference Deleted by FSARCR 06-MP2-016.
22 Qualification of Exxon Nuclear Fuel for Extended Burnup, XN-NF-82-06(P)(A)
Revision 1 and Supplements 2, 4 and 5, Exxon Nuclear Company, October 1986.
23 EMF-2664, Rev. 0, Millstone Unit 2 Thermal Hydraulic Compatibility Analysis, January 2002.
3.5-16                                  Rev. 35
 
Design and Operating Parameters                            Value re Rated Power                                      2700 MWt ction of Heat Generated in Fuel                    0.975 mary System Pressure                                2250 psia re Inlet Temperature                                549&deg;F actor Coolant Flow (Minimum)                        360,000 gpm a sembly Pitch                                        8.18 inches pass Flow Fraction (Best Estimate)                  0.0303 erage Linear Heat Rate                              6.206 kW/ft tal Number of Assemblies                            217 Flow reductions to 349,200 gpm are compensated for by reductions in the FrT and linear heat rate limits.
l Parameters Design and Operating Parameters                            Value el Rod OD                                            0.440 inches ide Tube OD (above dashpot)                        1.115 inches d Array                                              14 by 14 d Pitch                                              0.580 inches mber of Fuel Rods/Assembly                          176 mber of Guide Tubes/Assembly                        5 tive Fuel Length                                    136.7 inches tal Fuel Rod Assembly Length                        146.25 inches mber of Spacers                                    9 3.5-17                                  Rev. 35
 
Component                              Load Description                                      Load Value Core Support Barre            Radial pressure differential directed inward opposite inlet 40 psi duct Core Support Barrel and Upper Uplift load                                                480,000 pounds Guide Structure Flow Skirt                    Radial pressure differential directed inward                6.0 psi average, 10.2 psi maximum, over 40&deg; sector Bottom Plate                  Pressure differential load directed upward                  43,400 pounds Core Support Plate            Pressure differential load directed upward                  43,100 pounds Fuel Assembly                Uplift load                                                1194 lbs at 120% flow Core Shroud                  Radial load directed outward                                20.8 psi at bottom, 0.0 psi at top Upper Guide Structure        Pressure differential load directed upward                  148,000 pounds Fuel Alignment Plate          Pressure differential load directed upward                  89,600 pounds Upper Guide Plate            Pressure differential load directed downward                132,000 pounds CEA Shrouds                  Lateral drag load                                          4,200 pounds (dual CEA) 1,100 pounds (single CEA) 3.5-18                                                          Rev
 
BLE 3.5-3 UNCERTAINTY SOURCES FOR DNBR CALCULATIONS (DELETED) 3.5-19                    Rev. 35
 
.1 SEISMIC ANALYSIS
.1.1 Introduction amic analyses of the reactor vessel internals for both horizontal and vertical seismic itation were conducted to provide further bases for assessing the adequacy of their seismic gn. These analyses were performed in conjunction with the dynamic seismic analyses of the tor coolant system (RCS) which is discussed in Appendix 4.A. The following paragraphs vide a discussion of the analytical procedures used for the reactor internals, including a cription of the mathematical models. Significant results are listed and compared to the results ined from application of the design loads.
.1.2 Method of Analysis
.1.2.1 General procedure used in conducting the seismic analysis of the reactor internals consisted basically hree steps. The first step involved the formulation of a mathematical model. The natural uencies and mode shape of the model were determined during the second step. The response he model to the seismic excitation was determined in the third step. In this analysis, the zontal and vertical components of the seismic excitation were considered separately and the imum responses added to obtain conservative results.
.1.2.2 Mathematical Models the dynamic analysis of the reactor internals, equivalent multi-mass mathematical models e developed to represent the system. Since the seismic input excitation of the reactor internals obtained in the form of acceleration time history of the reactor vessel flange, only the rnals are included in the model. The coupling effect of the internals response on the vessel ge acceleration was accounted for by including a simplified representation of the reactor rnals with the model of the RCS. This is discussed in Appendix 4.A. Since the horizontal and ical responses were treated as uncoupled, separate horizontal and vertical models were eloped to more efficiently account for the structural differences in these directions. A sketch of internals showing the relative node locations for the horizontal model is presented in ure 3.A-1. Figures 3.A-2 and 3.A-3 show the idealized horizontal and vertical models. Since structural details provide for no vertical load transfer between the upper guide structure (UGS) core or core shroud, the vertical response of the UGS is independent of the rest of the rnals. Consequently, the vertical model was divided into two submodels. Model I consists of core support barrel/thermal shield (CSB/TS), lower support structure, core shroud and core s; Model II consists of the UGS.
mathematical models of the internals are constructed in terms of lumped masses and elastic m elements. At appropriate locations within the internals, points (nodes) are chosen to lump weights of the structure. Between these nodes, properties are calculated for moments of 3.A-1                                    Rev. 35
 
.1.2.2.1    Hydrodynamic Effects dynamic analysis of reactor internals presents some special problems due to their immersion confined fluid. It has been shown both analytically and experimentally (Reference 3.A-1) that ersion of a body in a dense fluid medium lowers its natural frequency and significantly alters vibratory response as compared to that in air. The effect is more pronounced where the fining boundaries of the fluid are in close proximity to the vibrating body as is the case for the tor internals. The method of accounting for the effects of a surrounding fluid on a vibrating em has been to ascribe to the system additional or hydrodynamic mass.
s hydrodynamic mass decreases the frequencies of the system, but is not directly involved in inertia force effects. The hydrodynamic mass of an immersed system is a function of the ensions of the real mass and the space between the real mass and confining boundary.
rodynamic mass effects for moving cylinders in a water annulus are discussed in References
-1 and 3.A-2. The results of these references are applied to the internals structures to obtain total (structural plus hydrodynamic) mass matrix which was then used in the evaluation of the ral frequencies and mode shapes for the model.
.1.2.2.2    Fuel Assemblies the horizontal model, the fuel assemblies are treated as vibrating in unison. The member perties for the beam elements representing the fuel assemblies were derived from the results of erimental tests of the fuel assembly load deflection characteristics and natural frequency.
.1.2.2.3    Core Support Barrel Flanges obtain accurate lateral and vertical stiffnesses of the upper and lower flanges, finite element lyses of these two regions were performed. As shown in Figures 3.A-4 and 3.A-5, the flanges e modeled with quadrilateral and triangular ring elements. Asymmetric loads, equivalent to ral shear loads and bending moments, and symmetric axial loads were applied and the lting displacements calculated. These results were then used to derive the equivalent member perties for the flanges.
.1.2.2.4    Control Element Assembly Shrouds the horizontal model, the control element assembly (CEA) shrouds are treated as vibrating in on and are modeled as guided cantilever beams in parallel. To account for the decreased ral stiffness of the UGS due to local bending of the fuel alignment plate, a short member with perties approximating the local bending stiffness of the fuel alignment plate is included at the om of the CEA shrouds. Since the stiffness of the UGS support plate is large compared to that he shrouds, the CEA shrouds are assumed to be rigidly connected to the UGS support plate.
3.A-2                                      Rev. 35
 
the horizontal model, the thermal shield supports are modeled as horizontal members. The mber properties of the beam elements representing the positioning pins were based on the al stiffness of the circumferential set of pins. Likewise, the properties of the beam member esenting the support lugs were based on the tangential stiffness of the circumferential set of
. For the vertical model, the equivalent cross-section area of the bar element representing the port lugs was based on the axial bending stiffness of the circumferential set of lugs. For both horizontal and vertical models, the stiffness of the thermal shield supports includes the effect ocal deformation of the core support barrel.
.1.2.2.6    Upper Guide Structure Support Plate and Lower Support Structure Grid Beams se grid beam structures were modeled as plane grids. Displacements due to vertical (out of e) loads applied at the beam junctions were calculated through the use of the STRUDL puter code (Reference 3.A-3). Average stiffness values based on these results yielded ivalent member cross-section areas for the vertical model.
.1.2.3 Natural Frequencies and Normal Modes mass and beam element properties of the models were utilized in STAR, a computer program m the MRI/STARDYNE Analysis System programs (Reference 3.A-4) to obtain the natural uencies and mode shapes. This system utilizes the stiffness matrix method of structural lysis. The natural frequencies and mode shapes are extracted from the system of equations.
[K-Wn2 M]n = 0 where:
K = Model stiffness matrix M = Model mass matrix Wn = Natural circular frequency for the nth mode n = Normal mode shape matrix for nth mode mass matrix, M, includes the hydrodynamic and structural masses.
natural frequencies and mode shapes calculated for the first 3 modes for the horizontal model presented in Figures 3.A-6 through 3.A-8. The natural frequencies calculated for the vertical del are presented in Table 3.A-1. The modal data shown is typical and is presented for strative purposes. The effect of additional higher modes was included in the response analyses.
3.A-3                                    Rev. 35


The Batch T and later HTP fuel assemblies have a lower total pr essure drop than the previous bimetallic fuel assemblies (i.e
.1.2.4.1      Horizontal Direction time history analysis technique was utilized to obtain the response of the internals for the zontal seismic excitation. The horizontal excitation was specified as the acceleration time ory of the reactor vessel flange, resulting from the operational basis earthquake (OBE) (OBE
., Batch S and prior). A thermal hy draulic compatibility analysis was performed in Reference 3.5-23 for HTP fuel assemblies co-resident with bimetallic fuel assemblies in the Millstone Unit 2 core. This analysis demonstrates that the two fuel assembly types are compatible. Of note is that the core pressure drop would decrease by approximately 1.5% from the all bimetallic core (Cycle 14) to an all HTP core. The core pressure drop decrease from Cycle 14 to Cycle 15 will be approximately 0.69% since the Cycle 15 core has 80 HTP (Batch T) fuel assemblies and 137 bimetallic (Batch N, P, and R) fuel assemblies. Use of the HMP spacer in the lowermost position (R eload Y and later) has a negligible effect on core differential pressure.
.09g ground acceleration). The flange excitation resulting from the design basis earthquake E) (DBE = 0.17g ground acceleration) was conservatively specified as 0.17/0.09 times that the OBE.
3.5.3.3.2 Guide Tube Bypass Flow and Heating Analysis The guide tube thermal-hydrauli c design calculations are perfor med to demonstrate adequate cooling of the CEA fingers an d to ensure that bypass flow through the guide tubes does not unduly reduce core flow.
time history response analysis was performed utilizing the MRI STARDYNE System/
NRE 1 Computer Program. This program utilizes the Normal Mode Method to obtain time ory response of linear elastic structure. Details of the program and the Normal Mode hod are presented in References 3.A-4, 3.A-5 and 3.A-6.
ut to DYNRE 1 consisted of the modal data as determined in Section 3A.2.3, the modal ping factors, and the forcing function time history. This analysis used the modal data for all des with frequencies below 100 cps. This included the first 14 modes. Contributions from her modes are negligible.
modal damping factors were obtained by the method of Mass Mode Weighting which es:
M i  in  i n = -------------------------
M i  in where:
n = Modal damping factor Mi = Structural mass of mass node i lil = Absolute value of the mode shape as mass mode i i = Damping associated with pass point i damping factor assigned to the nodes representing the fuel assemblies was 5 percent. This is a servative value derived from proprietary experimental results. A value of 1 percent was used the other nodes.
output from the DYNRE 1 code consists of the nodal displacement, velocity, and acceleration e history relative to the base. The member bending moments and shears were obtained from STAR code (Reference 3.A-5) and were derived from the DYNRE 1 nodal displacement tors at the times of peak response.
3.A-4                                  Rev. 35


Flow enters the guide tube thro ugh the weep holes and cap screw and exits through the top of the guide tube. In the Millstone Unit 2 core, there are 81 assemblies under CEA positions. Of these, 73 assemblies are under ac tive CEA positions. The CEA fingers extend a short distance into the guide tube in these 73 assemblies at the all-rods-out (ARO) posit ion which provides a substantial reduction in the guide tube bypass flow. The remaining eight assemb lies were originally under the part length CEAs which have be en removed. In these eight assemblies, the flow is unimpeded, since the last flow plugging devices were removed in Cycle 12. The assembly guide tubes of 91 assemblies project a short distan ce into close fitting sockets in the upper alignment plate. The resulting flow annulus represen ts a significant resistance to gu ide tube bypass flow in these assemblies. The remaining 45 core locations are instrument tube locations. In these locations, the peripheral guide tubes also proj ect a short distance into clos e fitting sockets in the upper alignment plate. The center guide tube contains instrumentation which produces a flow annulus which in turn reduces the flow in the center guide tubes.
response of the reactor internals to the vertical excitation was obtained by the response ctrum technique. Because of the high natural frequencies and resulting low levels of responses the vertical direction, the more conservative spectrum response analysis results were used ead of time history results. The response spectrum utilized was derived from the vertical eleration time history at the reactor vessel flange. The spectrum curve is presented in ure 3.A-9.
The guide tube model employed in the flow and heating calculations uses loss coefficients to determine the guide tube flow path hydraulic losse
acceleration level corresponding to the natural frequency of each mode was selected from the ctrum curve. The response spectrum technique uses these acceleration values to determine the tia forces, accelerations, and displacements of each mode. The results for each mode were servatively combined on the basis of absolute values. For the vertical models, the first seven des were included in the results.
: s. The core pressure drop at rated power and flow is employed as the driv ing force for flow through the guide tube. The model permits calculation of the guide tube configurations described above.
.1.3 Results mbined results for the horizontal and vertical dynamic seismic analyses are presented in le 3.A-2 in terms of stresses at critical locations in the reactor internals for the DBE.
The guide tube thermal model includes the effects of coolant heating by gamma deposition and neutron deceleration. The effects of heating due to neutron abso rption and gamma deposition in the inserted control rod are MPS2 UFSAR3.5-13Rev. 35 evaluated. Heat transfer through the guide tube wa ll to the coolant in the surrounding assembly is accounted for in the model.
le 3.A-2 also lists the seismic stresses which result from application of the design loads cified for the DBE. A comparison shows the results of the dynamic analysis to be less severe.
Calculations were performed to assess th e maximum expected guide tube bypass flow (Reference 3.5-6). At hot full pow er (HFP), ARO configuration was se lected as that resulting in the greatest bypass flow. The total core bypass flow, including flow through the guide tubes in this instance, was determined to be le ss than 4.0 percent of vessel flow. The result conf irms that guide tube bypass flow does not unduly reduce core flow.To assess the adequacy of guide tube cooli ng, a simulation was also performed for a single assembly with the CEA fully inserted at HF P conditions. The fully inserted CEA fingers substantially increase the hydraulic resistance in the guide t ube, and also represent a significant heat source. The exit coolant temp erature is well below saturation.
.1.4 Conclusion concluded that the seismic loads specified for the design of the internals are adequate. All mic loads calculated by the dynamic seismic analysis are less than the design loads specified he DBE.
Heat transfer through the guide tube wall provides a significant part of the cooling.
.2 NORMAL OPERATING ANALYSIS ign analyses were performed on the reactor internals for normal operating conditions to onstrate that the mechanical design bases were satisfied. These design calculations included ropriate vibration analyses of the component assemblies. The flow induced vibration of the B/TS, during normal operation, was characterized as a forced response to deterministic and dom pressure fluctuations in the coolant. Methods were developed for predicting the response omponents to the hydraulic forcing functions.
phasis was placed on analysis and design of those components which were particularly critical susceptible to vibratory excitation, such as the thermal shield. Using a top supported, as osed to a bottom supported, thermal shield design improves stability as it eliminates a free e in the flow path. Increasing the number of upper supports and lower jackscrews, in the cific manner chosen, provides a much stiffer structure and the use of an all-welded shield inates local flexibilities and relative motion at bolted joints. Analytical studies show the mal shield to be stable on its support system when exposed to the axial annular flow ountered during normal operation. The snubber design is based upon limiting the motion of core support barrel under conditions of hydraulically induced vibrations. The snubbers are at 3.A-5                                    Rev. 35


Based on the results desc ribed above, it is concluded that ample guide tube cooling is afforded by the current design, and that bypass flow remains within acceptable limits.
ribution of snubbers assures restraint regardless of the direction of response.
random hydraulic forcing function was developed by analytical and experimental methods.
analytical expression was developed to define the turbulent pressure fluctuation for fully eloped flow. This expression was modified, based upon the result of scale model testing, to ount for the fact that flow in the downcomer was not fully developed. Based upon test results, expression was developed to define the spatial dependency of the turbulent pressure tuations. In addition, experimentally adjusted analytical expressions were developed to ne; the peak value of the pressure spectral density associated with the turbulence and; the imum area of coherence, in terms of the boundary layer displacement, across which the dom pressure fluctuations are in phase.
natural frequencies and mode shapes of the CSB/TS system were obtained using the ymmetric shell finite element computer program, ASHSD (Reference 3.A-7). This computer gram is capable of obtaining natural frequencies and mode shapes of complex axisymmetric ls; e.g., arbitrary meridional shape, varying thickness, branches, multi-materials, orthotropic erial properties, etc. To employ the ASHSD code, the CSB/TS were modeled as a series of ical shell frustrums joined at their nodal point circles. The length of each element, throughout ASHSD model, was a fraction of the shell decay length. Since rapid changes in the stress ern occur in regions of structural discontinuity, the nodal point circles were more closely ced in such regions. The finite element model of the CSB/TS system included representation he core support barrel upper and lower flanges, sections of different wall thickness, and mal shield support lugs and jackscrews. Elements with orthotropic material properties were zed to provide equivalent axisymmetric models of the structural stiffness and constraints to tive motion between the core support barrel and thermal shield provided by the thermal shield port lugs and jackscrews. Those modes which reflect the mass of the lower support structure, shroud and fuel were simulated by the addition of concentrated masses at specific nodes in core support barrel flange finite element model.
lying Hamiltons Variational Principle to the conical shell elements an equation of motion formulated for each degree of freedom of the system. An inverse iteration technique was zed in the program to obtain solutions to the characteristic equation, which was based on a onalized form of a consistent mass matrix and stiffness matrix developed using the finite ment method. Four degrees of freedom  radial displacement, circumferential displacement, ical displacement, and meridional rotation  were taken into account in the analysis, giving to coupled mode shapes and corresponding frequencies. Evaluation of the reduction of these uencies for the system immersed in coolant was made by means of the virtual mass method ined in Reference 3.A-2.
random response analysis considers the response of the CSB/TS system to the turbulent ncomer flow during steady-state operation. The random forcing function is assumed to be a e-band stationary random process with a pressure spectral density equal to the peak value ciated with the turbulence. The rms vibration level of the CSB/TS system was obtained based n a damped, single degree of freedom analysis assuming the rms random pressure fluctuation 3.A-6                                      Rev. 35


3.5.3.3.3 Control Element Assembly Insertion Time AnalysisA large data base of CEA insertion time measurements has been obtained at a CE plant similar to Millstone Unit 2, with fuel identical in pertinent guide tube design characteristics to the Millstone Unit 2 AREVA reload fuel. The measurements sp an a time period during which reload quantities of AREVA fuel resided in the core. Statistical analysis (Reference 3.5-
eloped by a Combustion Engineering (CE) consultant using the random loads discussed ve. Modeling the reactor vessel snubbers and core support barrel system as a single degree of dom spring-mass system, the number and magnitude of snubber, core support barrel impacts calculated based upon the response of the system to random excitation. The snubbers were gned, based upon this loading requirement, to meet the cyclic strength requirements specified ection III of the ASME Boiler and Pressure Vessel Code.
: 6) of this data indicates that the CEA 90 percent insertion time is equal to or less than 2.5 seconds, which is well below the maximum acceptable 90 percent insertion time of 2.75 seconds specified in the Technical Specifications.
forced response of the reactor internals to deterministic loading was evaluated by classical lytical methods, using lumped mass and continuous elastic structural models. These calculated onses were used to verify the structural integrity of the reactor vessel internals to normal rating vibratory excitation. Components were design analyzed to assure that there were no erse effects from dominant excitation frequencies, such as pump rotational and blade passing uencies.
Over 500 CEA insertion time measurements from nine different tests were analyzed. The measurements reflect the time required to reach 90 percent inserti on from interrupti on of power to the CEA drive mechanism. Approximately six stan dard deviations separate the mean of the measured CEA insertion time data from the 2.75 second maxi mum allowable for Millstone Unit 2.
.3 LOSS OF COOLANT ACCIDENT ANALYSIS
With over 500 data points, higher order statistics may also be applied to the data to conclude that the rod drop time will be equal to or less than the greatest time measured in the tests with a probability of 99 percent at a 99 percent confidence level. The gr eatest rod drop time in the tests, as noted above, was 2.50 seconds. The AREVA assembli es are, therefore, expected to conform to the maximum CEA 90 percent insertion time of 2.75 seconds with a substantial margin.
.3.1 Discussion ynamic analysis (Reference 3.A-8) has been performed to determine the structural response of reactor vessel internals to the transient loss of coolant accident (LOCA) loading. The analysis rmined the shell, beam and rigid body motions of the internals using established puterized structural response analyses. The finite-element computer code, ASHSD ference 3.A-7) was used to calculate the time-dependent beam and shell response of the CSB/
3.5.3.3.4 Fuel Assembly Liftoff The hydraulic lift force on the fuel assembly wa s calculated (Reference 3.5-6) using the drag coefficient for a 14 by 14 fuel assembly with bimeta llic spacer grids. This value differed slightly for Reload Batches M, N, and P (Cycles 10, 11, and 12). The replacement of a bimetallic spacer with a debris resistant Inconel HTP spacer increased the drag while the thermal rounding of the MPS2 UFSAR3.5-14Rev. 35 leading edges of the remaining bimetallic spacers decreased the drag. The overall effect was a slight increase in drag force. The total of th e buoyancy and hydraulic lift fo rces was calculated to be 1194 pounds. The assembly we ight and spring force totals 1801 lbs, thus providing a 607 pound holddown margin. This margin, wh ich is more than half of th e worst case steady state lift force, will envelope any minor variation due to the spacer modifications. It will also provide holddown during and after a 20% pump overspeed re sulting in a 44% lift force increase. For Reload Batch R (Cycle 13) and Batch S, the fuel assembly wei ght increased by approximately 40 pounds and a bimetallic sp acer replaced the Inconel HTP spacer, increasing the margin to liftoff.
system to the transient LOCA loading. The finite-element computer code SAMMSOR-NASOR (Reference 3.A-9) was used to evaluate the core support barrels potential for kling when loaded by a net external radial pressure resulting from an outlet line break. The ctural response of the reactor internals to vertical and transverse loads resulting from inlet and et breaks, was determined using the spring-mass computer code, SHOCK (Reference 3.A-10).
A similar analysis was performed for the Relo ad T design. The use of HMP spacers beginning with Reload Y has a negligible effect on lift.
time and space dependent pressure loads used in the above analysis were the result of a iled hydraulic blowdown analysis. The pressure fluctuations were determined for each node he hydraulic model for inlet and outlet line breaks. The pressure time histories at these nodal tions were then decomposed into the Fourier harmonics which define the circumferential sure distribution at the nodal elevations. Where the hydraulic model nodes did not correspond hose of the structural model, the hydraulic model pressure components were interpolated to vide the required loading information.
The maximum shear stress of 84,062 psi in the holddown springs occurs in the cold reactor condition. This is below the de sign criterion of 100,000 psi. The stress at reactor operating conditions is 74,188 psi, which is below the cr iterion of 90,000 psi at operating temperature.
finite element computer code, ASHSD, was used to calculate the dynamic response of the B/TS to transient LOCA loading resulting from an inlet break. To employ the ASHSD code, CSB/TS were modeled as a series of conical shell frustrums (elements) joined at their nodal nt circles. Applying Hamiltons Variational Principle to the conical shell elements a damped ation of motion was formulated for each degree of freedom of the system. Four degrees of dom  radial displacement, circumferential displacement, vertical displacement and idional rotation  were taken into account in the analysis, giving rise to coupled modes. The 3.A-7                                    Rev. 35
Irradiation may cause some stress relaxation of the Inc onel X-750 holddown springs while causing irradiation induced growth of the fuel assemblies. The as sembly growth results in higher spring deflection which offsets any radiation induced relaxation of the springs. The springs are partially shrouded in spring cups, which minimi ze flow-induced vibrati on of the springs and prevent potential fretting wear.


====3.5.4 TESTS====
h that it is small compared to the shortest period of the finite element system. The model eloped for the CSB/TS system is shown in Figure 3.A-10. The length of each element, ughout the analytical model, was a fraction of the shell decay length. Since rapid changes in stress pattern occur in regions of structural discontinuity, the nodal point circles were more ely spaced in such regions. The finite element model of the CSB/TS system included esentation of the core support barrel upper and lower flanges, sections of different wall kness, and thermal shield support lugs and jackscrews. Elements with orthotropic material perties were utilized to provide equivalent axisymmetric models of the structural stiffness and straints to relative motion between the core support barrel and thermal shield provided by the mal shield support lugs and jackscrews. Those modes which reflect the mass of the lower port structure, core shroud and fuel were stimulated by the addition of concentrated masses at cific nodes in the core support barrel flange finite element model.
AND INSPECTIONS 3.5.4.1 Reactor Testing Thermal-hydraulic design cr iteria are verified during plant star tup testing. This is accomplished by measuring the primary intrinsi c parameters (e.g., levels, pres sures, temperatures, flows, neutron fluence and diff erential pressures) and calculating th e non-measurable and extrinsic parameters (e.g., power level, core peaking factors). During the operating cy cle, various thermal-hydraulic parameters are pe riodically monitored to ensure compliance with the Technical Specifications.
erforming the dynamic analysis of the CSB/TS system, the transient load harmonics were lied in two successive phases to account for time-dependent boundary conditions at the bbers. The first phase used those harmonics which excite the beam modes, whereas the second se used those harmonics which excite the shell modes. During the first phase, the lower end of core support barrel was unrestrained. Within a very few milliseconds, the clearances between core support barrel and reactor vessel snubbers were closed and for the remainder of the CA transient, the core support barrel was restrained radially at the snubber level. Transient onses were computed throughout each loading phase.
3.5.4.2 AREVA DNB and Hydraulic Testing 3.5.4.2.1 DNB Testing Details of the testing supporting the HTP DNB correlation are contained in Reference 3.5-3.
ASHSD code computed the nodal point displacement, resultant shell forces, shell stresses maximum principle stresses as functions of time. The maximum principle stresses at the rnal and external surfaces of the CSB/TS were determined from the bending and membrane ponents during each phase of transient loading. Stress intensity levels calculated from the ciple stresses were combined with normal operating and seismic induced stresses for parison with design criteria.
3.5.4.2.2 Fuel Assembly Hydraulic Testing Single-phase hydraulic characteristics of the AREVA Millstone Unit 2 fuel assembly were experimentally determined by hydraulic tests (R eference 3.5-5) performed in AREVA's Portable Hydraulic Test Facility (PHTF).
urate representation and analysis of the CSB/TS shell structures was obtained through use of finite element code ASHSD. Accurate representation of the remainder of the internals (i.e.,
The pressure drop testing characte rized the component loss/flow coef ficients of the lower tie plate (including the inlet hardware), spacers, and th e upper tie plate (includi ng the exit hardware).
, core shroud, CEAs, UGS, lower support structure, etc.) was obtained using the SHOCK e.
MPS2 UFSAR3.5-15Rev. 35Differential pressure meas urements were taken over a range of Reynolds Numbers (N Re). These data were used to drive empirica l relationships, which describe th e single-phase pre ssure drops of the Millstone Unit 2 fuel assembly and its components.These test data from Reference 3.5-5 were used to calculate the Batch M, N, and P lower tie plate, spacer, and upper tie plate pressure drop coefficients, and the bare rod friction factor. Additional test data and analyses were used to determine the Batch R lower tie plate pressure drop coefficient correlations. The loss/flow coefficients derived from these tests and calculations are all referenced to the bare rod Reynolds Number.
SHOCK code determines the response of structures which are represented as lumped-mass ems and subjected to arbitrary loading functions. The code solves the differential equations of ion for each mass by a numerical step-integration procedure. The lumped mass model can esent a vertically or laterally responding system subject to arbitrary loading functions and al conditions. Options are available for describing steady state loads, preloads, input elerations, linear and nonlinear springs (including tension and compression only springs) gaps, structural and viscous damping.
3.
reactor internals were developed in terms of a spring-mass system for both vertical and lateral ctions; see Figures 3.A-11 and 3.A-12. For both models, the spring rates were generally 3.A-8                                    Rev. 35


==5.5 REFERENCES==
del analyses. The lumped mass weights were generally based upon the mass distribution of the orm support structures, but included at appropriate nodes, local masses such as snubber cks, fuel end fittings, thermal shield lugs, etc. The net result was a lumped-mass system having same distribution of mass as the actual structure. To simulate the effect associated with the rnals oscillating laterally in the water filled vessel, a distributed virtual mass was calculated ed upon the procedure outlined in Reference 3.A-8 (which includes the annulus effect) and added to the structural lumped-mass system, to provide an analytical model with a dynamic onse quantitatively similar to the actual internals. In the case of the vertical model, the raulic effect is notably one of reducing the effective weight of the reactor internals and this ct was included in the structural lumped-mass system.
3.5-1XN-NF-81-58(P)(A), Revision 2, and Supplements 1 and 2, "RODEX2 Fuel Rod Thermal-Mechanical Response Evaluation Model," March 1984.3.5-2XN-NF-75-21(P)(A), Revision 2, "XCOBRA-IIIC: A Comput er Code to Determine the Distribution of Coolant During Steady-State and Transient Core Operation," January 1986.3.5-3EMF-92-153(P)(A) Rev. 1, "H TP: Departure From Nucleate Boiling Correlation for High Thermal Performance Fuel," Siem ens Power Corporation, January 2005.3.5-4XN-75-32(P)(A), Supplements 1, 2, 3, and 4, "Computational Procedure for Evaluating Fuel Rod Bowing," October 1983.3.5-5ANF-89-018(P), "Single-Phase Hydraulic Flow Te st of ANF Millstone-2 Fuel Assembly," January 1989.3.5-6ANF-88-088(P), Revision 1, "Design Report for Millst one Point Unit 2, Reload ANF-1," August 1988.3.5-7BNWL-1695, "COBRA-IIIC: A Digital Computer Program for Steady-State and Transient Thermal-Hydraulic Analysis of Rod Bundle Nuclear Fuel Elements," March 1973.3.5-8XN-NF-82-21(P)(A), Revision 1, "Appl ication of Exxon Nuclear Company PWR Thermal Margin Methodology to Mixed Co re Configurations," September 1983.
SHOCK code provided excellent facility for modeling clearances, preloads and component rfaces. In the lateral model, the core support barrel, reactor vessel snubber clearance was ulated by a nonlinear spring which accounted for the increased resistance to core support el motion when snubbing occurred. In the vertical model, nonlinear springs in the form of pression only springs, were used extensively to simulate preload and interface conditions, h as exist between the UGS support plate and core support barrel upper flange; at the fuel d-down spring; at the fuel, core support plate interface and at the core shroud, core support e interface. Tension only springs were used to simulate the effect of the core shroud tie rods.
3.5-9 EMF-2135, Revision 0, "Millstone Unit 2 Cycle 13 Extended Shutdown Safety Analysis Report," January 1999.3.5-10J. A. Boure, A. E. Bergles, and L. S. Tong, "Review of Two-Phase Flow Instability," ASME Paper 71-HT-42, August 1971.
oth the vertical and lateral SHOCK models, damping was varied throughout the system to ulate structural and hydraulic frictional effects within the reactor internals. The effect of raulic drag in the vertical model was simulated by a force time-history applied to the fuel er end-fitting. Vertical loads were used directly from the detailed hydraulic analysis, whereas ral loads were obtained by integrating those harmonics which excite the beam modes to obtain net lateral load on the CSB/TS system.
MPS2 UFSAR3.5-16Rev. 353.5-11S. Kakac et. al., "Sustained and Transient Boiling Flow Instabilities in a Cross-Connected Four-Parallel-Channel Upfl ow System," Fifth International Heat Transfer Conf., pp. 235-239, Tokyo, Japan (September 1974).3.5-12S. A. Eide, "Instability Study for LOFT for L2-1, L2-2 and L2-3 Pretest Steady State Operating Conditions," RE-A-78-096, Idaho National Engineering Laboratory, November 1978.3.5-13J. March-Leuba, "Density-Wave Instabilities in Boiling Wa ter Reactors," Oak Ridge National Laboratory Report ORNL/TM-12130 (September 1992).3.5-14Proceedings of the International Workshop on Boiling Water Reactor Stability, Committee on the Safety of Nuclear Reactors Installations, OECD Nuclear Energy Agency, Holtsville , NY (October 1990).3.5-15H. S. Kao, C. D. Morgan, and W. B. Parker , "Prediction of Flow Oscillation in Reactor Core Channel," Trans. ANS Vol. 16, pp. 212-213 (1973).3.5-16A. E. Bergles and M. Suo, "Investigation of Boiling Wa ter Fl ow Regimes at High Pressure," Dynatech Corp. NYO-3304-8 (February 1966).3.5-17E. R. Hosler, "Flow Patterns in High Pressure Two-Phase (Steam-Water) Flow with Heat Addition," 9th National Heat Transfer Conferrence, Chemical Engineering Progress Symposium Series, Number 82, Vol. 64, pp. 54-66 (August 1967).3.5-18Weisman et. al., "Experiment al Determination of the Depa rture from Nucleate Boiling in Large Rod Bundles at High Pressure," 9th National Heat Transfer Conference, Chemical Engineering Progress Symposium Series, Number 82, Vol. 64, pp. 114-125 (August 1967).3.5-19Reference Deleted3.5-20Letter, R. I. Wescott (SPC) to C. H. Wu (NU), "T ransmitt al of Bases for New Uncertainties in the Setpoint Analysis for Millstone Unit 2," RIW:97:049, February 27, 1998.3.5-21Reference Deleted by FSARCR 06-MP2-016.3.5-22"Qualification of Exxon Nuclear Fuel for Extended Burnup,"
SHOCK code calculated the vertical and lateral response of the system in terms of lacements, velocities and accelerations and internal force, moments and shears as related to h model. These quantities were sufficient to permit calculation of membrane and where ropriate bending stresses for comparison with design criteria.
XN-NF-82-06(P)(A) Revision 1 and Supplements 2, 4 and 5, Exxon Nuclear Company, October 1986.
finite-element code SAMMSOR-DYNASOR was used to determine the dynamic response of core support barrel, with initially imperfect geometry, to a net external radial pressure lting from an outlet line break. The above analysis has the capability of determining the linear dynamic response of axisymmetric shells with initial imperfections subjected to trarily varying load configurations.
3.5-23EMF-2664, Rev. 0, "Millstone Unit 2 Therma l Hydraulic Compatibility Analysis,"
ce SAMMSOR-DYNASOR is a finite-element program, a model was developed, Figure 3.A-of the core support barrel using axisymmetric finite-elements similar to those used for the HSD analysis. As was for the ASHSD model, the SAMMSOR-DYNASOR finite-element ths were considerably less than the decay length of the core support barrel. The boundary dition at the core support barrel flange was considered fixed, whereas at the core support el lower flange radial displacements were restrained. These boundary conditions represented 3.A-9                                   Rev. 35
January 2002.
MPS2 UFSAR3.5-17Rev. 35TABLE 3.5-1  NOMINAL REACTOR AND FUEL DESIGN PARAMETERS Design and Operating ParametersValue Core Rated Power 2700 MWt Fraction of Heat Ge nerated in Fuel 0.975 Primary System Pressure 2250 psiaCore Inlet Temperature 549&deg;F Reactor Coolant Flow (Minimum) 360,000 gpm aa.Flow reductions to 349,200 gpm are comp ensated for by reductions in the F r T and linear heat rate limits.Assembly Pitch8.18 inches Bypass Flow Fraction (Best Estimate)0.0303Average Linear Heat Rate6.206 kW/ftTotal Number of Assemblies217 Fuel Parameter sDesign and Operating ParametersValueFuel Rod OD0.440 inches Guide Tube OD (above dashpot)1.115 inches Rod Array14 by 14Rod Pitch0.580 inchesNumber of Fuel Rods/Assembly176 Number of Guide Tubes/Assembly5Active Fuel Length136.7 inchesTotal Fuel Rod Assembly Length146.25 inches Number of Spacers9 MPS2 UFSARMPS2 UFSAR3.5-18Rev. 35TABLE 3.5-2  DESIGN OPERATING HYDRAULIC LOADS ON VESSEL INTERNALS ComponentLoad DescriptionLoad Value Core Support BarreRadial pressure differ ential directed inward opposite inlet duct 40 psi Core Support Barrel and Upper Guide Structure Uplift load 480,000 pounds Flow SkirtRadial pressure differential directed inward 6.0 psi average, 10.2 psi maximum, over 40&deg; sector Bottom PlatePressure differential load directed upward 43,400 pounds Core Support PlatePressure differential load directed upward 43,100 pounds Fuel Assembly Uplift load1194 lbs at 120% flow Core Shroud Radial load directed out ward 20.8 psi at bottom, 0.0 psi at topUpper Guide StructurePressure differential load directed upward 148,000 pounds Fuel Alignment PlatePressure diff erential load directed upward 89,600 pounds Upper Guide PlatePressure differen tial load directed downward 132,000 pounds CEA Shrouds Lateral drag load 4,200 pounds (dual CEA) 1,100 pounds (single CEA)
MPS2 UFSAR3.5-19Rev. 35TABLE 3.5-3  UNCERTAINTY SOURCES FOR DNBR CALCULATIONS (DELETED)
MPS2 UFSAR3.A-1Rev. 35 3.A ANALYSIS OF REACTOR VESSEL INTERNALS 3.A.1 SEISMIC ANALYSIS 3.A.1.1 IntroductionDynamic analyses of the reacto r vessel internals for both hori zontal and vertical seismic excitation were conducted to prov ide further bases for assessing the adequacy of their seismic design. These analyses were perf ormed in conjunction with the dyna mic seismic analyses of the reactor coolant system (RCS) which is discus sed in Appendix 4.A. The following paragraphs provide a discussion of the anal ytical procedures used for th e reactor internals, including a description of the mathematical models. Significant results are listed and compared to the results obtained from applicati on of the design loads.
3.A.1.2 Method of Analysis 3.A.1.2.1 General The procedure used in conducting th e seismic analysis of the reacto r internals consisted basically of three steps. The first step involved the formulation of a ma thematical model. The natural frequencies and mode sh ape of the model were determined during th e second step. The response of the model to the seismic excitation was determined in the third step. In this analysis, the horizontal and vertical components of the seismic excitation were considered separately and the maximum responses added to ob tain conservative results.
3.A.1.2.2 Mathematical Models For the dynamic analysis of the re actor internals, equivalent mu lti-mass mathematical models were developed to represent the system. Since the seismic input ex citation of the reactor internals was obtaine d in the form of acceleration time history of the reactor vessel flange, only the internals are included in the model. The coupling eff ect of the internals' response on the vessel flange acceleration was accounted for by including a simplified representation of the reactor internals with the model of the RCS. This is discussed in Appendix 4.A.
Since the horizontal and vertical responses were treated as uncoupled, separate horizont al and vertical models were developed to more efficiently account for the structural differences in these directions. A sketch of the internals showing the relati ve node locations for the horiz ontal model is presented in Figure 3.A-1. Figures 3.A-2 and 3.A-3 show the ideal ized horizontal and vertical models. Since the structural details provide fo r no vertical load transfer betw een the upper guide structure (UGS) and core or core shroud, the vert ical response of the UGS is in dependent of the rest of the internals. Consequently, the vert ical model was divided into two submodels. Model I consists of the core support barrel/thermal shield (CSB/TS), lower support structure, core shroud and core mass; Model II consists of the UGS.
The mathematical models of the internals are constructed in terms of lumped masses and elastic beam elements. At appropriate lo cations within the internals, poi nts (nodes) are chosen to lump the weights of the structure.
Between these nodes, properties ar e calculated for moments of MPS2 UFSAR3.A-2Rev. 35inertia, cross-section areas, effective shear areas, and le ngths. The salient detail s of the models are discussed below.
3.A.1.2.2.1 Hydrodynamic Effects The dynamic analysis of r eactor internals presents some spec ial problems due to their immersion in a confined fluid. It has been shown both analytically and experimentally (Reference 3.A-1) that immersion of a body in a dense flui d medium lowers its natural freq uency and significantly alters its vibratory response as compar ed to that in air. The effect is more pronounced where the confining boundaries of the fluid are in close proximity to the vibrating body as is the case for the reactor internals. The method of accounting for the effects of a surroundi ng fluid on a vibrating system has been to ascribe to the system additional or "hydrodynamic mass."
This "hydrodynamic mass" decreases the frequencies of the system, but is not directly involved in the inertia force effects. Th e hydrodynamic mass of an immersed system is a f unction of the dimensions of the real mass and the space between the real mass and confining boundary.Hydrodynamic mass effects for mo ving cylinders in a water annulus are discussed in References 3.A-1 and 3.A-2. The results of th ese references are applied to the internals structures to obtain the total (structural plus hydrodynamic) mass matrix which was then used in the evaluation of the natural frequencies and m ode shapes for the model.
3.A.1.2.2.2 Fuel Assemblies For the horizontal model, the fu el assemblies are treated as vibrating in unison. The member properties for the beam elements representing the fuel assemblies we re derived from the results of experimental tests of the fuel assembly load deflection characteristics and natural frequency.
3.A.1.2.2.3 Core Support Barrel FlangesTo obtain accurate lateral and vertical stiffnesses of the upper a nd lower flanges, finite element analyses of these two regions were performed. As s hown in Figures 3.A-4 and 3.A-5, the flanges were modeled with quadrilateral and triangular ring elements. Asym metric loads, equivalent to lateral shear loads and bending moments, and symmetric axial loads were applied and the resulting displacements ca lculated. These results were then used to derive the e quivalent member properties for the flanges.
3.A.1.2.2.4 Control Element Assembly Shrouds For the horizontal model, the control element asse mbly (CEA) shrouds are tr eated as vibrating in unison and are modeled as guided cantilever beams in parallel. To acc ount for the decreased lateral stif fness of the UGS due to local bending of the fuel alignment plate, a short member with properties approximating the local bending stiffness of the fuel ali gnment plate is included at the bottom of the CEA shrouds. Since the stiffness of the UGS support plate is large compared to that of the shrouds, the CEA shrouds are assumed to be rigidly connected to the UGS support plate.
MPS2 UFSAR3.A-3Rev. 35 3.A.1.2.2.5 Thermal Shield Supports For the horizontal model, the th ermal shield supports are modele d as horizontal members. The member properties of the beam elements repr esenting the positioning pins were based on the radial stif fness of the circumfe rential set of pins. Likewise, th e properties of the beam member representing the support lugs were based on the tangential stiffness of the circumferential set of lugs. For the vertical model, th e equivalent cross-section area of the bar element representing the support lugs was based on the axial bending stiffness of the circumferential set of lugs. For both the horizontal and vertical models, the stiffness of the thermal shield supports includes the effect of local deformation of the core support barrel.
3.A.1.2.2.6 Upper Guide Structure Support Plate and Lower Support Structure Grid BeamsThese grid beam structures were modeled as plane grids. Displace ments due to vert ical (out of plane) loads applied at the be am junctions were calculated through the use of the STRUDL computer code (Reference 3.A-3). Average stiffness values based on these results yielded equivalent member cross-secti on areas for the vertical model.
3.A.1.2.3 Natural Frequencies and Normal ModesThe mass and beam element properties of the models were utilized in STAR, a computer program from the MRI/STARDYNE Analysis System programs (Reference 3.A-4) to obtain the natural frequencies and mode shapes. This system utilizes the "stiffness matrix" method of structural analysis. The natural frequencies and mode shapes are extracted from th e system of equations.
[K-W n 2 M]n = 0 where: K = Model stiffness matrix M = Model mass matrix W n = Natural circular frequency for the n th moden = Normal mode shape matrix for n th mode The mass matrix, M , includes the hydrodynamic and structural masses.
The natural frequencies and mode shapes calculated for the first 3 modes for the horizontal model are presented in Figures 3.A-6 through 3.A-8. The natural frequencie s calculated for the vertical model are presented in Table 3.A
-1. The modal data shown is typical and is presented for illustrative purposes. The effect of additional higher modes was included in the response analyses.
MPS2 UFSAR3.A-4Rev. 35 3.A.1.2.4  Response Calculations 3.A.1.2.4.1 Horizontal Direction The time history analysis technique was utilized to obtain th e response of the internals for the horizontal seismic excitation. The horizontal excitation was specifi ed as the acceleration time history of the reactor vessel fl ange, resulting from the operationa l basis earthquake (OBE) (OBE = 0.09g gr ound acceleration). The fla nge excitation resulting from the design basis earthquake (DBE) (DBE = 0.17g ground acceleration) was conservatively specified as 0.17/0.09 times that for the OBE.
The time history response analysis was perfo rmed utilizing the MRI STARDYNE System/
DYNRE 1 Computer Program. This program utilizes the "Normal Mode Method" to obtain time history response of linear elas tic structure. Detail s of the program a nd the "Normal Mode Method" are presented in Refe rences 3.A-4, 3.A-5 and 3.A-6.
Input to DYNRE 1 consisted of the modal data as determined in Section 3A.2.3, the modal damping factors, and the forcing function time history. This analysis used the modal data for all modes with frequencies below 100 cps. This in cluded the first 14 modes. Contributions from higher modes are negligible.
The modal damping factors were obtained by the method of "Mass Mode Weighting" which gives: where:n = Modal damping factor M i = Structural mass of mass node i lil = Absolute value of the mode shape as mass mode ii = Damping associated with pass point i The damping factor assigned to the nodes representing the fuel assemblies was 5 percent. This is a conservative value derived from proprietary experimental results.
A value of 1 percent was used for the other nodes.
The output from the DYNRE 1 code consists of the noda l displacement, velocity, and acceleration time history relative to the base. The member bendi ng moments and shears were obtained from the STAR code (Reference 3.A-5) and were derived from the DYNRE 1 nodal displacement vectors at the times of peak response.nM iiniM iin--------------
-----------
-=
MPS2 UFSAR3.A-5Rev. 35 3.A.1.2.4.2 Vertical DirectionThe response of the reactor internals to the vertical excitation was obtained by the response spectrum technique. Because of th e high natural frequencies and resu lting low levels of responses for the vertical direction, the more conservative spectrum response analysis results were used instead of time history results. The response spectrum utilized was derived from the vertical acceleration time history at th e reactor vessel flange. The sp ectrum curv e is presented in Figure 3.A-9.
An acceleration level corresponding to the natural frequency of each mode was selected from the spectrum curve. The response spect rum technique uses these acceler ation values to determine the inertia forces, accelerations, and displacements of each mode. The results for each mode were conservatively combined on the basis of absolute values. For the vertical models, the first seven modes were included in the results.


3.A.1.3 Results Combined results for the horizontal and vertical dynamic seismic analyses are presented in Table 3.A-2 in terms of stresses at critical locations in the reactor internals for the DBE. Table 3.A-2 also lists the seis mic stresses which result from application of the design loads specified for the DBE. A comparison shows the resu lts of the dynamic analysis to be less severe.
alignment plate, core shroud and core support plate were neglected.
3.A.1.4 ConclusionIt is concluded that the seismic loads specified for the design of the inte rnals are adequate. All seismic loads calculated by the dynamic seismic analysis are less than th e design loads specified by the DBE.
ce the basic phenomenon in buckling is nonlinear instability, the initial deviation of the cture from a perfect geometry greatly affects its response. The initial imperfection was applied he core support barrel by means of a pseudo-load so developed to provide the maximum erfection over each of the desired number of circumferential harmonics. The actual transient ing in terms of its harmonics was applied to the initially imperfect geometry core support el and the response obtained for each of the imperfection harmonics for the combined loading monics.
3.A.2 NORMAL OPERATING ANALYSIS Design analyses were performe d on the reactor internals for normal operating conditions to demonstrate that the mechan ical design bases were satisfied. These design calculations included appropriate vibration anal yses of the component assemblies.
.3.2 Analysis Codes HSD (Reference 3.A-7) is a structural finite-element computer code developed at the versity of California, Berkeley, and supported in part by the National Science Foundation. It orms dynamic analyses of complex axisymmetric structures subjected to arbitrary dynamic ings or base accelerations. The frequencies of free vibrations as calculated by ASHSD pare well to those calculated by the equations of Hermann-Mirshy and Flugge, erences 3.A-11 and 3.A-12, respectively. The authors also make comparisons with available erimental results (Reference 3.A-13) of free vibrations of cylindrical shells. The resulting parison is good. Comparison of the numerical solution (Reference 3.A-14) of the dynamic onse of a shell to suddenly applied loads and the finite-element (ASHSD) solution of the e problem are in good agreement. The response of a shell to a moving axisymmetric pressure was evaluated by ASHSD and analytically (Reference 3.A-15) with the results being in good ement.
The flow induced vibration of the CSB/TS, during normal operation, was characterized as a forced response to deterministic and random pressure fluctuations in the coolant. Me thods were developed fo r predicting the response of components to the hydraulic forcing functions.
MMSOR-DYNASOR (Reference 3.A-9) is a finite-element computer code developed at Texas M University and supported in part by a NASA grant from the Manned Spacecraft Center, ston, Texas. This code has the capability of determining the nonlinear dynamic response of ymmetric shells subjected to arbitrary dynamic loads. Asymmetrical dynamic buckling can be stigated using this program. The program has been extensively tested, using problems the tions to which have been reported by other researchers, in order to establish the validity of the es. Among these are a shallow shell with axisymmetric loading as described in Reference 3.A-Identical results are obtained with those of Reference 3.A-17 for the analytical evaluation of t loadings on a cylindrical shell. Calculations made by SAMMSOR-DYNASOR for the metric buckling of a shallow spherical cap is in good agreement with the analyses of erences 3.A-18 and 3.A-19 and the experimental data of References 3.A-20 and 3.A-21.
BOR DRASTIC, (Reference 3.A-22) is a structural finite-element computer code eloped at the Aeroelastic and Structures Research Laboratory, Department of Aeronautics at Massachusetts Institute of Technology. The work was administered by the Air Force Systems mmand with technical monitoring by the Aerospace Corp. SABOR 5 - DRASTIC is the end lt of combining a finite-difference solution procedure and a finite-element program to permit dicting the transient response of complex shells of revolution which are subjected to arbitrary sient loadings. Comparisons with reliable independent analytical predictions (notably finite-3.A-10                                    Rev. 35


Emphasis was placed on analysis and design of those components which were particularly critical and susceptible to vibratory ex citation, such as the thermal sh ield. Using a top supported, as opposed to a bottom supported, thermal shield design improves stability as it eliminates a free edge in the flow path. Increas ing the number of upper supports and lower jackscrews, in the specific manner chosen, provides a much stiffer structure and the use of an all-welded shield eliminates local flexibilities and relative motio n at bolted joints. Analytical studies show the thermal shield to be stable on its support system when exposed to the axial annular flow encountered during normal operation. The snubber design is based upon limiting the motion of the core support barrel under condi tions of hydraulically induced vi brations. The snubbers are at MPS2 UFSAR3.A-6Rev. 35 the position of maximum amplitude for the funda mental lateral bending mode of the barrel, thereby restricting motion of the barrel at the most efficien t position. The ci rcumferential distribution of snubbers assures restraint regardless of the di rection of response.
lysis were performed by the Aerospace Corp. (Reference 3.A-23) to verify the ability of the e to account for a complex geometry shell of revolution subjected to transient asymmetric
The random hydraulic forcing f unction was developed by analyti cal and experimental methods.
: s. Loads were applied by means of well-defined explosive charges. Based upon the results of amic strain measurements made on the test structure, it is evident that the SABOR 5 -
An analytical expression was de veloped to define the turbulent pressure fluctuation for fully developed flow. This expression was modified, based upon the result of scale model testing, to account for the fact that flow in the downcomer was not fully de veloped. Based upon test results, an expression was developed to define the spatial dependenc y of the turbulent pressure fluctuations. In addition, experi mentally adjusted analytical expressions were developed to define; the peak value of the pressure spectral density associated with the turbulence and; the maximum area of coherence, in terms of the boundary layer displacement, across which the random pressure fluctu ations are in phase.
ASTIC code is capable of solving complex dynamic shell structure problems successfully.
The natural frequencies and mode shapes of the CSB/TS sy stem were obtained using the axisymmetric shell finite elem ent computer program, ASHSD (Reference 3.A-7). This computer program is capable of obtaining natural frequencies and mode sh apes of complex axisymmetric shells; e.g., arbitrary meri dional shape, varying thickness, bran ches, multi-materials, orthotropic material properties, etc. To em ploy the ASHSD code, the CSB/TS were modeled as a series of conical shell frustrums joined at their nodal point circles. The length of each element, throughout the ASHSD model, was a fraction of the shell de cay length. Since rapid changes in the stress pattern occur in regions of structural discontinuity, the nodal poi nt circles were more closely spaced in such regions. The finite element model of the CSB/TS system included representation of the core support barrel upper and lower flanges, sections of different wall thickness, and thermal shield support lugs and j ackscrews. Elements with orthot ropic material properties were utilized to provide equiva lent axisymmetric models of the structural stif fness and constraints to relative motion between the core support barrel and thermal shield provided by the thermal shield support lugs and jackscrews. Those modes which re flect the mass of the lower support structure, core shroud and fuel were simulated by the addition of concentrated masses at specific nodes in the core support barrel fla nge finite element model.Applying Hamilton's Variational Principle to the conical shell el ements an equation of motion was formulated for each degree of freedom of the system. An inverse iterati on technique was utilized in the program to obtain solutions to the charact eristic equation, wh ich was based on a diagonalized form of a consistent mass matrix and stiffness matr ix developed using the finite element method. Four degrees of freedom - radial di splacement, circumfere ntial displacement, vertical displacement, and meridi onal rotation - were taken into account in the analysis, giving rise to coupled mode shapes and corresponding frequencies. Eval uation of the reduction of these frequencies for the system immersed in coolant was made by means of the "virtual mass" method outlined in Reference 3.A-2.
eveloping the above finite-element computer codes, (i.e., ASHSD, SAMMSOR-DYNASOR, BOR 5 - DRASTIC) the authors have independently verified their codes with respect to the lts of other established structural programs, classical solutions and as possible to experimental
The random response analysis cons iders the response of the CSB/
. The correlations demonstrate that the above programs are capable of solving complex amic shell structure problems successfully and that the finite-element method of modeling vides accurate representation of the structural phenomena. The SABOR 5 - DRASTIC code, ch has had extensive and successful analytical and experimental correlation (Reference 3.A-6) transient (explosive) asymmetric loading, was used to analyze a core support barrel structure h short-term loading. The results of this well-verified program are identical to these of the te-element codes ASHSD and SAMMSOR-DYNASOR (which are used in the LOCA lysis) for the same core support barrel problem, demonstrating the ability of these programs to quately represent and evaluate the effect of a transient load on an axisymmetric structure like core support barrel.
TS system to the turbulent downcomer flow during steady-st ate operation. The random forcing function is assumed to be a wide-band stationary random pro cess with a pressure spectral density equal to the peak value associated with the turbulence.
.4 EFFECTS OF THERMAL SHIELD REMOVAL owing the discovery of the thermal shield support degradation at the end of Cycle 5 in July, 3, the thermal shield was removed. A detailed inspection of the core barrel revealed damage at thermal shield support lug locations. Repairs to the core barrel comprised of drilling crack stor holes at the ends of through-wall cracks and removal by machining of non through-wall ks.
The rms vibration level of the CS B/TS system was obtained based upon a damped, single degree of freedom analysis assuming the rms random pressure fluctuation MPS2 UFSAR3.A-7Rev. 35 to be spatially invariant. The analysis demonstrates that the anticipated rms response of the CSB&#xda;TS system is low. Snubber load s were derived using an anal ytical technique originally developed by a Combustion Engin eering (CE) consultant usi ng the random loads discussed above. Modeling the reactor vessel snubbers and core support barrel system as a single degree of freedom spring-mass system, the number and magnitude of snubber, core support ba rrel impacts was calculated based upon the res ponse of the system to random excitation. The sn ubbers were designed, based upon this loading re quirement, to meet the cyclic strength requirements specified in Section III of the ASME Boiler and Pressure Vessel Code.
lytical evaluations and assessments were performed to demonstrate continued structural quacy of the reactor internals without the thermal shield for all design loading conditions.
The forced response of the reactor internals to deterministic loading was evaluated by classical analytical methods, using lumped mass and continuous elastic struct ural models. These calculated responses were used to verify the structural integrity of the reactor ve ssel internals to normal operating vibratory excitation. Components were de sign analyzed to assure that there were no adverse effects from dominant ex citation frequencies, such as pump rotational and blade passing frequencies.
cial attention was paid to the core barrel to justify the repairs. A description of the repairs to core barrel, analyses, and significant results is given in Reference 3.A-24.
3.A.3 LOSS OF COOLANT ACCIDENT ANALYSIS 3.A.3.1 DiscussionA dynamic analysis (Reference 3.A-8) has been perf ormed to determine the structural response of the reactor vessel internals to th e transient loss of c oolant accident (LOCA) loading. The analysis determined the shell, beam and rigid body mo tions of the internal s using established computerized structural response analyses. The finite-element computer code, ASHSD (Reference 3.A-7) was used to ca lculate the time-dependent beam and shell response of the CSB/TS system to the transient LOCA loading. The finite-element computer code SAMMSOR-DYNASOR (Reference 3.A-9) was used to eval uate the core support barrel's potential for buckling when loaded by a net external radial pressure resulting from an outlet line break. The structural response of the reactor internals to vertical and transver se loads resulting from inlet and outlet breaks, was determined using the spring-mass computer code, SHOCK (Reference 3.A-10).
conclusion, there was no significant change in the loads and the stresses in the internal ctures remained within the ASME Code allowables.
The time and space depende nt pressure loads used in the above analysis were the result of a detailed hydraulic blowdown analysis. The pressure fluctuations we re determined for each node in the hydraulic model for inlet a nd outlet line breaks. The pressure time histories at these nodal locations were then decomposed into the Four ier harmonics which define the circumferential pressure distribution at the nodal elevations. Where the hydrau lic model nodes did not correspond to those of the structural mode l, the hydraulic model pressure co mponents were interpolated to provide the required loading information.
.5 LEAK-BEFORE-BREAK ANALYSIS k-Before-Break (LBB) analyses for the reactor coolant system (RCS) main coolant loops, for pressurizer surge line, and unisolable RCS portions of the safety injection and shutdown ling piping, which demonstrated that the probability of fluid system piping rupture was emely low, was reviewed and approved by the commission. (See References 3.A-25 through
The finite element computer c ode, ASHSD, was used to calcul ate the dynamic response of the CSB/TS to transient LOCA loading resulting from an inlet break. To employ the ASHSD code, the CSB/TS were modeled as a series of conical shell frustrums (elements) joined at their nodal point circles. Applying Hamilton's Variational Principle to the conical shell elements a damped equation of motion was formulated for each degree of freedom of the system. F our degrees of freedom - radial displacement , circumferential displacement, vertical displacement and meridional rotation - were taken into account in the analysis, giving rise to coupled modes. The MPS2 UFSAR3.A-8Rev. 35differential equations of motions were solved numerically using a step integration procedure. To ensure computational stability of the numerical solution, the inte gration time step was chosen such that it is small compared to the shortest period of the finite el ement system. The model developed for the CSB/TS system is shown in Figure 3.A-10. Th e length of each element, throughout the analytical model, was a fraction of the shell decay length.
-29.) Subsequent to the commission review and approval, weld overlays were applied to imilar metal welds (DMWs) at the shutdown cooling, the safety injection and the pressurizer 3.A-11                                    Rev. 35
Since rapid changes in the stress pattern occur in regions of structural discontinuity, the nodal po int circles were more closely spaced in such regions. The finite el ement model of the CSB/TS system included representation of the core support barrel upper and lower flanges, sections of different wall thickness, and thermal shield support lugs and j ackscrews. Elements with orthotropic material properties were utilized to provide equivalent axis ymmetric models of the structural stiffness and constraints to relative motion be tween the core support barrel a nd thermal shield provided by the thermal shield support lugs and jackscrews. Those modes which re flect the mass of the lower support structure, core shroud a nd fuel were stimulated by the addition of conc entrated masses at specific nodes in the core support barrel flange finite element model.
In performing the dynamic analysis of the CSB/TS system, the transient load harmonics were applied in two successive phas es to account for time-depe ndent boundary conditions at the snubbers. The first phase used th ose harmonics which excite the beam modes, whereas the second phase used those harmonics which excite the shell modes. During th e first phase, the lower end of the core support barrel was unrestrained. Within a very few milliseconds, the clearances between the core support barrel and reac tor vessel snubbers were closed and for the remainder of the LOCA transient, the core support barrel was restrained radially at the snubber level. Transient responses were computed throughout each loading phase.


The ASHSD code computed the nodal point displ acement, resultant shell forces, shell stresses and maximum principle stresses as functions of time. The maximum principle stresses at the internal and external surfaces of the CSB/TS were determined from the bending and membrane components during each phase of transient loading. Stress intens ity levels calcul ated from the principle stresses were comb ined with normal operating and seismic induced stresses for comparison with design criteria.
ve piping segments, including the effects of pipe whipping and discharging fluids have been luded from the design basis of the following reactor vessel and reactor internals components:
Accurate representation and analysis of the CSB/
Core barrel snubbers, core barrel stabilizer blocks Reactor vessel core support ledge Reactor Cavity Seal Plate, Neutron Shielding
TS shell structures wa s obtained through use of the finite element code ASHSD. Accurate representation of the remainder of the internals (i.e., fuel, core shroud, CEAs, UGS, lower support st ructure, etc.) was obtained using the SHOCK code.The SHOCK code determines the response of structures which ar e represented as lumped-mass systems and subjected to arbitrary loading functions. The code solves the differential equations of motion for each mass by a numerical step-integration procedure.
.6 REFERENCES
The lumped mass model can represent a vertically or laterally responding system subject to arbitrary loading functions and initial conditions. Options are available for de scribing steady state loads, preloads, input accelerations, linear a nd nonlinear springs (incl uding tension and compressi on only springs) gaps, and structural and viscous damping.
-1    Fritz, R. J., and Kiss, E., The Vibration Response of a Cantilevered Cylinder Surrounded by an Annular Fluid, KAPL-M-6539, February 1966.
-2    Kiss, E., Analysis of the Fundamental Vibration Frequency of a Radial Vane Internal Steam Generator Structure, ANL-7685, Proceedings of Conference on Flow-Induced Vibrations in Reactor System Components, May 1970, Argonne National Laboratory, Argonne, IL.
-3    ICES STRUDL-II, The Structural Design Language Engineering Users Manual.
-4    MRI/STARDYNE - Static and Dynamic Structural Analysis System: User Information Manual, Control Data Corporation, June 1, 1970.
-5    MRI/STARDYNE User Manual, Computer Methods Department, Mechanics Research, Inc., Los Angeles, California, January 1, 1970.
-6    Hurty, W. C., and Rubinstein, M. F., Dynamics of Structures, Chapter 8, Prentice Hall, Inc., Englewood Cliffs, New Jersey, 1964.
-7    Ghosh, S., Wilson, E., Dynamic Stress Analysis of Axisymmetric Structures Under Arbitrary Loading, Dept. No. EERC 69-10, University of California, Berkeley, September 1969.
-8    CENPD-42, Topical Report on Dynamic Analysis of Reactor Vessel Internals Under Loss of Coolant Accident Conditions with Application of Analysis to C-E 800 Mw(e)
Class Reactors, August 1972.
-9    Tillerson, J. R., Haisler, W. E., SAMMSOR II - A Finite Element Program to Determine Stiffness and Mass Matrices of Shells-of- Revolution, Texas A&M University, TEES-RPT-70-18, October 1970. DYNASOR II - A Finite Element Program for the Dynamic Nonlinear Analysis of Shells-of-Revolution, Texas A&M University, TEES-RPT-70-19, October 1970.
-10 Gabrielson, V. K., SHOCK - A Computer Code for Solving Lumped-Mass Dynamic Systems, SCL-DR-65-34, January 1966.
3.A-12                                  Rev. 35


The reactor internals were developed in terms of a spring-mass system for both vertical and lateral directions; see Figures 3.A-1 1 and 3.A-12. For both models, the spring rates were generally MPS2 UFSAR3.A-9Rev. 35 evaluated using strength of material techniques. However, in complex areas such as at the core support barrel flanges and UGS support flange, the stiffness was derived from finite element model analyses. The lumped mass weights were generally based upon the mass distribution of the uniform support structures, but included at appr opriate nodes, local ma sses such as snubber blocks, fuel end fittings , thermal shield lugs, etc. The net re sult was a lumped-mass system having the same distribution of mass as the actual structure. To simulate the effect associated with the internals oscillating laterally in the water filled vessel, a distributed virtual mass was calculated based upon the procedure outlined in Reference 3.A-8 (which includes the annulus effect) and was added to the structural lumped-mass system, to provide an analytical model with a dynamic response quantitatively similar to the actual internals. In the case of the vertical model, the hydraulic effect is notably one of reducing the effe ctive weight of the reactor internals and this effect was included in the structural lumped-mass system.
78, P. 563-568, 1956.
The SHOCK code provided excellent facility for modeling cleara nces, preloads and component interfaces. In the lateral model, the core supp ort barrel, reactor vess el snubber clearance was simulated by a nonlinear spring wh ich accounted for the increased resistance to core support barrel motion when snubbing occurred. In the vertical model, nonlinear springs in the form of compression only springs, were used extensively to simulate preload and interface conditions, such as exist between the UGS support plate and co re support barrel upper flange; at the fuel hold-down spring; at the fuel, core support plate interface and at the co re shroud, core support plate interface. Tension only springs were used to simulate the effect of the core shroud tie rods.
-12 Flugge, W., Stresses in Shells, Third Printing, Springer-Verlag, New York, 1966.
In both the vertical and latera l SHOCK models, damping was vari ed throughout the system to simulate structural and hydraulic frictional effects within the reactor internals. The effect of hydraulic drag in the vertical mo del was simulated by a force time-history applied to the fuel lower end-fitting. Vertical loads were used directly from the de tailed hydraulic analysis, whereas lateral loads were obtained by inte grating those harmonics which excite the beam modes to obtain the net lateral load on the CSB/TS system.The SHOCK code calculated the vertical and lateral response of the system in terms of displacements, velocities and accelerations and internal force, moments and shear s as related to each model. These quantities were sufficient to permit calculation of membrane and where appropriate bending stresses for comparison with design criteria.The finite-element code SAMMSOR-DYNASOR was used to dete rmine the dynamic response of the core support barrel, with initially imperfect geometry, to a net external radial pressure resulting from an outlet line break. The above an alysis has the capability of determining the nonlinear dynamic response of axisymmetric shells with initial imperf ections subjected to arbitrarily varying load configurations.Since SAMMSOR-DYNASOR is a finite-element program, a model was developed, Figure 3.A-13, of the core support barrel usi ng axisymmetric finite-elements similar to those used for the ASHSD analysis. As was for the ASHSD model, the SAMMSOR-DYNASOR finite-element lengths were considerably less than the decay length of the core support barrel. The boundary condition at the core support barrel flange was considered fixed, whereas at the core support barrel lower flange radial displacements were restrained. These boundary conditions represented MPS2 UFSAR3.A-10Rev. 35 the restraint due to the expansi on compensating ring and pressure vessel head at the top and the snubbers and lower support structure at the bottom. For conservatism, the stiffening effects of the fuel alignment plate, core shroud an d core support plate were neglected.
-13 Koval, L. R., Cranch, E. I., On the Free Vibrations of Thin Cylindrical Shells Subjected to Initial Torque, Proceedings of the U. S. National Congress of Applied Mechanics, P.
Since the basic phenomenon in buckling is nonlinear instability, the init ial deviation of the structure from a perfect geometry greatly affects its response. The initial imperfection was applied to the core support barrel by means of a pse udo-load so developed to provide the maximum imperfection over each of the de sired number of circumferentia l harmonics. The actual transient loading in terms of its harmonics was applied to the initially "i mperfect" geometry core support barrel and the response obtained fo r each of the imperfection harm onics for the combined loading harmonics.
11, 1962.
3.A.3.2 Analysis CodesASHSD (Reference 3.A-7) is a structural finite-element computer code developed at the University of California, Berkeley, and supporte d in part by the Nationa l Science Foundation. It performs dynamic analyses of co mplex axisymmetric structures subjected to arbitrary dynamic loadings or base accelerations. The frequencies of free vibrations as calculated by ASHSD compare well to those calculated by th e equations of Herma nn-Mirshy and Flugge, References 3.A-11 and 3.A-12, respectively. The authors also make comparisons with available experimental results (Reference 3.A
-14 Reismann, H., and Padloy, J., Forced, Axisymmetric Motions of Cylindrical Shells, Journal of the Franklin Institute, Vol. 284, Number 5, November 1967.
-13) of free vibrat ions of cylindrical shells. The resulting comparison is good. Comparison of the numerical solution (Reference 3.A-14) of the dynamic response of a shell to suddenly applied loads a nd the finite-element (A SHSD) solution of the same problem are in good agreement. The response of a shell to a moving axisymmetric pressure load was evaluated by ASHSD and analytically (Reference 3.A-15) with the results being in good agreement.SAMMSOR-DYNASOR (Reference 3.A-9) is a finite-element computer code developed at Texas A&M University and supported in part by a NA SA grant from the Manned Spacecraft Center, Houston, Texas. This code has the capability of determining th e nonlinear dynami c response of axisymmetric shells subjected to arbitrary dynamic loads. Asymme trical dynamic buckling can be investigated using this program. The program has been extensiv ely tested, using problems the solutions to which have been reported by other researchers, in order to estab lish the validity of the codes. Among these are a shallow shell with axisym metric loading as described in Reference 3.A-
-15 Tang, Sing-Chih, Response of a Finite Tube to Moving Pressure, Journal Engineering Mechanics Division, ASCE, Vol. 93, Number EM3, June 1967.
: 16. Identical results are obtained with those of Reference 3.A-17 for the analytical evaluation of blast loadings on a cylindrical shell. Calc ulations made by SAMMSOR-DYNASOR for the symmetric buckling of a shallow spherical cap is in good agreement with the analyses of References 3.A-18 and 3.A-19 a nd the experimental data of References 3.A-20 and 3.A-21.
-16 Klein, S., and Sylvester, R. J., The Linear Elastic Dynamic Analysis of Shells of Revolution by the Matrix Displacement Method, Air Force Slight Dynamics Laboratory, TR-66-80, 1966, P. 299-329.
SABOR DRASTIC, (Reference 3.A-22) is a structural fini te-element computer code developed at the Aeroelastic and Structures Research Laboratory, Department of Aeronautics at the Massachusetts Institute of Technology. The work was administ ered by the Air Force Systems Command with technical monitoring by the Aerospace Corp. SABOR 5 - DRASTIC is the end result of combining a finite-dif ference solution procedure and a fi nite-element program to permit predicting the transient response of complex shells of revolution wh ich are subjected to arbitrary transient loadings. Comparisons w ith reliable independent analytic al predictions (notably finite-MPS2 UFSAR3.A-11Rev. 35difference transient response solutions submitted by AVCO) co nfirm the accuracy and reliability of the SABOR 5 -DRASTIC dynamic response pr edictions. An experiment and accompanying analysis were performed by the Aerospace Corp. (Reference 3.A-23) to verify the ability of the code to account for a complex ge ometry shell of revolution subj ected to transient asymmetric loads. Loads were applied by means of well-defined explosive charges.
-17 Johnson, D. E., Grief, R., Dynamic Response of a Cylindrical Shell: Two Numerical Methods, AIAA Journal, Vol. 4, Number 3, March 1966, P. 486-494.
Based upon the results of dynamic strain measurements made on the test st ructure, it is evident that the SABOR 5 -
-18 Huang, N. C., Axisymmetric Dynamic Snap-through of Elastic Clamped Shallow Spherical Shells, AIAA Journal, Vol. 7, Number 2, February 1969, P. 215-220.
DRASTIC code is capable of solving complex dynamic shell structure problems successfully.
-19 Stephen, W. B., and Fulton, R. E., Axisymmetric Static and Dynamic Buckling of Spherical Caps due to Centrally Distributed Pressures, Paper 69-89, AIAA Journal, 1969.
In developing the above finite-element com puter codes, (i.e., ASHSD, SAMMSOR-DYNASOR, SABOR 5 - DRASTIC) the authors have independently verified their codes with respect to the results of other established struct ural programs, classical solutions and as possible to experimental data. The correlations demonstr ate that the above programs ar e capable of solving complex dynamic shell structure problems successfully and that the fini te-element method of modeling provides accurate representation of the st ructural phenomena. The SABOR 5 - DRASTIC code, which has had extensive and successful analytical and experimental correlation (Reference 3.A-6) for transient (explosive) asymmetr ic loading, was used to analyze a core support barrel structure with short-term loading. The resu lts of this well-verified program are identical to these of the finite-element codes ASHSD and SAMMSOR-DYNASOR (which are used in the LOCA analysis) for the same core suppor t barrel problem, demonstrating th e ability of these programs to adequately represent and evaluate the effect of a transient load on an axisymmetric structure like the core support barrel.
-20 Lock, M. H., Okrebo, S., and Whittier, J. S., Experiment of the Snapping of a Shallow Dome Under a Step Pressure Loading, AIAA Journal, Vol. 6, No. 7, July 1968, P. 1320-1326.
3.A.4 EFFECTS OF THERMAL SHIELD REMOVAL Following the discovery of the thermal shield s upport degradation at the end of Cycle 5 in July, 1983, the thermal shield was removed. A detailed insp ection of the core barrel revealed damage at two thermal shield support lug locations. Repairs to the core barrel comp rised of drilling crack arrestor holes at the ends of through-wall cracks and rem oval by machining of non through-wall cracks.Analytical evaluations and asse ssments were performed to dem onstrate continued structural adequacy of the reactor intern als without the thermal shield for all design loading conditions.
-21 Stricklin, J. A., and Martinez, J. E., Dynamic Buckling of Clamped Spherical Caps Under Step Pressure Loadings, AIAA Journal, Vol. 7, Number 6, June 1969, P. 1212-1213.
Special attention was paid to the co re barrel to justify the repairs.
-22 Kotanchik, J. J., et al., The Transient Linear Elastic Response Analysis of Complex Thin Shells of Revolution Subjected to Arbitrary External Loadings, by the Finite-Element Program SABOR 5 - DRASTIC, AD-709-189, Massachusetts Institute of Technology, April 1970.
A description of the repairs to the core barrel, analyses, and significant results is given in Reference 3.A-24.
-23 Klein, S., A Static and Dynamic Finite Element Shell Analysis with Experimental Verification, International Journal for Numerical Methods in Engineering, Vol. 3, P.
In conclusion, there was no significant change in the loads and the stresses in the internal structures remained within the ASME Code allowables.
3.A.5 LEAK-BEFORE-BREAK ANALYSIS Leak-Before-Break (LBB) analyses for the reactor coolant system (RCS) main coolant loops, for the pressurizer surge line, and unisolable RCS portions of the safety injection and shutdown cooling piping, which demonstrated that the pr obability of fluid syst em piping rupture was extremely low, was reviewed and approved by the commission. (S ee References 3.A-25 through 3.A-29.) Subsequent to the com mission review and approval, we ld overlays were applied to dissimilar metal we lds (DMWs) at the shutdown cooling, the safety inje ction and the pressurizer MPS2 UFSAR3.A-12Rev. 35surge nozzles. A revised LBB analysis was performed for these nozzles (see Reference 3.A-30). Accordingly, pursuant to revised GDC 4, the dynamic effects associated with pipe ruptures in the above piping segments, including the effects of pipe whipping and discharging fluids have been excluded from the design basis of the following reactor vessel a nd reactor internals components:
Core barrel snubbers, core barrel stabilizer blocks Reactor vessel core support ledge Reactor Cavity Seal Plate, Neutron Shielding 3.A.6 REFERENCES3.A-1Fritz, R. J., and Kiss, E., "The Vibration Response of a Cantilevered Cylinder Surrounded by an Annular Fluid," KAPL-M-6539, February 1966.3.A-2Kiss, E., "Analysis of the Fundamental Vibration Frequency of a Radial Va ne Internal Steam Generator Structure," ANL-7685, Proc eedings of Conference on Flow-Induced Vibrations in Reactor System Components, May 1970, Argonne National Laboratory, Argonne, IL.3.A-3ICES STRUDL-II, The Structural De sign Language Engineering Users' Manual.3.A-4"MRI/STARDYNE - Static and Dynamic St ructural Analysis System: User Information Manual," Control Data Corporation, June 1, 1970.3.A-5MRI/STARDYNE User Manual, Computer Methods Department, Mechanics Research, Inc., Los Angeles, California, January 1, 1970.
3.A-6Hurty, W. C., and Rubinstein, M. F., "Dynamics of Structures," Chapter 8, Prentice Hall, Inc., Englewood Cliffs, New Jersey , 1964.3.A-7Ghosh, S., Wilson, E., "Dynamic Stress Analysis of Axisymmetric St ructures Under Arbitrary Loading," Dept. No. EERC 69-10, University of California, Berkeley, September 1969.3.A-8CENPD-42, "Topical Report on Dynamic Analysis of Reactor Ve ssel Internals Under Loss of Coolant Accident Conditions with Application of Analys is to C-E 800 Mw(e) Class Reactors," August 1972.3.A-9Tillerson, J. R., Haisler, W. E., "SA MMSOR II - A Finite Element Program to Determine Stiffness and Mass Matrices of Shells-of- Revolution," Texas A&M University, TEES-RPT-70-18, October 1970. "DYNASOR II - A Finite Element Program for the Dynamic Nonlinear Analysis of Shells-of-Revolution," Texas A&M University, TEES-RPT-70-19, October 1970.3.A-10Gabrielson, V. K., "SHOCK - A Computer Code for Solving Lumped-Mass Dynamic Systems," SCL-DR-65-34, January 1966.
MPS2 UFSAR3.A-13Rev. 353.A-11Hermann, G., Mirshy, I., "Three Dimensi onal Shell Theory Analysis of Axially Symmetric Motions of Cylinders
," Journal of Applied Mechanics, Trans. ASME, Vol. 78, P. 563-568, 1956.3.A-12Flugge, W., "Stresses in Shells," Third Printing, Springer-Verlag, New York, 1966.3.A-13Koval, L. R., Cranch, E. I., "On the Free Vibrations of Thin Cylindrical Shells Subjected to Initial Torque," Proceedings of the U. S. National Congress of Applied Mechanics, P. 11, 1962.3.A-14Reismann, H., and Padloy, J., "Forced, Axis ymmetric Motions of Cylindrical Shells,"
Journal of the Franklin Institute, Vol. 284, Number 5, November 1967.
3.A-15Tang, Sing-Chih, "Response of a Finite Tube to Moving Pr essure," Journal Engineering Mechanics Division, ASCE, V ol. 93, Number EM3, June 1967.3.A-16Klein, S., and Sylvester, R. J., "The Li near Elastic Dynamic Analysis of Shells of Revolution by the Matrix Displacement Method," Air Force Slight Dynamics Laboratory, TR-66-80, 1966, P. 299-329.3.A-17Johnson, D. E., Grief, R., "Dynamic Response of a Cylindrical Shell: Tw o Numerical Methods," AIAA Journal, Vol. 4, Number 3, March 1966, P. 486-494.3.A-18Huang, N. C., "Axisymmetric Dynamic Snap-through of Elastic Clamped Shallow Spherical Shells," AIAA Journal, Vol. 7, Number 2, February 1969, P. 215-220.3.A-19Stephen, W. B., and Fulton, R. E., "Axisymmetric St at ic and Dynamic Buckling of Spherical Caps due to Centrally Distributed Pressures," Paper 69-89, AIAA Journal, 1969.3.A-20Lock, M. H., Okrebo, S., and Whittier, J.
S., "Experiment of the Snapping of a Shallow Dome Under a St ep Pressure Loading," AIAA Journal, Vol. 6, No. 7, July 1968, P. 1320-1326.3.A-21Stricklin, J. A., and Martinez, J. E., "Dynamic Buckling of Clamped Spherical Caps Under St ep Pressure Loadings," AIAA Journal, Vol. 7, Number 6, June 1969, P. 1212-1213.3.A-22Kotanchik, J. J., et al., "The Transient Linear Elastic Response Analysis of Complex Thin Shells of Revolution Subjected to Ar bitrary External Loadings, by the Finite-Element Program SABOR 5 - DRASTIC," AD-709-189, Massa chusetts Institute of Technology, April 1970.3.A-23Klein, S., "A Static and Dynamic Finite Element Shell Analysis with Experimental Ve rification," International Journal for Numerical Methods in Engineering, Vol. 3, P.
299-315, 1971.
299-315, 1971.
MPS2 UFSAR3.A-14Rev. 353.A-24"Thermal Shield Damage Recovery Program - Final Report," Northeast Nuclear Energy Company, Millstone Nuclear Power Station, Unit Number 2, Docket No. 50-336, License No. DPR-65, December, 1983.3.A-25NRC Letter from D. G. McDonald, Jr. to M. L. Bowling, Jr., "Rev ised Evaluation of the Primary Cold Leg Piping Leak - Before-Break Analysis for the Millstone Nuclear Power Station, Unit Number 2," dated November 9, 1998.3.A-26NRC Letter from D. G. McDonald, Jr. to M. L. Bowling, Jr., "Application of Leak -
3.A-13                                  Rev. 35
Before-Break Status to the Portions of the Safety Injection and Shutdown Cooling System for the Millstone Nuclear Power Sta tion, Unit Number 2," dated November 9, 1998.3.A-27NRC Letter from B. Eaton to R. P. Necci, "Staff Review of the Submittal by Northeast Nuclear Energy Company to Apply Leak-Before-Break Status to the Pressurizer Surge Line, Millstone Nuclear Power Station, Unit 2," dated May 4, 1999.3.A-28NRC Letter from G.S. Vissing to J.F. Opeka, "Application of Reactor Coolant System Leak-Before-Break Analysis," dated September 1, 1992.3.A-29Federal Register/Vol ume 53, No. 66/April 6, 1988, "10 CFR Part 50 Leak Before Break Te chnology Solicitation of Public Co mment on Additiona l Applications."3.A-30Structural Integrity Associates Report: 0901238.401, Revision 0, dated: December 2010, Updated Leak-Before-Break Evaluation of Weld Overlaid Hot Leg Surge, Shutdown Cooling and Safety Injection Nozzles for Millstone Nuclear Power Station, Unit 2.
 
MPS2 UFSAR3.A-15Rev. 35TABLE 3.A-1  NATURAL FREQUENCIES FOR VERTICAL SEISMIC ANALYSIS MATHEMATICAL MODELMode NumberSub-Model I Frequency, cpsSub-Model II Frequency, cps121.6072.98267.75404.093124.59-MPS2 UFSAR3.A-16Rev. 35TABLE 3.A-2  SEISMIC STRESSES IN CRITICAL REACTOR INTERNALS COMPONENTS FOR THE DESIGN BASIS EARTHQUAKE Structural ComponentLocationStress Mode Design Load StressDynamic Analysis Stress Core Support BarrelUpper Section of BarrelTension & Bending1,129 psi746 psiLower Core SupportBeam FlangeBending5,278 psi929 psi CEA Shrouds:  
License No. DPR-65, December, 1983.
-25 NRC Letter from D. G. McDonald, Jr. to M. L. Bowling, Jr., Revised Evaluation of the Primary Cold Leg Piping Leak - Before-Break Analysis for the Millstone Nuclear Power Station, Unit Number 2, dated November 9, 1998.
-26 NRC Letter from D. G. McDonald, Jr. to M. L. Bowling, Jr., Application of Leak -
Before-Break Status to the Portions of the Safety Injection and Shutdown Cooling System for the Millstone Nuclear Power Station, Unit Number 2, dated November 9, 1998.
-27 NRC Letter from B. Eaton to R. P. Necci, Staff Review of the Submittal by Northeast Nuclear Energy Company to Apply Leak-Before-Break Status to the Pressurizer Surge Line, Millstone Nuclear Power Station, Unit 2, dated May 4, 1999.
-28 NRC Letter from G.S. Vissing to J.F. Opeka, Application of Reactor Coolant System Leak-Before-Break Analysis, dated September 1, 1992.
-29 Federal Register/Volume 53, No. 66/April 6, 1988, 10 CFR Part 50 Leak Before Break Technology Solicitation of Public Comment on Additional Applications.
-30 Structural Integrity Associates Report: 0901238.401, Revision 0, dated: December 2010, Updated Leak-Before-Break Evaluation of Weld Overlaid Hot Leg Surge, Shutdown Cooling and Safety Injection Nozzles for Millstone Nuclear Power Station, Unit 2.
3.A-14                                Rev. 35
 
MATHEMATICAL MODEL ode Number Sub-Model I Frequency, cps  Sub-Model II Frequency, cps 1          21.60                      72.98 2          67.75                      404.09 3          124.59                       -
3.A-15                            Rev. 35
 
COMPONENTS FOR THE DESIGN BASIS EARTHQUAKE Dynamic Structural                                             Design Load Analysis Component              Location          Stress Mode      Stress    Stress re Support Barrel Upper Section of     Tension & Bending 1,129 psi    746 psi Barrel wer Core         Beam Flange          Bending          5,278 psi   929 psi pport A Shrouds:        End of Shroud        Tension & Bending 3,548 psi    1,295 psi gle A Shrouds: Dual End of Shroud          Tension & Bending 2,762 psi    697 psi per Grid Beams    Center of Beam      Bending          1,652 psi    127 psi per Guide        Junction of Flange & Tension & Bending 2,823 psi    146 psi ucture Flange    Barrel Cylinder 3.A-16                          Rev. 35
 
MPS-2 FSAR FIGURE 3A-1 REPRESENTATIVE NODE LOCATIONS - HORIZONTAL MATHEMATICAL MODEL
* MASS NOPE o MASSLESS NODE April 1990                    Rev. 26.2
 
MPS-2 FSAR FIGURE 3A-2 MATHEMATICAL MODEL - HORIZONTAL SEISMIC ANALYSIS R.V. FLANGE
* MASS NODE MASSLESS NODE 48 0                                            13 RIGID CONNECTING LINK                        UPPER        12 l' HINGED                        GUIDE CONNECTION          4S      STRUCTURE 44 43                  11 . CORE 4 - - SUPPORT 42        33        10 BARREL 34 32      *9
                                  . 31        8 30            19 29        7 CORE                              28 SHROUD              FUlL ASSY's      'l1            17 THERMAL 26                SHIELD 16 25 24            15 21          23            14 22 20 LOWER SUPPORT      ~
STRUCTURE April 1990                    Rev. 26.2
 
MPS-2 FSAR FIGURE 3A-3 MATHEMATICAL MODEL - VERTICAL SEISMIC ANALYSIS
* MASS NODE o MASSLES S NODE                        4f' c::::::;, RIG 10 CON NECTING      51      ;2      ~  50 LINK 4  4CJ              '48 4  47 46 MODEL II
                                    ~  45 44 UPPER                  43        MODEL I
                *GUIDE                                    ~  42    CORE STRUCT URE                                  III    SUPPORT
* 41 FUEL                    BARREL 37      ~ 39'AlIGNMENT        40 PLATE          35 4 33                                    30 31 428 26      Z7 24 I
MODEL I                                    22      23 20                                            THERMAL 16      17      SHltLD
* 18 14            FUEL MASS 12  4  13 AND FUEL 9        /ALIGNMENT          11 PLATE                4~ 8 6                  5 LOWER SUPPORT STRUCTUR E.
                                ... I  4 3
42 1
April 1990                        Rev. 26.2
 
MPS-2 FSAR FIGURE 3A-4 CORE SUPPORT BARREL UPPER FLANGE - FINITE ELEMENT MODEL R
                                          ~Z April 1990                    Rev. 26.2
 
MPS-2 FSAR FIGURE 3A-5 CORE SUPPORT BARREL LOWER FLANGE - FINITE ELEMENT MODEL f-
                  ,\
                  \      \
                          \
                    \, \ \
                      \      \
                                            .Jz R
                              \
                        , \\. \
                                                      ~-I-
                                                      ~i-  ...
                                                      ~--
1---...
                                                      ~-
                                                      ~--
                                                      ~--
                                                      ~--
1'/ i'\/ ~      ---
v              ---
                                      ~
v        "      --........
                                                      ~
    \.
Y
    \
      \
April 1990                    Rev. 26.2
 
MPS-2 FSAR FIGURE 3A-6 LATERAL SEISMIC MODEL - MODE 1, 3.065 CPS MODE 1 FREQUENCY 3.065 cps April 1990                      Rev. 26.2
 
MPS-2 FSAR FIGURE 3A-7 LATERAL SEISMIC MODEL - MODE 2, 5.118 CPS MOOE 2 FREQUENCY 5.118 cps                                I I
J I
                                                *I t
I I
f I                                        I T                                        1 I                                      *I I          1 f                                      I          I I                                      ;            ,
      *I                                        I I
                                                                \,
                                                                \
J I                                                  I 1
      ;                                      *I          1 I
I                                        I          I
    +                                        1          I I                                      T I                                        I iL.
t I                      1 I
I                        I L_                      *f April 1990                        Rev. 26.2
 
MPS-2 FSAR FIGURE 3A-8 LATERAL SEISMIC MODEL - MODE 2, 5.118 CPS MODE 3 FREQUENCY 8.166 cps f
r I
      *I I
        ~
I I
I April 1990                      Rev. 26.2
 
MPS-2 FSAR FIGURE 3A-9 REACTOR VESSEL FLANGE VERTICAL RESPONSE SPECTRUM (1% DAMPING) 0.30 0.25
        *0.20
~
  ~
D::    0.15
  ~
LIJ U
U 4:
0.10 o.m 0.10          1.0 FREQUENCY. CDS April 1990                              Rev. 26.2
 
MPS-2 FSAR FIGURE 3A-10 ASHSD FINITE ELEMENT MODEL OF THE CORE SUPPORT BARREL/THERMAL SHIELD SYSTEM 60        80  100 120  140  160  180    200    220    240    260      280    300  320  320.5 o pz,2f      4~      I        I    I    I    I      I    I      I      I      I      I      -I      I      I I I. in.
* IS 20          25          30    35  40 SO 60        70 74.0 --fl.'      I            I                  I    I    f      I I i I /'                                                                        .
8~.S I
85 l~Det.i1A                    csa R,In.
Snubber Elevation NoJe            LaNer
: 6. 125                                          ~ff1bers      .Jocksc; reN' 74.0 -                .;.,.1 9 '0  11 8.**rl                                                                          c::=      -69.0 7-                                                                                    .- 128 6-                                                                                127.
5-                                                                                  12y 4-3-                                                                                125
: 2.                                                                                          1-75* 75 85.S- .......--~
I 321. 125            328.5 Detoll A                                                                    Detail 8 Upper Flang~                                                                lower Flange April 1990                                                          Rev. 26.2
 
MPS-2 FSAR FIGURE 3A-11 VERTICAL SHOCK MODEL
:t Expansion
                          ~U;7;~:p:::-;:~-:---_--i Compensating ~J*ng
~
~
Ten.sion Only                    er Guide Support ....__...
                                                            .~
1\
Sprang                      S  ructure  Plate FI ange                          ~CSB
.t                                Control                    L      Upper Flange
                            ."" Element                    ~
}Cs0rTlpresslo n Only            Shroud
~ pring                          Assemblies Fu.eJ                                    ~    Core 5upport Alignment                                      Barrel Plale Support L~gs
                    \      i                .l
                ~                                                      Thermal I                                                      .Shield
                ~,
                &#xa3;    Core
                ~    Shroud 1.
:t
                ~
                          ~
Cor S        ~ Lower Suppo*rt Plat~ upport 1 ~ structure April 1990                            Rev. 26.2


SingleEnd of ShroudTension & Bending3,548 psi1,295 psiCEA Shrouds: DualEnd of ShroudTension & Bending2,762 psi697 psiUpper Grid BeamsCenter of BeamBending1,652 psi127 psi Upper Guide Structure Flange Junction of Flange &
MPS-2 FSAR FIGURE 3A-12 LATERAL SHOCK MODE Upper Guide Sup~rt Structure Plate Assy Control Element        Core Support Shroud Assemblies      Barrel lugs Fuel Alignment Plate Core Shroud Core Support Plate legend                        Lower Support Struet168 o Mass Point Connected By Lateral Springs April 1990                Rev. 26.2


Barrel CylinderTension & Bending2,823 psi146 psi MPS-2 FSAR April 1990 Rev. 26.2FIGURE 3A-1REPRESENTATIVE NODE LOCATIONS - HORIZONTAL MATHEMATICAL MODEL MPS-2 FSAR April 1990 Rev. 26.2FIGURE 3A-2MATHEMATICAL MODEL - HORIZONTAL SEISMIC ANALYSIS MPS-2 FSAR April 1990 Rev. 26.2FIGURE 3A-3MATHEMATICAL MODEL - VERTICAL SEISMIC ANALYSIS MPS-2 FSAR April 1990 Rev. 26.2FIGURE 3A-4CORE SUPPORT BARREL UPPER FLANGE - FINITE ELEMENT MODEL MPS-2 FSAR April 1990 Rev. 26.2FIGURE 3A-5CORE SUPPORT BARREL LOWER FLANGE -
MPS-2 FSAR FIGURE 3A-13 SAMMSOR DYNASOR FINITE ELEMENT MODEL OF CORE SUPPORT BARREL Upper Barrel
FINITE ELEMENT MODEL MPS-2 FSAR April 1990 Rev. 26.2FIGURE 3A-6LATERAL SEISMI C MODEL - MODE 1, 3.065 CPS MPS-2 FSAR April 1990 Rev. 26.2FIGURE 3A-7LATERAL SEISMIC MODEL - MODE 2, 5.118 CPS MPS-2 FSAR April 1990 Rev. 26.2FIGURE 3A-8LATERAL SEISMIC MODEL - MODE 2, 5.118 CPS MPS-2 FSAR April 1990 Rev. 26.2FIGURE 3A-9REACTOR VESSEL FLANGE VERTIC AL RESPONSE SPECTRUM (1% DAMPING)
                        /       Core Region Lower Barrel April 1990                     Rev. 26.2}}
MPS-2 FSAR April 1990 Rev. 26.2FIGURE 3A-10ASHSD FINITE ELEMENT MODEL OF THE CORE SUPPORT BARREL/THERMAL SHIELD SYSTEM MPS-2 FSAR April 1990 Rev. 26.2FIGURE 3A-11 VERTICAL SHOCK MODEL MPS-2 FSAR April 1990 Rev. 26.2FIGURE 3A-12LATERAL SHOCK MODE MPS-2 FSAR April 1990 Rev. 26.2FIGURE 3A-13SAMMSOR DYNASOR FINITE ELEMENT MODEL OF CORE SUPPORT BARREL}}

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Final Safety Analysis Report, Rev. 35, Chapter 3, Reactor
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Millstone Power Station Unit 2 Safety Analysis Report Chapter 3

Table of Contents tion Title Page

SUMMARY

DESCRIPTION.............................................................................. 3.1-1 1 References................................................................................................... 3.1-3 DESIGN BASES ................................................................................................. 3.2-1 1 Mechanical Design Bases ........................................................................... 3.2-1 1.1 Fuel Assembly Design Bases...................................................................... 3.2-1 1.2 AREVA Fuel Rod Cladding Design Bases................................................. 3.2-2 1.3 Control Element Assembly Design Bases .................................................. 3.2-2 1.4 Reactor Internals Design Bases .................................................................. 3.2-3 1.5 CEDM/RVLMS (HJTC) Pressure Housing Design Bases ......................... 3.2-5 2 Nuclear Design Bases ................................................................................. 3.2-6 3 Thermal and Hydraulic Design Basis ......................................................... 3.2-8 4 References................................................................................................... 3.2-8 MECHANICAL DESIGN ................................................................................... 3.3-1 1 Core Mechanical Design............................................................................. 3.3-1 1.1 AREVA Fuel Rod ....................................................................................... 3.3-1 1.1.1 Fuel Rod Mechanical Criteria..................................................................... 3.3-1 1.1.2 Fuel Rod Design Analyses.......................................................................... 3.3-3 1.2 (Deleted) ..................................................................................................... 3.3-6 1.3 AREVA Fuel Assembly.............................................................................. 3.3-6 1.3.1 Design Summary......................................................................................... 3.3-6 1.3.2 Fuel Assembly Mechanical Criteria ........................................................... 3.3-9 1.4 Fuel Assembly Holddown Device ............................................................ 3.3-12 1.5 Control Element Assembly....................................................................... 3.3-12 1.6 Neutron Source Design ............................................................................. 3.3-13 1.7 In-Core Instruments .................................................................................. 3.3-13 1.8 Heated Junction Thermocouples............................................................... 3.3-14 2 Reactor Internal Structures ....................................................................... 3.3-14 2.1 Core Support Assembly ............................................................................ 3.3-15 2.2 Core Support Barrel .................................................................................. 3.3-15 2.3 Core Support Plate and Support Columns ................................................ 3.3-16 2.4 Core Shroud .............................................................................................. 3.3-16 2.5 Flow Skirt ................................................................................................. 3.3-16 2.6 Upper Guide Structure Assembly ............................................................. 3.3-16 3 Control Element Drive Mechanism .......................................................... 3.3-17 3.1 Design ....................................................................................................... 3.3-17 3-i Rev. 35

tion Title Page 3.2 Control Element Drive Mechanism Pressure Housing ............................. 3.3-18 3.2.1 Heated Junction Thermocouple Pressure Boundary ................................. 3.3-19 3.3 Magnetic Jack Assembly .......................................................................... 3.3-19 3.4 Position Indication .................................................................................... 3.3-19 3.5 Control Element Assembly Disconnect .................................................... 3.3-20 3.6 Test Program............................................................................................. 3.3-20 4 References................................................................................................. 3.3-20 NUCLEAR DESIGN AND EVALUATION ...................................................... 3.4-1 1 General Summary ....................................................................................... 3.4-1 2 Core Description ......................................................................................... 3.4-1 3 Nuclear Core Design................................................................................... 3.4-1 3.1 Analytical Methodology ............................................................................. 3.4-2 3.2 Physics Characteristics ............................................................................... 3.4-2 3.2.1 Power Distribution Considerations ............................................................. 3.4-2 3.2.2 Control Rod Reactivity Requirements ........................................................ 3.4-2 3.2.3 Moderator Temperature Coefficient Considerations .................................. 3.4-3 4 Post-Reload Startup Testing ....................................................................... 3.4-3 5 Reactor Stability ......................................................................................... 3.4-4 5.1 General........................................................................................................ 3.4-4 5.2 Detection of Oscillations ............................................................................ 3.4-4 5.3 Control of Oscillations................................................................................ 3.4-5 5.4 Operating Experience ................................................................................. 3.4-6 5.5 Method of Analysis..................................................................................... 3.4-6 5.5.1 Radial Xenon Oscillations .......................................................................... 3.4-7 5.5.2 Azimuthal Xenon Oscillations.................................................................... 3.4-7 5.5.3 Axial Xenon Oscillations............................................................................ 3.4-7 6 References................................................................................................... 3.4-8 THERMAL-HYDRAULIC DESIGN.................................................................. 3.5-1 1 Design Bases............................................................................................... 3.5-1 1.1 Thermal Design........................................................................................... 3.5-1 1.2 Hydraulic Stability ...................................................................................... 3.5-1 1.3 Coolant Flow Rate, Distribution and Void Fraction................................... 3.5-1 2 Thermal and Hydraulic Characteristics of the Design................................ 3.5-2 2.1 Fuel Temperatures ...................................................................................... 3.5-2 2.1.1 Fuel Cladding Temperatures....................................................................... 3.5-2 2.1.2 Fuel Pellet Temperatures ............................................................................ 3.5-2 2.1.3 UO2 Thermal Conductivity ........................................................................ 3.5-3 3-ii Rev. 35

tion Title Page 2.1.4 Gap Conductance ........................................................................................ 3.5-3 2.2 Departure from Nucleate Boiling Ratio...................................................... 3.5-3 2.2.1 Departure from Nucleate Boiling ............................................................... 3.5-3 2.2.2 Hot Channel Factors ................................................................................... 3.5-3 2.2.3 Effects of Rod Bow on DNBR ................................................................... 3.5-5 2.3 Void Fraction and Distribution ................................................................... 3.5-5 2.4 Coolant Flow Distribution .......................................................................... 3.5-5 2.4.1 Coolant Flow Distribution and Bypass Flow.............................................. 3.5-5 2.4.2 Core Flow Distribution ............................................................................... 3.5-6 2.5 Pressure Losses and Hydraulic Loads ........................................................ 3.5-6 2.5.1 Pressure Losses ........................................................................................... 3.5-6 2.5.2 Hydraulic Loads.......................................................................................... 3.5-7 2.6 Correlation and Physical Data .................................................................... 3.5-7 2.7 Plant Parameters for Thermal-Hydraulic Design........................................ 3.5-7 2.8 Summary of Thermal and Hydraulic Parameters ....................................... 3.5-8 3 Thermal And Hydraulic Evaluation............................................................ 3.5-8 3.1 Analytical Techniques and Uncertainties ................................................... 3.5-8 3.1.1 XCOBRA-IIIC DNBR Analyses ................................................................ 3.5-8 3.1.2 Parameter Uncertainties .............................................................................. 3.5-8 3.2 Hydraulic Instability Analysis .................................................................... 3.5-8 3.3 Core Hydraulics ........................................................................................ 3.5-11 3.3.1 Fuel Assembly Pressure Drop Coefficients .............................................. 3.5-11 3.3.2 Guide Tube Bypass Flow and Heating Analysis ...................................... 3.5-12 3.3.3 Control Element Assembly Insertion Time Analysis ............................... 3.5-13 3.3.4 Fuel Assembly Liftoff............................................................................... 3.5-13 4 Tests And Inspections ............................................................................... 3.5-14 4.1 Reactor Testing ......................................................................................... 3.5-14 4.2 AREVA DNB and Hydraulic Testing ...................................................... 3.5-14 4.2.1 DNB Testing ............................................................................................. 3.5-14 4.2.2 Fuel Assembly Hydraulic Testing ............................................................ 3.5-14 5 References................................................................................................. 3.5-15 ANALYSIS OF REACTOR VESSEL INTERNALS ........................................ 3.A-1

.1 Seismic Analysis........................................................................................ 3.A-1

.1.1 Introduction................................................................................................ 3.A-1

.1.2 Method of Analysis.................................................................................... 3.A-1

.1.2.1 General....................................................................................................... 3.A-1

.1.2.2 Mathematical Models ................................................................................ 3.A-1

.1.2.3 Natural Frequencies and Normal Modes ................................................... 3.A-3

.1.2.4 Response Calculations .............................................................................. 3.A-4 3-iii Rev. 35

tion Title Page

.1.3 Results........................................................................................................ 3.A-5

.1.4 Conclusion ................................................................................................. 3.A-5

.2 Normal Operating Analysis ....................................................................... 3.A-5

.3 Loss of Coolant Accident Analysis ........................................................... 3.A-7

.3.1 Discussion .................................................................................................. 3.A-7

.3.2 Analysis Codes ........................................................................................ 3.A-10

.4 Effects of Thermal Shield Removal......................................................... 3.A-11

.5 Leak-Before-Break Analysis ................................................................... 3.A-11

.6 References................................................................................................ 3.A-12 3-iv Rev. 35

List of Tables mber Title 1 Stress Limits for Reactor Vessel Internal Structures 1 Mechanical Design Parameters

  • 2 Pressurized Water Reactor Primary Coolant Water Chemistry Recommended Specifications 1 Fuel Characteristics for a Representative Reload Core 2 Neutronics Characteristics for a Representative Reload Core 3 Representative Shutdown Margin Requirements 1 Nominal Reactor and Fuel Design Parameters 2 Design Operating Hydraulic Loads on Vessel Internals 3 Uncertainty Sources for DNBR Calculations (DELETED)

-1 Natural Frequencies for Vertical Seismic Analysis Mathematical Model

-2 Seismic Stresses in Critical Reactor Internals Components for the Design Basis Earthquake 3-v Rev. 35

List of Figures ure Title 1 Reactor Vertical Arrangement 2 Reactor Core Cross Section 1 Fuel Rod Assembly 2A AREVA - Reload Fuel Assembly Batch S and Prior 2B AREVA - Reload Fuel Assembly Batch T and Later 3A AREVA - Reload Fuel Assembly Components Batch S and Prior 3B AREVA - Reload Fuel Assembly Components Batch T and Later 4A Bi-Metallic Fuel Spacer Assembly 4B HTP Fuel Space Assembly 5 Fuel Assembly Hold Down Device

-6 Control Element Assembly 7 Control Element Assembly Materials 8 Control Element Assemblies Group and Number Designation 9 Core Orientation 10 In-Core Instrumentation Assembly

-11 Reactor Internals Assembly 12 Pressure Vessel-Core Support Barrel Snubber Assembly 13 Core Shroud Assembly 14 Upper Guide Structure Assembly 15 Control Element Drive Mechanism (Magnetic Jack) 16 (Left Blank Intentionally) 17 Heated Junction Thermocouple Probe Pressure Boundary Installation 18 Typical Heated Junction Thermocouple Probe Assembly Installation 19 Placement of Natural Uranium Replacement Fuel Rods and Fuel Assembly Orientation Relative to the Core Baffle for Cycle 19 1 Representative Full Core Loading Pattern 2 Representative Quarter Core Loading Pattern 3-vi Rev. 35

NOTE: REFER TO THE CONTROLLED PLANT DRAWING FOR THE LATEST REVISION.

ure Title 3 Representative BOC and EOC Exposure Distribution 4 Representative Boron Letdown, HFP, ARO 5 Representative Normalized Power Distributions, Hot Full Power, Equilibrium Xenon, 150 MWD/MTU 6 Representative Normalized Power Distribution, Hot Full Power, Equilibrium Xenon, 18,020 MWD/MTU

-1 Representative Node Locations - Horizontal Mathematical Model

-2 Mathematical Model - Horizontal Seismic Analysis

-3 Mathematical Model - Vertical Seismic Analysis

-4 Core Support Barrel Upper Flange - Finite Element Model

-5 Core Support Barrel Lower Flange - Finite Element Model

-6 Lateral Seismic Model - Mode 1, 3.065 CPS

-7 Lateral Seismic Model - Mode 2, 5.118 CPS

-8 Lateral Seismic Model - Mode 2, 5.118 CPS

-9 Reactor Vessel Flange Vertical Response Spectrum (1% Damping)

-10 ASHSD Finite Element Model of the Core Support Barrel/Thermal Shield System

-11 Vertical Shock Model

-12 Lateral Shock Mode

-13 SAMMSOR DYNASOR Finite Element Model of Core Support Barrel 3-vii Rev. 35

SUMMARY

DESCRIPTION reactor is of the pressurized water type using two reactor coolant loops. A vertical cross ion of the reactor is shown in Figure 3.1-1. The reactor core is composed of 217 fuel mblies, 73 control element assemblies (CEA) and four neutron source assemblies. The fuel mblies are arranged to approximate a right circular cylinder with an equivalent diameter of inches and an active length of 136.7 inches. The fuel assemblies are comprised of a structure fuel and poison rods. The structure, which provides for 176 rod positions, consists of five de tubes attached to spacer grids and is enclosed at the top and bottom by end fittings. Each of guide tubes replaces four fuel rod positions and provides a channel which guides the control ment over its entire length of travel. In selected fuel assemblies the central guide tube houses ore instrumentation. The reactor is currently fueled by assemblies produced by AREVA.

fuel is low enrichment UO2 in the form of ceramic pellets and encapsulated in zircaloy tubes.

se tubes are seal welded as hermetic enclosures.

ure 3.1-2 shows a view of the reactor core cross section and some dimensional relations ween fuel assemblies, fuel rods and CEA guide tubes.

reactor internals support and orient the fuel assemblies and CEAs, absorb the static and amic loads and transmit the loads to the reactor vessel flange, provide a passage way for the tor coolant, and guide in-core instrumentation.

internals will safely perform their function during normal operating, upset and emergency ditions. The internals are designed to safely withstand the forces due to dead weight, pressure erential, flow impingement, temperature differential, vibrations and seismic acceleration. All tor components are considered category 1 for seismic design. The reactor internals design ts deflection where required by function. Where necessary, components have been subjected atigue analysis. Where appropriate, the effect of neutron irradiation on the materials concerned cluded in the design evaluation. The effects of shock loadings on the internals is included in design analysis.

ctivity control is provided by two independent systems: The control element drive system DS) and the chemical and volume control system (CVCS). The CEDS controls short term tivity changes and is used for rapid shutdown. The CVCS is used to compensate for long term tivity changes and can make the reactor subcritical without the benefit of the CEDS. The gn of the core and the reactor protective system (RPS) prevents fuel damage limits from being eeded for any single malfunction in either of the reactivity control systems.

CEAs consist of five poison rods (control elements) assembled in a square array, with one rod he center. The rods are connected to a spider casting which is coupled to the control element e mechanism (CEDM) shaft. There are a total of 73 CEAs. Some CEAs are mechanically nected in pairs and are known as dual CEAs.

3.1-1 Rev. 35

maximum reactivity worth of the CEAs and the associated reactivity addition rate are limited ore, CEA and CEDS design to prevent sudden large reactivity increases. The design restraints such that reactivity increases will not result in violation of the fuel damage limits, rupture of reactor coolant pressure boundary (RCPB), or disruption of the core or other internals icient to impair the effectiveness of emergency cooling.

three-batch fuel management scheme is employed, where approximately 40 percent of the is replaced at each refueling. Sufficient margin is provided to ensure that peak burnups of the vidual fuel assemblies are within acceptable limits.

nuclear design of the core will ensure that the combined response of all reactivity coefficients n increase in reactor thermal power yields a net decrease in reactivity and that CEAs are ved in groups to satisfy the requirements of shutdown, power level changes and operational euvering. The control systems are designed to produce power distributions that are within the eptable limits on overall nuclear heat flux factor (FNQ) and departure from nucleate boiling o (DNBR). The RPS and administrative controls ensure that these limits are not exceeded.

reactor coolant enters the upper section of the reactor vessel through four inlet nozzles, flows nward between the reactor vessel shell and the core barrel, and passes through the flow skirt into the lower plenum where the flow distribution is equalized. The coolant then flows ard through the core removing heat from the fuel rods, exits from the reactor vessel through outlet nozzles and passes through the tube side of the vertical U tube steam generators re heat is transferred to the secondary system. The reactor coolant pumps (RCPs) return the lant to the reactor vessel.

principal objective of the thermal-hydraulic design is to avoid fuel damage during normal ration and anticipated transients. It is recognized that there is a small probability of limited damage in certain situations as discussed in Chapter 14.

rder to meet the objective of the thermal-hydraulic design the following design limits are blished, but violation of either is not necessarily equivalent to fuel damage:

a. There is a high confidence level that departure from nuclear boiling (DNB) is avoided during normal operation and anticipated transients. This is achieved by confirming the minimum DNBR calculated according to the HTP correlation (Reference 3.1-1) is greater than the 95/95 limit for the correlation;
b. The melting point of the UO2 fuel is not reached during normal operation or anticipated transients.

RPS and the reactor control system (RCS) provide for automatic reactor trip or corrective ons before these design limits are exceeded.

3.1-2 Rev. 35

raulic design is discussed in Section 3.5.

1 REFERENCES 1 EMF-92-153(P)(A) Rev. 1, HTP: Departure From Nucleate Boiling Correlation for High Thermal Performance Fuel, Siemens Power Corporation, January 2005.

3.1-3 Rev. 35

MPS-2 FSAR FIGURE 3.1-1 REACTOR VERTICAL ARRANGEMENT

'N'-CORE tNSTRUMENT ATI ON ASSEMBLY CONTROl ELEMENT ASSEMBLY EX'PANSION CFUllYWITHDRAWNl RI~

ALIGNMENT PIN

~ UPPER 40' 11 7h" GUIDE

~

.....--INLET NOZZLE Nt--CORE

..---==--!J-J~~.",MM SUPPORT BARREL*

n5-7i!OH ~...,..,.- FUEL A,CTIVE ASSEMBLY CORE UENGTH ""-CORE SHROUD CORE CORE: STOP SUPPORT ASSEMBLY

.t FlCNY SKIRT April 1998 Rev. 26.2

MPS-2 FSAR FIGURE 3.1-2 REACTOR CORE CROSS SECTION CORE REACTOR

~~UIVALENT VESSEl.

13~ETER.

CORE SUPPORT BARREl----.;;~ CORE SHROUD FUEl ROD GUIDE TUBE 0.440' OD o!~..,

~TT TT

"";~_' "TT r T' '.'

0.S8O".........

~~~nOD

~ .. 1-0.1;:11- ~ Oltr?P~E

..tHua RODS J

0.200 11 WATER GAP April 1998 Rev. 26.2

full power thermal rating of the core is 2,700 MWt. The physics and thermal and hydraulic rmation presented in this section is based on this core power level.

1 MECHANICAL DESIGN BASES 1.1 Fuel Assembly Design Bases design bases for evaluating the structural integrity of AREVA fuel assemblies are:

Fuel Assembly Handling The fuel assembly is evaluated for dynamic axial loads of approximately 2.5 times the fuel assembly weight.

For All Applied Loads for Normal Operation and Anticipated Operational Events Fuel assembly component strength is evaluated against either prototype testing or elastic stress analysis. When the stress analysis method is used, the stress limits presented in the ASME Boiler and Pressure Vessel Code,Section III, Division 1, are used as a guide.

stress design limits for structural components are:

Pm 1.0Sm Pm + Pb 1.5Sm P + Q 3.0Sm re:

Pm is the primary membrane stress intensity Pb is the primary bending stress intensity P is the primary stress intensity Q is the secondary stress intensity design stress, Sm is identified in the ASME Boiler and Pressure Vessel Code for austenitic nless steel as a function of temperature. In the case of Zircaloy, which is not specifically tified in the ASME Boiler and Pressure Vessel Code, the design stress is identified as the er of two-thirds the yield stress, Sy, or one-third the ultimate stress, Su.

ASME Boiler and Pressure Vessel Code defines the stress intensity based on the maximum ar stress theory. The stress intensity is equal to one-half the largest algebraic difference ween two principal stresses.

3.2-1 Rev. 35

ary stress is that it is not self-limiting. If a primary stress exceeds the yield strength of the erial through the entire wall thickness, the prevention of failure is entirely dependent on the in-hardening properties of the material.

ondary stresses are developed by the self-constraint of a structure. It must satisfy an imposed in pattern rather than being in equilibrium with an external load. The basic characteristic of a ondary stress is that it is self-limiting. Local yielding and minor distortions can satisfy the ontinuity conditions due to thermal expansions which cause the stress to occur.

Loads during Postulated Accidents lection or failure of components shall not interfere with reactor shutdown or emergency ling of the fuel rods.

fuel assembly structural component stresses under faulted conditions are evaluated using arily the methods outlined in Appendix F of the ASME Boiler and Pressure Vessel Code, tion III. The current methods utilize the limits provided for elastic system analysis.

design stress intensity value (Sm) is defined the same as for normal operating conditions.

cer grid crush load strength is based on the 95% confidence level on the true mean as taken m test measurements on unirradiated production grids at (or corrected to) operating perature.

1.2 AREVA Fuel Rod Cladding Design Bases iscussion of the AREVA fuel rod cladding is given as part of the AREVA fuel rod discussion ection 3.3.1.1.

1.3 Control Element Assembly Design Bases CEA has been designed to ensure that the stress intensities in the individual structural ponents do not exceed the allowable limits for the appropriate material established in tion III of the ASME Boiler and Pressure Vessel Code. The exceptions to this criterion are that he Inconel 625 cladding is permitted to sustain plastic strain up to 3 percent due to irradiation uced expansion of the filler materials, and (b) because the ASME Code does not apply to ngs, the allowable stresses for the CEA springs are based on values which have been proven in tice.

CEA stress analyses consider the following load sources:

a. Internal pressure build up due to the effect of irradiation on B4C (production of helium).

3.2-2 Rev. 35

assumed).

c. Dynamic stresses produced by seismic loading.
d. Dynamic loads produced by stepping motion of the magnetic jack.
e. Mechanical and hydraulic loads produced during SCRAM.
f. Cladding loads produced by differential expansion between clad and filler materials.

addition to the comparison of calculated stress levels with allowable stresses, the fatigue age produced by significant cyclic stresses is also determined. It is a design requirement that calculated cumulative damage factor for any location may not be equal to or greater than 1.0.

fatigue usage factor calculations are based on the fatigue curves (stress range vs. number of les) contained in Section III of the ASME Boiler and Pressure Vessel Code.

1.4 Reactor Internals Design Bases reactor vessel internals are designed to meet the loading conditions and the design limits cified below. The materials used in fabrication of the reactor internal structures are primarily e 304 stainless steel. The flow skirt is fabricated from Inconel. Welded connections are used re feasible; however, in locations where mechanical connections are required, structural eners are used which are designed to remain captured in the event of a single failure.

ctural fastener material is typically a high strength austenitic stainless steel; however, in less cal applications, Type 316 stainless steel is employed. Hardfacing, of Stellite material, is used ear points. The effect of irradiation on the properties of the materials is considered in the gn of the reactor internal structures.

Categorization and Combination of Loadings

1. Normal Operating and Upset Conditions The reactor vessel internals are designed to perform their functions safely without shutdown. The combination of design loadings for these conditions are the following:

Normal operating temperature differences Normal operating pressure differences Low impingement loads Weights, reactions and superimposed loads 3.2-3 Rev. 35

Shock loads (including OBE)

Transient loadings of frequent occurrences not requiring shutdown Handling loads

2. Emergency Conditions The internals are designed to permit an acceptable amount of local yielding while experiencing the loadings listed above with the SSE load replacing the OBE load.
3. Faulted Conditions Permanent deformation of the reactor internal structures is permitted. The loadings for these conditions include all the loadings listed for emergency conditions plus the loadings resulting from the postulated LOCA.

Design Limits Reactor internal components are designed to ensure that the stress levels and deflections are within an acceptable range. The stress values for core support structures are not greater than those given in the May 1972 draft of Section III of the ASME Boiler and Pressure Vessel Code, Subsection NG, including Appendix F, Rules for Evaluation of Faulted Conditions. Stress limits for the reactor vessel core support structures are presented in Table 3.2-1. In addition, to properly perform their functions, the reactor internal structures will satisfy the deformation limits listed below.

1. Under design loadings plus operating basis earthquake forces or normal operating loadings plus SSE forces, deflections will be limited so that the CEAs can function and adequate core cooling is preserved.
2. Under normal operating loadings plus SSE forces plus pipe rupture loadings resulting from a break of the largest line connect to the primary system piping, deflections will be limited so that the core will be held in place, adequate core cooling is preserved, and all CEAs can be inserted. Those deflections which would influence CEA movement will be limited to less than 80 percent of the deflections required to prevent CEA insertion.
3. Under normal operating loadings plus SSE forces plus the maximum pipe rupture loadings resulting from the full spectrum of pipe breaks, deflections will be limited so that the core will be held in place and adequate core cooling is preserved.

Although CEA insertion is not required for a safe and orderly shutdown for break sizes greater than the largest line connected to the primary system piping, calculations show that the CEAs will be insertable for larger breaks except for a 3.2-4 Rev. 35

1.5 CEDM/RVLMS (HJTC) Pressure Housing Design Bases control element drive mechanism and Reactor Vessel Level Monitoring System (RVLMS) sure housings form part of the reactor coolant boundary and are, therefore, designed to meet stress requirements consistent with those of the reactor vessel closure head. The limiting sses in the CEDMs and RVLMS pressure boundary components due to the design, Level A, el B, Level C, Level D and Test conditions satisfy ASME Boiler Pressure Vessel Code, tion III, Subsection NB plus Appendix 1 and Section II, Part D, 1998 Edition through 2000 enda, including Code Case N-4-12 for the CEDM motor housing material.

CEDMs and the RVLMs are designed to function normally during and after exposure to mal operating conditions plus the design basis earthquake (DBE). Under normal operating ditions, plus DBE, plus pipe rupture loadings, deflections of the CEDM will be limited so that CEAs can be inserted after exposure to these conditions. Those deflections, which could uence CEA movement, will be limited to less than 80 percent of the deflections required to vent CEA movement. The RVLMS and the adjacent CEDMs do not contact each other with imum lateral displacement of the pressure housings.

Loading Combinations ASME Code Subsection sign Condition Pm Sm NB-3221 P1 1.5Sm P1 + Pb < 1.5Sm vel A and Level B P1 + Pb + Q 3Sm NB-3222 and NB3223 ormal and Upset) U1 vel C Condition Pm greater of [1.2Sm, Sy] NB-3224 mergency) P1 + Pb greater of [1.8Sm, 1.5Sy]

vel D Condition Pm lesser of [2.4Sm, 0.7Su] Paragraph F-1330 or F-1340, Appendix F ulted) P1 + Pb lesser of [3.6Sm, 1.05Su]

t Conditions Pm 0.9Sy NB-3226 Pm + Pb 1.35Sy when Pm 0.67Sy or Pm + Pb (2.15 Sy - 1.2Pm) when 0.67Sy

< Pm 0.9Sy sign Condition Pm Sm NB-3221 3.2-5 Rev. 35

ar Stress 0.6Sm NB-3227.2 ere Pm = General primary membrane stress intensity P1 = Primary local membrane stress intensity P1 + Pb = Primary membrane plus bending stress intensity P1 + Pb + Q = Primary plus secondary stress intensity Sm = Design stress intensity Sy = Yield strength Su = Tensile strength U = Cumulative fatigue usage factor 2 NUCLEAR DESIGN BASES initial full power thermal rating of the core is 2700 MWt. It is upon this power level that the sics and thermal and hydraulic information presented in this section are based. The design s for the nuclear design of the fuel and reactivity control systems are:

a. Excess Reactivity and Fuel Burnup The excess reactivity provided for each cycle is based on the depletion characteristics of the fuel and burnable poison and the desired burnup for each cycle. The desired burnup is based on the economic analysis of both the fuel cost and the projected operating load demand cycle for the plant. The average burnup in the core is chosen so as to insure that the peak assembly burnup is not greater than 56,000 MWD/MTU for Batch N, 52,500 MWD/MTU for Batch P, and 57,400 MWD/MTU for Batch R and later.
b. Core Design Lifetime and Fuel Replacement Program The core design lifetime and fuel replacement program are based on a three region core with approximately 40 percent of the fuel assemblies replaced at each refueling.
c. Negative Reactivity Feedback and Reactivity Coefficients The negative reactivity feedback provided by the design is based on the requirement of General Design Criterion (GDC) 11. In the power operating range, the inherent combined response of the reactivity feedback characteristics (fuel temperature coefficient (FTC), moderator temperature coefficient (MTC),

moderator void coefficient (MVC), and moderator pressure coefficient (MPC)) to an increase in reactor thermal power will be a decrease in reactivity.

3.2-6 Rev. 35

The burnable poison reactivity worth provided in the design will be sufficient to ensure that moderator coefficients of reactivity have magnitudes and algebraic signs consistent with the requirements for negative reactivity feedback and acceptable consequence in the event of postulated accidents or anticipated operational occurrences, viewed in conjunction with the supplied protective equipment.

e. Stability Criteria The design of the reactor and the instrumentation and control systems is based on meeting the requirements of GDC 12 with respect to spatial oscillations and stability. Sufficient CEA rod worth will be available to suppress xenon-induced power oscillations.
f. Maximum Controlled Reactivity Insertion Rates The maximum reactivity addition rates are limited by core, CEA, and reactor regulating system (RRS) design based on preventing increases in reactivity which would result in the violation of specified acceptable fuel design limits, damage to the reactor pressure boundary, or disruption of the core or other internals sufficient to impair the effectiveness of emergency core cooling.
g. Power Distribution Control Acceptable operation of the reactor in the absence of an accidental transient depends on maintaining a relationship among many parameters, some of which depend on the power distribution. In the absence of an accidental transient the power distribution is controlled such that in conjunction with other controlled parameters, limiting conditions of operation (LCO) are not violated. LCO are not less conservative than the initial conditions used in the accident analyses in Chapter 14. LCO and limiting safety system settings (LSSS) are determined such that specified acceptable fuel design limits are not violated as a result of anticipated operational occurrences and such that specified predicted acceptable consequence are not exceeded for other postulated accidents.
h. Shutdown Margins and Stuck Rod Criteria The amount of reactivity available from insertion of withdrawn CEAs is required to be sufficient, under all power operating conditions, to ensure that the reactor can be brought to at least 3.6 percent subcritical from the existing condition, including the effects of cooldown to an average coolant temperature of 532°F, even when the highest worth CEA fails to insert. This criteria is exclusive of any safety allowance and is consistent with the most pessimistic analysis in Chapter 14.

3.2-7 Rev. 35

The chemical and volume control system (CVCS) (Section 9.2) is used to adjust dissolved boron concentration in the moderator. After a reactor shutdown, this system is able to compensate for the reactivity changes associated with xenon decay and reactor coolant temperature decrease to ambient temperature. It also provides adequate shutdown margin during refueling. This system also has the capability of controlling long term reactivity changes due to fuel burnup, and reactivity changes during xenon transients resulting from changes in reactor load independently of the CEAs. In particular, any xenon transient may be accommodated at any time in the fuel cycle.

3 THERMAL AND HYDRAULIC DESIGN BASIS idance of thermally induced fuel damage during normal steady state and anticipated transient ration is the principal thermal and hydraulic design basis. It is recognized that there is a small bability of limited fuel damage in certain unlikely accident situations discussed in Chapter 14.

following corollary design basis are established, but violation of them is not necessarily ivalent to fuel damage.

a. A limit corresponding to 95% probability with 95% confidence (Reference 3.2-1) is set on the departure from nucleate boiling ratio (DNBR) during normal operation and any anticipated transients as calculated according to the HTP correlation.
b. The peak temperature of the fuel will be less than the melting point during normal operation and anticipated transients.

reactor control and protection system will provide for automatic reactor trip or other ective action before these design limits are exceeded.

core hydraulic resistance was considered in establishing the operational limits curves vided in Figures 4.5-4 and 4.5-5, and the Low Temperature Overpressure Protection (LTOP) tem described in Section 7.4.8. As fuel design changes, effects on the flow resistance will be luated to determine the impact.

4 REFERENCES 1 EMF-92-153(P)(A) Rev. 1, HTP: Departure From Nucleate Boiling Correlation for High Thermal Performance Fuel, Siemens Power Corporation, January 2005.

3.2-8 Rev. 35

Operating Conditions Stress Categories and Limits of Stress Intensities Normal and Upset Figure NG 3221.1 including notes Emergency Figure NG 3224.1 including notes Faulted Appendix F, Rules for Evaluating Faulted Conditions 3.2-9 Rev. 35

reactor core and internals are shown in Figure 3.3-1. A cross section of the reactor core and rnals is shown in Figure 3.1-2. Mechanical design features of the reactor internals, the control ment drive mechanisms (CEDM) and the core are described below. Mechanical design meters are listed in Table 3.3-1.

1 CORE MECHANICAL DESIGN core approximates a right circular cylinder with an equivalent diameter of 136 inches and an ve height of 136.7 inches. It is made up of Zircaloy-4 clad fuel rods containing slightly ched uranium in the form of sintered UO2 pellets and UO2-Gd2O3 pellets. The fuel rods are uped into 217 assemblies.

rt term reactivity control is provided by 73 control element assemblies (CEA). The CEAs are ded within the core by the guide tubes which are integral parts of the fuel assemblies.

1.1 AREVA Fuel Rod detailed fuel rod design (see Figure 3.3-1) establishes such parameters as pellet diameter and th, density, cladding-pellet diametral gap, fission gas plenum size, and rod pre-pressurization

l. The design also considers effects and physical properties of fuel rod components which y with burnup.

integrity of the fuel rods is ensured by designing to prevent excessive fuel temperatures, essive internal rod gas pressures, and excessive cladding stresses and strains. This end is ieved by designing the fuel rods to satisfy the design criteria during normal operation and cipated operational occurrences over the fuel lifetime. For each design criteria, the ormance of the most limiting fuel rod shall not exceed the specified limits.

l rods are designed to function throughout the design life of the fuel based upon the reactor rating conditions designated below without loss of mechanical integrity, significant ensional distortion, or release of fuel or fission products.

assemblies were evaluated for a peak assembly burnup of 56,000 MDW/MTU for Batch N, 00 MWD/MTU for Batch P, and 57,400 MWD/MTU for Batch R and later.

Millstone Unit 2 Cycle 19 reload core included four fuel assemblies with natural uranium acement fuel rods with an anti-rotation feature designed to prevent spinning of the rod during rations. The four assemblies containing replacement rods, and the conditions under which were evaluated for use, are discussed in Section 3.3.1.3.1, "Design Summary".

1.1.1 Fuel Rod Mechanical Criteria cladding primary and secondary stresses shall meet the 1977 ASME Boiler and Pressure sel Code Section III (Reference 3.3-1) requirements summarized below:

3.3-1 Rev. 35

Stress Intensity Limits (Parameter) Yield Strength Ultimate Tensile Strength mary Membrane (Pm) < 2/3 Sy < 1/3 Su mary Membrane Plus Primary Bending (Pm + Pb) < 1.0 Sy < 0.5 Su mary Plus Secondary (P + Q) < 2.0 Sy < 1.0 Su mary stresses are developed by the imposed loading which is necessary to satisfy the laws of ilibrium between external and internal forces and moments. The basic characteristic of a ary stress is that it is not self-limiting. If a primary stress exceeds the yield strength of the erial through the entire thickness, the prevention of failure is entirely dependent on the strain-dening properties of the material.

ondary stresses are developed by the self constraint of a structure. It must satisfy an imposed in pattern rather than being in equilibrium with an external load. The basic characteristic of a ondary stress is that it is self-limiting. Local yielding and minor distortions can satisfy the ontinuity conditions due to thermal expansions which cause the stress to occur.

dding circumferential strain shall not exceed the design limit through end-of-life (EOL).

total uniform strain, elastic and plastic shall not exceed the design limit during a transient.

strain analysis was performed with the RODEX2 (Reference 3.3-2) RAMPEX codes chmarked to available power ramp test data, i.e., INTERRAMP, OVERRAMP, and PERRAMP.

fuel rod shall be designed such that at a rod average burnup when substantial axial solidation has occurred, the total clad creep deformation shall not exceed the initial minimum metral fuel cladding gap. This will prevent pellet hangups allowing the plenum spring to close l gaps until densification is substantially complete, thus preventing the formation of pellet mn gaps of sufficient size for clad flattening.

fuel rod pressure at EOL shall not exceed the criteria approved by the NRC (Ref. 3.3-3). A ew of departure from nucleate boiling ratio (DNBR) limits for condition III or IV postulated dents events is required for fuel rods that exceed nominal system pressure. When fuel rod sure is predicted to exceed system pressure, the pellet-cladding gap shall not increase for dy or increasing power conditions. Analysis approved by the NRC has shown that the fuel rod pressure can safely exceed system pressure without causing any damage to the cladding.

al cladding wall thinning due to generalized external and internal corrosion shall not exceed a e which will impair mechanical performance over the projected fuel rod design lifetime under most adverse projected power conditions within coolant chemistry limits recommendations of le 3.3-2. It will also assure that the metal/oxide interface temperature will remain well below 3.3-2 Rev. 35

cumulative usage factor for cyclic stresses for all important cyclic loading conditions shall exceed the design limit.

clearance between the upper and lower tie plate shall be able to accommodate the maximum erential fuel rod and fuel assembly growth to the designed burnup.

centerline temperature of the hottest pellet shall be below the melting temperature. Fuel terline temperature is calculated at overpower conditions to verify that fuel pellet overheating s not occur during normal operation and anticipated operational occurrences.

1.1.2 Fuel Rod Design Analyses h design analysis was performed with AREVA methodology which involves a well defined ction of appropriate data and parameters, and the latest approved versions of computer codes.

s methodology, as required, has been submitted to the Nuclear Regulatory Commission (NRC) approved. The analysis is performed in accordance with the methods described in AREVAs alification of Exxon Nuclear Fuel For Extended Burnup (Reference 3.3-3).

cladding steady state stress analysis was performed by considering primary and secondary mbrane and bending stresses due to hydrostatic pressure, flow-induced vibration, spacer tact, pellet cladding interaction (PCI), thermal and mechanical bow and thermal gradients.

sses were calculated for the various combinations of the following conditions:

a. beginning of life (BOL) and EOL
b. cold and hot conditions
c. at mid-span and at spacer locations
d. at both the inner and outer surfaces of the cladding analysis was performed for the various sources of stress, including pressure, thermal, spacer tact, PCI, and rod bow. The applicable stresses at each orthogonal direction were combined to ulate the maximum stress intensities which are compared to the ASME design criteria. The lts of the analysis indicate that all stress values are within acceptable design limits for both L and EOL, hot and cold conditions. The EOL stresses have ample margin for both the hot and condition stresses.

cladding steady state strain is evaluated with the RODEX2 code, which has been approved by NRC (Reference 3.3-2). The code considers the thermal-hydraulic environment at the ding surface, the pressure inside the cladding, and the thermal, mechanical and compositional e of the fuel and cladding. Pellet density, swelling, densification, and fission gas release or orption models, and cladding and pellet diameters are input to RODEX2 to provide the most 3.3-3 Rev. 35

a. Radial Thermal Conduction and Gap Conductance
b. Fuel Swelling, Densification, Cracking, and Crack Healing
c. Gas Release and Absorption
d. Cladding Creep Deformation and Irradiation-Induced Growth
e. Cladding Corrosion
f. PCI
g. Free Rod Volume and Gas Pressure calculations are performed on a time incremental basis with conditions updated at each ulated increment so that the power history and path dependent processes can be modeled. The l dependence of the power and burnup distributions are handled by dividing the fuel rod into a ber of axial and radial regions. Power distributions can be changed at any desired time, and coolant and cladding temperatures are readjusted in all the regions. All the performance dels, e.g., giving the deformations of the fuel and cladding and gas release, are calculated at cessive times during each period of assumed constant power generation. The calculated ding strain is reviewed throughout the life of the fuel and both the maximum circumferential in and the maximum strain increment are compared with the design criteria. The calculated in did not exceed the strain limit. Both the maximum strain and the positive strain increment below the design limit strain.

ramping strain and the fatigue evaluation of the fuel rod were evaluated. The ramps are med to occur anytime during the irradiation and may reach the maximum peaking factor wed by the limits of operation. The ramps are analyzed either from cold shutdown or from a ety of hot powered starting conditions. The approach to rated power at the beginning of each tor cycle is performed to satisfy the AREVA maneuvering and conditioning mmendations. The clad response during ramping power changes is calculated with the MPEX code. This code calculates the PCI during a power ramp for one axial node at a time.

initial conditions are obtained from RODEX2 output. The RAMPEX code considers the mal condition of the rod in its flow channel, and the mechanical interactions that result from and cladding creep at any desired axial section in the rod during the power ramp. As pared to RODEX2, RAMPEX additionally models the pellet cladding axial stress interaction, mary creep with strain hardening, the effects of pellet chips, and localized stresses due to ing.

RAMPEX code provides the hoop stress and the stress intensity. The stress results of the ping analysis are used to evaluate the cladding fatigue damage through life due to the cyclic 3.3-4 Rev. 35

MPEX over the power cycling range, are compared with this curve to determine the allowed les for each stress range. This result is combined with the projected number of duty cycles to rmine a fatigue usage factor. All of the reactor cycle (startup) ramp stresses were within the gn limit.

ep collapse calculations are performed with RODEX2 and COLAPX codes. The RODEX2 e determines the cladding temperature and internal pressure history based on a model which ounts for changes in fuel rod volumes, fuel densification and swelling, and fill gas absorption.

reactor coolant, fuel rod internal temperature, and pressure histories generated by the DEX2 analysis are input to the COLAPX code along with a conservative statistical estimate of al cladding ovality and the fast flux history. The COLAPX code calculates, by large deflection ry, the ovality of the cladding as a function of time while the uniform cladding creepdown is ined by the RODEX2 analysis. The cladding ovality increase and creepdown are summed, at d average burnup when substantial axial consolidation has occurred, to show that they remain than the initial minimum pellet clad gap. Measurements of highly densifying irradiated fuel e demonstrated that pellet densification is essentially complete by the time the fuel has ined this burnup so that further creepdown after this phase will not result in significant pellet ellet gaps. The combined radial creepdown was shown to meet the design criteria. This will vent pellet hangups due to cladding creep, allowing the plenum spring to close axial gaps until sification is substantially complete, and thus assures that clad collapse will not occur. The h of the plenum spring is less than the spacing calculated for stiffening rings in a cylindrical l under external pressure which will prevent clad collapse in the plenum area.

culation of the gas pressure within a fuel rod is performed with the RODEX2 code. The initial gas is found by calculating the initial free volume and using the ideal gas law, along with input es for fill gas pressure and reference fill gas temperature. The free gaseous fission product d is calculated for each axial region and the total yield obtained by summing the axial region tributions. The power of each history used was multiplied for each cycle by a factor required the highest projected rod power to reach the Fr limit plus uncertainties. The calculations show for all power histories analyzed, the rod internal gas pressure will remain below the criteria roved by the NRC (Reference 3.3-3) for use in extended burnup gas pressure analysis.

waterside corrosion of fuel rods is evaluated with the MATPRO-11 (Reference 3.3-5) elation. The MATPRO-11 model is a two-stage corrosion rate model which is cubic in endence on oxide thickness until a transition to a subsequent linear dependence occurs. To ulate the rate changes as a function of both oxide thickness and the operating conditions of the rod, the MATPRO model is incorporated into AREVAs RODEX2 fuel performance code.

RODEX2 code determines the temperature increase of the water along the fuel rod assuming t balance within a channel for the prescribed mass flow and inlet temperature. The radial perature drops are evaluated successively between the water, the oxide surface, the metal/

de interface, and the inside of the cladding using RODEX2 correlations and methods. To ount for the change in corrosion rate due to the changing oxide layer and thermal conditions, code includes an update in cladding temperature at every calculation step. This is an iterative cess due to the continuously changing oxide thickness. Conditions are also revised at times 3.3-5 Rev. 35

mblies in seven separate reactors. Each data point represents the maximum thickness sured along a rod length. The enhancement factor is based on a best fit regression analysis of data. A final multiplier is also applied which envelopes the data. The waterside corrosion in cladding was evaluated with RODEX2 for the steady state strain analysis. A best-fit corrosion lification factor was applied to the MATPRO model along with a final multiplier to bound the sured data on AREVA standard cladding. The maximum calculated oxide thickness was w the design limit.

l rod and fuel assembly growth projected to occur during irradiation was based on servative design curves established from measured irradiation growth data. The rod growth us the assembly growth plus tolerances was compared with the clearance within the assembly fuel rod growth. Differential thermal expansion between the fuel rods and guide tubes was considered. There is space between the upper and lower tie plates to accommodate the imum differential growth out to a rod burnup of 62,000 MWd/MTU.

pellet centerline temperature calculation was performed with the RODEX2 code. Fuel pellet terline temperatures were calculated at overpower conditions. The high power cycle of each er history was modified to include a spike in each cycle. This spike increased the maximum er of a pellet in the rod up to FTQ. Pellet melting temperature is a function of burnup.

sidering a conservative peak pellet burnup to determine the minimum pellet melting perature at EOL, the maximum pellet centerline temperature is well below both BOL and L limits.

1.2 (Deleted) 1.3 AREVA Fuel Assembly 1.3.1 Design Summary AREVA fuel assemblies are 14 by 14 arrays containing 176 fuel rods in a cage structure of 5 de tubes and 9 spacer grids. Both the guide tubes and the fuel rod cladding are made of aloy-4 for low neutron absorption and high corrosion resistance. The fuel assembly upper tie es are stainless steel castings with Inconel holddown springs. The fuel assembly upper tie e is mechanically locked to the guide tubes and may be easily removed to allow inspection of diated fuel rods. For Reload T (Cycle 15) and beyond, lower tie plates are the ELGUARD' debris resistant design.

eloads M, N, and P (Cycles 10-12), eight of the nine spacers in each fuel assembly are made Zircaloy-4 structure with Inconel-718 springs (i.e., bi-metallic spacer). The ninth spacer, ted just above the lower tie plate, is made of Inconel-718 and, using features of the AREVA h Thermal Performance (HTP) spacer design, has been adapted to provide fuel assembly ris resistance.

3.3-6 Rev. 35

ger end cap serves to raise the fuel rod cladding above the debris trapping region of the ninth tom) spacer.

Reloads T through X (Cycles 15-18), the High Thermal Performance (HTP) fuel assembly gn was implemented in which all nine spacers are of the Zircaloy-4 HTP design. This design ined the longer, solid fuel rod lower end cap.

fuel assembly design for Reload Y (Cycle 19) and later utilized eight Zircaloy-4 HTP spacers replaces the ninth, bottom spacer with an Inconel High Mechanical Performance (HMP) cer. The HMP spacer is similar to the HTP spacer, except that it is constructed of Inconel-718 the flow channels are parallel to the fuel.

wings of the AREVA fuel assemblies are given in Figure 3.3-2A and Figure 3.3-3A. Fuel mbly drawings for Reload T (Cycle 15) and beyond are included in Figures 3.3-2B and 3.3-analysis has shown that the AREVA reload fuel assemblies will meet the design criteria:

a. The maximum steady state cladding strain is well below the design limit.
b. The maximum steady state cladding stress meets the ASME Boiler and Pressure Vessel Code requirements.
c. The transient strain is within the circumferential limit.
d. Cladding creep collapse is precluded.
e. The fuel rod internal pressure at the EOL remains below the criteria approved by the NRC (Ref. 3.3-3).
f. The maximum clad oxidation is below the design limit.
g. The cladding fatigue usage factor is well below the design limit.
h. There is space between the upper and lower tie plate to accommodate fuel rod growth.
i. Pellet centerline temperatures remain below the design criteria.
j. The fuel assembly growth is within the space available between the upper and lower core plates in the reactor core.
k. The assembly holddown springs will prevent bundle lift-off.

3.3-7 Rev. 35

aloy-4 tubular cladding. Zircaloy-4 end caps are welded to each end to give a hermetic seal.

fuel rod upper plenum contains a high strength alloy compression spring to prevent fuel mn separation during fabrication and shipping, and during incore operation. The rods are surized with helium to improve heat transfer and reduce clad creep ovality.

fuel assembly structure consists of an upper tie plate assembly, lower tie plate, guide tubes spacer grids, which together provide the support for the fuel rods.

lower tie plate is a machined stainless steel casting which provides the lower end support for guide tubes. The Zircaloy guide tubes are attached to the lower tie plate by means of Inconel screws. The FUELGUARDTM lower tie plate, included in Reload T and beyond provides ection to the fuel from debris in the primary coolant.

upper tie plate assembly latches to and provides the upper end support for the guide tubes.

upper tie plate assembly consists of a stainless steel grid structure and reaction plate taining five Inconel X-750 holddown springs. The springs are located around Inconel locking and sleeves which mechanically attach to the guide tubes and pilot into the reactor alignment

e. The springs are partially shrouded on the outside diameter by stainless steel cups to prevent induced spring vibration.

guide tubes, in conjunction with the spacers and tie plates, form the structural skeleton of the assembly and provide channels for insertion of the control rods. The guide tubes are icated from Zircaloy-4 tubing and are fully annealed. The center tube is of uniform diameter reas the outer four guide tubes have a reduced diameter section at the bottom which produces shpot action to decelerate the dropped CEAs.

end plug is welded to the lower end of the guide tube and is drilled and threaded to accept the er cap screws. At the upper end, the guide tube is crimped into an external stainless steel ing sleeve which engages the upper tie plate assembly. The upper tie plate assembly is locked he guide tube end fittings and can be unlocked for reconstitution or for fuel examination using cial tools.

ainless steel sleeve assembly with a chrome plated inside diameter is inserted in the top end of guide tube assembly. This sleeve protects the guide tube from control rod fretting and wear n the rod is in the withdrawn/ready position. The sleeve is mechanically captured by the upper late.

l rod pitch and position is maintained by nine spacer grids. The spacers are axially positioned hat the assemblies will be compatible with existing fuel assemblies.

bi-metallic spacers used in Reloads M through S (Cycles 10-14) are formed by an rlocking rectangular grid of Zircaloy-4 structural strips (see Figure 3.3-4A). Inconel-718 ng strips are mechanically secured within these strips. The Zircaloy-4 structural strips are ded at all intersections and to the side plates. Dimples formed in the structural strips center the 3.3-8 Rev. 35

eloads M, N, and P (Cycles 10-12), the debris resistant Inconel HTP spacer grid in the ninth, om location is located just above the lower tie plate. It is formed by an interlocking angular grid of Inconel-718 strips. The strips are welded at all intersections and to the side es. The spacer is positioned on top of the lower tie plate with the strip intersections directly ve the tie plate flow holes. This reduces the size of debris that may pass through the flow holes eby reducing the possibility of fretting against the cladding. Reloads R and S (Cycles 13 and use a similar debris resistant concept with the Inconel HTP spacer replaced by a bimetallic cer coupled with a longer lower end cap on the fuel rods.

HTP spacers for Reloads T through X (Cycles 15-18) are all Zircaloy-4 (Figure 3.3-4B). The ps are welded at the intersections and side plates. The structure of the Zircaloy-4 strips vides the rod support.

Reload Y (Cycle 19) and later, all Zircaloy-4 HTP spacers are used in eight locations. The onel-718 HMP spacer is used in the ninth, bottom location. The Inconel-718 HMP bottom cer is similar in design to the HTP spacers except for the flow channels, which are not canted.

Millstone Unit 2 Cycle 19 reload core included four fuel assemblies with natural uranium acement fuel rods with an anti-rotation feature designed to prevent spinning of the rod during rations. The four assemblies containing replacement rods were installed in symmetric, pheral core locations against the baffle as shown in Figure 3.3-19 (Reference 3.3-9). The core tions into which the assemblies were placed where P-1, A-8, H-21, and Y-14 (see ure 3.4-1). The replacement rods installed under these conditions were evaluated against blished mechanical, nuclear, and thermal/hydraulic design criteria for Millstone Unit 2 fuel, were determined to be compliant with their design and licensing bases (Reference 3.3-10).

1.3.2 Fuel Assembly Mechanical Criteria structural integrity of the fuel assemblies is assured by setting design limits on stresses and ormations due to various handling operational and accident loads. These limits are applied to design and evaluation of upper and lower tie plates, grid spacers, guide tubes, holddown ngs, and locking hardware.

design bases for evaluating the structural integrity of the fuel assemblies are:

a. Fuel Assembly Handling - Dynamic axial loads approximately 2.5 times assembly weight.
b. For All Applied Loads for Normal Operation and Anticipated Operational Events -

The fuel assembly component structural design criteria are established for the two primary material categories, austenitic stainless steels (tie plates), and Zircaloy (guide tubes, grids, spacer sleeves). The stress categories and strength theory for 3.3-9 Rev. 35

Steady state stress limits are given in FSAR Section 3.3.1.1.1. Stress nomenclature is per the ASME Boiler and Pressure Vessel Code,Section III.

c. Loads During Postulated Accidents - Deflection or failure of components shall not interfere with reactor shutdown or emergency cooling of the fuel rods during postulated seismic and loss of coolant accident (LOCA) occurrences.

The assembly structural component stresses under faulted conditions are evaluated using primarily the methods outlined in Appendix F of the ASME Boiler and Pressure Vessel Code,Section III.

design basis for the guide tube wear sleeves is that the fuel assembly shall not be damaged by A induced fretting-wear. Flow tests at reactor conditions of prototypic fuel and guide tube r sleeve assemblies have been used in establishing the performance of the CEA wear sleeve bination.

holddown springs, as compressed by the upper core plate during reactor operation, shall vide a net positive downward force during steady state operation, based on the most adverse bination of component dimensional and material property tolerances. In addition, the ddown springs are designed to accommodate the additional load associated with a pump rspeed transient (resulting in possible temporary liftoff of the fuel assemblies), and to continue nsure fuel assembly holddown following such an occurrence.

fuel assembly growth plus BOL length shall not exceed the minimum space between the er and lower core plates in the reactor cold condition (70°F). The reactor cold condition is ting since the expansion coefficient of the stainless steel core barrel is greater than the fficient of expansion of the Zircaloy guide tubes.

spacer assembly is designed to withstand the thermal and irradiation induced differential ansion between the fuel rods and guide tubes and to withstand the design handling and dent loads discussed above. The debris resistant Inconel-718 HTP spacer used in the ninth, om location for reloads M, N and P (Cycles 10-12) was positioned such that the internal strip rsections are directly above the lower tie plate flow holes, thus reducing the size of debris ch could pass through the lower tie plate.

Reloads R and S (Cycles 13 and 14), the Inconel-718 HTP spacer grid at the ninth, bottom tion was replaced with a bimetallic spacer which is raised off the upper surface of the lower plate. The gap between the upper surface of the lower tie plate and the lower surface of the etallic spacer is spanned by a long fuel rod end cap of solid Zircaloy-4.

Zircaloy-4 HTP spacer grid is used in all nine locations in Reloads T through X (Cycles 15-This design is typically referred to as the HTP Fuel Assembly. This spacer grid design 3.3-10 Rev. 35

eload Y (Cycle 19) and later, the Zircaloy-4 HTP spacer grid is used in eight locations and an onel HMP spacer grid is used in the ninth, bottom location. This design retains the long fuel lower end cap and is typically referred to as the HTP+HMP Fuel Assembly. The HTP+HMP gn has improved structural strength, and fretting resistance compared to the HTP design.

design analysis is based upon reactor operating conditions. Typically, these conditions are:

Nominal Core Thermal Power = 2700 MW Nominal Coolant Pressure = 2250 psia Maximum Flow for Fuel Assembly Liftoff = 422,466 gallons per minute (at 480°F)

Maximum Core Coolant Inlet Temperature at Nominal Power = 549°F Total Average Linear Power = 6.206 kW/ft power histories used in the design analysis are designed to achieve a peak assembly burnup 6,000 MWD/MTU for Batch N, 52,500 MDW/MTU for Batch P, and 57,400 MDW/MTU for ch R and later.

servative rod local peaking factors are used which result in a peak rod burnup of 62,000 d/MTU. Each of the rod design histories follows the single hottest rod in the first cycle ration, the hottest rod in second cycle operation, etc.

l assembly components must be able to withstand anticipated seismic and LOCA forces.

se may result from bundle vibration and impact due to a seismic or LOCA event. An analysis performed for the previous reloads to determine the maximum bundle displacements and the imum spacer grid forces expected during postulated accidents for Millstone 2. The loads and lacements analysis, which was performed by CE (Reference 3.3-6), considered the safe tdown earthquake (SSE) and limiting Branch Line LOCA events, and the dynamic properties he AREVA reload fuel assemblies. The resulting fuel assembly displacements and the bined seismic and LOCA grid spacer impact forces were provided to AREVA.

loads and displacements were conservatively adjusted for the Batch R design due to the mization of the fuel rod. The fuel weight was increased and the assembly stiffness was reased. The spacer impact loads and the fuel assembly maximum deflections were servatively recalculated from the reference analysis values. The spacer strength margin, the de tube stresses, and the fuel rod stresses were calculated for the adjusted loads.

culated stresses at the appropriate deflections were combined with the steady state stresses and pared with the ASME Design Criteria for faulted conditions. This limit is 0.7 times the mate strength for the primary stress combinations as compared to 0.5 times ultimate for steady e loadings. This criteria was met for both the fuel rods and the guide tubes.

3.3-11 Rev. 35

imum projected one-sided impact load and the maximum through grid load. The maximum wable crushing load is the 95 percent confidence lower limit of the true mean of the ribution of crush test measurements. The allowable through grid strength is well above the imum through grid load. It is also above the maximum one-sided impact load. For Reload R beyond, the seismic/LOCA calculations were reviewed and determined to be bounding.

1.4 Fuel Assembly Holddown Device uel assembly holddown device has been incorporated to prevent the possibility of lifting the assembly by hydraulic forces under all normal flow conditions with temperature greater than

°F. The holddown device consists of a spring-loaded plate which is integral to the fuel mbly. The springs are compressed as the upper guide structure is lowered into place. The ed spring load, together with the weight of the fuel assembly, prevents possible axial motion he fuel assembly during operating conditions.

holddown device is incorporated into the upper end fitting and features a movable holddown e which acts on the underside of the fuel alignment plate (refer to Figure 3.3-5). The movable e is loaded by coil springs which are located around the upper end fitting posts. The springs positioned at the upper end of the assembly so that the spring load combines with the mbly weight in counteracting the upward hydraulic forces. The springs are sized and the ng preload selected, such that a net downward force will be maintained for all normal and cipated transient flow and temperature conditions. It should be noted that the movable plate serves as the lifting surface during handling of the fuel assembly.

embly holddown was previously analyzed in Section 3.6.1 of Reference 3.3-8. The analysis been reperformed for Batch T and beyond fuel and is conservative.

1.5 Control Element Assembly As are provided by Combustion Engineering (CE) and AREVA. The CEA (shown in ure 3.3-6) is comprised of five Inconel tubes 0.948 inch in diameter. All tubes contain neutron on materials with the distribution of the poison materials as depicted in Figure 3.3-7. Each is sealed by welded end caps. A gas expansion space is provided to limit maximum tube ss due to internal pressure developed by the release of gas and moisture from the boron ide. The overall length of the CEA is provided in Table 3.3-1. Four tubes are assembled in a are array around the centrally located fifth tube. The tubes are welded to an upper end fitting.

upper end fittings are attached to a spider hub which couples the CEA to the drive mechanism ugh the extension shaft.

chanical reactivity control is achieved by operational maneuvering of groups of single CEAs.

dual CEA is made up of two single CEAs connected to separate grippers attached to single nsion shaft. The arrangement of the CEAs in the core is shown in Figures 3.3-8 and 3.3-9.

3.3-12 Rev. 35

uffer (deceleration dashpot) system is used for slowing down the CEAs at the end of a reactor

. The buffering action is accomplished by guide tubes which have a reduced diameter in the er section. When the tip of a CEA falls into the buffer region, the pressure buildup in the lower de tube supplies the force to slow down the CEA. The velocity is decreased to a level which minimize impact. The final impact is further cushioned by a coil spring arrangement mounted und the center CEA finger.

four outer guide tubes have the reduced diameter lower section (dashpot). There is no dashpot he center guide tube. There are four bleed holes above the dashpot region for the four outer de tubes. For the center guide tube, these four bleed holes are at a lower elevation. For all de tubes, there is a small drain hole at the bottom. The CEA tip is filled with a Silver-Indium-mium alloy. This replaces the B4C to avoid the change of buffer characteristics that B4C ation-induced swelling might bring about.

design parameters have been optimized to establish the best combination of buffer stroke and er annulus. A significant analytical effort has shown that the pressure buildup and the impact s are not damaging to the system. In addition, a test program has confirmed the feasibility of system. It has demonstrated that the buffer will work under the worst expected tolerance dition.

1.6 Neutron Source Design Cycle 18 and beyond, the reactor core will not utilize neutron sources. It has been determined during startups without neutron sources, there will continue to be a sufficient neutron count at each of the four Wide Range (WR) Excore fission detectors due to the high burnup fuel mblies that will be positioned on the core periphery.

Cycle 17 and earlier, four neutron sources were installed in the reactor core. They were held acant CEA guide tubes by means of an externally loaded spring reacting between the upper alignment plate and the top of the fuel assembly. The cladding of the neutron source rods is of ee standing design. The internal pressure is always less than reactor operating pressure.

rnal gaps and clearances are provided to allow for differential expansion between the source erial and cladding.

1.7 In-Core Instruments in-core instruments (refer to Section 7.5.4) are located in the in-core instrumentation mbly (Figure 3.3-10). The in-core instrumentated thimble support frame and guide tubes are ported by the upper guide structure (UGS) assembly. The tubes are conduits which protect the ore instruments and guide them during removal and insertion operations. The thimble support e supports the 43 in-core thimble assemblies and acts as an elevator to lift the thimbles from core into the UGS during the refueling operation.

3.3-13 Rev. 35

heated junction thermocouple (HJTC) system is composed of two channels of HJTC ruments. Each HJTC instrument channel is manufactured into a probe assembly consisting of t HJTC sensors, a seal plug, and electrical connectors (Figure 7.5-6). The eight HJTC sensors physically independent and located at eight levels from the reactor vessel head to the fuel nment plate.

probe assembly is housed in a stainless steel support tube structure that protects the sensors m flow loads and serves as the guide path for the sensors. Figure 3.3-18 describes the locations he HJTC probe assemblies.

C Probes and Support Tubes in Upper Guide Structure HJTC probes and support tubes are installed inside two-part length CEA shrouds which ect the support tubes from normal operating cross-flow loads as well as blowdown loads. The port tubes are latched to the bottom of the CEA shroud and permanently tensioned by means threaded spanner nut at the top. Operating loads are far less than the preload developed by the ioning operation. Therefore, the support tubes will not be affected by thermal or flow loads.

support tubes are designed to account for all tolerance conditions so that proper clearances be assured. Physically, the support tubes are similar in mass and size to a typical control ment assembly drive shaft, which would reside in the same area of the upper guide structure.

presence or absence of the HJTC probes within the support tubes will in no way affect the grity of the support tubes, the UGS, the pressure boundary, and will have no significant effect n the hydraulic conditions within the reactor vessel head.

2 REACTOR INTERNAL STRUCTURES reactor internals are designed to support and orient the reactor core fuel assemblies and As, absorb the CEA dynamic loads and transmit these and other loads to the reactor vessel ge, provide flow paths for the reactor coolant, and guide in-core instrumentation.

internals are designed to safely perform their function during all steady state conditions and ng normal operating transients. The internals are designed to safely withstand the forces due eadweight, handling, system pressure, flow impingement, temperature differential, vibration seismic acceleration. All reactor components are considered Class 1 for seismic design. The tor internals design limits deflection where required by function. In most cases the design of tor internals components is limited by stress, not deflection. For the CEA shroud which is the t limiting internal component for deflection, the allowable design deflection limit is 0.5 inch.

s limit is two-thirds of the conservatively established loss-of-function deformation limit, 0.75 and applies to a break whose equivalent diameter is no larger than the largest line connected he primary coolant line. The structural components satisfy stress values given in Section III of ASME Pressure Vessel Code. Certain components have been subjected to a fatigue analysis.

ere appropriate, the effect of neutron irradiation on the materials concerned is included in the gn evaluation.

3.3-14 Rev. 35

CEA shrouds, the in-core instrumentation guide tubes and the HJTC support tubes). The flow t, although functioning as an integral part of the coolant flow path is separate from the rnals and is affixed to the bottom head of the pressure vessel. These components are shown in ure 3.1-1 and 3.3-11. The in-core instrumentation is described in Section 7.5.4.

amic system analysis methods and procedures which have been used to determine dynamic onses of reactor internals have been provided in CE, Report CENPD-42, Topical Report of amic Analysis of Reactor Vessel Internals under Loss-of-Coolant Accident Conditions with lication of Analysis to CE 800 MWe Class Reactors.

2.1 Core Support Assembly major support member of the reactor internals is the core support assembly. This assembled cture consists of the core support barrel, the lower support structure, and the core shroud. The or materials for the assembly is Type 304 stainless steel.

core support assembly is supported at its upper end by the upper flange of the core support el which rests on a ledge in the reactor vessel flange.

lower flange of the core support barrel supports and positions the lower support structure.

lower support structure provides support for the core by means of a core support plate ported by columns resting on beam assemblies. The core support plate provides support and ntation for the fuel assemblies. The core shroud which provides lateral support for the fuel mblies is also supported by the core support plate. The lower end attaches the core barrel to pressure vessel.

2.2 Core Support Barrel core support barrel is a right circular cylinder with a nominal inside diameter of 148 inches a minimum wall thickness of 1.75 inch. It is suspended by a 4 inch thick flange from a ledge he pressure vessel. The core support barrel, in turn, supports the lower support structure upon ch the fuel assemblies rest. Press fitted into the flange of the core support barrel are four nment keys located 90 degrees apart. The reactor vessel, closure head and upper guide cture assembly flanges are slotted in locations corresponding to the alignment key locations to vide proper alignment between these components in the vessel flange region.

ce the core support barrel is over 27 feet long and is supported only at its upper end, it is sible that coolant flow could induce vibrations in the structure. Therefore, amplitude limiting ices, or snubbers are installed on the outside of the core support barrel near the bottom end.

snubbers consist of six equally spaced double lugs around the circumference and are the oves of a tongue-and groove assembly; the pressure vessel lugs are the tongues. Minimizing clearance between the two mating pieces limits the amplitude of any vibration. During mbly, as the internals are lowered into the vessel, the pressure vessel tongues engage the core port grooves in an axial direction. With this design, the internals may be viewed as a beam 3.3-15 Rev. 35

sure vessel tongues have bolted, lock welded Inconel X shims and the core support barrel oves are hardfaced with Stellite to minimize wear. The snubber assembly is shown in ure 3.3-12.

2.3 Core Support Plate and Support Columns core support plate is a 147 inch diameter, 2 inch thick, Type 304 stainless steel plate into ch the necessary flow distributor holes for the fuel assemblies have been machined. Fuel mbly locating pins (four for each assembly) are shrunk-fit into this plate. Columns and port beams are located between this plate and the bottom of the core support barrel in order to vide support for this plate and transmit the core load to the bottom flange of the core support el.

2.4 Core Shroud core shroud provides an envelope for the core and limits the amount of coolant bypass flow.

shroud (Figure 3.3-13) consists of two Type 304 stainless steel ring sections, aligned by ns of radial shear pins and attached to the core support plate by Type 348 stainless steel tie

s. A gap is maintained between the core shroud outer perimeter and the core support barrel in er to provide some coolant flow upward between the core shroud and core support barrel, eby minimizing thermal stresses in the core shroud and eliminating stagnant pockets.

2.5 Flow Skirt Inconel flow skirt is a right circular cylinder, perforated with 2-11/16 inch diameter holes, reinforced at the top and bottom with stiffening rings. The flow skirt is used to reduce ualities in core inlet flow distributions and to prevent formation of large vortices in the lower um. The skirt provides a nearly equalized pressure distribution across the bottom of the core port barrel. The skirt is supported by nine equally spaced machined sections which are welded he bottom of the pressure vessel.

2.6 Upper Guide Structure Assembly s assembly (Figure 3.3-14) consists of the upper support structure, 69 CEA shrouds, a fuel mbly alignment plate and an expansion compensating ring. The UGS assembly aligns and rally supports the upper end of the fuel assemblies, maintains the CEA spacing, prevents fuel mblies from being lifted out of position during a severe accident condition and protects the As from the effect of coolant crossflow in the upper plenum. The UGS is handled as one unit ng installation and refueling.

upper end of the assembly is a structure consisting of a support plate welded to a grid array of nch deep beams and a 24 inch deep cylinder which encloses and is welded to the ends of the ms. The periphery of the plate contains four accurately machined and located alignment ways, equally spaced at 90 degree intervals, which engage the core barrel alignment keys. The 3.3-16 Rev. 35

ure head. The grid aligns and supports the upper end of CEA shrouds.

CEA shrouds extend from the fuel assembly alignment plate to an elevation about three feet ve the UGS support plate. There are 57 single-type shrouds. These consist of cylindrical upper ions welded to integral bottom sections, which are shaped to provide flow passages for the lant passing through the alignment plate while shrouding the CEAs from cross-flow. There are 12 dual-type shrouds which in configuration consist of two single-type shrouds connected by ctangular section shaped to accommodate the dual CEAs. The shrouds are bolted to the fuel mbly alignment plate. At the UGS support plate, the single shrouds are connected to the plate spanner nuts which permit axial adjustment. The spanner nuts are tightened to proper torque lockwelded. The dual shrouds are attached to the upper plate by welding.

fuel assembly alignment plate is designed to align the upper ends of the fuel assemblies and upport and align the lower ends of the CEA shrouds.

cision machined and located holes in the fuel assembly alignment plate align the fuel mblies. The fuel assembly alignment plate also has four equally spaced slots on its outer edge ch engage with Stellite hardfaced pins protruding from the core shroud to limit lateral motion he UGS assembly during operation. The fuel alignment plate bears the upward force of the assembly holddown devices. This force is transmitted from the alignment plate through the A shrouds to the UGS support plate and hence to the expansion compensating ring.

expansion compensating ring bears on the flange at the top of the assembly to accommodate l differential thermal expansion between the core barrel flange, UGS flange and pressure sel flange support edge and head flange recess.

UGS assembly also supports the in-core instrumentation thimble support frame, guide tubes, HJTC support tubes.

integral connections in the reactor internals are designed within the stress intensity limits d in Tables N-422 and N-416.1 of Section III of the ASME code for normal and upset ditions. For emergency and faulted conditions, the design limits are as given in Table 3.2-1.

3 CONTROL ELEMENT DRIVE MECHANISM 3.1 Design CEDM is of the magnetic jack type drive. Each CEDM is capable of withdrawing, inserting, ding or tripping the CEA from any point within its 137-inch stroke. The design of the CEDM hown in Figure 3.3-15 and is identical to that for Maine Yankee (AEC Docket Number 50-

) and Calvert Cliffs Units 1 and 2 (AEC Docket Numbers. 50-317 and 50-318).

CEDM drives the CEA within the reactor core and indicates the position of the CEA with ect to the core. The speed at which the CEA is inserted or withdrawn from the core is 3.3-17 Rev. 35

nergized, allowing the CEA and the supporting CEDM components to drop into the core by vity. The CEA drop time is 2.75 seconds, where drop time is defined as the interval between time power is removed from the CEDM coils and the time the CEA has reached 90 percent of fully inserted position. The reactivity is reduced during such a drop at a rate sufficient to trol the core under any operating transient or accident condition. The CEA accelerates to about t/sec and is decelerated at the end of the drop by the buffer section of the CEA guide tubes.

ve down capability following a reactor trip is not required for safety purposes. The safety lyses of Chapter 14 assume the CEA of highest reactivity worth sticks in the fully withdrawn ition. A drive down feature would introduce the possibility of a failure which would prevent er from being removed from the CEDMs during a trip, which would lead to a reduction in t safety.

re are 69 CEDM nozzles on top of the reactor vessel closure head. Eight of the 69 nozzles e used for the part length CEAs in Cycle 1, six of which are no longer used, and two of which used for HJTC/RVLMS instrumentation. There are 61 CEDMs in current use. The six spare zles are capped with adapters. Each CEDM is connected to a CEA by a locked coupling. The ght of the CEAs and CEDMs is carried by the vessel head.

CEDM is designed to handle dual, single or part length CEAs. The maximum operating ed capability of the CEDMs is 40 inches per minute for single CEAs and 20 inches per minute dual CEAs.

3.2 Control Element Drive Mechanism Pressure Housing CEDM housing is attached to the reactor vessel head nozzle by means of a threaded joint and welded. The CEDM nozzles are made of Inconel Alloy 690 to minimize Primary Water Stress rosion Cracking. The CEDM pressure housings including the magnetic coil jack assemblies e replaced as part of the replacement reactor vessel closure head project.

CEDM upper housing design and fabrication conform to the requirements of the ASME ler and Pressure Vessel Code,Section III, 1998 Edition through 2000 Addenda. The housing is gned for steady state conditions as well as all anticipated pressure and thermal transients.

ce the CEDM housing is seal welded to the head nozzle, it need not be removed since all icing of the CEDM is performed from the top of the CEDM housing. This opening is closed means of an upper housing and an omega seal weld. The CEDM pressure housing is capable of g vented after major coolant refills of the reactor coolant system (RCS), such as after a eling and after reactor coolant pump (RCP) maintenance. However, venting of the CEDM sure housing is no longer necessary after major refills of the Reactor Coolant System (RCS),

e a vacuum refill method is used. The vacuum refill process involves a partial vacuum in the S while at mid-loop level and then slowly refilling the RCS.

3.3-18 Rev. 35

HJTC probe assemblies are located at the two original locations (CEDMs 11 and 13) on the acement reactor vessel closure head. The HJTC pressure boundary also known as the Reactor sel Level Monitoring System (RVLMS) pressure housing assembly consists of upper pressure sing tube, upper flange type Grayloc connection and lower housing. The lower housing is ed to the reactor vessel head nozzle by means of a threaded joint and an omega seal weld. The sure boundary at the top of the RVLMS pressure housing is maintained by a quick disconnect yloc type flange (See Figure 3.3-17). The components are designed to ASME Section III, PV Code 1998 Edition through 2000 Addenda.

pressure and thermal loads associated with normal operation and transient conditions have n included in stress analyses performed in accordance with ASME BPVC criteria. All stresses within allowable limits.

3.3 Magnetic Jack Assembly magnetic jack motor assembly is an integral unit which fits into the CEDM housing through opening in the top of the housing. This unit carries the motor tube, lift and hold pawls and nets. The drive power is supplied by electrical coils positioned around the CEDM housing.

CEDMs are cooled by air supplied at 900 CFM at 95°F (maximum) to each CEDM. The gn of the control element drive mechanism is such that loss of cooling air will not prevent the DM from releasing the CEA. The ability of the CEDM to release the rods is not dependent on cooling flow provided by the CEDM Cooling System. Cooling function is only to ensure ability of the CEDM coil stack. Following insertion of the CEDM motor assembly, the upper sure housing is threaded into the CEDM motor housing and seal welded. This upper pressure sing encloses the CEDM extension shaft and supports the shroud assembly. The reed switch mbly is supported by the shroud assembly.

lifting operation consists of magnetically operated step movements. Two sets of mechanical hes (one holding, one lifting) are utilized engaging a notched drive shaft. To prevent excessive h wear, a means has been provided to unload the lifting latches during the engaging and ngaging operations.

magnetic force is obtained from large DC magnet coils mounted on the outside of the motor er for the electromagnets is obtained from one of two separate supplies. A control grammer actuates the stepping cycle and obtains the CEA location by a forward or reverse ping sequence. CEDM hold for shutdown and regulating CEAs is obtained by energizing a d coil at a reduced current while all other coils are deenergized. The full length CEAs are ped upon interruption of electrical power to all coils.

3.4 Position Indication ee separate means are provided for transmitting CEA position indication.

3.3-19 Rev. 35

vide an output voltage proportional to CEA position. The third method utilizes three pairs of switches spaced at discrete locations within a position transmitter assembly. A permanent net built into the drive shaft actuates the reed switches one at a time as it passes by them. CEA ition instrumentation is discussed in detail in Section 7.5.3.

3.5 Control Element Assembly Disconnect CEA is connected to the drive shaft extension with an internal collet-type coupling at its er end. (Coupling is performed before the vessel head is installed). In order to disengage the A from the drive shaft extension, a tool is attached to the top end of the drive shaft when the tor vessel head has been removed.

pulling up on the spring-loaded operating rod in the center of the drive shaft, a tapered plunger ithdrawn from the center of the collet-type gripper causing it to collapse due to axial pressure m the CEA, thus permitting removal of the coupler from the CEA. Releasing the operating rod nger after the coupler has been withdrawn from the CEA expands the coupler to a diameter prevents recoupling to the CEA.

3.6 Test Program est program has been conducted to verify the adequacy of the magnetic jack CEDM. The gram is described in Section 1.5.4.

4 REFERENCES 1 ASME Boiler and Pressure Vessel Code,Section III, 1977 Edition, ASME New York, NY.

2 K. R. Merckx, RODEX2 - Fuel Rod Thermal-Mechanical Response Evaluation Model, XN-NF-81-58 (NP)(A), Revision 2, March 1985 and Supplements.

3 Qualification of Exxon Nuclear Fuel for Extended Burnup (PWR), XN-NF-82-06 (NP)(A), Revision 1, Supplements 2, 4, 5, October 1985.

4 W. J. O'Donnel and B. F. Langer, Fatigue Design Bases for Zircaloy Components, Nuclear Science and Engineering, Volume 20, January 1964.

5 MATPRO Version, A Handbook of Material Properties for Use in the Analysis of Light Water Reactor Fuel Rod Behavior, TREE-NUREG 1008, December 1976.

6 J. C. Winslow (CE) to T. J. Honan (NU), CE Letter, Seismic and Branch Line LOCA Analysis of SPC Reload Fuel for Millstone 2, NU-88-043 (March 31, 1988).

3.3-20 Rev. 35

8 ANF-88-88(P), Rev. 1, Design Report for Millstone Point Unit 2 Reload ANF-1, August 29, 1988.

9 AREVA Contract Requirements Document Number 89-9070921-001-AREVA Contract No. J37MIL219B, January 28, 2008.

10 AREVA Document 51-9074000-000, Compliance Document - Replacement Fuel Rod -

Millstone 2 Fuel Failure Mitigation, March 5, 2008.

3.3-21 Rev. 35

TABLE 3.3-1 MECHANICAL DESIGN PARAMETERS

  • l Assembly Geometry 14 by 14 Assembly Pitch, inches 8.180 Assembly Envelope, inches 8.160 Rod Pitch, inches 0.580 Number of Grids per Assembly 9 Approximate Assembly Weight, lb. 1280/1313
  • Fuel Rod to Fuel Rod Outside Dimension, inches 7.980 l Rod and Pellet Clad OD, inches 0.440 Clad thickness, inches 0.031/0.028
  • Pellet Diameter, inches 0.3700/0.3770
  • Pellet Length, inches 0.425/0.435
  • Pellet Density (% Theoretical) 94.0/95.0/95.35 **

Active Stack Length, Cold, inches 136.7 trol Rod Guide Tube Number per assembly 4 Tube ID, above dashpot, inches 1.035 Wall Thickness, inches 0.040 rumentation Tube Number per Assembly 1 Tube ID, inches 1.035 Wall Thickness, inches 0.040 cer Grid Material Zircaloy-4 / Inconel-718 3.3-22 Rev. 35

9/0 for Batch R, S 9/0 for Batch T - X 8/1 Batch Y and beyond ves (Wear)

Material SS/Chrome Plate nable Poison Rod Active Length, inches 124.7 + UO2 blankets Material Gd2O3 / U02 Pellet Diameter, inches 0.3700/0.3770 Clad Material Zircaloy-4 Clad ID, inches 0.378/0.384 Clad OD, inches 0.440 Clad Thickness, (nominal) inches 0.031/0.028 Diametral Gap, (cold, nominal), inches 0.008/0.007 Pellet Length, inches 0.545 trol Element Assembly Number 73 Number of Absorber Elements per Assembly 5 Type Cylindrical Rods Clad Material Inconel 625 Clad Thickness, inches 0.036 Clad OD, inches 0.948 Poison Material B4C & Ag-In-CD Corner Element Pitch, inches 4.64 Total CEA Length, inches 161.31- CE / 161.25 - AREVA Poison Length, inches 132 -CE / 133.5 - AREVA CEA Dry Weight, lb. 95 - CE / 85 - AREVA 3.3-23 Rev. 35

Single 210 - CE / 200 - AREVA Dual 334 - CE / 314 AREVA e Arrangement Number of Fuel Assemblies in Core Total 217 Number of Single CEAs 49 Number of Dual CEAs 12 CEA Pitch, minimum, inches 11.57 Spacing Between Fuel Assemblies, Fuel Rod Surface to Surface, inches 0.200 Spacing, Outer Fuel Rod Surface to Core Shroud, inches 0.18 Hydraulic Diameter, Nominal Channel, feet 0.04445 Total Flow Area (Excluding Guide Tubes), square feet 53.5 Total Core Area, square feet 101.1 Core Equivalent Diameter, inches 136 Core Circumscribed Diameter, inches 143.1 Core Volume, liters 32,526 Total Fuel Loading, MTU (Typical) 83.65 Total Heat Transfer Area, square feet 50,117 Applicable to Batches N, P/applicable to Batch R and subsequent Batches.

Applicable to Batches N, P/applicable to Batches R, S/applicable to Batch T and subsequent Batches.

3.3-24 Rev. 35

ABLE 3.3-2 PRESSURIZED WATER REACTOR PRIMARY COOLANT WATER CHEMISTRY RECOMMENDED SPECIFICATIONS ductivity (S/cm at 25°C) Relative to Lithium and Boron concentration.

at 25°C Determined by the concentration of boric acid and lithium present. Consistent with the Primary Chemistry Control Program.(4) solved Oxygen, at power < 0.1 ppm (1) (2) (3) oride < 0.15 ppm oride < 0.10 ppm rogen 25-50 cc (STP)/KgH2O pended Solids 0.35 ppm prior to reactor startup Consistent with the Primary Chemistry Control Program.(4) on, as boric acid 0-2620 ppm (5)

TES:

) The temperature at which the Oxygen limit applies is > 250°F.

) The at power operation residual Oxygen concentration control value is 0.005 ppm.

) During plant startup, Hydrazine may be used to control dissolved Oxygen concentration at 0.1 ppm.

) During power operation lithium is coordinated with boron to maintain a pH(t) of 7.0, but 7.4, consistent with the Primary Chemistry Control Program. Lithium is added to the RCS during plant startup, but prior to reactor criticality, and is in specification per the Primary Chemistry Control Program within 72 hours8.333333e-4 days <br />0.02 hours <br />1.190476e-4 weeks <br />2.7396e-5 months <br /> after criticality. Lithium may be removed from the reactor coolant immediately before, or during, shutdown periods to aid in the cleanup of corrosion products. By evaluation, a maximum lithium concentration of 4.5 ppm is permissible with a target lithium concentration of 4.3 ppm for 100% power operations.

) RCS boron concentration is maintained as necessary to ensure core reactivity or shutdown margin requirements are met. Although the RCS and related auxiliary systems containing reactor coolant are designed for a maximum concentration of 2620 ppm boron, it should be noted the design basis for the TSP baskets in the containment sump assumes the RCS, SITs, and RWST are at a maximum boron concentration of 2400 ppm.

3.3-25 Rev. 35

MPS-2 FSAR FIGURE 3.3-1 FUEL ROD ASSEMBLY UPPER END CAP PLENUM SPRING DISHED PELLETS FUEL CLADDING 136.70 ACTIVE FUEL LENGTH 146.25

.440 CLADDING OD

.3770 PELLET DIAMETER

.028 CLADDING WALL 136.70 ACTIVE FUEL LENGTH April 1998 Rev. 24.8

MPS-2 FSAR FIGURE 3.3-2A AREVA - RELOAD FUEL ASSEMBLY BATCH "S" AND PRIOR 9X TYPICAL SPACER UPPER TIE PLATE ASSEMBLY LOWER TIE PLATE Rev. 24.8

MPS-2 FSAR FIGURE 3.3-2B AREVA - RELOAD FUEL ASSEMBLY BATCH "T" AND LATER UPPER TIE PLATE LOWER TIE PLATE Rev. 24.8

MPS-2 FSAR FIGURE 3.3-3A AREVA - RELOAD FUEL ASSEMBLY COMPONENTS BATCH "S" AND PRIOR 9X TYPICAL SPACER UPPER TIE PLATE ASSEMBLY LOWER TIE PLATE Rev. 24.8

MPS-2 FSAR FIGURE 3.3-3B AREVA - RELOAD FUEL ASSEMBLY COMPONENTS BATCH "T" AND LATER FUEL ROD 136.70 ACTIVE FUEL LENGTH UPPER TIE PLATE SPACER LOWER TIE PLATE Rev. 24.8

MPS-2 FSAR FIGURE 3.3-4A BI-METALLIC FUEL SPACER ASSEMBLY FUEL ROD GUIDE TUBE LOCATION SPACER SIDEPLATE SPRING STRIP Rev. 21

MPS-2 FSAR FIGURE 3.3-4B HTP FUEL SPACER ASSEMBLY GUIDE TUBE FUEL ROD Rev. 21

MPS-2 FSAR FIGURE 3.3-5 FUEL ASSEMBLY HOLD DOWN DEVICE FUEL ALIGNMENT PLATE LOCKING NUT UPPER SPRING REACTION PLATE UPPER TIE PLATE Rev. 26.2

MPS-2 FSAR FIGURE 3.3-6 CONTROL ELEMENT ASSEMBLY SPIDER IDENTIFICATION CEA Serial Number SPRING POISON ROD ASSEMBLY POISON MATERIAL 132 Inches Total - CE 133.5 Inches Total - AREVA Rev. 30.2

MPS-2 FSAR FIGURE 3.3-7 CONTROL ELEMENT ASSEMBLY MATERIALS

~ 8" - CE 0" ~ 12.5" - AREVA ZONE A ZONE A ZONE B MATERIALS NUMBER ZONE A ZONE B Ag Ag B B 73 Ag B Ag Ag B B B B4 C Ag Ag In Cd Rev. 30.2

MPS-2 FSAR FIGURE 3.3-8 CONTROL ELEMENT ASSEMBLIES GROUP AND NUMBER DESIGNATION NORTH ABC D E F GHJKLMNPR S T V W X Y

! II I I II II 1 'I I t t 1 I 2 a-S3 3-64 3 A~5 7-40 4-64 A4

, * "-53 ,"

f' 4 7.Q A~ ,.~ 1-30 ~...wi '-IS A..c4 2*21 12-22 S y, A"

"I' ~

6 I~ &-15 ~1a M7

...52 2-20 J.7 5-4 1*1 2*23 4-55 1 7-- ~

8- *fJ

~ 3081 8-8 U6 I---

"I'8-1 9- '-28 "'1 10- ~

i--- 7*:tg 5-3 ,., $-5 7~' 270 0

-ll -- ~

J.~1 ,..--

13 - ~*SO '*21 [1-6 1*1 '*32 14- " 1\

15 -- '-51 2-'11 B~ " 5-2 8-9 2*24 4-51 ""--

16 lA-Q &1-4 11.,,",

,V ... "

~

17 A~ 2*" 2*25 "lA-48 18 t7-S9 If..c4 ",25 1-33 iA~ t7.u

.- I' 19 ~ ~50 7*38 4.07 A-G 20 ~-5B lUi 21

-1 I 1 I l e SOURCE: LOCATION April 1998 Rev. 26.2

MPS-2 FSAR FIGURE 3.3-9 CORE ORIENTATION Outlet Nozzle Alignment Key 4 Equally Spaced Inlet Nozzle See Figure 3.3-8 for Identification of Core Arrangement and CEA Groups Fuel Assembly CEDM Core Support Reactor Barrel Vessel CEA Building North Elevation View April 1998 Rev. 26.2

MPS-2 FSAR FIGURE 3.3-10 IN-CORE INSTRUMENTATION ASSEMBLY 90 0

180 270 April 1998 Rev. 26.2

MPS-2 FSAR FIGURE 3.3-11 REACTOR INTERNALS ASSEMBLY Expansion Upper Guide Structure Compensating Ring Support Plate Alignment Key CEA Shroud In-Core Outlet Nozzle Instrumentation Guide Tube Aignment Pins Core Support Barrel Fuel Aignment Plate Core Shroud Core Support Plate Core Support Assembly Snubber April 1990 Rev. 26.2

MPS-2 FSAR FIGURE 3.3-12 PRESSURE VESSEL-CORE SUPPORT BARREL SNUBBER ASSEMBLY CENTER SUPPORT BARREL HARD-FACED SURFACE CORE STABILIZING LUG BOLT (12 REQ'D PER SNUBBER SPACER BLOCK ASSEMBLY)

SHIM (2 REQ'D PER ASSEMBLY)

PIN (4 REQ'D PER ASSEMBLY)

BOLT (4 REQ'D PER ASSEMBLY)

PRESSURE VESSEL April 1990 Rev. 24.8

MPS-2 FSAR FIGURE 3.3-13 CORE SHROUD ASSEMBLY Upper Segment

.J Lower Segment April 1990 Rev. 24.8

MPS-2 FSAR FIGURE 3.3-14 UPPER GUIDE STRUCTURE ASSEMBLY EXPANSION COMPENSAtING RING

  • CEA SHROUD GRID ASSEMBLY CEA SHROUDS "FUEL ASSEMBLY AUGNMENT PLATE April 1990 Rev. 26.2

MPS-2 FSAR FIGURE 3.3-15 CONTROL ELEMENT DRIVE MECHANISM (MAGNETIC JACK)

U~PPER 1iT~--ES:XTENS ION ItW,PrltTCH HAFT MAGNfT NG PIRESSURE HOUSING UPPER I .--~---- O'PERATI NG ROD SHROUD--~

ELECTRICAl CO~DUIT LIFT COil DRJVING LATCH '

COIL DRIVING LATCHES PULlDOWN COil MOTOR TUBE LOAD

~ANSFER OIL HOLDING U~TCH HOLDING COIL LATCHES PRESSURE HOUSING LOWER '

April 1990 Rev. 26.2

MPS-2 FSAR FIGURE 3.3-16 (LEFT BLANK INTENTIONALLY)

April 1998 Rev. 26.2

MPS-2 FSAR FIGURE 3.3-17 HEATED JUNCTION THERMOCOUPLE PROBE PRESSURE BOUNDARY INSTALLATION Pressure Housing Shroud Assembly Grayloc Coupling Detail W Rev. 23.3

MPS-2 FSAR FIGURE 3.3-18 TYPICAL HEATED JUNCTION THERMOCOUPLE PROBE ASSEMBLY INSTALLATION P\.CEOM NOZ2~:

L =~ CS EX rr' ::~.~ w

- - - S~l.. ~~WCS iN I.C.!,

NOZZl.E (C*~ i't.ANiSi l.C.I. Pt.Ai !

L.OWe~!O PCStT10N

-U.G.5. SUPPC RT ~\.Ai'a

~ROBe ASS!M8LY VACANT PART l!NG1'H--.I C!A SHROUt) ASS!MSI. Y c£. --------..i-HOT LSG HJ"iC ScNSCR .

t.OC*~*nCN CIA P1.UG FUEL ALIGNMENT PUTi m

April 1990 Rev. 26.2

MPS-2 FSAR FIGURE 3.3-19 PLACEMENT OF NATURAL URANIUM REPLACEMENT FUEL RODS AND FUEL ASSEMBLY ORIENTATION RELATIVE TO THE CORE BAFFLE FOR CYCLE 19 CORE BAFFLE NATURAL URANIUM REPLACEMENT FILLER ROD ENRICHED FUEL ROD CEA GUIDE TUBE Rev. 27.4

1 GENERAL

SUMMARY

s section summarizes the nuclear characteristics of the core and discusses the design ameters which are of significance to the performance of the core in normal transient and steady e operational conditions. A discussion of the nuclear design methods employed and parisons with experiments which support the use of these methods is included.

numerical values presented are based on a representative core design. Sufficient analyses are pleted each cycle to ensure that actual reload batches keep operating parameters within gn limits, accommodate essential reactivity requirements with the control system provided, meet other requirements for safe operation.

2 CORE DESCRIPTION Millstone Unit 2 reactor consists of 217 assemblies, each having a 14 by 14 fuel rod array.

assemblies are composed of up to 176 fuel rods, four control rod guide tubes, and one center trol rod guide tube/instrument tube. The fuel rods consist of slightly enriched UO2 or 2-Gd2O3 pellets inserted into Zircaloy tubes. The control rod guide tubes and instrument tubes also made of Zircaloy. Each AREVA assembly contains nine spacers. A description of the EVA supplied fuel design and design methods is contained in References 3.4-1, 3.4-2 and 3.4-epresentative loading pattern is shown in Figure 3.4-1 and is expressed in terms of previous le core locations and fuel assembly identifiers. A summary of fuel characteristics for a esentative core design is presented in Table 3.4-1. Figure 3.4-2 presents representative rter core assembly movements. Representative beginning of cycle (BOC) and end of cycle C) assembly exposures are shown in a quarter core representation in Figure 3.4-3.

epresentative low radial leakage fuel management plan results in scatter loading of the fresh throughout the core. Some fresh assemblies loaded in the core interior contain gadolinia-ring fuel in order to control power peaking and reduce the initial boron concentration to ntain the moderator temperature coefficient (MTC) within its Technical Specification limit.

exposed fuel is also scatter loaded in the center in a manner to control the power peaking.

3 NUCLEAR CORE DESIGN nuclear design bases for core design are as follows:

a. The design shall permit operation within the Technical Specifications for Millstone Unit 2 Nuclear Plant.
b. The design Cycle length (EFPD) shall be determined on the basis of an estimated Cycle energy and previous Cycle energy window.

3.4-1 Rev. 35

1. The peak linear heat rate (LHR) and the peaking factor Fr shall not exceed Technical Specifications limits in any single fuel rod throughout the cycle under nominal full power operating conditions.
2. The SCRAM worth of all rods minus the most reactive rod shall exceed the shutdown requirement.

neutronic design methods used to ensure the above requirements are consistent with those cribed in Reference 3.4-4.

3.1 Analytical Methodology neutronics methods used in the core analysis are described in Reference 3.4-4. The neutronic gn analysis for each reload core is performed using the PRISM reactor simulator code. Full-depletion calculations performed with PRISM are used to determine the core wide power ribution in three dimensions and to reconstruct the individual rod power and burnup ributions. Thermal-hydraulic feedback and axial exposure distribution effects are explicitly ounted for in the PRISM calculations. The CASMO/MICBURN assembly depletion model is d to generate the microscopic cross section input to the PRISM code.

3.2 Physics Characteristics neutronics characteristics of a representative reload core are presented in Table 3.4-2. The ty analysis for each cycle is applicable for a specified previous cycle energy window. A esentative HFP letdown curve is shown in Figure 3.4-4.

3.2.1 Power Distribution Considerations resentative calculated power maps are shown in Figures 3.4-5 and 3.4-6 for BOC uilibrium xenon), and EOC conditions, respectively. The power distributions were obtained m a three-dimensional neutronics model with moderator density and Doppler feedback effects rporated. The Technical Specification limits on Fr and LHR are 1.69 and 15.1 kW/ft, ectively.

3.2.2 Control Rod Reactivity Requirements epresentative shutdown margin evaluation is given in Table 3.4-3. The Millstone Unit 2 hnical Specifications require a minimum shutdown margin of 3,600 pcm.

3.4-2 Rev. 35

Technical Specifications require that the MTC be less than +7 pcm/°F at or below 70 percent ated thermal power, less than +4 pcm/°F above 70 percent power and greater than -32 pcm/°F 00 percent of rated thermal power. Representative MTC calculation results are presented in le 3.4-2.

4 POST-RELOAD STARTUP TESTING tup tests will be performed at the beginning of each reload cycle to obtain the as-built core racteristics and to verify Technical Specification and core physics design parameters. The ad startup physics test program is based on ANSI-19.6-1 (Reference 3.4-9). The Startup Test ivity Reduction (STAR) Program (Reference 3.4-10) provides an alternative to the ANSI-6-1 test program provided that specific criteria for the reload core design and construction are sfied. The STAR Program criteria are established in station procedures and include additional licability requirements for core design, fuel and control element assembly (CEA) fabrication, A lifetime monitoring, refueling and startup testing.

reload startup physics test program shall consist of the following:

a. Critical Boron Concentration - HZP, Control Rods Withdrawn.
b. Critical Boron Concentration - HZP, Control Rod Group(s) of at least 1%

reactivity are fully inserted in the core. 1

c. Control Rod Group Worths - HZP, two or more control rod groups shall be measured which are well distributed radially and represent a predicted total worth of at least 3% reactivity. 1
d. Isothermal Temperature Coefficient - HZP.
e. Flux Symmetry - between 0 and 30% of full power.
f. Power Distribution - between 40 and 75% of full power.
g. Isothermal Temperature Coefficient - greater than 70% of full power.
h. Power Distribution - greater than 90% of full power.
i. Critical Boron Concentration - greater than 90% of full power.
j. HZP to full power reactivity difference.

his test may be eliminated if performing the STAR Program per Reference 3.4-10.

3.4-3 Rev. 35

5.1 General on induced spatial oscillations on the Millstone Unit 2 core fall into three classes or modes.

se are referred to as axial oscillations, azimuthal oscillations, and radial oscillations. An axial llation is one in which the axial power distribution periodically shifts to the top and bottom of core. An azimuthal oscillation is one in which the X-Y power distribution periodically shifts m one side of the core to the other. A radial oscillation is one in which the X-Y power ribution periodically shifts inward and outward from the center of the core to the periphery.

on stability analyses indicate that a number of general statements can be made:

a. The time scale on which the oscillations occur is long, and any induced oscillations typically exhibit a period of 25 to 30 hours3.472222e-4 days <br />0.00833 hours <br />4.960317e-5 weeks <br />1.1415e-5 months <br />.
b. As long as the initial power peaking associated with the perturbation initiating the oscillation is within the limiting conditions for operation, specified acceptable fuel design limits will not be approached for a period of hours allowing an operator time to decide upon and take appropriate remedial action prior to the time when allowable peaking factors would be exceeded.
c. The core will be stable to radial mode oscillations at all times in the burnup cycle.
d. The core will be stable to azimuthal mode oscillations at all times in the burnup cycle.
e. All possible modes of undamped oscillations can be detected by both exactor and in-core instrumentation as discussed below.

5.2 Detection of Oscillations mary reliance for the detection of any xenon oscillations is placed on the exactor flux nitoring instrumentation. The power range excore neutron detectors (one axial pair per drant) are used to monitor the symmetry of power distributions and are located at distinct muthal and axial positions. These detectors are sensitive primarily to the power density ations produced by peripheral fuel assemblies in the vicinity of the detectors. All possible on induced spatial oscillations will affect the power densities of the peripheral fuel assemblies he core.

ddition, the in-core instrumentation provides information which will be used in the early es of cycle operation to confirm predicted correlations between indications from the excore ctors and the space-dependent flux distribution within the core. Later on, during normal ration, the in-core detector system provides information which may be used to supplement that ilable from the excore detectors.

3.4-4 Rev. 35

ce the reactor will not be operated under conditions that imply instability with respect to muthal xenon oscillation, no special protective system features are needed to accommodate muthal mode oscillations. Regardless, a maximum azimuthal power tilt is prescribed in the hnical Specifications along with prescribed operating restrictions in the event that the muthal power tilt limit is exceeded.

described earlier, the power range excore neutron detectors are used to monitor the azimuthal metry of the power distributions since they are located at distinct locations in the X-Y plane.

uld the excore detectors indicate different readings in the azimuthal direction, a tilt in the core er distribution would be indicated. When the tilt exceeds a preset magnitude an alarm will ur. In the event of an alarm, the orientation of the tilt will be determined and, on the basis of ntation, the proper CEAs will be manually adjusted to reduce the magnitude of the tilt.

features provided for azimuthal xenon oscillation control are:

a. instrumentation for monitoring azimuthal power tilt.
b. administrative limits on azimuthal power tilt.

excore detectors are used to monitor the axial power distribution and to detect deviations m the equilibrium distribution such as those which would occur during an axial xenon llation. This is done by monitoring variations in the external axial shape index, a parameter ved from the excore detector readings which is related to the axial power distribution. Control xial xenon oscillation is accomplished utilizing Regulating Bank 7. When it is determined that axial shape index may exceed the boundaries of a specified control band about the equilibrium e, this bank is slowly inserted and eventually withdrawn over a period of several hours. The is then stabilized until a new oscillation develops.

features provided for axial xenon control and protection are:

a. equipment for monitoring axial shape index.
b. administrative limits on axial power distribution, external axial shape index.
c. an axial shape index reactor trip (local power density - high).
d. use of Regulating Bank 7 for control of axial power distribution.

3.4-5 Rev. 35

ent core designs for Millstone Unit 2 (Cycles 10 and beyond) have been developed to include ger fuel cycles along with low radial leakage fuel management. These current designs scatter fresh fuel assemblies throughout the interior of the core with the highest burnup fuel mblies being loaded along the core periphery. Core designs prior to Cycle 10 operation were of a low radial leakage design due to the loading of fresh fuel assemblies along the core phery.

h respect to xenon oscillations in the radial and azimuthal directions, studies indicate that core gns of a low radial leakage design (i.e., highest burnup assemblies loaded on the core phery with fresh fuel assemblies scatter loaded about the core interior) are more stable than e designs which load fresh fuel assemblies along the core periphery. Therefore, the clusions regarding xenon oscillations in the radial and azimuthal directions, which are ented in Section 3.4.5.5, remain applicable to current plant operations.

h regard to axial xenon oscillations, the core near end-of-cycle may be naturally unstable in absence of any control rod action even if low leakage core designs are utilized. But axial on oscillations are sufficiently slow (the period of oscillation being 25 to 30 hours3.472222e-4 days <br />0.00833 hours <br />4.960317e-5 weeks <br />1.1415e-5 months <br />) so that e would be sufficient time to control the oscillations. In addition, automatic protection is vided if operator action is not taken to remedy the situation. Regulating Bank 7 CEAs are zed for controlling axial xenon oscillations.

5.5 Method of Analysis classic method for assessing spatial xenon oscillations is that developed by Randall and St.

n (Reference 3.4-5) which consists of expanding small perturbations of the flux and xenon centrations about equilibrium values in eigenfunctions of the system with equilibrium xenon ent. However, it is necessary to extend this simple linear analysis to treat cores which are uniform because of fuel zoning, depletion, and CEA patterns, for example. Such extensions e been worked out and are reported in References 3.4-6 and 3.4-8. In this extension, the nvalue separations between the excited state of interest and the fundamental are computed erically for symmetrical flux shapes. For nonsymmetrical flux shapes, the eigenvalue aration can usually be obtained indirectly from the dominance ratio 1/0, computed during iteration cycle of the spatial calculation.

merical space time calculations are performed in the required number of spatial dimensions for various modes as checkpoints for the predictions for the extended Randall-St. John treatment cribed above.

3.4-6 Rev. 35

confirm that the radial oscillation mode is extremely stable, a space-time calculation was run a reflected, zoned core 11 feet in diameter without including the damping effects of the ative power coefficient. The initial perturbation was a poison worth of 0.4 percent in reactivity ed in the central 20 percent in the core for 1 hour1.157407e-5 days <br />2.777778e-4 hours <br />1.653439e-6 weeks <br />3.805e-7 months <br />. Following removal of the perturbation, the lting oscillation was followed in 4-hour time steps for a period of 80 hours9.259259e-4 days <br />0.0222 hours <br />1.322751e-4 weeks <br />3.044e-5 months <br />. Results show that oscillation died out very rapidly with a damping factor of about minus 0.06 per hour. When damping coefficient is corrected for a finite time mesh by the formula in Reference 3.4-7, it is e strongly convergent. On this basis, it is concluded that radial oscillation instability will not ur.

s conclusion is of particular significance because it means that there is no type of oscillation re the inner portions of the core act independently of the peripheral portions of the core whose avior is most closely followed by the excore flux detectors. Radial mode oscillations, even ugh highly damped, would be manifested as periodic variation in the excore flux power signal le the delta-T power signals remained constant. Primary reliance is placed on the excore flux ctors for the detection of any xenon oscillations.

5.5.2 Azimuthal Xenon Oscillations lyses indicate that the eigenvalue separation between the first asimuthal harmonic and the damental is about 0.86 percent in . The calculated damping coefficient for the first azimuthal de is minus 0.016 per hour, and the higher modes will be even more strongly damped.

thermore, the Doppler coefficient applicable to the Millstone Unit 2 reactor is calculated to be roximately minus 1.36 x 10-3 /(kW/ft) which is sufficiently negative to ensure stability of he azimuthal modes.

5.5.3 Axial Xenon Oscillations checkpoints for the predictions for the modified Randall-St. John approach, numerical spatial e calculations have been performed for the axial case at both beginning and end-of-cycle. The and poison burnup distributions were obtained by depletion with soluble boron control so that power distribution was strongly flattened. Spatial Doppler feedback was included in these ulations. The initial perturbation used to excite the oscillations was a 50 percent insertion into top of the core of a 1.5 percent reactivity CEA bank for 1 hour1.157407e-5 days <br />2.777778e-4 hours <br />1.653439e-6 weeks <br />3.805e-7 months <br />. The damping factor for this was calculated to be about +0.02 per hour; however, when corrected for finite time mesh rvals by the methods of Reference 3.4-7, the damping factor is increased to approximately

04. When this damping factor is plotted at the appropriate eigenvalue separation for this mode nd-of-cycle, it is apparent that good agreement is obtained with the modified Randall-St. John diction.

3.4-7 Rev. 35

s result suggests that the constant power condition which applies to the axial oscillations lts in a very weak moderator feedback since the moderator density distribution is fixed at the and bottom of the core and only the density distribution in between can change.

the calculated Doppler coefficient of minus 1.36 x 10-3 /(kW/ft), the damping factor toward end of the burnup cycle is positive. Thus, within the uncertainties in predicting power fficients and uncertainties in the analyses, there is a prediction of unstable axial xenon llations in the absence of any control action. These oscillations are sufficiently slow (the od of oscillation being 25 to 30 hours3.472222e-4 days <br />0.00833 hours <br />4.960317e-5 weeks <br />1.1415e-5 months <br />) so that there would be sufficient time to control the llations. In addition, automatic protection is provided if operator action is not taken to remedy situation. Regulating Bank 7 CEAs are utilized for controlling axial xenon oscillations.

6 REFERENCES 1 Generic Mechanical Design Report Exxon Nuclear 14 x 14 Fuel Assemblies for Combustion Engineering Reactors, XN-NF-82-09(A), Exxon Nuclear Company, Richland, WA 99352, November 1982.

2 Design Report for Millstone Point Unit 2 Reload ANF-1, ANF-88-088(P), Rev. 1, Advanced Nuclear Fuels Corporation, Richland, WA 99352, August 1988.

3 Millstone Unit 2 Mechanical Design Report for Increased Peaking EMF-91-245(P),

Siemens Nuclear Power Corporation, January 1992.

4 EMF-96-029(P)(A) Volumes 1 and 2, Reactor Analysis System for PWRs Volume 1 -

Mehodology Description, Volume 2 - Benchmarking Results, Siemens Power Corporation, January 1997.

5 Randall, D., Xenon Spatial Oscillations, Nucleonics, 16, 3, pages 82-86 (1958).

6 Stacey, Jr., W. M., Linear Analysis of Xenon Spatial Oscillations, Nuclear Sci. Eng.,

30, pages 453-455 (1967).

7 Poncelet, C. G., The Effect of a Finite Time Step Length on Calculated Spatial Xenon Stability Characteristics in Large PWR's Trans. ANS, 10, 2, page 571 (1967).

8 CEND-TP-26., Diatch, P.B.

9 ANSI/ANS-19.6-1 Reload Startup Physics Tests for Pressurized Water Reactors, 2005.

10 WCAP-16011-P-A, Revision 0, Startup Test Activity Reduction Program, February 2005.

3.4-8 Rev. 35

Fuel Types N1 N2 N3 N4 P1 P2 P3 P4 P5 R1 R2 R3 R4 R5 R Central Zone Assem- 3.94 3.90 3.87 3.82 3.87 3.86 3.84 3.81 3.76 4.49 4.49 4.47 4.39 4.33 4.42 bly Average Enrich-ment (w/o)

Number Gadolinia 0 6 12 16 0 4 8 12 16 0 4 8 12 16 12 Bearing Rods Nominal Density (% 94 94 94 94 94 94 94 94 94 95 95 95 95 95 95 TD)

Pellet OD (inches) 0.370 0.370 0.370 0.370 0.370 0.370 0.370 0.370 0.370 0.377 0.377 0.377 0.377 0.377 0.37 Clad OD (inches) 0.440 0.440 0.440 0.440 0.440 0.440 0.440 0.440 0.440 0.440 0.440 0.440 0.440 0.440 0.44 Diametral Gap 0.0080 0.0080 0.0080 0.0080 0.0080 0.0080 0.0080 0.0080 0.0080 0.007 0.007 0.007 0.007 0.007 0.00 (inches)

Clad Thickness 0.031 0.031 0.031 0.031 0.031 0.031 0.031 0.031 0.031 0.028 0.028 0.028 0.028 0.028 0.02 (inches)

Rod Pitch (inches) 0.580 0.580 0.580 0.580 0.580 0.580 0.580 0.580 0.580 0.580 0.580 0.580 0.580 0.580 0.58 Spacer Material Bime- Bime- Bime- Bime- Bime- Bime- Bime- Bime- Bime- Bime- Bime- Bime- Bime- Bime- Bim tallic tallic tallic tallic tallic tallic tallic tallic tallic tallic tallic tallic tallic tallic talli Fuel Supplier AREVA AREVA AREVA AREVA AREVA AREVA AREVA AREVA AREVA AREVA AREVA AREVA AREVA AREVA AR Fuel Stack Height 136.7 136.7 136.7 136.7 136.7 136.7 136.7 136.7 136.7 136.7 136.7 136.7 136.7 136.7 136.

Nominal (inches)

Number of Assem- 8 20 8 25 8 8 12 8 36 8 8 8 8 48 4 blies Regionwise Loading 3.04 7.60 3.03 9.43 3.04 3.04 4.55 3.03 13.58 3.19 3.19 3.19 3.17 18.98 1.59 (MTU) 3.4-9 Rev

RELOAD CORE

<characteristic> BOC EOC tical Boron (ppm): HZP, ARO, No Xenon 1453 ---

tical Boron (ppm): HFP, ARO, Equilibrium 1024 0 non derator Temperature Coefficient (pcm/°F): +2.0 -10.4 P

derator Temperature Coefficient (pcm/°F): -6.0 -23.3 P

ppler Coefficient (pcm/°F) -1.17 -1.33 ron Worth (pcm/ppm): HZP -8.8 -10.8 ron Worth (pcm/ppm): HFP -8.4 -10.4 R (kW/ft) HFP (a) 12.8 11.6 layed Neutron Fraction 0.0064 0.0054 P, PDIL Worth (pcm) 157 241 1 Rod Worth, HZP (pcm) 6271 7696 cess Shutdown Margin (pcm): HFP 124 323 cess Shutdown Margin (pcm): HZP 140 751 Including uncertainties.

3.4-10 Rev. 35

TABLE 3.4-3 REPRESENTATIVE SHUTDOWN MARGIN REQUIREMENTS trol Rod Worth (pcm)

BOC: BOC: EOC: EOC:

<parameter> HZP HFP HZP HFP ARI 9315 9315 10450 10450 N-1 6271 6271 7696 7696 PDIL 2116 157 2862 241

[(N-1) - PDIL]

  • 0.9 3740 5503 4351 6710 ctivity Insertion (pcm)

BOC: BOC: EOC: EOC:

<parameter> HZP HFP HZP HFP Power Defect 0 1507 0 2515 Void 0 50 0 50 Flux Redistribution 0 222 0 222 Total Requirements 0 1779 0 2787 tdown Margin (pcm)

BOC: BOC: EOC: EOC:

<parameter> HZP HFP HZP HFP

[(N-1)

  • PDIL]
  • 0.9 - Total 3740 3724 4351 3923 Required Shutdown 3600 3600 3600 3600 Excess Shutdown Margin 140 124 751 323 3.4-11 Rev. 35

MNPS-2 FSAR MPS-2 FSAR FIGURE 3.4-1 REPRESENTATIVE FULL CORE LOADING PATTERN 1 2 3 4 s I 7

  • 11 13 15 16 17 18 18 20 21 y

I=J N2B1 N191 N32/

0.16 e-os R-13 x H2O R01 P64 Rl0 N11 R15 P45 ROO N27 S..()5 F-04 L*'S F-,S S.17 w N55 R2D P05 R39 P33 Ra1 P36 R80 P04 R23 N57

$-07 X-16 C-07 0.15 X-06 8-15 v N58 P09 P21 R65 N65 R49 N52 R55 NIt R70i P26 P1. N54 R.oEi W-OS J.02 V*17 l-20 V..os J.2O T*19 Re16 T N22 R20i P1I R70 P63 R59 P52 R31 P.... A62 P40 R69 P17 "R19 N25 T~ X*13 "N-04 E-07 E-15 N-1S X-09 T*1' s R07 P03 R75 P65 P19 P57 R43 N43 R4Ei PSl P27 P47 R68 P02 R04 F.(J2 V.()9 V.04 8-13 T-09 S.Q9 V*18 V-l3 F*20 R P66 R77 N62 R63 P58 P42 N03 R27 N06 P38 P71 ASS N68 R38 P54

~

N29 V*1S E..oc J.06 -'-13 N-03 N*19 J.09 J.16 E*l8 V.()6

-N36 p N N-U1 R1S P29 RS6 P37 R47 N07 R35 Pl0 R34 N02 R42 P68 R52 P32 R09 N*'~

f--- I'"'-

N1. R-19 R-17 W-09 T.03 W*13 R..05 R43 N13 M L F*'9 N26 R82 N48 R32 N44 R28 P12 ~ P13 R26 N42 R30 N51 R84 N1E F.oJ f---

N23 R*l1 X*11 J.05 C-05 0.11 W*17 B-11 ~11 H-"

-N24 K J S-19 Rl1 P31 RSO P48 R44 N04 R36 P11 R33 N05 R45 P51 R54 P30 R14 S..()J f--- 10-N34 6-19 G-1, e-09 ~.19 0.13 (;.OS G-03 NJ, H G .un P62 R40 N66 AGO P72 P39 NOB R25 N01 P43 P56 AS1 N64 R79 P67 J.1S D-1E T-G4 N-06 N-13 J.Q3 J.19 N..()9 N*1S T-18 o-ee f R02 P08 R66 PSO P22 P60 R48 N41 R41 P59 P20 P53 R73 P06 R05 S-02 D-09 [).04 F-13 E*13 F.09 [).18 0-13 S-2O E H15 R17 P24 R71 P70 R64 P49 R29 P41 R57 P55 R72 P23 R22 N12 E..06 8-13 J.04 T-07 T*15 J.18 8.09 E-16 D N56 P16 P2S R76 N63 R53 N50 ASl N67 R67 P28 P15 N60 G-OE E~ N~ 0.17 L.Q2 [).()5 N*2(l 0-17 G-1E c N59 R21 P01 R78 P35 RaJ P34 R37 P07 R18 N53 F-UT 8-16 W-UT W-15 B.o& F-1S 8 N17 ROB P46 R13 N21 R12 P69 R03 N10 Assembly 1.0.

F-OS S..()4 l.07 $-18 F-17 Previous Cycle L ocatlon A

N18 N3S/

G-09 W-1S w~ ~13 I~Itail il

  • 10 12 14 Figure 3.4-1 Representative Full Core Loading Pattern June 2000 June 2000

MPS-2 FSAR FIGURE 3.4-2 REPRESENTATIVE QUARTER CORE LOADING PATTERN MNPS-2 FSAR 11 13 15 16 17 18 19 20 21 1 N4 2 P2 3 R4 4 N4 5 R4 6 N4 7 R6 8 HZ L C-l1 [-19 E-13 L-20 L-15 9 N2 90 90 270 270 F-19 K 10 P2 11 R5 12 HI 13 R5 14 P5 IS RS 16 P4 17 R2 180 J E-19 C.. 13 [-15 e-15 18 H3 270 270 J-1S H 19 R4 20 Nl 21 P5 22 PS 23 RS 24 N4 2S R5 26 P5 G J-19 J-13 F-13 D-17 F-18 180 90 90 270 27 "N4 28 R5 29 P5 30 P3 31 P5 32 R5 33 PI 34 RI F E-13 J-16 D-18 D-13 B*16 270 90 35 R4 36 P5 37 RS 38 PS 39 R5 40 P3 41 R3 42 H2

[ G-17 J-18 J*20 [-16 90 270 43 N4 44 RS 45 N4 46 R5 47 P3 48 P2 49 H4 D B-l1 [-18 B*13 e-17 G*16 270 270 90 SO R6 51 P4 52 RS 53 PI 54 R3 55 N4 t G-19 F-20 F-15 90 270 56 HZ 57 R2 58 P5 S9 RI 60 H2 Region No./Subbatch Type B G-11 0-16 F-17 Previous Cycle 1/4 Core Loca tion 270 90 Rotation (Deg. CCW) 61 H2 62 H3 A e*16 G*13 180 12 14 FIGURE 3.4-2 REPRESENTATIVE QUARTER CORE LOADING PATTERN 00-.)3 June 2000 June 2000

MPS-2 FSAR MNPS*2 FSAR FIGURE 3.4-3 REPRESENTATIVE BOC AND EOC EXPOSURE DISTRIBUTION 11 13 15 16 17 18 19 20 21 1 "4 2 P2 3 a4 4 1'4 5 R4 6 114 7 1t6 8 112 34.036 16.585 0.000 35.125 0.000 30.580 0.000 32.683 9 112 50.652 37.623 23.655 52.112 24.328 48.977 23.811 46.522 32.000 10 '2 11 ItS 12 .1 13 ItS 14 '5 15 as 16 P4 17 R2 38.656 J 16.585 0.000 31.361 0.000 20.938 0.000 20.147 0.000 18 .3 37.623 23.096 48.812 22.853 Cl.311 23.C95 39.970 20.063 34.801 B 19 1t4 20 .1 21 PS 22 .5 23 as 24 .4 25 as 26 '5 39.877 c 0.000 31.375 18.933 20.637 0.000 32.250 0.000 20.199 23.655 48.835 37.652 39.981 23.397 49.790 22.002 34.127 27 .4 28 ItS 29 P5 30 P3 31 '5 32 1t5 33 PI 34 Itl 35.125 0.000 20.590 16.156 20.581 0.000 12.918 0.000 52.112 22.843 39.941 36.614 40.659 23.238 33.445 16.106 35 1t4 36 '5 37 itS 38 PS 39 itS 40 .3 41 R3 42 .2

  • 0.000 24.328 20.946 41.288 0.000 23.383 20.562 40.643 0.000 23.301 16.'42 36.317 0.000 19.103 33.040 40.028 43 N4 .. 4 JlS 45 114 46 RS 47 '3 48 P2 49 .. 4 D 30.580 0.000 32.250 0.000 16.938 16.589 34.067 48.977 23.419 49.774 23.235 36.317 31.056 40.643 50 1t6 51 P4 52 :R5 53 PI 54 R3 55 Mol Region No. I .atch t ype e 0.000 20.156 0.000 12.924 0.000 33.971 BOC Exposure {GWdl MTU) 23.811 39.964 21.995 33.449 19.107 40.559 00-.;13 EOC Exposure (Gwdl MTU) 56 "2 57 R2 58 . '5 59 1t1 60 112
  • 32.683 46.522 0.000 20.109 20.166 3 ....107 0.000 16.108 33.070 40.056 61 112 62 113 31.943 34.813 38.624 39.900 12 14 FIGURE 3.4-3 REPRESENTATIVE BOC AND EOC EXPOSURE DISTRIBUTION

-- ~

June 2000 June 2000

MPS-2 FSAR FIGURE 3.4-4 REPRESENTATIVEMNPS-2 FSAR BORON LETDOWN, HFP, ARO 1~

1100

- 1(XX) '" I'..

E -,

a..

-Q.

!IX)

"~

(

0 fa)

.J

.J ~,

0 -

700 t

.J

( -,

e 0

c 600 0 SOO 0

"~

(

L 0

0 400 '" "\

r\.

r\

Q)

m ~

1"\

X()

i\.

100

" \.

o r\

o 2 3 t 5 6 1 8 9 10 II 12 13 Ii IS 16 11 18 19 C9cle Exposure lGWd/MTU)

Figure 3.4-4 Representative Boron Letdown, HFP, ARO Millstone Unit 2 June 2000 June 2000

MPS-2 FSAR FIGURE 3.4-5 REPRESENTATIVE NORMALIZED POWER DISTRIBUTIONS, HOT FULL POWER, EQUILIBRIUM XENON, 150 MWD/MTU MNPS-2 FSAR 11 13 15 16 17 18 19 20 21 i M4 2 P2 3 R4 4 1f4 5 R4 6 ..4 7 R6 8 112 L 0.932 1.197 1.307 0.880 1.317 0.971 1.327 0.765 , .2

. 0 . 981 1.313 1.600 0.933 1.5'4 1.047 1.577 0.'45 0.347 10 P2 11 JtS 12 51 13 ItS 14 P5 15 ItS 16 Pot 17 R2 0.626 1.197 1.261 0.950 1.214 1.085 1.241 1.108 1.212 18 53 1.313 1.570 1.048 1.517 1.155 1.543 1.179 1.573 0.263 B l' R4 20 .1 21 75 22 P5 23 ItS 24 114 25 as 26 P5 0.545 o 1.307 0.950 1.058 1.089 1.259 0.926 1.211 0.781 1.600 1.049 1.173 1.186 1.57" 0.981 1.5"2 1.045 27 .... 28 ItS 29 P5 30 P3 31 P5 32 ItS 33 PI 34 JU p 0.880 1.214 1.089 1.215 1.130 1.278 1.212 1.027 0.933 1.516 1.187 1.376 1.222 1.596 1.31" 1.574 35 R4 36 PS 37 ItS 38 P5 39 ItS ..0 P3 41 R3 .. 2 112

  • 1.317 1.564 1.083 1.151 1.258 1.573 1.130 1.222 1.302 1.587 1.115 1.218 1.120 1.574 0.385 0.846 43 1f4 44 as 45
  • N4 4' Jl5 4' P3 48 P2 49 114 D 0.971 1.236 0.924 1.277 1.115 0.823 0.346 1.047 1.536 0.geO 1.595 1.218 1.068 0.684 50 R6 51 P4 52 R5 53 PI 54 R3 55 114 e 1.327 1.107 1.210 1.212 1.120 0.347 Re9ioft 110. I ruel 1.577 1.1'7 1.541 1.313 1.574 0.686 A**-=bly Average Power 56 K2
  • 57 R2 58 P5 59 JU 60 .2 A** embly Peak Pin Power
  • 0.765 0.945 1.216 1.578 0.782 1.046 1.027 1.57" 0.385 0.845 Pr
  • 1.600 (e 11) 61 If2 62 1f3 PEAK LRR (kW/ft)
  • 12.8 (1 6 B 16) 0.349 0.264 0.630 0.547 12 ~.

FIG~RE 3.4-5 REPRESENTATIVE NORMALIZED POWER DISTRIBUTIONS 00*

.13 HFP, EQUILIBRIUM XENON, 150 -'IVi-WD/MTli- -- - ---

June 2000 June 2000

MPS-2 FSAR FIGURE 3.4-6 REPRESENTATIVE NORMALIZED MNPS..2 FSAR POWER DISTRIBUTION, HOT FULL POWER, EQUILIBRIUM XENON, 18,020 MWD/MTU

~1 13 15 16 17 18 19 20 21 1f4 2 P2 3 a4 4 )14 5 1t4 6 .4 7 1t6 8 .2

~

L 0.905 1.131 1.401 0.969 1.425 1.005 1.313 0.736 52 0.955 1.228 1.495 1.008 1.510 1.066 1.449 0.883 0.383 Jt 10 P2 11 as 12 Jfl 13 as 14 P5 15 as 16 Pol 17 R2 0.619 J 1.131 1.386 0.982 1.388 1.112 1.402 1.038 1.062 18 .3 1.228 1.503 1.017

  • 1 . 514 1.161 1.51" 1.139 1.300 0.293 B 19 It" 20 Xl 21 P5 22 P5 23 ItS 24 .4 25 R5 26 PS 0.546 c 1.401 0.982 0.993 1.027 1.395 0.t81 1.300 0.740 1.495 1.017 1.053 1.129 1.517 1.032 1.459 0.984 27 .4 28 ItS 29 P5 30 P3 31 P5 32 as 33 P1 34 1t1

., 0.969 1.388 1.027 1.051 1.060 1.363 1.057 0.859 1.008 1.5106 1.130 1.183 1.125 1.478 1.140 1.261 35 1t4 36 P5 37 RS 38 PS 39 itS 40 P3 41 R3 42 .2 1.425 1.111 1.396 1.061 1.349 1.011 1.036 0.386 1.510 1.161 1.517 1.126 1.0663 1.104 1.356 0.759 43 Mol 44 itS 45 114 46 ItS 4' P3 48 P2 49 N4 D 1.005 1.400 0.981 1.363 1.011 0.771 0.371 1.066 1.513 1.031 1.478 1.104 0.976 0.677 50 Jl6 51 Pol S2 as 53 PI 54 Jl3 55 1f4 c 1.313 1.038 1.300 1.057 1.036 0.371 Region No. I Fuel 1.449 1.138 1.461 1.141 1.357 0.679 Aaaembly Average Power 56 .2 57 R2 58 P5 S9 III 60 52 A** embly P.ak Pin Power

  • 0.736 0.883 1.063 1.302 0.'40 0.985 0.859 1.262 0.386 0.758 "r
  • 1.517 (E 15) 61 112 62 .3 PEAX LHR (kW/ft)
  • 11.6 (21 E 15) 0.384 0.293 0.621 0.547 12 14 FIGURE 3.4-6 REPRESENT:ATIV~_~O~~~_I..I:l:~D~9WER DISTRIBUTION HFP, EQUILIBRIUM XENON, EOC Rev. 32 June 2000

s section presents thermal and hydraulic analysis of the reactor core, analytical methods zed, and experimental work supporting the analytical techniques. The prime objective of the mal and hydraulic design of the reactor is the assurance that the core can meet normal steady e and anticipated transient performance requirements without exceeding the design bases. A mary of the significant reactor and fuel parameters used in the thermal and hydraulic design analysis is presented in Table 3.5-1.

1 DESIGN BASES 1.1 Thermal Design idance of thermally induced fuel damage during any normal steady state and anticipated sient operation is the principal thermal and hydraulic design basis. The following limits are blished, but violation of them will not necessarily result in fuel damage. The Reactor tection System will provide for automatic reactor trip or other corrective action before these gn limits are exceeded.

a. Avoidance of departure from nucleate boiling (DNB) for the limiting rod in the core with 95 percent probability at a 95 percent confidence level.
b. Limitation of the peak temperature of the fuel to less than the melting point during normal operation and anticipated transients.

ce the departure from nucleate boiling ratio (DNBR) criterion ensures that the cladding perature remains close to the coolant temperature, no additional criteria for cladding perature are required for normal operation and anticipated transients. For design basis dent conditions (loss of coolant accidents (LOCA)), under which the DNBR criterion does not ly, cladding temperatures are calculated to ensure that they remain below 2200°F, which is the k clad temperature criterion of 10 CFR 50 Appendix K. For other postulated accidents, fuel ure is assumed to occur if the calculated DNBR is below the DNB correlation 95/95 limit.

1.2 Hydraulic Stability rating conditions shall not lead to flow instability during normal steady state and anticipated sient operation.

1.3 Coolant Flow Rate, Distribution and Void Fraction ower limit on the total primary coolant flow rate, called design flow, is set to assure that the is adequately cooled when uncertainties in system resistance, pump head, and core bypass are taken in the adverse direction. By design of the reactor internal flow passages, this flow istributed to the core such that the core is adequately cooled with all permissible core power ributions. The hydraulic loads for the design of the internals are based on the upper limit of the 3.5-1 Rev. 35

nsure that sufficient coolant flow reaches the fuel, the amount of coolant flow which bypasses core through the guide tubes must not excessively reduce the active core flow. The guide tube lant flow must, however, be sufficient to ensure that coolant in the guide tubes will not boil ensure adequate cooling of the CEA fingers. The CEA drop time in the guide tubes must also t the criterion of 90 percent insertion within 2.75 seconds to ensure that scram performance is ccordance with plant Technical Specifications.

hough the coolant velocity, its distribution, and the coolant voids affect the thermal margin, gn limits need not be applied to these parameters because they are not themselves limiting h respect to thermal margin. These parameters are included in the thermal margin analyses and affect the thermal margin to the design limits.

2 THERMAL AND HYDRAULIC CHARACTERISTICS OF THE DESIGN 2.1 Fuel Temperatures RODEX2 code (Reference 3.5-1) incorporates models to describe the thermal and hanical behavior of the fuel rod in a flow channel including the gas release, swelling, sification, and cracking in the pellet; the gap conductance; the radial thermal conduction; the volume and gas pressure internal to the fuel rod; the fuel and cladding deformations; and the ding corrosion as a function of burnup. The calculations are performed on a time-incremental s with conditions being updated at each calculated increment.

2.1.1 Fuel Cladding Temperatures RODEX2 thermal-hydraulic model (Reference 3.5-1) calculates the lowest cladding surface perature based on one of two heat transfer regimes; i.e., forced convection and fully developed leate boiling. The forced convection and fully developed nucleate boiling heat transfer elations in RODEX2 were developed by Kays and Thom et al., respectively.

2.1.2 Fuel Pellet Temperatures RODEX2 radial temperature distribution model begins with the standard differential equation eat conduction (Poisson Equation) for an isotropic solid with internal heat generation. The ation is written in cylindrical coordinates assuming that the thermal conductivity of the fuel is nction of fuel temperature, but is independent of position. With additional assumptions of l symmetry, negligible heat conduction in the axial direction, and steady state conditions, a

-dimensional (i.e., radial) steady state form of the equation is derived and employed.

minimum power level required to produce centerline melt in Zircaloy clad uranium fuel rods efined as the Fuel Centerline Melt Linear Heat Rate (FCMLHR) limit and is expressed in kW/

This FCMLHR is determined using the methodology of Reference 3.5-22. A conservative le specific FCMLHR limit is used for Millstone Unit 2. The maximum LHR for normal 3.5-2 Rev. 35

temperature than an all-uranium-bearing fuel rod. Gadolinia rods are specifically analyzed to terline melt criteria.

2.1.3 UO2 Thermal Conductivity ns expression for thermal conductivity of the fuel is used in RODEX2. Two corrections are lied: one for density and one to account for the gadolinia content in the fuel.

2.1.4 Gap Conductance RODEX2 gap conductance model is based on that proposed by Kjaerheim and Rolstad. The l gap conductance has three components: (1) gas conductance, (2) radiation, and (3) fuel/

ding solid-to-solid contact.

2.2 Departure from Nucleate Boiling Ratio BRs are calculated using approved correlations. An approved core thermal-hydraulic puter code is used to determine the flow and enthalpy distribution in the core and the local ditions in the hot channel for use in the DNB correlation.

2.2.1 Departure from Nucleate Boiling XCOBRA-IIIC (Reference 3.5-2) computer code is employed to evaluate the thermal-raulic conditions in the various assemblies and in the subchannels of the limiting assembly.

t, mass, and momentum fluxes between the inter-rod flow channels are explicitly calculated.

l and reactor design conditions employed in these calculations are given in Table 3.5-1.

calculations include a statistically determined engineering factor to account for ufacturing tolerances, thermal expansion and densification effects. The engineering factor is lied to the local heat flux in the calculation of DNBR.

eactor densification results in a shortening of the fuel column. At power levels typical of BR-limiting rods, thermal expansion tends to offset the densification effect. The XCOBRA-model does not specifically model changes in stack length due to thermal expansion and sification.

HTP DNB correlation, demonstrated to be applicable to the AREVA 14 by 14 reload fuel mblies for CE reactors, is described in Reference 3.5-3. A minimum allowable limit esponding to 95% probability with 95% confidence is set on the DNBR during normal ration and any anticipated transients.

2.2.2 Hot Channel Factors channel factors for heat flux and enthalpy rise, Fq and Fr:

3.5-3 Rev. 35

rage ratios of these quantities. The heat flux hot channel factor (Fq) considers the local imum linear heat generation rate at a point (the hot spot), and the enthalpy rise hot channel or (Fr) involves the maximum integrated linear heat generation rate along a channel (the hot nnel).

ineering hot channel factor, FE:

engineering hot channel factor is used to evaluate the maximum linear heat generation rate in core. This subfactor is determined by statistically combining the fabrication uncertainties for pellet diameter, density, and enrichment, as well as the effect of densification. A conservative e of 1.03 is used. The effect of variations in fabrication tolerances is considered in the lysis. To account for manufacturing uncertainties and densification, the peak rod heat flux is eased by 3% in the calculation of DNBR.

2.2.2.1 Nuclear Peaking Factors embly and rod peaking factors and axial power distributions are input into the XCOBRA-IIIC

e. Departure from nucleate boiling is dependent on the local rod heat flux and the local fluid ditions within the channel.

effect of asymmetries in core power distribution (specifically azimuthal power tilt) is not ctly taken into account in the XCOBRA-IIIC thermal-hydraulic calculations. The effects of muthal power tilt are accounted for in the generation (verification) of the TM/LP trip and LPD monitoring setpoints through the measurement of radial peaking factors.

2.2.2.2 Rod Bowing Factor the fuel assembly burnup increases, the gaps between fuel rods change. Decreased rod-to-rod s can occur, which can reduce the DNB ratio. Penalties are calculated as a function of burnup applied to the DNBR or peak linear power as appropriate.

2.2.2.3 Inlet Flow Distribution Factor t flow maldistribution is treated in the XCOBRA-IIIC model by applying a generic inlet flow alty to the limiting assembly and its crossflow neighbors.

2.2.2.4 Flow Mixing Factor effects of both pressure-driven and turbulent flow mixing between channels on the hot nnel enthalpy rise are calculated by the XCOBRA-IIIC computer code. The turbulent flow ing is modeled empirically and is based on the reduction of the data from hot mixing tests g XCOBRA-IIIC.

3.5-4 Rev. 35

2.2.3 Effects of Rod Bow on DNBR ccordance with AREVA rod bow methodology (Reference 3.5-4), the magnitude of rod bow assemblies of the type used in Millstone Unit 2 has been estimated. Significant impact on the BR due to rod bow does not occur until the gap closures exceed 50 percent. The maximum gn exposure for AREVA reload fuel in Millstone Unit 2 is significantly less than that at which percent closure occurs; therefore, rod bow does not significantly impact the minimum DNBR DNBR). A further consequence of the small amount of rod bow for AREVA fuel is that total er peaking is not significantly impacted.

2.3 Void Fraction and Distribution XCOBRA-IIIC model calculates the local thermal and hydraulic conditions for input to the B correlation. While local conditions of enthalpy, quality, flow rate and pressure are ciated with a code-calculated local void fraction, the void fraction is not input to the DNB elation. The DNB correlation is approved over a local quality range, but it is not a direct ction of void fraction. Therefore, there is no explicit limit set on average or local void fraction ond that implied in the test conditions used to develop the DNB correlation.

2.4 Coolant Flow Distribution 2.4.1 Coolant Flow Distribution and Bypass Flow minimum primary coolant flow rate at full power conditions is given in Table 3.5-1.

cing the coolant flow path in Figure 3.1-1, the coolant enters the four inlet nozzles and flows the annular plenum between the reactor vessel and core support barrel. It then flows down the ulus between the reactor vessel and core barrel and up through the flow skirt to the plenum w the core lower support structure. The skirt and lower support structure help to even out the t flow distribution to the core. The coolant passes through the openings in the lower core plate flows axially through the fuel assemblies. A portion of the coolant passes through the lower plate and into the guide tubes in the fuel assemblies. The fuel assembly alignment plate is not led through in guide tube locations without CEAs; therefore, core bypass flow is limited in e guide tubes. After passing through the core, the coolant flows into the region outside the trol element assembly shrouds. From this region, the coolant flows across the control element mbly shrouds and passes out through the outlet sleeves on the core barrel to the outlet nozzles.

coolant which does not contact any fuel rods is termed core bypass coolant. The following are principal core bypass routes:

a. Direct inlet to outlet coolant flow at the joint between the core support barrel sleeve and reactor vessel nozzle.

3.5-5 Rev. 35

c. Coolant flow in the region between the core support barrel and core shroud.
d. Coolant flow from the inlet nozzle region through the alignment keyways to the vessel head region.

le 3.5-1 gives the best estimate value for the core bypass flow rate as a fraction of the total ary flow rate. Taking into account the core bypass flow rate, the core flow rate, which is the ctive flow rate for heat transfer, can be calculated from the total primary coolant flow rate.

2.4.2 Core Flow Distribution core flow distribution (CFD) analysis is performed to assess cross flow between assemblies he core for use in subsequent MDNBR subchannel analyses. A full core model provides cross-boundary conditions to a full assembly model at the assembly boundaries. MDNBRs are puted from a full assembly simulation.

he analysis, each fuel assembly in the Millstone Unit 2 core is modeled as a hydraulic channel.

calculations are performed with the XCOBRA-IIIC computer code (Reference 3.5-2). Cross between adjacent assemblies in the open lattice core is directly modeled. The single-phase coefficients are used in the CFD analyses to hydraulically characterize the assemblies in the s computational procedure is designed to evaluate thermal-hydraulic conditions during boiling non-boiling conditions. One-dimensional, two phase separated, slip flow is assumed in the OBRA-IIIC calculation. These assumptions are valid only if the cross flow between necting channels is small compared to the axial velocities in the individual channels. Because ll cross flow does exist, mathematical models have to be postulated for both turbulent and ersion cross-flow mixing. Models of the two-phase state are also defined in terms of void tion, which is a function of enthalpy, flow rate, heat flux, pressure, and axial position. This putational procedure is not applicable when large blockages exist in the fuel bundles since leads to considerable cross flow which cannot be adequately represented by the

-dimensional analysis.

le 3.5-1 summarizes the reactor and fuel design parameters used in these CFD calculations subsequent MDNBR analyses.

2.5 Pressure Losses and Hydraulic Loads 2.5.1 Pressure Losses fuel assembly irrecoverable pressure losses have been calculated using standard loss fficient methods and results from model tests. The pressure loss across the AREVA fuel mbly was determined based on the results of Reference 3.5-5 and analyses.

3.5-6 Rev. 35

2.5.2.1 Hydraulic Loads on Vessel Internal Components design hydraulic loads for the internal components for steady state operating conditions are d in Table 3.5-2. These loads were derived from analysis and from reactor flow model and ponent test results. All hydraulic loads in Table 3.5-2 are based on the maximum expected em flow rate and a coolant temperature of 500°F. When these hydraulic loads are used in the ctural analysis, they are adjusted for coolant temperature. The worst condition (i.e., coolant perature) is not necessarily the same for each internal component; therefore, the loads are sted to reflect the difference in coolant temperature. This is done to ensure the design raulic stresses are acceptable during start-up and during power operation.

types of loads considered in the analysis are: (1) steady-state drag and impingement loads, (2) fluctuating loads induced by pump pressure pulsations, turbulence, and vortex shedding.

of these loads are not exerted on each internal component, but each component sees at least of the loads. Table 3.5-2 lists the components and type of loads that are exerted on them.

2.5.2.2 Core Hydraulic Loads/Fuel Assembly Liftoff holddown spring force and the assembly weight force prevent the fuel assembly from lifting the core support plate during reactor steady-state operation, based on the most adverse bination of component dimensional and material property tolerances. In addition, the ddown springs are designed to accommodate the additional load associated with a pump rspeed transient (resulting in possible temporary liftoff of the fuel assemblies), and to continue nsure fuel assembly holddown following such occurrences. The limiting reactor steady-state ditions are the 4th pump startup conditions. These correspond to the minimum temperature and imum pressure and coolant flow for reactor startup. Thermal expansion of the reactor vessel fuel assembly is also considered.

2.6 Correlation and Physical Data erence 3.5-1 describes the correlations and physical data employed in heat transfer ulations performed by RODEX2. Reference 3.5-7 describes the correlations and physical data loyed in the hydraulic calculations performed by XCOBRA-IIIC. Reference 3.5-3 describes correlations and physical data employed in the DNB correlation.

2.7 Plant Parameters for Thermal-Hydraulic Design plant parameters considered include total primary coolant flow rate, vessel inlet temperature, mary pressure, and core thermal power. Two sets of thermal-hydraulic conditions are defined:

inal conditions and design conditions. Nominal plant conditions represent the best estimate the primary coolant flow rate, pressure, and vessel inlet temperature and do not include wances for instrument errors. Design plant conditions represent the lower limit on primary rate when uncertainties in system resistance and pump head are included, and represent the er limit on vessel inlet temperature when design margins on steam generator performance are 3.5-7 Rev. 35

h the design plant parameters. During steady state operation, the possible variations in these meters define an operating envelope. One combination of these parameters gives the NBR, and this combination is utilized in Chapter 14 as the initial conditions in transient and dent analysis. Table 3.5-1 lists the nominal plant parameters.

2.8 Summary of Thermal and Hydraulic Parameters thermal and hydraulic parameters for the reactor are listed in Table 3.5-1.

3 THERMAL AND HYDRAULIC EVALUATION 3.1 Analytical Techniques and Uncertainties 3.1.1 XCOBRA-IIIC DNBR Analyses thermal-hydraulic simulations employed to evaluate the MDNBR were performed in ordance with AREVAs Nuclear Regulatory Commission (NRC) approved thermal-hydraulic hodology for mixed cores (Reference 3.5-8).

MDNBR performance of the core during anticipated transients will be demonstrated to meet thermal-hydraulic design criterion on DNBR through the performance of transient analysis of limiting events. The results of this analysis are included in Chapter 14.

3.1.2 Parameter Uncertainties les 14.0.7-2 through 14.0.7-5 identify parameter uncertainties included in the AREVA thermal hydraulic and DNB methodology. Plant instrument calibration procedures and related cification requirements are designed so that these uncertainties do not increase.

3.2 Hydraulic Instability Analysis ling flows may be susceptible to thermohydrodynamic instabilities. These instabilities are esirable in reactors since they may cause a change in thermohydraulic conditions that may to a reduction in the DNB heat flux or to undesired forced vibrations of core components.

wever, unlike in Boiling Water Reactors (BWRs), hydraulic stability is not a concern in PWR

s. This statement, which is discussed below, is supported by the literature and the state of the on instabilities occurring in two-phase flow systems.

abilities in vertical up-flow of a two-phase mixture in a heated channel can be broadly sified into several categories. Of these, the following relevant instabilities are discussed.

1. Flow Excursion 3.5-8 Rev. 35

curve (internal characteristic) becomes smaller than the slope of the loop supply pressure drop-flow rate curve (external characteristic), i.e.,

( P -) internal < d--------------

d-------------- ( P -) ----- external dG dG where P is the pressure drop and G is the mass flow rate.

In this manner, a negative flow perturbation will be amplified as the internal pressure drop becomes larger than the external at the perturbed flow and the flow decelerates further until a stable point is reached.

If the core is considered as a single average channel, the external pressure and flow characteristics as seen by the core exhibit

( P -) ----- external < 0 d--------------

dG due to the pump characteristics. This negative slope is stabilizing.

On the other hand, considering flow in a single limiting bundle, the other parallel flow paths impose a flat pressure drop versus flow relation where d(P)/dG = 0.

While this situation is less stable than the average core assumption, it is mitigated by the cross flow and mixing between this limiting bundle and the neighboring bundles. Ref. 3.5-11 shows experimentally a definite stabilizing influence of cross flow mixing.

The internal pressure drop versus flow characteristics were shown to satisfy the Ledinegg stability criterion

( P -) ----- internal > 0 d--------------

dG for a wide range of conditions in the LOFT reactor (Ref. 3.5-12) which closely approximates a PWR core during nominal and worst case operating conditions.

Therefore, in conclusion, Ledinegg Instability is not a concern in PWR cores.

2. Density Wave Instability Dynamic instabilities may occur even when the static stability criterion is satisfied (pressure drop increases when flow increases). For a density wave dynamic 3.5-9 Rev. 35

increase is delayed. In the case of a sinusoidal inlet flow perturbation of particular frequency, the lagging pressure drop response is such that its instantaneous value supports the growth of the initial perturbation (Ref. 3.5-13). Such unstable behavior requires the delayed portion of the total pressure drop (in the two-phase region) to be large compared with the single-phase pressure drop. The onset of this instability depends on the operating conditions and the distribution of pressure drop along the channel, as well as the external loop characteristics. A vast body of literature and several computer programs for the analysis of density waves exists mainly for BWR concerns (see for example the collection of papers in Ref.

3.5-14). Inferences from BWR experience are drawn to dismiss the possibility of density wave instabilities in a PWR core:

  • Unlike a BWR, there is no riser section contributing significantly to the 2-phase pressure drop.
  • For a single limiting channel with a constant pressure drop boundary condition, the cross flow in a PWR core has a stabilizing effect.
  • Density wave oscillations are known to be stabilized with increasing pressure (decreasing enthalpy and density difference between the two phases). No unstable density wave oscillations could be obtained for pressures higher than 1200 psia (Ref. 3.5-15).
  • BWR oscillations occur when the saturated boiling boundary is low (elevation <<4 feet). For a PWR, such boiling boundary can be achieved at nominal flow rates by more than doubling the power, which leaves a considerable stability margin even for the worst case transient.
  • Considering the nuclear coupling, the void-reactivity coefficient in a PWR is reduced when the coolant is borated. Such reduction in the void-reactivity coefficient is stabilizing to this mode of oscillation.
  • For a density wave coupled with an out-of-phase neutron flux oscillation mode, the large subcritical reactivity of the first flux harmonic stabilizes this mode of hydraulic-neutronic oscillation. This is due to the PWR core being small compared with typical BWR cores.

The LOFT reactor stability study also addressed the density wave oscillations and concluded that these are not likely (Ref. 3.5-12).

In conclusion, Density Wave Instability is not a concern in PWR cores.

3. Flow Pattern Transition Instability 3.5-10 Rev. 35

channel experiences a succession of high void and low void flows as a vapor slug passes through (Ref. 3.5-12). As a vapor slug clears the channel exit, the average void content in the channel is temporarily reduced and vice versa resulting in pressure drop and flow rate oscillations. In a worst case condition in a PWR, slug flow may occur in a small number of channels near the exit. No significant oscillatory response is expected, particularly since the slug formation is limited to a short segment near the exit of the hot channels.

The more common meaning of the Flow Pattern Transition Instability refers to unstable transitions between bubbly and annular flow (Ref. 3.5-10). A flow rate perturbation decreasing the flow rate and increasing the void fraction will result in flow transition from bubbly-slug to annular pattern. The annular flow is characterized with lower pressure drop, which results in accelerating the flow. The increase in flow rate brings the void fraction back below the value required to support annular flow. Thus the transition back to bubbly-slug regime takes place.

Extensive work has been done on flow pattern transition (see for example Ref.

3.5-16). Most work was limited to pressures of 1000 psia and below where these transitions are more distinct. At higher pressures, Hosler (Ref. 3.5-17) notes for 1400 and 2000 psia, that the flow appears more homogeneous with no reliable observation of pattern transition.

Weisman et. al. (Ref. 3.5-18) observed no premature DNB due to bubbly-to-slug flow transition which they expected as the range of tested void fractions covers the transition range. Hosler (Ref. 3.5-17), on the other hand, noted that CHF occurred via a film dryout mechanism in established annular flow, which is far from the transition boundary to bubbly-slug pattern.

In conclusion, Flow Pattern Transition Instability is not a concern in PWR cores.

3.3 Core Hydraulics 3.3.1 Fuel Assembly Pressure Drop Coefficients ssure drop coefficients for the AREVA reload fuel presented are derived from pressure drop s performed in AREVAs portable hydraulic test facility (Reference 3.5-5). The pressure drop fficients are for the liquid phase and are referenced to the bare rod flow area.

reload Batches M (Cycle 10), N (Cycle 11), and P (Cycle 12) the pressure drop coefficient for lower tie plate/spacer combination includes the effects of a debris resistant spacer. The reload ch R (Cycle 13) and S (Cycle 14) fuel assemblies implemented an alternative debris resistant gn which has a slightly lower pressure drop across the lower tie plate/spacer combination pared to the Batch P and prior fuel assemblies. As a result, for Cycles 13 and 14, this results 3.5-11 Rev. 35

to crossflow effects, the decreased flow will equilibrate with adjacent assemblies within the t one or two spacers. Limiting MDNBRs occur toward the top of the core. Therefore, the slight stribution in the inlet flows, due to the new lower tie plate and adjacent spacer, will not affect ulated MDNBRs.

Batch T and later HTP fuel assemblies have a lower total pressure drop than the previous etallic fuel assemblies (i.e., Batch S and prior). A thermal hydraulic compatibility analysis performed in Reference 3.5-23 for HTP fuel assemblies co-resident with bimetallic fuel mblies in the Millstone Unit 2 core. This analysis demonstrates that the two fuel assembly es are compatible. Of note is that the core pressure drop would decrease by approximately

% from the all bimetallic core (Cycle 14) to an all HTP core. The core pressure drop decrease m Cycle 14 to Cycle 15 will be approximately 0.69% since the Cycle 15 core has 80 HTP tch T) fuel assemblies and 137 bimetallic (Batch N, P, and R) fuel assemblies. Use of the HMP cer in the lowermost position (Reload Y and later) has a negligible effect on core differential sure.

3.3.2 Guide Tube Bypass Flow and Heating Analysis guide tube thermal-hydraulic design calculations are performed to demonstrate adequate ling of the CEA fingers and to ensure that bypass flow through the guide tubes does not uly reduce core flow.

w enters the guide tube through the weep holes and cap screw and exits through the top of the de tube. In the Millstone Unit 2 core, there are 81 assemblies under CEA positions. Of these, assemblies are under active CEA positions. The CEA fingers extend a short distance into the de tube in these 73 assemblies at the all-rods-out (ARO) position which provides a substantial uction in the guide tube bypass flow. The remaining eight assemblies were originally under the length CEAs which have been removed. In these eight assemblies, the flow is unimpeded, e the last flow plugging devices were removed in Cycle 12. The assembly guide tubes of 91 mblies project a short distance into close fitting sockets in the upper alignment plate. The lting flow annulus represents a significant resistance to guide tube bypass flow in these mblies. The remaining 45 core locations are instrument tube locations. In these locations, the pheral guide tubes also project a short distance into close fitting sockets in the upper nment plate. The center guide tube contains instrumentation which produces a flow annulus ch in turn reduces the flow in the center guide tubes.

guide tube model employed in the flow and heating calculations uses loss coefficients to rmine the guide tube flow path hydraulic losses. The core pressure drop at rated power and is employed as the driving force for flow through the guide tube. The model permits ulation of the guide tube configurations described above. The guide tube thermal model udes the effects of coolant heating by gamma deposition and neutron deceleration. The effects heating due to neutron absorption and gamma deposition in the inserted control rod are 3.5-12 Rev. 35

culations were performed to assess the maximum expected guide tube bypass flow ference 3.5-6). At hot full power (HFP), ARO configuration was selected as that resulting in greatest bypass flow. The total core bypass flow, including flow through the guide tubes in this ance, was determined to be less than 4.0 percent of vessel flow. The result confirms that guide bypass flow does not unduly reduce core flow.

assess the adequacy of guide tube cooling, a simulation was also performed for a single mbly with the CEA fully inserted at HFP conditions. The fully inserted CEA fingers stantially increase the hydraulic resistance in the guide tube, and also represent a significant t source. The exit coolant temperature is well below saturation. Heat transfer through the guide wall provides a significant part of the cooling.

ed on the results described above, it is concluded that ample guide tube cooling is afforded by current design, and that bypass flow remains within acceptable limits.

3.3.3 Control Element Assembly Insertion Time Analysis rge data base of CEA insertion time measurements has been obtained at a CE plant similar to lstone Unit 2, with fuel identical in pertinent guide tube design characteristics to the Millstone t 2 AREVA reload fuel. The measurements span a time period during which reload quantities REVA fuel resided in the core. Statistical analysis (Reference 3.5-6) of this data indicates that CEA 90 percent insertion time is equal to or less than 2.5 seconds, which is well below the imum acceptable 90 percent insertion time of 2.75 seconds specified in the Technical cifications.

r 500 CEA insertion time measurements from nine different tests were analyzed. The surements reflect the time required to reach 90 percent insertion from interruption of power to CEA drive mechanism. Approximately six standard deviations separate the mean of the sured CEA insertion time data from the 2.75 second maximum allowable for Millstone Unit h over 500 data points, higher order statistics may also be applied to the data to conclude that rod drop time will be equal to or less than the greatest time measured in the tests with a bability of 99 percent at a 99 percent confidence level. The greatest rod drop time in the tests, oted above, was 2.50 seconds. The AREVA assemblies are, therefore, expected to conform to maximum CEA 90 percent insertion time of 2.75 seconds with a substantial margin.

3.3.4 Fuel Assembly Liftoff hydraulic lift force on the fuel assembly was calculated (Reference 3.5-6) using the drag fficient for a 14 by 14 fuel assembly with bimetallic spacer grids. This value differed slightly Reload Batches M, N, and P (Cycles 10, 11, and 12). The replacement of a bimetallic spacer h a debris resistant Inconel HTP spacer increased the drag while the thermal rounding of the 3.5-13 Rev. 35

1194 pounds. The assembly weight and spring force totals 1801 lbs, thus providing a 607 nd holddown margin. This margin, which is more than half of the worst case steady state lift e, will envelope any minor variation due to the spacer modifications. It will also provide ddown during and after a 20% pump overspeed resulting in a 44% lift force increase. For oad Batch R (Cycle 13) and Batch S, the fuel assembly weight increased by approximately 40 nds and a bimetallic spacer replaced the Inconel HTP spacer, increasing the margin to liftoff.

imilar analysis was performed for the Reload T design. The use of HMP spacers beginning h Reload Y has a negligible effect on lift.

maximum shear stress of 84,062 psi in the holddown springs occurs in the cold reactor dition. This is below the design criterion of 100,000 psi. The stress at reactor operating ditions is 74,188 psi, which is below the criterion of 90,000 psi at operating temperature.

diation may cause some stress relaxation of the Inconel X-750 holddown springs while sing irradiation induced growth of the fuel assemblies. The assembly growth results in higher ng deflection which offsets any radiation induced relaxation of the springs. The springs are ially shrouded in spring cups, which minimize flow-induced vibration of the springs and vent potential fretting wear.

4 TESTS AND INSPECTIONS 4.1 Reactor Testing rmal-hydraulic design criteria are verified during plant startup testing. This is accomplished measuring the primary intrinsic parameters (e.g., levels, pressures, temperatures, flows, tron fluence and differential pressures) and calculating the non-measurable and extrinsic meters (e.g., power level, core peaking factors). During the operating cycle, various mal-hydraulic parameters are periodically monitored to ensure compliance with the Technical cifications.

4.2 AREVA DNB and Hydraulic Testing 4.2.1 DNB Testing ails of the testing supporting the HTP DNB correlation are contained in Reference 3.5-3.

4.2.2 Fuel Assembly Hydraulic Testing gle-phase hydraulic characteristics of the AREVA Millstone Unit 2 fuel assembly were erimentally determined by hydraulic tests (Reference 3.5-5) performed in AREVAs Portable raulic Test Facility (PHTF).

pressure drop testing characterized the component loss/flow coefficients of the lower tie plate luding the inlet hardware), spacers, and the upper tie plate (including the exit hardware).

3.5-14 Rev. 35

were used to drive empirical relationships, which describe the single-phase pressure drops of Millstone Unit 2 fuel assembly and its components.

se test data from Reference 3.5-5 were used to calculate the Batch M, N, and P lower tie plate, cer, and upper tie plate pressure drop coefficients, and the bare rod friction factor. Additional data and analyses were used to determine the Batch R lower tie plate pressure drop coefficient elations. The loss/flow coefficients derived from these tests and calculations are all referenced he bare rod Reynolds Number.

5 REFERENCES 1 XN-NF-81-58(P)(A), Revision 2, and Supplements 1 and 2, RODEX2 Fuel Rod Thermal-Mechanical Response Evaluation Model, March 1984.

2 XN-NF-75-21(P)(A), Revision 2, XCOBRA-IIIC: A Computer Code to Determine the Distribution of Coolant During Steady-State and Transient Core Operation, January 1986.

3 EMF-92-153(P)(A) Rev. 1, HTP: Departure From Nucleate Boiling Correlation for High Thermal Performance Fuel, Siemens Power Corporation, January 2005.

4 XN-75-32(P)(A), Supplements 1, 2, 3, and 4, Computational Procedure for Evaluating Fuel Rod Bowing, October 1983.

5 ANF-89-018(P), Single-Phase Hydraulic Flow Test of ANF Millstone-2 Fuel Assembly, January 1989.

6 ANF-88-088(P), Revision 1, Design Report for Millstone Point Unit 2, Reload ANF-1, August 1988.

7 BNWL-1695, COBRA-IIIC: A Digital Computer Program for Steady-State and Transient Thermal-Hydraulic Analysis of Rod Bundle Nuclear Fuel Elements, March 1973.

8 XN-NF-82-21(P)(A), Revision 1, Application of Exxon Nuclear Company PWR Thermal Margin Methodology to Mixed Core Configurations, September 1983.

9 EMF-2135, Revision 0, Millstone Unit 2 Cycle 13 Extended Shutdown Safety Analysis Report, January 1999.

10 J. A. Boure, A. E. Bergles, and L. S. Tong, Review of Two-Phase Flow Instability, ASME Paper 71-HT-42, August 1971.

3.5-15 Rev. 35

Transfer Conf., pp. 235-239, Tokyo, Japan (September 1974).

12 S. A. Eide, Instability Study for LOFT for L2-1, L2-2 and L2-3 Pretest Steady State Operating Conditions, RE-A-78-096, Idaho National Engineering Laboratory, November 1978.

13 J. March-Leuba, Density-Wave Instabilities in Boiling Water Reactors, Oak Ridge National Laboratory Report ORNL/TM-12130 (September 1992).

14 Proceedings of the International Workshop on Boiling Water Reactor Stability, Committee on the Safety of Nuclear Reactors Installations, OECD Nuclear Energy Agency, Holtsville, NY (October 1990).

15 H. S. Kao, C. D. Morgan, and W. B. Parker, Prediction of Flow Oscillation in Reactor Core Channel, Trans. ANS Vol. 16, pp. 212-213 (1973).

16 A. E. Bergles and M. Suo, Investigation of Boiling Water Flow Regimes at High Pressure, Dynatech Corp. NYO-3304-8 (February 1966).

17 E. R. Hosler, Flow Patterns in High Pressure Two-Phase (Steam-Water) Flow with Heat Addition, 9th National Heat Transfer Conferrence, Chemical Engineering Progress Symposium Series, Number 82, Vol. 64, pp. 54-66 (August 1967).

18 Weisman et. al., Experimental Determination of the Departure from Nucleate Boiling in Large Rod Bundles at High Pressure, 9th National Heat Transfer Conference, Chemical Engineering Progress Symposium Series, Number 82, Vol. 64, pp. 114-125 (August 1967).

19 Reference Deleted 20 Letter, R. I. Wescott (SPC) to C. H. Wu (NU), Transmittal of Bases for New Uncertainties in the Setpoint Analysis for Millstone Unit 2, RIW:97:049, February 27, 1998.

21 Reference Deleted by FSARCR 06-MP2-016.

22 Qualification of Exxon Nuclear Fuel for Extended Burnup, XN-NF-82-06(P)(A)

Revision 1 and Supplements 2, 4 and 5, Exxon Nuclear Company, October 1986.

23 EMF-2664, Rev. 0, Millstone Unit 2 Thermal Hydraulic Compatibility Analysis, January 2002.

3.5-16 Rev. 35

Design and Operating Parameters Value re Rated Power 2700 MWt ction of Heat Generated in Fuel 0.975 mary System Pressure 2250 psia re Inlet Temperature 549°F actor Coolant Flow (Minimum) 360,000 gpm a sembly Pitch 8.18 inches pass Flow Fraction (Best Estimate) 0.0303 erage Linear Heat Rate 6.206 kW/ft tal Number of Assemblies 217 Flow reductions to 349,200 gpm are compensated for by reductions in the FrT and linear heat rate limits.

l Parameters Design and Operating Parameters Value el Rod OD 0.440 inches ide Tube OD (above dashpot) 1.115 inches d Array 14 by 14 d Pitch 0.580 inches mber of Fuel Rods/Assembly 176 mber of Guide Tubes/Assembly 5 tive Fuel Length 136.7 inches tal Fuel Rod Assembly Length 146.25 inches mber of Spacers 9 3.5-17 Rev. 35

Component Load Description Load Value Core Support Barre Radial pressure differential directed inward opposite inlet 40 psi duct Core Support Barrel and Upper Uplift load 480,000 pounds Guide Structure Flow Skirt Radial pressure differential directed inward 6.0 psi average, 10.2 psi maximum, over 40° sector Bottom Plate Pressure differential load directed upward 43,400 pounds Core Support Plate Pressure differential load directed upward 43,100 pounds Fuel Assembly Uplift load 1194 lbs at 120% flow Core Shroud Radial load directed outward 20.8 psi at bottom, 0.0 psi at top Upper Guide Structure Pressure differential load directed upward 148,000 pounds Fuel Alignment Plate Pressure differential load directed upward 89,600 pounds Upper Guide Plate Pressure differential load directed downward 132,000 pounds CEA Shrouds Lateral drag load 4,200 pounds (dual CEA) 1,100 pounds (single CEA) 3.5-18 Rev

BLE 3.5-3 UNCERTAINTY SOURCES FOR DNBR CALCULATIONS (DELETED) 3.5-19 Rev. 35

.1 SEISMIC ANALYSIS

.1.1 Introduction amic analyses of the reactor vessel internals for both horizontal and vertical seismic itation were conducted to provide further bases for assessing the adequacy of their seismic gn. These analyses were performed in conjunction with the dynamic seismic analyses of the tor coolant system (RCS) which is discussed in Appendix 4.A. The following paragraphs vide a discussion of the analytical procedures used for the reactor internals, including a cription of the mathematical models. Significant results are listed and compared to the results ined from application of the design loads.

.1.2 Method of Analysis

.1.2.1 General procedure used in conducting the seismic analysis of the reactor internals consisted basically hree steps. The first step involved the formulation of a mathematical model. The natural uencies and mode shape of the model were determined during the second step. The response he model to the seismic excitation was determined in the third step. In this analysis, the zontal and vertical components of the seismic excitation were considered separately and the imum responses added to obtain conservative results.

.1.2.2 Mathematical Models the dynamic analysis of the reactor internals, equivalent multi-mass mathematical models e developed to represent the system. Since the seismic input excitation of the reactor internals obtained in the form of acceleration time history of the reactor vessel flange, only the rnals are included in the model. The coupling effect of the internals response on the vessel ge acceleration was accounted for by including a simplified representation of the reactor rnals with the model of the RCS. This is discussed in Appendix 4.A. Since the horizontal and ical responses were treated as uncoupled, separate horizontal and vertical models were eloped to more efficiently account for the structural differences in these directions. A sketch of internals showing the relative node locations for the horizontal model is presented in ure 3.A-1. Figures 3.A-2 and 3.A-3 show the idealized horizontal and vertical models. Since structural details provide for no vertical load transfer between the upper guide structure (UGS) core or core shroud, the vertical response of the UGS is independent of the rest of the rnals. Consequently, the vertical model was divided into two submodels. Model I consists of core support barrel/thermal shield (CSB/TS), lower support structure, core shroud and core s; Model II consists of the UGS.

mathematical models of the internals are constructed in terms of lumped masses and elastic m elements. At appropriate locations within the internals, points (nodes) are chosen to lump weights of the structure. Between these nodes, properties are calculated for moments of 3.A-1 Rev. 35

.1.2.2.1 Hydrodynamic Effects dynamic analysis of reactor internals presents some special problems due to their immersion confined fluid. It has been shown both analytically and experimentally (Reference 3.A-1) that ersion of a body in a dense fluid medium lowers its natural frequency and significantly alters vibratory response as compared to that in air. The effect is more pronounced where the fining boundaries of the fluid are in close proximity to the vibrating body as is the case for the tor internals. The method of accounting for the effects of a surrounding fluid on a vibrating em has been to ascribe to the system additional or hydrodynamic mass.

s hydrodynamic mass decreases the frequencies of the system, but is not directly involved in inertia force effects. The hydrodynamic mass of an immersed system is a function of the ensions of the real mass and the space between the real mass and confining boundary.

rodynamic mass effects for moving cylinders in a water annulus are discussed in References

-1 and 3.A-2. The results of these references are applied to the internals structures to obtain total (structural plus hydrodynamic) mass matrix which was then used in the evaluation of the ral frequencies and mode shapes for the model.

.1.2.2.2 Fuel Assemblies the horizontal model, the fuel assemblies are treated as vibrating in unison. The member perties for the beam elements representing the fuel assemblies were derived from the results of erimental tests of the fuel assembly load deflection characteristics and natural frequency.

.1.2.2.3 Core Support Barrel Flanges obtain accurate lateral and vertical stiffnesses of the upper and lower flanges, finite element lyses of these two regions were performed. As shown in Figures 3.A-4 and 3.A-5, the flanges e modeled with quadrilateral and triangular ring elements. Asymmetric loads, equivalent to ral shear loads and bending moments, and symmetric axial loads were applied and the lting displacements calculated. These results were then used to derive the equivalent member perties for the flanges.

.1.2.2.4 Control Element Assembly Shrouds the horizontal model, the control element assembly (CEA) shrouds are treated as vibrating in on and are modeled as guided cantilever beams in parallel. To account for the decreased ral stiffness of the UGS due to local bending of the fuel alignment plate, a short member with perties approximating the local bending stiffness of the fuel alignment plate is included at the om of the CEA shrouds. Since the stiffness of the UGS support plate is large compared to that he shrouds, the CEA shrouds are assumed to be rigidly connected to the UGS support plate.

3.A-2 Rev. 35

the horizontal model, the thermal shield supports are modeled as horizontal members. The mber properties of the beam elements representing the positioning pins were based on the al stiffness of the circumferential set of pins. Likewise, the properties of the beam member esenting the support lugs were based on the tangential stiffness of the circumferential set of

. For the vertical model, the equivalent cross-section area of the bar element representing the port lugs was based on the axial bending stiffness of the circumferential set of lugs. For both horizontal and vertical models, the stiffness of the thermal shield supports includes the effect ocal deformation of the core support barrel.

.1.2.2.6 Upper Guide Structure Support Plate and Lower Support Structure Grid Beams se grid beam structures were modeled as plane grids. Displacements due to vertical (out of e) loads applied at the beam junctions were calculated through the use of the STRUDL puter code (Reference 3.A-3). Average stiffness values based on these results yielded ivalent member cross-section areas for the vertical model.

.1.2.3 Natural Frequencies and Normal Modes mass and beam element properties of the models were utilized in STAR, a computer program m the MRI/STARDYNE Analysis System programs (Reference 3.A-4) to obtain the natural uencies and mode shapes. This system utilizes the stiffness matrix method of structural lysis. The natural frequencies and mode shapes are extracted from the system of equations.

[K-Wn2 M]n = 0 where:

K = Model stiffness matrix M = Model mass matrix Wn = Natural circular frequency for the nth mode n = Normal mode shape matrix for nth mode mass matrix, M, includes the hydrodynamic and structural masses.

natural frequencies and mode shapes calculated for the first 3 modes for the horizontal model presented in Figures 3.A-6 through 3.A-8. The natural frequencies calculated for the vertical del are presented in Table 3.A-1. The modal data shown is typical and is presented for strative purposes. The effect of additional higher modes was included in the response analyses.

3.A-3 Rev. 35

.1.2.4.1 Horizontal Direction time history analysis technique was utilized to obtain the response of the internals for the zontal seismic excitation. The horizontal excitation was specified as the acceleration time ory of the reactor vessel flange, resulting from the operational basis earthquake (OBE) (OBE

.09g ground acceleration). The flange excitation resulting from the design basis earthquake E) (DBE = 0.17g ground acceleration) was conservatively specified as 0.17/0.09 times that the OBE.

time history response analysis was performed utilizing the MRI STARDYNE System/

NRE 1 Computer Program. This program utilizes the Normal Mode Method to obtain time ory response of linear elastic structure. Details of the program and the Normal Mode hod are presented in References 3.A-4, 3.A-5 and 3.A-6.

ut to DYNRE 1 consisted of the modal data as determined in Section 3A.2.3, the modal ping factors, and the forcing function time history. This analysis used the modal data for all des with frequencies below 100 cps. This included the first 14 modes. Contributions from her modes are negligible.

modal damping factors were obtained by the method of Mass Mode Weighting which es:

M i in i n = -------------------------

M i in where:

n = Modal damping factor Mi = Structural mass of mass node i lil = Absolute value of the mode shape as mass mode i i = Damping associated with pass point i damping factor assigned to the nodes representing the fuel assemblies was 5 percent. This is a servative value derived from proprietary experimental results. A value of 1 percent was used the other nodes.

output from the DYNRE 1 code consists of the nodal displacement, velocity, and acceleration e history relative to the base. The member bending moments and shears were obtained from STAR code (Reference 3.A-5) and were derived from the DYNRE 1 nodal displacement tors at the times of peak response.

3.A-4 Rev. 35

response of the reactor internals to the vertical excitation was obtained by the response ctrum technique. Because of the high natural frequencies and resulting low levels of responses the vertical direction, the more conservative spectrum response analysis results were used ead of time history results. The response spectrum utilized was derived from the vertical eleration time history at the reactor vessel flange. The spectrum curve is presented in ure 3.A-9.

acceleration level corresponding to the natural frequency of each mode was selected from the ctrum curve. The response spectrum technique uses these acceleration values to determine the tia forces, accelerations, and displacements of each mode. The results for each mode were servatively combined on the basis of absolute values. For the vertical models, the first seven des were included in the results.

.1.3 Results mbined results for the horizontal and vertical dynamic seismic analyses are presented in le 3.A-2 in terms of stresses at critical locations in the reactor internals for the DBE.

le 3.A-2 also lists the seismic stresses which result from application of the design loads cified for the DBE. A comparison shows the results of the dynamic analysis to be less severe.

.1.4 Conclusion concluded that the seismic loads specified for the design of the internals are adequate. All mic loads calculated by the dynamic seismic analysis are less than the design loads specified he DBE.

.2 NORMAL OPERATING ANALYSIS ign analyses were performed on the reactor internals for normal operating conditions to onstrate that the mechanical design bases were satisfied. These design calculations included ropriate vibration analyses of the component assemblies. The flow induced vibration of the B/TS, during normal operation, was characterized as a forced response to deterministic and dom pressure fluctuations in the coolant. Methods were developed for predicting the response omponents to the hydraulic forcing functions.

phasis was placed on analysis and design of those components which were particularly critical susceptible to vibratory excitation, such as the thermal shield. Using a top supported, as osed to a bottom supported, thermal shield design improves stability as it eliminates a free e in the flow path. Increasing the number of upper supports and lower jackscrews, in the cific manner chosen, provides a much stiffer structure and the use of an all-welded shield inates local flexibilities and relative motion at bolted joints. Analytical studies show the mal shield to be stable on its support system when exposed to the axial annular flow ountered during normal operation. The snubber design is based upon limiting the motion of core support barrel under conditions of hydraulically induced vibrations. The snubbers are at 3.A-5 Rev. 35

ribution of snubbers assures restraint regardless of the direction of response.

random hydraulic forcing function was developed by analytical and experimental methods.

analytical expression was developed to define the turbulent pressure fluctuation for fully eloped flow. This expression was modified, based upon the result of scale model testing, to ount for the fact that flow in the downcomer was not fully developed. Based upon test results, expression was developed to define the spatial dependency of the turbulent pressure tuations. In addition, experimentally adjusted analytical expressions were developed to ne; the peak value of the pressure spectral density associated with the turbulence and; the imum area of coherence, in terms of the boundary layer displacement, across which the dom pressure fluctuations are in phase.

natural frequencies and mode shapes of the CSB/TS system were obtained using the ymmetric shell finite element computer program, ASHSD (Reference 3.A-7). This computer gram is capable of obtaining natural frequencies and mode shapes of complex axisymmetric ls; e.g., arbitrary meridional shape, varying thickness, branches, multi-materials, orthotropic erial properties, etc. To employ the ASHSD code, the CSB/TS were modeled as a series of ical shell frustrums joined at their nodal point circles. The length of each element, throughout ASHSD model, was a fraction of the shell decay length. Since rapid changes in the stress ern occur in regions of structural discontinuity, the nodal point circles were more closely ced in such regions. The finite element model of the CSB/TS system included representation he core support barrel upper and lower flanges, sections of different wall thickness, and mal shield support lugs and jackscrews. Elements with orthotropic material properties were zed to provide equivalent axisymmetric models of the structural stiffness and constraints to tive motion between the core support barrel and thermal shield provided by the thermal shield port lugs and jackscrews. Those modes which reflect the mass of the lower support structure, shroud and fuel were simulated by the addition of concentrated masses at specific nodes in core support barrel flange finite element model.

lying Hamiltons Variational Principle to the conical shell elements an equation of motion formulated for each degree of freedom of the system. An inverse iteration technique was zed in the program to obtain solutions to the characteristic equation, which was based on a onalized form of a consistent mass matrix and stiffness matrix developed using the finite ment method. Four degrees of freedom radial displacement, circumferential displacement, ical displacement, and meridional rotation were taken into account in the analysis, giving to coupled mode shapes and corresponding frequencies. Evaluation of the reduction of these uencies for the system immersed in coolant was made by means of the virtual mass method ined in Reference 3.A-2.

random response analysis considers the response of the CSB/TS system to the turbulent ncomer flow during steady-state operation. The random forcing function is assumed to be a e-band stationary random process with a pressure spectral density equal to the peak value ciated with the turbulence. The rms vibration level of the CSB/TS system was obtained based n a damped, single degree of freedom analysis assuming the rms random pressure fluctuation 3.A-6 Rev. 35

eloped by a Combustion Engineering (CE) consultant using the random loads discussed ve. Modeling the reactor vessel snubbers and core support barrel system as a single degree of dom spring-mass system, the number and magnitude of snubber, core support barrel impacts calculated based upon the response of the system to random excitation. The snubbers were gned, based upon this loading requirement, to meet the cyclic strength requirements specified ection III of the ASME Boiler and Pressure Vessel Code.

forced response of the reactor internals to deterministic loading was evaluated by classical lytical methods, using lumped mass and continuous elastic structural models. These calculated onses were used to verify the structural integrity of the reactor vessel internals to normal rating vibratory excitation. Components were design analyzed to assure that there were no erse effects from dominant excitation frequencies, such as pump rotational and blade passing uencies.

.3 LOSS OF COOLANT ACCIDENT ANALYSIS

.3.1 Discussion ynamic analysis (Reference 3.A-8) has been performed to determine the structural response of reactor vessel internals to the transient loss of coolant accident (LOCA) loading. The analysis rmined the shell, beam and rigid body motions of the internals using established puterized structural response analyses. The finite-element computer code, ASHSD ference 3.A-7) was used to calculate the time-dependent beam and shell response of the CSB/

system to the transient LOCA loading. The finite-element computer code SAMMSOR-NASOR (Reference 3.A-9) was used to evaluate the core support barrels potential for kling when loaded by a net external radial pressure resulting from an outlet line break. The ctural response of the reactor internals to vertical and transverse loads resulting from inlet and et breaks, was determined using the spring-mass computer code, SHOCK (Reference 3.A-10).

time and space dependent pressure loads used in the above analysis were the result of a iled hydraulic blowdown analysis. The pressure fluctuations were determined for each node he hydraulic model for inlet and outlet line breaks. The pressure time histories at these nodal tions were then decomposed into the Fourier harmonics which define the circumferential sure distribution at the nodal elevations. Where the hydraulic model nodes did not correspond hose of the structural model, the hydraulic model pressure components were interpolated to vide the required loading information.

finite element computer code, ASHSD, was used to calculate the dynamic response of the B/TS to transient LOCA loading resulting from an inlet break. To employ the ASHSD code, CSB/TS were modeled as a series of conical shell frustrums (elements) joined at their nodal nt circles. Applying Hamiltons Variational Principle to the conical shell elements a damped ation of motion was formulated for each degree of freedom of the system. Four degrees of dom radial displacement, circumferential displacement, vertical displacement and idional rotation were taken into account in the analysis, giving rise to coupled modes. The 3.A-7 Rev. 35

h that it is small compared to the shortest period of the finite element system. The model eloped for the CSB/TS system is shown in Figure 3.A-10. The length of each element, ughout the analytical model, was a fraction of the shell decay length. Since rapid changes in stress pattern occur in regions of structural discontinuity, the nodal point circles were more ely spaced in such regions. The finite element model of the CSB/TS system included esentation of the core support barrel upper and lower flanges, sections of different wall kness, and thermal shield support lugs and jackscrews. Elements with orthotropic material perties were utilized to provide equivalent axisymmetric models of the structural stiffness and straints to relative motion between the core support barrel and thermal shield provided by the mal shield support lugs and jackscrews. Those modes which reflect the mass of the lower port structure, core shroud and fuel were stimulated by the addition of concentrated masses at cific nodes in the core support barrel flange finite element model.

erforming the dynamic analysis of the CSB/TS system, the transient load harmonics were lied in two successive phases to account for time-dependent boundary conditions at the bbers. The first phase used those harmonics which excite the beam modes, whereas the second se used those harmonics which excite the shell modes. During the first phase, the lower end of core support barrel was unrestrained. Within a very few milliseconds, the clearances between core support barrel and reactor vessel snubbers were closed and for the remainder of the CA transient, the core support barrel was restrained radially at the snubber level. Transient onses were computed throughout each loading phase.

ASHSD code computed the nodal point displacement, resultant shell forces, shell stresses maximum principle stresses as functions of time. The maximum principle stresses at the rnal and external surfaces of the CSB/TS were determined from the bending and membrane ponents during each phase of transient loading. Stress intensity levels calculated from the ciple stresses were combined with normal operating and seismic induced stresses for parison with design criteria.

urate representation and analysis of the CSB/TS shell structures was obtained through use of finite element code ASHSD. Accurate representation of the remainder of the internals (i.e.,

, core shroud, CEAs, UGS, lower support structure, etc.) was obtained using the SHOCK e.

SHOCK code determines the response of structures which are represented as lumped-mass ems and subjected to arbitrary loading functions. The code solves the differential equations of ion for each mass by a numerical step-integration procedure. The lumped mass model can esent a vertically or laterally responding system subject to arbitrary loading functions and al conditions. Options are available for describing steady state loads, preloads, input elerations, linear and nonlinear springs (including tension and compression only springs) gaps, structural and viscous damping.

reactor internals were developed in terms of a spring-mass system for both vertical and lateral ctions; see Figures 3.A-11 and 3.A-12. For both models, the spring rates were generally 3.A-8 Rev. 35

del analyses. The lumped mass weights were generally based upon the mass distribution of the orm support structures, but included at appropriate nodes, local masses such as snubber cks, fuel end fittings, thermal shield lugs, etc. The net result was a lumped-mass system having same distribution of mass as the actual structure. To simulate the effect associated with the rnals oscillating laterally in the water filled vessel, a distributed virtual mass was calculated ed upon the procedure outlined in Reference 3.A-8 (which includes the annulus effect) and added to the structural lumped-mass system, to provide an analytical model with a dynamic onse quantitatively similar to the actual internals. In the case of the vertical model, the raulic effect is notably one of reducing the effective weight of the reactor internals and this ct was included in the structural lumped-mass system.

SHOCK code provided excellent facility for modeling clearances, preloads and component rfaces. In the lateral model, the core support barrel, reactor vessel snubber clearance was ulated by a nonlinear spring which accounted for the increased resistance to core support el motion when snubbing occurred. In the vertical model, nonlinear springs in the form of pression only springs, were used extensively to simulate preload and interface conditions, h as exist between the UGS support plate and core support barrel upper flange; at the fuel d-down spring; at the fuel, core support plate interface and at the core shroud, core support e interface. Tension only springs were used to simulate the effect of the core shroud tie rods.

oth the vertical and lateral SHOCK models, damping was varied throughout the system to ulate structural and hydraulic frictional effects within the reactor internals. The effect of raulic drag in the vertical model was simulated by a force time-history applied to the fuel er end-fitting. Vertical loads were used directly from the detailed hydraulic analysis, whereas ral loads were obtained by integrating those harmonics which excite the beam modes to obtain net lateral load on the CSB/TS system.

SHOCK code calculated the vertical and lateral response of the system in terms of lacements, velocities and accelerations and internal force, moments and shears as related to h model. These quantities were sufficient to permit calculation of membrane and where ropriate bending stresses for comparison with design criteria.

finite-element code SAMMSOR-DYNASOR was used to determine the dynamic response of core support barrel, with initially imperfect geometry, to a net external radial pressure lting from an outlet line break. The above analysis has the capability of determining the linear dynamic response of axisymmetric shells with initial imperfections subjected to trarily varying load configurations.

ce SAMMSOR-DYNASOR is a finite-element program, a model was developed, Figure 3.A-of the core support barrel using axisymmetric finite-elements similar to those used for the HSD analysis. As was for the ASHSD model, the SAMMSOR-DYNASOR finite-element ths were considerably less than the decay length of the core support barrel. The boundary dition at the core support barrel flange was considered fixed, whereas at the core support el lower flange radial displacements were restrained. These boundary conditions represented 3.A-9 Rev. 35

alignment plate, core shroud and core support plate were neglected.

ce the basic phenomenon in buckling is nonlinear instability, the initial deviation of the cture from a perfect geometry greatly affects its response. The initial imperfection was applied he core support barrel by means of a pseudo-load so developed to provide the maximum erfection over each of the desired number of circumferential harmonics. The actual transient ing in terms of its harmonics was applied to the initially imperfect geometry core support el and the response obtained for each of the imperfection harmonics for the combined loading monics.

.3.2 Analysis Codes HSD (Reference 3.A-7) is a structural finite-element computer code developed at the versity of California, Berkeley, and supported in part by the National Science Foundation. It orms dynamic analyses of complex axisymmetric structures subjected to arbitrary dynamic ings or base accelerations. The frequencies of free vibrations as calculated by ASHSD pare well to those calculated by the equations of Hermann-Mirshy and Flugge, erences 3.A-11 and 3.A-12, respectively. The authors also make comparisons with available erimental results (Reference 3.A-13) of free vibrations of cylindrical shells. The resulting parison is good. Comparison of the numerical solution (Reference 3.A-14) of the dynamic onse of a shell to suddenly applied loads and the finite-element (ASHSD) solution of the e problem are in good agreement. The response of a shell to a moving axisymmetric pressure was evaluated by ASHSD and analytically (Reference 3.A-15) with the results being in good ement.

MMSOR-DYNASOR (Reference 3.A-9) is a finite-element computer code developed at Texas M University and supported in part by a NASA grant from the Manned Spacecraft Center, ston, Texas. This code has the capability of determining the nonlinear dynamic response of ymmetric shells subjected to arbitrary dynamic loads. Asymmetrical dynamic buckling can be stigated using this program. The program has been extensively tested, using problems the tions to which have been reported by other researchers, in order to establish the validity of the es. Among these are a shallow shell with axisymmetric loading as described in Reference 3.A-Identical results are obtained with those of Reference 3.A-17 for the analytical evaluation of t loadings on a cylindrical shell. Calculations made by SAMMSOR-DYNASOR for the metric buckling of a shallow spherical cap is in good agreement with the analyses of erences 3.A-18 and 3.A-19 and the experimental data of References 3.A-20 and 3.A-21.

BOR DRASTIC, (Reference 3.A-22) is a structural finite-element computer code eloped at the Aeroelastic and Structures Research Laboratory, Department of Aeronautics at Massachusetts Institute of Technology. The work was administered by the Air Force Systems mmand with technical monitoring by the Aerospace Corp. SABOR 5 - DRASTIC is the end lt of combining a finite-difference solution procedure and a finite-element program to permit dicting the transient response of complex shells of revolution which are subjected to arbitrary sient loadings. Comparisons with reliable independent analytical predictions (notably finite-3.A-10 Rev. 35

lysis were performed by the Aerospace Corp. (Reference 3.A-23) to verify the ability of the e to account for a complex geometry shell of revolution subjected to transient asymmetric

s. Loads were applied by means of well-defined explosive charges. Based upon the results of amic strain measurements made on the test structure, it is evident that the SABOR 5 -

ASTIC code is capable of solving complex dynamic shell structure problems successfully.

eveloping the above finite-element computer codes, (i.e., ASHSD, SAMMSOR-DYNASOR, BOR 5 - DRASTIC) the authors have independently verified their codes with respect to the lts of other established structural programs, classical solutions and as possible to experimental

. The correlations demonstrate that the above programs are capable of solving complex amic shell structure problems successfully and that the finite-element method of modeling vides accurate representation of the structural phenomena. The SABOR 5 - DRASTIC code, ch has had extensive and successful analytical and experimental correlation (Reference 3.A-6) transient (explosive) asymmetric loading, was used to analyze a core support barrel structure h short-term loading. The results of this well-verified program are identical to these of the te-element codes ASHSD and SAMMSOR-DYNASOR (which are used in the LOCA lysis) for the same core support barrel problem, demonstrating the ability of these programs to quately represent and evaluate the effect of a transient load on an axisymmetric structure like core support barrel.

.4 EFFECTS OF THERMAL SHIELD REMOVAL owing the discovery of the thermal shield support degradation at the end of Cycle 5 in July, 3, the thermal shield was removed. A detailed inspection of the core barrel revealed damage at thermal shield support lug locations. Repairs to the core barrel comprised of drilling crack stor holes at the ends of through-wall cracks and removal by machining of non through-wall ks.

lytical evaluations and assessments were performed to demonstrate continued structural quacy of the reactor internals without the thermal shield for all design loading conditions.

cial attention was paid to the core barrel to justify the repairs. A description of the repairs to core barrel, analyses, and significant results is given in Reference 3.A-24.

conclusion, there was no significant change in the loads and the stresses in the internal ctures remained within the ASME Code allowables.

.5 LEAK-BEFORE-BREAK ANALYSIS k-Before-Break (LBB) analyses for the reactor coolant system (RCS) main coolant loops, for pressurizer surge line, and unisolable RCS portions of the safety injection and shutdown ling piping, which demonstrated that the probability of fluid system piping rupture was emely low, was reviewed and approved by the commission. (See References 3.A-25 through

-29.) Subsequent to the commission review and approval, weld overlays were applied to imilar metal welds (DMWs) at the shutdown cooling, the safety injection and the pressurizer 3.A-11 Rev. 35

ve piping segments, including the effects of pipe whipping and discharging fluids have been luded from the design basis of the following reactor vessel and reactor internals components:

Core barrel snubbers, core barrel stabilizer blocks Reactor vessel core support ledge Reactor Cavity Seal Plate, Neutron Shielding

.6 REFERENCES

-1 Fritz, R. J., and Kiss, E., The Vibration Response of a Cantilevered Cylinder Surrounded by an Annular Fluid, KAPL-M-6539, February 1966.

-2 Kiss, E., Analysis of the Fundamental Vibration Frequency of a Radial Vane Internal Steam Generator Structure, ANL-7685, Proceedings of Conference on Flow-Induced Vibrations in Reactor System Components, May 1970, Argonne National Laboratory, Argonne, IL.

-3 ICES STRUDL-II, The Structural Design Language Engineering Users Manual.

-4 MRI/STARDYNE - Static and Dynamic Structural Analysis System: User Information Manual, Control Data Corporation, June 1, 1970.

-5 MRI/STARDYNE User Manual, Computer Methods Department, Mechanics Research, Inc., Los Angeles, California, January 1, 1970.

-6 Hurty, W. C., and Rubinstein, M. F., Dynamics of Structures, Chapter 8, Prentice Hall, Inc., Englewood Cliffs, New Jersey, 1964.

-7 Ghosh, S., Wilson, E., Dynamic Stress Analysis of Axisymmetric Structures Under Arbitrary Loading, Dept. No. EERC 69-10, University of California, Berkeley, September 1969.

-8 CENPD-42, Topical Report on Dynamic Analysis of Reactor Vessel Internals Under Loss of Coolant Accident Conditions with Application of Analysis to C-E 800 Mw(e)

Class Reactors, August 1972.

-9 Tillerson, J. R., Haisler, W. E., SAMMSOR II - A Finite Element Program to Determine Stiffness and Mass Matrices of Shells-of- Revolution, Texas A&M University, TEES-RPT-70-18, October 1970. DYNASOR II - A Finite Element Program for the Dynamic Nonlinear Analysis of Shells-of-Revolution, Texas A&M University, TEES-RPT-70-19, October 1970.

-10 Gabrielson, V. K., SHOCK - A Computer Code for Solving Lumped-Mass Dynamic Systems, SCL-DR-65-34, January 1966.

3.A-12 Rev. 35

78, P. 563-568, 1956.

-12 Flugge, W., Stresses in Shells, Third Printing, Springer-Verlag, New York, 1966.

-13 Koval, L. R., Cranch, E. I., On the Free Vibrations of Thin Cylindrical Shells Subjected to Initial Torque, Proceedings of the U. S. National Congress of Applied Mechanics, P.

11, 1962.

-14 Reismann, H., and Padloy, J., Forced, Axisymmetric Motions of Cylindrical Shells, Journal of the Franklin Institute, Vol. 284, Number 5, November 1967.

-15 Tang, Sing-Chih, Response of a Finite Tube to Moving Pressure, Journal Engineering Mechanics Division, ASCE, Vol. 93, Number EM3, June 1967.

-16 Klein, S., and Sylvester, R. J., The Linear Elastic Dynamic Analysis of Shells of Revolution by the Matrix Displacement Method, Air Force Slight Dynamics Laboratory, TR-66-80, 1966, P. 299-329.

-17 Johnson, D. E., Grief, R., Dynamic Response of a Cylindrical Shell: Two Numerical Methods, AIAA Journal, Vol. 4, Number 3, March 1966, P. 486-494.

-18 Huang, N. C., Axisymmetric Dynamic Snap-through of Elastic Clamped Shallow Spherical Shells, AIAA Journal, Vol. 7, Number 2, February 1969, P. 215-220.

-19 Stephen, W. B., and Fulton, R. E., Axisymmetric Static and Dynamic Buckling of Spherical Caps due to Centrally Distributed Pressures, Paper 69-89, AIAA Journal, 1969.

-20 Lock, M. H., Okrebo, S., and Whittier, J. S., Experiment of the Snapping of a Shallow Dome Under a Step Pressure Loading, AIAA Journal, Vol. 6, No. 7, July 1968, P. 1320-1326.

-21 Stricklin, J. A., and Martinez, J. E., Dynamic Buckling of Clamped Spherical Caps Under Step Pressure Loadings, AIAA Journal, Vol. 7, Number 6, June 1969, P. 1212-1213.

-22 Kotanchik, J. J., et al., The Transient Linear Elastic Response Analysis of Complex Thin Shells of Revolution Subjected to Arbitrary External Loadings, by the Finite-Element Program SABOR 5 - DRASTIC, AD-709-189, Massachusetts Institute of Technology, April 1970.

-23 Klein, S., A Static and Dynamic Finite Element Shell Analysis with Experimental Verification, International Journal for Numerical Methods in Engineering, Vol. 3, P.

299-315, 1971.

3.A-13 Rev. 35

License No. DPR-65, December, 1983.

-25 NRC Letter from D. G. McDonald, Jr. to M. L. Bowling, Jr., Revised Evaluation of the Primary Cold Leg Piping Leak - Before-Break Analysis for the Millstone Nuclear Power Station, Unit Number 2, dated November 9, 1998.

-26 NRC Letter from D. G. McDonald, Jr. to M. L. Bowling, Jr., Application of Leak -

Before-Break Status to the Portions of the Safety Injection and Shutdown Cooling System for the Millstone Nuclear Power Station, Unit Number 2, dated November 9, 1998.

-27 NRC Letter from B. Eaton to R. P. Necci, Staff Review of the Submittal by Northeast Nuclear Energy Company to Apply Leak-Before-Break Status to the Pressurizer Surge Line, Millstone Nuclear Power Station, Unit 2, dated May 4, 1999.

-28 NRC Letter from G.S. Vissing to J.F. Opeka, Application of Reactor Coolant System Leak-Before-Break Analysis, dated September 1, 1992.

-29 Federal Register/Volume 53, No. 66/April 6, 1988, 10 CFR Part 50 Leak Before Break Technology Solicitation of Public Comment on Additional Applications.

-30 Structural Integrity Associates Report: 0901238.401, Revision 0, dated: December 2010, Updated Leak-Before-Break Evaluation of Weld Overlaid Hot Leg Surge, Shutdown Cooling and Safety Injection Nozzles for Millstone Nuclear Power Station, Unit 2.

3.A-14 Rev. 35

MATHEMATICAL MODEL ode Number Sub-Model I Frequency, cps Sub-Model II Frequency, cps 1 21.60 72.98 2 67.75 404.09 3 124.59 -

3.A-15 Rev. 35

COMPONENTS FOR THE DESIGN BASIS EARTHQUAKE Dynamic Structural Design Load Analysis Component Location Stress Mode Stress Stress re Support Barrel Upper Section of Tension & Bending 1,129 psi 746 psi Barrel wer Core Beam Flange Bending 5,278 psi 929 psi pport A Shrouds: End of Shroud Tension & Bending 3,548 psi 1,295 psi gle A Shrouds: Dual End of Shroud Tension & Bending 2,762 psi 697 psi per Grid Beams Center of Beam Bending 1,652 psi 127 psi per Guide Junction of Flange & Tension & Bending 2,823 psi 146 psi ucture Flange Barrel Cylinder 3.A-16 Rev. 35

MPS-2 FSAR FIGURE 3A-1 REPRESENTATIVE NODE LOCATIONS - HORIZONTAL MATHEMATICAL MODEL

  • MASS NOPE o MASSLESS NODE April 1990 Rev. 26.2

MPS-2 FSAR FIGURE 3A-2 MATHEMATICAL MODEL - HORIZONTAL SEISMIC ANALYSIS R.V. FLANGE

  • MASS NODE MASSLESS NODE 48 0 13 RIGID CONNECTING LINK UPPER 12 l' HINGED GUIDE CONNECTION 4S STRUCTURE 44 43 11 . CORE 4 - - SUPPORT 42 33 10 BARREL 34 32 *9

. 31 8 30 19 29 7 CORE 28 SHROUD FUlL ASSY's 'l1 17 THERMAL 26 SHIELD 16 25 24 15 21 23 14 22 20 LOWER SUPPORT ~

STRUCTURE April 1990 Rev. 26.2

MPS-2 FSAR FIGURE 3A-3 MATHEMATICAL MODEL - VERTICAL SEISMIC ANALYSIS

  • MASS NODE o MASSLES S NODE 4f' c::::::;, RIG 10 CON NECTING 51 ;2 ~ 50 LINK 4 4CJ '48 4 47 46 MODEL II

~ 45 44 UPPER 43 MODEL I

  • GUIDE ~ 42 CORE STRUCT URE III SUPPORT
  • 41 FUEL BARREL 37 ~ 39'AlIGNMENT 40 PLATE 35 4 33 30 31 428 26 Z7 24 I

MODEL I 22 23 20 THERMAL 16 17 SHltLD

  • 18 14 FUEL MASS 12 4 13 AND FUEL 9 /ALIGNMENT 11 PLATE 4~ 8 6 5 LOWER SUPPORT STRUCTUR E.

... I 4 3

42 1

April 1990 Rev. 26.2

MPS-2 FSAR FIGURE 3A-4 CORE SUPPORT BARREL UPPER FLANGE - FINITE ELEMENT MODEL R

~Z April 1990 Rev. 26.2

MPS-2 FSAR FIGURE 3A-5 CORE SUPPORT BARREL LOWER FLANGE - FINITE ELEMENT MODEL f-

,\

\ \

\

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.Jz R

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April 1990 Rev. 26.2

MPS-2 FSAR FIGURE 3A-6 LATERAL SEISMIC MODEL - MODE 1, 3.065 CPS MODE 1 FREQUENCY 3.065 cps April 1990 Rev. 26.2

MPS-2 FSAR FIGURE 3A-7 LATERAL SEISMIC MODEL - MODE 2, 5.118 CPS MOOE 2 FREQUENCY 5.118 cps I I

J I

  • I t

I I

f I I T 1 I *I I 1 f I I I  ; ,

  • I I I

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J I I 1

*I 1 I

I I I

+ 1 I I T I I iL.

t I 1 I

I I L_ *f April 1990 Rev. 26.2

MPS-2 FSAR FIGURE 3A-8 LATERAL SEISMIC MODEL - MODE 2, 5.118 CPS MODE 3 FREQUENCY 8.166 cps f

r I

  • I I

~

I I

I April 1990 Rev. 26.2

MPS-2 FSAR FIGURE 3A-9 REACTOR VESSEL FLANGE VERTICAL RESPONSE SPECTRUM (1% DAMPING) 0.30 0.25

  • 0.20

~

~

D:: 0.15

~

LIJ U

U 4:

0.10 o.m 0.10 1.0 FREQUENCY. CDS April 1990 Rev. 26.2

MPS-2 FSAR FIGURE 3A-10 ASHSD FINITE ELEMENT MODEL OF THE CORE SUPPORT BARREL/THERMAL SHIELD SYSTEM 60 80 100 120 140 160 180 200 220 240 260 280 300 320 320.5 o pz,2f 4~ I I I I I I I I I I I -I I I I I. in.

  • IS 20 25 30 35 40 SO 60 70 74.0 --fl.' I I I I f I I i I /' .

8~.S I

85 l~Det.i1A csa R,In.

Snubber Elevation NoJe LaNer

6. 125 ~ff1bers .Jocksc; reN' 74.0 - .;.,.1 9 '0 11 8.**rl c::= -69.0 7- .- 128 6- 127.

5- 12y 4-3- 125

2. 1-75* 75 85.S- .......--~

I 321. 125 328.5 Detoll A Detail 8 Upper Flang~ lower Flange April 1990 Rev. 26.2

MPS-2 FSAR FIGURE 3A-11 VERTICAL SHOCK MODEL

t Expansion

~U;7;~:p:::-;:~-:---_--i Compensating ~J*ng

~

~

Ten.sion Only er Guide Support ....__...

.~

1\

Sprang S ructure Plate FI ange ~CSB

.t Control L Upper Flange

."" Element ~

}Cs0rTlpresslo n Only Shroud

~ pring Assemblies Fu.eJ ~ Core 5upport Alignment Barrel Plale Support L~gs

\ i .l

~ Thermal I .Shield

~,

£ Core

~ Shroud 1.

t

~

~

Cor S ~ Lower Suppo*rt Plat~ upport 1 ~ structure April 1990 Rev. 26.2

MPS-2 FSAR FIGURE 3A-12 LATERAL SHOCK MODE Upper Guide Sup~rt Structure Plate Assy Control Element Core Support Shroud Assemblies Barrel lugs Fuel Alignment Plate Core Shroud Core Support Plate legend Lower Support Struet168 o Mass Point Connected By Lateral Springs April 1990 Rev. 26.2

MPS-2 FSAR FIGURE 3A-13 SAMMSOR DYNASOR FINITE ELEMENT MODEL OF CORE SUPPORT BARREL Upper Barrel

/ Core Region Lower Barrel April 1990 Rev. 26.2

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