ML21305A093
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DCPP UNITS 1 & 2 FSAR UPDATE Chapter 4 Reactor CONTENTS Section Title Page 4.1
SUMMARY
DESCRIPTION 4.1-1 4.
1.1 REFERENCES
4.1-3 4.2 MECHANICAL DESIGN 4.2-1 4.2.1 FUEL 4.2-2 4.2.1.1 Design Bases 4.2-3 4.2.1.2 Fuel Rods 4.2-3 4.2.1.3 Fuel Assembly Structure 4.2-14 4.2.1.4 Operational Experience 4.2-24 4.2.1.5 Safety Evaluation 4.2-24 4.2.1.6 Tests and Inspections 4.2-25 4.2.2 REACTOR VESSEL INTERNALS 4.2-26 4.2.2.1 Design Bases 4.2-26 4.2.2.2 Acceptance Criteria 4.2-27 4.2.2.3 Reactor Vessel Internals Description 4.2-27 4.2.2.4 Reactor Vessel Internals Design Evaluation 4.2-32 4.2.2.5 Safety Evaluation 4.2-34 4.2.3 REACTIVITY CONTROL SYSTEM 4.2-36 4.2.3.1 Design Bases 4.2-36 4.2.3.2 Reactivity Control System Acceptance Criteria 4.2-37 4.2.3.3 Reactivity Control System Description 4.2-39 4.2.3.4 Reactivity Control System Design Evaluation 4.2-47 4.2.3.5 Safety Evaluation 4.2-54 4.2.3.6 Tests and Inspections 4.2-56 4.2.3.7 Instrumentation Applications 4.2-57 4.
2.4 REFERENCES
4.2-58 4.3 NUCLEAR DESIGN 4.3-1 4.3.1 DESIGN BASES 4.3-2 4.3.1.1 General Design Criterion 10, 1971 - Reactor Design 4.3-2 4.3.1.2 General Design Criterion 11, 1971 - Reactor Inherent Protection 4.3-2 4.3.1.3 General Design Criterion 12, 1971 - Suppression of i Revision 24 September 2018
DCPP UNITS 1 & 2 FSAR UPDATE Chapter 4 Reactor CONTENTS Section Title Page Reactor Power Oscillations 4.3-2 4.3.1.4 General Design Criterion 25, 1971 - Protection System Requirements for Reactivity Control Malfunctions 4.3-2 4.3.1.5 General Design Criterion 26, 1971 - Reactivity Control System Redundancy and Capability 4.3-2 4.3.1.6 General Design Criterion 28, 1971 - Reactivity Limits 4.3-2 4.3.2 NUCLEAR DESIGN ACCEPTANCE CRITERIA 4.3-3 4.3.2.1 Fuel Burnup 4.3-3 4.3.2.2 Control of Power Distribution 4.3-3 4.3.2.3 Negative Reactivity Feedbacks (Reactivity Coefficients) 4.3-3 4.3.2.4 Stability 4.3-3 4.3.2.5 Maximum Controlled Reactivity Insertion Rate 4.3-3 4.3.2.6 Shutdown Margins 4.3-4 4.
3.3 DESCRIPTION
4.3-4 4.3.3.1 Nuclear Design Description 4.3-4 4.3.3.2 Power Distribution 4.3-5 4.3.3.3 Reactivity Coefficients 4.3-16 4.3.3.4 Control Requirements 4.3-19 4.3.3.5 Control 4.3-22 4.3.3.6 Control Rod Patterns and Reactivity Worths 4.3-24 4.3.3.7 Criticality of Fuel Assemblies 4.3-25 4.3.3.8 Stability 4.3-26 4.3.3.9 Vessel Irradiation 4.3-29 4.3.3.10 Analytical Methods 4.3-30 4.3.4 SAFETY EVALUATION 4.3-34 4.3.4.1 General Design Criterion 10, 1971 - Reactor Design 4.3-34 4.3.4.2 General Design Criterion 11, 1971 - Reactor Inherent Protection 4.3-35 4.3.4.3 General Design Criterion 12, 1971 - Suppression of Reactor Power Oscillations 4.3-36 4.3.4.4 General Design Criterion 25, 1971 - Protection System Requirements for Reactivity Control Malfunctions 4.3-37 4.3.4.5 General Design Criterion 26, 1971 - Reactivity Control System Redundancy and Capability 4.3-37 4.3.4.6 General Design Criterion 28, 1971 - Reactivity Limits 4.3-37 ii Revision 24 September 2018
DCPP UNITS 1 & 2 FSAR UPDATE Chapter 4 Reactor CONTENTS Section Title Page 4.
3.5 REFERENCES
4.3-38 4.4 THERMAL AND HYDRAULIC DESIGN 4.4-1 4.4.1 DESIGN BASES 4.4-1 4.4.1.1 General Design Criterion 10, 1971 - Reactor Design 4.4-1 4.4.1.2 General Design Criterion 12, 1971 - Suppression of Reactor Power Oscillations 4.4-1 4.4.2 THERMAL AND HYDRAULIC DESIGN ACCEPTANCE CRITERIA 4.4-2 4.4.2.1 Departure from Nucleate Boiling Acceptance Criteria 4.4-2 4.4.2.2 Fuel Temperature Acceptance Criteria 4.4-2 4.4.2.3 Core Flow Acceptance Criteria 4.4-2 4.4.2.4 Hydrodynamic Stability Acceptance Criteria 4.4-2 4.4.3 SYSTEM DESCRIPTION 4.4-2 4.4.3.1 Summary Comparison 4.4-2 4.4.3.2 Fuel Cladding Temperatures 4.4-2 4.4.3.3 Departure from Nucleate Boiling Ratio 4.4-6 4.4.3.4 Flux Tilt Considerations 4.4-12 4.4.3.5 Void Fraction Distribution 4.4-13 4.4.3.6 Core Coolant Flow Distribution 4.4-13 4.4.3.7 Core Pressure Drops and Hydraulic Loads 4.4-13 4.4.3.8 Correlation and Physical Data 4.4-14 4.4.3.9 Thermal Effects of Operation Transients 4.4-16 4.4.3.10 Uncertainties in Estimates 4.4-16 4.4.3.11 Plant Configuration Data 4.4-18 4.4.3.12 Core Hydraulics 4.4-19 4.4.3.13 Influence of Power Distribution 4.4-21 4.4.3.14 Core Thermal Response 4.4-22 4.4.3.15 Analytical Techniques 4.4-22 4.4.3.16 Hydrodynamic and Flow-Power Coupled Instability 4.4-26 4.4.3.17 Temperature Transient Effects Analysis 4.4-27 4.4.3.18 Potentially Damaging Temperature Effects During Transients 4.4-28 4.4.3.19 Energy Release During Fuel Element Burnout 4.4-29 4.4.3.20 Energy Release During Rupture of Waterlogged Fuel Elements4.4-29 4.4.3.21 Fuel Rod Behavior Effects from Coolant Flow Blockage 4.4-29 iii Revision 24 September 2018
DCPP UNITS 1 & 2 FSAR UPDATE Chapter 4 Reactor CONTENTS Section Title Page 4.4.4 THERMAL AND HYDRAULIC DESIGN EVALUATION 4.4-30 4.4.4.1 Departure from Nucleate Boiling 4.4-30 4.4.4.2 Fuel Temperature 4.4-31 4.4.4.3 Core Flow 4.4-32 4.4.4.4 Hydrodynamic Stability 4.4-32 4.4.5 SAFETY EVALUATION 4.4-32 4.4.5.1 General Design Criterion 10, 1971 - Reactor Design 4.4-32 4.4.5.2 General Design Criterion 12, 1971 - Suppression of Reactor Power Oscillations 4.4-32 4.4.6 TESTS AND INSPECTIONS 4.4-33 4.4.6.1 Testing Prior to Initial Criticality 4.4-33 4.4.6.2 Initial Power Plant Operation 4.4-33 4.4.6.3 Component and Fuel Inspections 4.4-33 4.4.7 INSTRUMENTATION APPLICATIONS 4.4-33 4.4.7.1 Incore Instrumentation 4.4-33 4.4.7.2 Overtemperature and Overpower T Instrumentation 4.4-34 4.4.7.3 Instrumentation to Limit Maximum Power Output 4.4-34 4.
4.8 REFERENCES
4.4-35 iv Revision 24 September 2018
DCPP UNITS 1 & 2 FSAR UPDATE Chapter 4 TABLES Table Title 4.1-1 Reactor Design Comparison 4.1-2 Analytical Techniques in Core Design 4.1-3 Design Loading Conditions for Reactor Core Components 4.2-1 Deleted in Revision 23 4.3-1 Nuclear Design Parameters (Typical) 4.3-2 Unit 1 - Reactivity Requirements for Rod Cluster Control Assemblies 4.3-3 Unit 2 - Reactivity Requirements for Rod Cluster Control Assemblies 4.3-4 Axial Stability Index PWR Core With a 12-ft Height (Historical) 4.3-5 Typical Neutron Flux Levels (n/cm2 sec) at Full Power 4.3-6 Comparison of Measured and Calculated Doppler Defects 4.3-7 Benchmark Critical Experiments 4.3-8 Saxton Core II Isotopics, Rod MY, Axial Zone 6 4.3-9 Critical Boron Concentrations, at HZP, BOL 4.3-10 Comparison of Measured and Calculated Rod Worth 4.3-11 Comparison of Measured and Calculated Moderator Temperature Coefficients at HZP, BOL (Historical) 4.4-1 Unit 1 - Void Fractions at Nominal Reactor Conditions with Design Hot Channel Factors 4.4-2 Unit 2 - Void Fractions at Nominal Reactor Conditions with Design Hot Channel Factors.
4.4-3 Comparison of THINC-IV and THINC-I Predictions with Data from Representative Westinghouse Two- and Three-loop Reactors (Historical) 4.4-4 Non-LOCA DNB Analysis Method v Revision 24 September 2018
DCPP UNITS 1 & 2 FSAR UPDATE Chapter 4 FIGURES Figure Title 4.2-1(a) Core Cross Section Outline 4.2-2(a) 17x17 Fuel Assembly Outline (LOPAR) 4.2-2A Deleted in Revision 23 4.2-3(a) Fuel Rod Assembly Outline (LOPAR) 4.2-3A Deleted in Revision 23 4.2-4 Typical Clad and Pellet Dimensions as a Function of Exposure 4.2-5 Representative Fuel Rod Internal Pressure and Linear Power Density for the Lead Burnup Rod as a Function of Time 4.2-6 Removable Rod Compared to Standard Rod (HISTORICAL) 4.2-7 Removable Fuel Rod Assembly Outline (HISTORICAL) 4.2-8 Location of Removable Rods Within an Assembly (HISTORICAL) 4.2-9 Unit 1 - Lower Core Support Assembly 4.2-10 Unit 2 - Lower Core Support Assembly 4.2-11 Unit 2 - Neutron Shield Pad Lower Core Support Structure 4.2-12 Unit 1 - Upper Core Support Structure 4.2-13 Unit 2 - Upper Core Support Structure 4.2-14 Plan View of Upper Core Support Structure 4.2-15 Rod Cluster Control and Drive Rod Assembly with Interfacing Components 4.2-16(a) Rod Cluster Control Assembly Outline 4.2-17(a) Absorber Rod 4.2-18 Deleted in Revision 23 vi Revision 24 September 2018
DCPP UNITS 1 & 2 FSAR UPDATE Chapter 4 FIGURES Figure Title 4.2-18A(a) Wet Annular Burnable Absorber 4.2-19 Deleted in Revision 23 4.2-20 Deleted in Revision 23 4.2-21(a) Secondary Source Assembly 4.2-21A Deleted in Revision 23 4.2-22(a) Thimble Plug Assembly 4.2-23 Deleted 4.2-23A Deleted 4.2-24(a) Control Rod Drive Mechanism Schematic 4.2-24A Deleted 4.2-25 Nominal Latch Clearance at Minimum and Maximum Temperature 4.2-26 Control Rod Drive Mechanism Latch Clearance Thermal Effect 4.3-1 Fuel Loading Arrangement 4.3-2 Production and Consumption of Higher Isotopes 4.3-3 Boron Concentration vs. Cycle Burnup With Burnable Absorber Rods 4.3-4 Burnable Absorber Rod Arrangement Within an Assembly 4.3-5 Typical Integral Fuel Burnable Absorber Rod Arrangement Within an Assembly 4.3-6 Burnable Absorber Loading Pattern 4.3-7 Normalized Power Density Distribution Near Beginning of Life (BOL),
Unrodded Core, Hot Full Power, No Xenon vii Revision 24 September 2018
DCPP UNITS 1 & 2 FSAR UPDATE Chapter 4 FIGURES Figure Title 4.3-8 Normalized Power Density Distribution Near BOL, Unrodded Core, Hot Full Power, Equilibrium Xenon 4.3-9 Unit 1 - Normalized Power Density Distribution Near BOL, Group D at Insertion Limit, Hot Full Power, Equilibrium Xenon 4.3-10 Unit 2 - Normalized Power Density Distribution Near BOL, Group D at Insertion Limit, Hot Full Power, Equilibrium Xenon 4.3-11 Normalized Power Density Distribution Near Middle of Life (MOL),
Unrodded Core, Hot Full Power, Equilibrium Xenon 4.3-12 Normalized Power Density Distribution Near End of Life (EOL),
Unrodded Core, Hot Full Power, Equilibrium Xenon 4.3-13 Rodwise Power Distribution in a Typical Assembly (G-10) Near BOL, Hot Full Power, Equilibrium Xenon, Unrodded Core 4.3-14 Rodwise Power Distribution in a Typical Assembly (G-10) Near EOL, Hot Full Power, Equilibrium Xenon, Unrodded Core 4.3-15 Possible Axial Power Shapes at BOL Due to Adverse Xenon Distributions 4.3-16 Possible Axial Power Shapes at MOL Due to Adverse Xenon Distributions 4.3-17 Possible Axial Power Shapes at EOL Due to Adverse Xenon Distributions 4.3-18 Deleted 4.3-19 Deleted 4.3-20 Deleted 4.3-21 Peak Power Density During Control Rod Malfunction Overpower Transients 4.3-22 Peak Linear Power During Boration/Dilution Overpower Transients viii Revision 24 September 2018
DCPP UNITS 1 & 2 FSAR UPDATE Chapter 4 FIGURES Figure Title 4.3-23 Maximum F QT x Power vs. Axial Height During Normal Operations 4.3-24 Deleted in Revision 23 4.3-25 Comparison Between Calculated and Measured Relative Fuel Assembly Power Distribution 4.3-26 Comparison of Calculated and Measured Axial Shape 4.3-27 Measured Values of F QT for Full Power Rod Configurations 4.3-28 Doppler Temperature Coefficient at BOL and EOL 4.3-29 Doppler Only Power Coefficient at BOL and EOL 4.3-30 Doppler Only Power Defect at BOL and EOL 4.3-31 Moderator Temperature Coefficient at BOL, No Rods 4.3-32 Moderator Temperature Coefficient at EOL 4.3-33 Moderator Temperature Coefficient as a Function of Boron Concentration at BOL, No Rods 4.3-34 Hot Full Power Moderator Temperature Coefficient for Critical Boron Concentration 4.3-35 Total Power Coefficient at BOL and EOL 4.3-36 Total Power Defect at BOL and EOL 4.3-37 Unit 1 - Rod Cluster Control Assembly Pattern 4.3-38 Unit 2 - Rod Cluster Control Assembly Pattern 4.3-39 Accidental Simultaneous Withdrawal of Two Control Banks EOL, HZP Banks B and D Moving in the Same Plane 4.3-40 Design - Trip Curve ix Revision 24 September 2018
DCPP UNITS 1 & 2 FSAR UPDATE Chapter 4 FIGURES Figure Title 4.3-41 Normalized Rod Worth vs. Percent Insertion, All Rods But One 4.3-42 Axial Offset vs. Time, PWR Core with a 12-ft Core Height and 121 Assemblies 4.3-43 XY Xenon Test Thermocouple Response Quadrant Tilt Difference vs.
Time 4.3-44 Calculated and Measured Doppler Defect and Coefficients at BOL, for a Two-loop Plant with a 12-ft Core Height and 121 Assemblies 4.3-45 Comparison of Calculated and Measured Boron Concentration for a Two-loop Plant with a 12-ft Core Height and 121 Assemblies 4.3-46 Comparison of Calculated and Measured Boron for a Two-loop Plant with a 12-ft Core Height and 121 Assemblies 4.3-47 Comparison of Calculated and Measured Boron in a Three-loop Plant with a 12-ft Core Height and 157 Assemblies 4.4-1 Peak Fuel Average and Surface Temperatures During Fuel Rod Lifetime vs. Linear Power Density 4.4-2 Peak Fuel Centerline Temperature During Fuel Rod Lifetime vs. Linear Power Density 4.4-3 Thermal Conductivity of UO2 (Data Corrected to 95% Theoretical Density) 4.4-4 Axial Variation of Average Clad Temperature for Rod Operating at 5.43 kW/ft 4.4-5 Probability Curves for W-3 and R Grid DNB Correlations 4.4-6 TDC vs. Reynolds Number for 26-inch Grid Spacing 4.4-7 Normalized Radial Flow and Enthalpy Distribution at 4-ft Elevation 4.4-8 Normalized Radial Flow and Enthalpy Distribution at 8-ft Elevation x Revision 24 September 2018
DCPP UNITS 1 & 2 FSAR UPDATE Chapter 4 FIGURES Figure Title 4.4-9 Normalized Radial Flow and Enthalpy Distribution at 12-ft Elevation Core Exit 4.4-10 Void Fraction vs. Thermodynamic Quality H-HSAT/HG-HSAT 4.4-11 PWR Natural Circulation Test 4.4-12 Comparison of a Representative W Two-loop Reactor Incore Thermocouple Measurements with THINC-IV Predictions 4.4-13 Comparison of a Representative W Three-loop Reactor Incore Thermocouple Measurements with THINC-IV Predictions 4.4-14 Hanford Subchannel Temperature Data Comparison With THINC-IV 4.4-15 Hanford Subcritical Temperature Data Comparison With THINC-IV 4.4-16 Unit 1 - Distribution of Incore Instrumentation 4.4-17 Unit 2 - Distribution of Incore Instrumentation 4.4-18 Improved Thermal Design Procedure Illustration 4.4-19 Measured Versus Predicted Critical Heat Flux-WRB-1 Correlation (HISTORICAL) 4.4-20 Measured Versus Predicted Critical Heat Flux-WRB-2 Correlation NOTE:
(a)
This figure corresponds to a controlled engineering drawing that is incorporated by reference into the UFSAR. Refer to Table 1.6-1 for the correlation between the UFSAR figure number and the corresponding controlled engineering drawing number.
xi Revision 24 September 2018
DCPP UNITS 1 & 2 FSAR UPDATE Chapter 4 REACTOR This chapter describes the design for the reactors at Diablo Canyon Power Plant (DCPP) Unit 1 and Unit 2, and evaluates their capabilities to function safely under all operating modes expected during their lifetimes.
4.1
SUMMARY
DESCRIPTION This chapter describes the following subjects: (a) the mechanical components of the reactor and reactor core, including the fuel rods and fuel assemblies, reactor vessel internals, and the control rod drive mechanisms (CRDMs), (b) the nuclear design, and (c) the thermal-hydraulic design.
The reactor core of each unit consists of VANTAGE+ fuel assemblies.
The significant mechanical design features of the VANTAGE+ design, as defined in Reference 1, include the following:
Integral fuel burnable absorber (IFBA)
Intermediate flow mixer (IFM) grids Protective grid assemblies (P-Grid)
Reconstitutable top nozzle (RTN)
Axial blanket Debris filter bottom nozzle (DFBN)
The core is cooled and moderated by light water at a nominal pressure of 2250 psia to preclude bulk boiling under normal operating conditions. The coolant uses boron as a neutron absorber. Boron concentration in the coolant is varied as required to control relatively slow reactivity changes, such as those associated with the fuel burnup.
Additional boron, in the form of IFBA or burnable absorber rods may be employed to limit the moderator temperature coefficient (MTC) and/or the local power peaking that can be achieved.
A fuel assembly consists of up to 264 mechanically joined fuel rods in a 17 x 17 square array. The fuel rods are supported at intervals along their length by grid assemblies that maintain the lateral spacing between the rods throughout the design life of the assembly. The grid assembly consists of an "egg-crate" arrangement of interlocked straps. The straps contain springs and dimples for maintaining fuel rod lateral and axial support, as well as mixing vanes at the top of the straps for coolant mixing. The fuel 4.1-1 Revision 24 September 2018
DCPP UNITS 1 & 2 FSAR UPDATE rods consist of enriched UO2 cylindrical pellets contained in zirconium alloy tubing that is plugged and seal-welded at the ends. To increase fatigue life, all fuel rods are pressurized with helium during fabrication to reduce stress and strain.
The center position of the fuel assembly contains an instrument tube; the remaining 24 positions in the array are equipped with guide thimbles joined to the grids and the top and bottom nozzles. Depending on assembly position in the core, the guide thimbles are used as core locations for rod cluster control assemblies (RCCAs), neutron source assemblies, and burnable absorber rods (if used).
The DFBN is a box-like structure that serves as a bottom structural element of the fuel assembly and directs the coolant flow to the assembly. The pattern and size of the flow holes in the bottom nozzle are designed to reduce the possibility of fuel rod damage due to debris-induced fretting. This feature is discussed in Section 4.2.1.3.2.1.
The top nozzle assembly functions as the upper structural element of the fuel assembly in addition to providing a partial protective housing for the RCCA or other components.
Each RCCA consists of a group of individual absorber rods fastened at the top end to a common hub or spider assembly.
The CRDMs for the RCCA are of the magnetic latch type. The latches are controlled by three magnetic coils. Upon a loss of power to the coils, the RCCA is released and falls by gravity to shut down the reactor.
Components of the reactor vessel internals are divided into three parts: (a) the lower core support structure (including the entire core barrel, the Unit 1 thermal shield, and the Unit 2 neutron shield pad assembly), (b) the upper core support structure, and (c) the incore instrumentation support structure. Reactor vessel internals support the core, maintain fuel alignment, limit fuel assembly movement, maintain alignment between fuel assemblies and CRDMs, direct coolant flow past the fuel assemblies to the pressure vessel head, provide gamma and neutron shielding, and provide guides for incore instrumentation.
The nuclear design analyses and evaluations establish physical locations for fuel assemblies, control rods, burnable absorber, and physical parameters such as fuel enrichments and boron concentration in the coolant. These characteristics, together with the reactor control and protection systems and the emergency core cooling system (ECCS), provide adequate reactivity control even if the RCCA with the highest reactivity worth is stuck in the fully withdrawn position ensuring that the reactor performance and safety criteria specified in Section 4.2 are met.
The thermal-hydraulic design analyses and evaluations establish coolant flow parameters that ensure adequate heat transfer between fuel cladding and reactor coolant. The thermal-hydraulic design takes into account local variations in dimensions, power generation, flow distribution, mixing and the IFM grids in the VANTAGE+ fuel 4.1-2 Revision 24 September 2018
DCPP UNITS 1 & 2 FSAR UPDATE assembly. The mixing vanes incorporated in the fuel assembly spacer grid design induce additional flow mixing between the various flow channels within a fuel assembly as well as between adjacent assemblies.
Instrumentation is provided in and out of the core to monitor the nuclear, thermal-hydraulic, and mechanical performance of the reactor, and to provide input signals to control and protection functions and to the plant computer.
Table 4.1-1 presents a comparison of the reactor design parameters for the DCPP Unit 1 and Unit 2 reactor cores fueled with VANTAGE+ fuel assemblies. The analysis techniques employed in the core design are tabulated in Table 4.1-2. Design loading conditions for reactor core components are tabulated in Table 4.1-3.
4.
1.1 REFERENCES
- 1. S. L. Davidson (Ed.), et al., VANTAGE+ Fuel Assembly Reference Core Report, WCAP-12610-P-A, April 1995.
4.1-3 Revision 24 September 2018
DCPP UNITS 1 & 2 FSAR UPDATE 4.2 MECHANICAL DESIGN For design purposes, the DCPP conditions are divided into four categories, in accordance with their anticipated frequency of occurrence and risk to the public, as follows:
(1) Condition I - Normal Operation (2) Condition II - Incidents of Moderate Frequency (3) Condition III - Infrequent Faults (4) Condition IV - Limiting Faults In general, Condition I occurrences are accommodated with margin between any plant parameter and the value of that parameter which would require either automatic or manual protective action. Condition II incidents are accommodated with, at most, a shutdown of the reactor with the plant capable of returning to operation after corrective action.
The release of radioactive material due to Condition III incidents should not be sufficient to interrupt or restrict public use of areas outside the exclusion area. Furthermore, a Condition III incident shall not, by itself, generate a Condition IV fault or result in a consequential loss of function of the reactor coolant system (RCS) or reactor containment barriers.
Condition IV occurrences are faults that are not expected to occur, but are defined as limiting faults that must be considered in design. Condition IV faults shall not cause a release of radioactive material that results in an undue risk to public health and safety.
The reactor is designed so that its components meet the following performance and safety criteria:
(1) The mechanical design of the reactor core components and their physical arrangement, together with corrective actions by the reactor control, protection, and emergency cooling systems (when applicable) ensure that:
(a) Fuel damage is not expected during Conditions I and II events, although a very small number of fuel rod failures is anticipated. This number of failures is within the capability of the plant cleanup system and is consistent with the plant design bases. Fuel damage is defined as penetration of the fission product barrier (i.e., the fuel rod cladding).
4.2-1 Revision 24 September 2018
DCPP UNITS 1 & 2 FSAR UPDATE (b) The reactor can be brought to a safe state following a Condition III event with only a small number of fuel rods damaged, although sufficient fuel damage may occur to preclude resumption of operation without considerable outage time.
(c) The reactor can be brought to a safe state and the core can be kept subcritical with acceptable heat transfer geometry, following transients arising from Condition IV events.
(2) The fuel assemblies are designed to withstand loads induced during shipping, handling, and core loading without exceeding the criteria of Section 4.2.1.3.2.5.
(3) The fuel assemblies are designed to accept control rod insertions to provide the reactivity control required for power operations and shutdown conditions (if in such core locations).
(4) All fuel assemblies have provisions for the insertion of the incore instrumentation necessary for plant operation (if in such core locations).
(5) The reactor internals, in conjunction with the fuel assemblies, direct reactor coolant through the core to achieve acceptable flow distribution and to restrict bypass flow so that the heat transfer performance requirements can be met for all modes of operation. In addition, internals provide core support and distribute coolant flow to the pressure vessel head. The distribution of flow into the vessel head minimizes axial and circumferential temperature gradients, thus precluding excessive rotation or warpage that could result in leakage past the O-ring gaskets during Conditions I and II operations. Required inservice inspections can be carried out since the internals are removable and provide access to the inside of the pressure vessel.
4.2.1 FUEL The fuel assembly and fuel rod design data are listed in Table 4.1-1. U.S. Nuclear Regulatory Commission (NRC) approval of the VANTAGE+ design is given in Reference 29. Figure 4.2-1 shows a cross-section of the fuel assembly array, and Figure 4.2-2 shows a fuel assembly full-length outline. The fuel rods are loaded into the fuel assembly structure so that there is clearance between the fuel rod ends and the top and bottom nozzles.
Each fuel assembly is installed vertically in the reactor vessel and stands upright on the lower core plate, which is fitted with alignment pins to locate and orient the assembly.
After all fuel assemblies are set in place, the upper support structure is installed.
Alignment pins, built into the upper core plate, engage and locate the upper ends of the 4.2-2 Revision 24 September 2018
DCPP UNITS 1 & 2 FSAR UPDATE fuel assemblies. The upper core plate then bears downward against the fuel assemblies' top nozzles, via the holddown springs, to hold the fuel assemblies in place.
4.2.1.1 Design Bases 4.2.1.1.1 General Design Criterion 2, 1967 - Performance Standards The reactor core is designed to withstand, without fuel and/or clad damage that could interfere with continued effective core cooling, the effects of, or is protected against, natural phenomena such as earthquakes.
4.2.1.1.2 General Design Criterion 10, 1971 - Reactor Design The fuel is designed with appropriate margin to assure that specified acceptable fuel design limits (SAFDLs) are not exceeded during any condition of normal operation, including the effects of anticipated operational occurrences.
4.2.1.1.3 Safety Function Requirements (1) Loads During Handling The fuel assemblies are designed to accommodate conditions expected to exist as a result of handling during assembly, inspection, and refueling operations, as well as shipping loads.
4.2.1.1.4 10 CFR 50.46(b)(4) - Coolable Geometry The fuel is designed with appropriate margin to assure that following a postulated loss-of-coolant accident (LOCA) the calculated core geometry remains amenable to cooling by the ECCS. The calculated core geometry allows adequate ECCS cooling to assure the peak cladding temperature does not exceed 2200F and the maximum cladding oxidation nowhere exceeds 0.17 times the total cladding thickness before the start of significant oxidation during the event.
4.2.1.2 Fuel Rods 4.2.1.2.1 Fuel Rods Acceptance Criteria To meet GDC 10, 1971 and ensure their integrity, fuel rods are designed to prevent excessive fuel temperatures, excessive internal gas pressures due to fission gas buildup, and excessive cladding stresses and strains. To this end, the following conservative design bases are adopted for Condition I and Condition II events:
(1) Fuel Pellet Temperatures - The center temperature of the hottest pellet is to be below the melting temperature of the UO2. The melting point of un-irradiated UO2 is 5080°F (Reference 1). Irradiation reduces the melting 4.2-3 Revision 24 September 2018
DCPP UNITS 1 & 2 FSAR UPDATE point of UO2 by 58°F per 10,000 megawatt days/metric ton of uranium (MWD/MTU). While a limited amount of center melting can be tolerated, the design conservatively precludes center melting. A calculated centerline fuel temperature of 4700°F has been selected as the overpower limit. This provides sufficient margin for uncertainties.
(2) Internal Gas Pressure - The fuel rod internal gas pressure remains below the value that can cause the fuel-cladding diametral gap to increase due to outward cladding creep during steady state operation. Rod pressure is also limited so that extensive departure from nucleate boiling (DNB) propagation does not occur during normal operation and accident events (Reference 14). Also, cladding flattening (Reference 15) will not occur during the fuel rod incore life.
(3) Cladding Stress and Strain - The design limit for the fuel rod clad strain is the total plastic tensile creep strain due to:
uniform clad creep, uniform cylindrical fuel pellet expansion due to swelling, and thermal expansion is less than 1% from the un-irradiated condition.
The design limit for the fuel rod clad stress is that the volume average effective stress calculated with the von Mises equation considering interference due to uniform cylindrical pellet-clad contact, caused by:
pellet thermal expansion, pellet swelling and uniform clad creep, and pressure differences is less than the ZIRLO 0.2% offset yield stress, with due consideration to temperature and irradiation effects under Condition I and Condition ll events.
While the clad has some capability for accommodating plastic strain, the yield stress has been established as a conservative design limit.
(4) Cladding Tensile Strain - The total tensile strain due to uniform cylindrical pellet thermal expansion during a transient is less than 1% from the pre-transient value.
(5) Strain Fatigue - The fuel system will not be damaged due to excessive clad fatigue. The fatigue life usage factor is limited to less than 1.0 to prevent reaching the material fatigue limit.
4.2-4 Revision 24 September 2018
DCPP UNITS 1 & 2 FSAR UPDATE (6) Fuel Clad Oxidation and Hydriding - Fuel rod damage will not occur due to excessive clad oxidation and hydriding. In order to limit metal-oxide formation to acceptable values, the ZIRLO metal-oxide interface temperature is limited for Condition I and Condition II events. The clad and structural component hydrogen pickup is limited at end of life (EOL) to preclude loss of ductility due to hydrogen embrittlement by the formation of zirconium hydride platelets.
(7) Fuel Clad Wear - The fuel system will not be damaged due to fuel rod clad fretting. A design wall thickness reduction of 10% is a general guide in evaluating clad imperfections including fretting wear marks.
(8) Fuel Clad Flattening - Fuel clad flattening is the long-term creep collapse of the fuel rod into the axial gap between fuel pellet columns. No clad flattening occurs in Westinghouse fuel designs (Reference 33).
(9) Fuel Rod Axial Growth - The fuel rods will be designed with adequate clearance between the fuel rod ends and the top and bottom nozzles to accommodate the differences in the growth of fuel rods and the growth of the fuel assembly.
The preceding fuel rod acceptance criteria and other supplementary fuel design criteria/limits are given in Reference 29. Reference 25 provides the methodology for peak rod burnup in excess of 50,000 MWD/MTU. The above requirements impact design parameters such as pellet size and density, cladding-pellet diametral gap, gas plenum size, and helium pre-pressurization. The design also considers effects such as fuel density changes, fission gas release, cladding creep, and other physical properties that vary with burnup.
An extensive irradiation testing and fuel surveillance operational experience program has been conducted to verify the adequacy of the fuel performance and design bases.
This program is discussed in Section 4.2.1.4.
4.2.1.2.2 Fuel Rods Description The fuel rods consist of zirconium alloy tubing that is plugged and seal-welded at the ends. The VANTAGE+ fuel rods contain enriched uranium dioxide fuel pellets and may also include axial blanket (natural or enriched uranium dioxide) pellets, and/or an IFBA coating on some of the enriched fuel pellets. The VANTAGE + design is capable of achieving extended burnup operation. Fuel rod schematics are shown in Figure 4.2-3.
The fuel pellets are right circular cylinders consisting of uranium dioxide powder that has been compacted by cold pressing and then sintered to the required density. The ends of each pellet are dished slightly to allow greater axial expansion at the center of the pellets.
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DCPP UNITS 1 & 2 FSAR UPDATE The axial blanket reduces power at the ends of the rods and causes a slight increase in axial power peaking. Axial blankets reduce neutron leakage and improve fuel utilization.
Certain fuel assemblies utilize annular fuel pellets in the axial blanket region of the fuel rods. The use of this feature provides additional margin to the fuel rod internal pressure design limits. The axial blankets typically are a nominal 6 inches of natural or enriched fuel pellets at each end of the fuel rod pellet stack. However, the option exists to increase the top axial blanket length to a nominal 7 inches. The axial blankets utilize pellets, which are physically different from the non-axial blanket pellets to prevent accidental mixing during manufacturing.
The IFBA coated fuel pellets are identical to the enriched uranium dioxide pellets except for the addition of a thin ZrB2 coating on the pellet cylindrical surface. Coated pellets may occupy the central portion of the fuel column. The number and pattern of IFBA rods within an assembly may vary depending on specific application. An evaluation and test program for the IFBA design features is given in Section 2.5 in Reference 26.
The IFBA flattens the axial power distribution and reduces the local peaking as it burns out during irradiation. The net result of axial blankets and IFBA is a slight increase in power peaking during core operation.
The new fuel regions incorporate assemblies whose non-IFBA rods contain fully enriched solid fuel pellets in the blanket region.
All fuel rods are internally pressurized with helium during the welding process to minimize compressive cladding stresses and creep due to coolant operating pressures.
Fuel rod pressurization depends on the planned fuel burnup, as well as other fuel design parameters and fuel characteristics. To avoid overstressing of the cladding or seal welds, void volume and clearances are provided within the rods to accommodate fission gases released from the fuel, differential thermal expansion between the cladding and the fuel, and fuel density changes during burnup. Shifting of the fuel within the cladding during handling or shipping prior to core loading is prevented by a helical spring within the fuel rod that bears on top of the fuel.
4.2.1.2.2.1 Materials - Fuel Cladding VANTAGE+ fuel has ZIRLO fuel cladding. Reference 29 provides additional details on ZIRLO fuel cladding.
Metallographic examination of irradiated commercial fuel rods has shown occurrences of fuel/cladding chemical interaction. Reaction layers of 1 mil in thickness have been observed between fuel and cladding at limited points around the circumference. These data give no indication of propagation of the layer and eventual cladding penetration.
Stress corrosion cracking is another postulated phenomenon related to fuel/cladding chemical interaction. Out-of-reactor tests have shown that in the presence of high cladding tensile stresses, large concentrations of iodine can chemically attack the fuel 4.2-6 Revision 24 September 2018
DCPP UNITS 1 & 2 FSAR UPDATE cladding and lead to eventual cladding cracking. Westinghouse has no evidence that this mechanism is operative in commercial fuel.
4.2.1.2.2.2 Materials - Fuel Pellets Sintered, high-density UO2 reacts only slightly with the cladding at core operating temperatures and pressures. In the event of cladding defects, the high resistance of uranium dioxide to attack by water protects against fuel deterioration, although limited fuel erosion can occur. Operating experience and extensive experimental work reveal that the thermal design parameters conservatively account for changes in the thermal performance of the fuel elements due to pellet fracture that may occur during power operation. The consequences of defects in the cladding are greatly reduced by the ability of uranium dioxide to retain fission products, including those that are gaseous or highly volatile.
Improvements in fuel fabrication techniques, based on extensive analytical and experimental work (References 9 and 33), have eliminated or minimized the fuel pellet densification effect that had previously been observed in fuel irradiated in operating Westinghouse pressurized water reactors (PWRs) (References 5 and 8).
Fuel densification is considered in the nuclear and thermal-hydraulic design of the reactor, as described in Sections 4.3.3.2.5 and 4.4.3.2.1, respectively.
Some fuel pellets are fabricated with a thin boride coating on the pellet outside surface for reactivity control (refer to Section 4.2.1.2.2).
4.2.1.2.2.3 Materials - Strength Considerations One of the most important limiting factors in fuel element duty is the mechanical interaction of fuel and cladding. This fuel-cladding interaction produces cyclic stresses and strains in the cladding, and these in turn consume cladding fatigue life. To reduce fuel-cladding interaction, which is a principal goal of design, and enhance the cyclic operational capability of the fuel rod, pre-pressurized fuel rods are used.
Pre-pressurized fuel rods partially offset the effect of the coolant external pressure and reduce the rate of cladding creep toward the surface of the fuel. Fuel rod pre-pressurization delays the time at which substantial fuel-cladding interaction and hard contact occur. This significantly reduces the number and extent of cyclic stresses and strains experienced by the cladding, both before and after fuel-cladding contact. These factors increase the fatigue life margin of the cladding and lead to greater cladding reliability. If gaps should form in the fuel stacks, cladding flattening will be prevented by the rod pre-pressurization so that the flattening time will be greater than the fuel core life.
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DCPP UNITS 1 & 2 FSAR UPDATE To minimize fuel-cladding interaction during startup, following handling of irradiated fuel assemblies during a refueling, or a cold shutdown, limitations in power increase rates are instituted.
4.2.1.2.2.4 Steady State Performance Evaluation In the calculation of the steady state performance of a nuclear fuel rod, the following interacting factors must be considered:
(1) Cladding creep and elastic deflection (2) Pellet density changes, thermal expansion, gas release, and thermal properties as a function of temperature and fuel burnup (3) Internal pressure as a function of fission gas release, rod geometry, and temperature distribution These effects are evaluated using the fuel rod design model of Reference 34. The model modifications for time-dependent fuel densification are given in Reference 34.
The model determines fuel rod performance characteristics for a given rod geometry, power history, and axial power shape. In particular, internal gas pressure, fuel and cladding temperatures, and cladding deflections are calculated. The fuel rod is divided lengthwise into several sections and radially into a number of annular zones. Fuel density changes, cladding stresses, strains and deformations, and fission gas releases are calculated separately for each segment. These effects are then integrated to obtain the total internal pressure.
Subject to the design criteria of Section 4.2.1.2.1, the initial rod internal pressure is selected to delay fuel-cladding mechanical interaction and to avoid the potential for flattened rod formation.
The gap conductance between the pellet surface and the cladding inner diameter is calculated as a function of the composition, temperature, and pressure of the gas mixture, and the gap size or contact pressure between cladding and pellet. After computing the fuel temperature for each pellet's annular zone, the fractional fission gas release is calculated based on local fuel temperature and burnup. Finally, the gas released is summed over all zones and the pressure is calculated.
The PAD 4.0 code shows good agreement in fit for a variety of published and proprietary data on fission gas release, fuel temperatures, and cladding deflection (Reference 34). Included in this spectrum are variations in power, time, fuel density, and geometry.
Typical fuel cladding inner diameter and the fuel pellet outer diameter as a function of exposure are presented in Figure 4.2-4. The cycle-to-cycle changes in the pellet outer diameter represent the effects of power changes as the fuel is moved into different 4.2-8 Revision 24 September 2018
DCPP UNITS 1 & 2 FSAR UPDATE positions during refueling. The gap size at any time is given by the difference between cladding inner radius and pellet outer radius. Total cladding-pellet surface contact occurs between 600 and 800 effective full power days (EFPD). Figure 4.2-4 represents hot fuel dimensions for a fuel rod operating at the power level shown in Figure 4.2-5.
Figure 4.2-5 also illustrates representative fuel rod internal gas pressure and linear power for the lead burnup rod versus irradiation time. In addition, it outlines the typical operating range of internal gas pressures that is applicable to the total fuel rod population within a region. The plenum height of the fuel rod, in conjunction with other characteristics, is designed to ensure that the maximum internal pressure of the fuel rod remains below the value that causes the fuel-cladding diametral gap to increase due to outward cladding creep (Reference 29).
Cladding stresses during steady state operation are low. Compressive stresses are created by the pressure differential between the coolant pressure and the rod internal gas pressure.
The design fuel rod internal pressure limit which precludes gap increase is up to 1150 psi above system pressure for ZIRLO rods, based on the ZIRLO creep rate.
Stresses due to the temperature gradient are not included because their contribution to the cladding volume average stress is small and decreases with time during steady state operation due to stress relaxation. The stress due to pressure differential is highest in the minimum power rod at the beginning of life (BOL) (due to low internal gas pressure), and the thermal stress is highest in the maximum power rod (due to the steep radial temperature gradient).
Tensile stresses could be created once the cladding comes in contact with the pellet.
These stresses would be induced by the fuel pellet swelling during irradiation. As shown in Figure 4.2-4, there is very limited cladding pushout after pellet-cladding contact. Fuel swelling can result in small cladding strains (<1% for expected discharge burnups but the associated cladding stresses are very low because of cladding creep thermal and irradiation-induced creep). The 1% strain criterion is extremely conservative for fuel-swelling driven cladding strain because the strain rate associated with solid fission products swelling is very slow.
4.2.1.2.2.5 Transient Evaluation Method The PAD 4.0 code (Reference 34) is the principal design tool for fuel rod performance evaluations. PAD 4.0 iteratively calculates the interrelated effects of temperature, pressure, cladding elastic and plastic behavior, fission gas release, and fuel densification and swelling as a function of time and linear power as a function time and linear power.
Westinghouse uses the PAD 4.0 code to show that the VANTAGE+ fuel design meets pellet cladding interaction (PCI) acceptance criteria, 1) less than 1% transient-induced cladding strain, and 2) no centerline fuel melting. The NRC has found PAD 4.0 acceptable for application to the 4.2-9 Revision 24 September 2018
DCPP UNITS 1 & 2 FSAR UPDATE VANTAGE+ design with changes in the cladding creep model for ZIRLO cladding and changes in the rod growth model for the rod pressure calculation (Reference 34).
Pellet thermal expansion due to power increases in a fuel rod is considered the only mechanism by which significant stresses and strains can be imposed on the cladding.
Such power increases in commercial reactors can result from fuel shuffling, reactor power escalation following extended reduced power operation, and control rod movement. In the mechanical design model, depletion of lead rods is calculated using best estimate power histories as determined from core physics calculations. During the depletion, the diametral gap closure is evaluated using the pellet expansion-cracking model, cladding creep model, and fuel swelling model. At various times during the depletion, the power is increased locally on the rod to the burnup-dependent attainable power density, as determined by core physics calculation. The radial, tangential, and axial cladding stresses resulting from the power increase are combined into a volume average effective cladding stress.
The von Mises' criterion, described in Section 4.2.1.2.3, is used to determine if the cladding yield stress has been exceeded. The yield stress correlation is that for irradiated cladding, since fuel-cladding interaction occurs at high burnup. Furthermore, the effective stress is increased by an allowance that accounts for stress concentrations in the cladding adjacent to radial cracks in the pellet, prior to the comparison with the yield stress. This allowance was evaluated using a two-dimensional (r, ) finite element model.
Since slow transient power increases can result in large cladding strains without exceeding the cladding yield stress due to cladding creep and stress relaxation, a criterion on allowable cladding positive strain is necessary. Based on high strain rate burst and tensile test data for irradiated tubing, 1 percent strain was adopted as the lower limit on irradiated cladding ductility.
It is recognized that a possible limitation to the satisfactory behavior of the fuel rods in a reactor that is subjected to daily load follow is the failure of the cladding by low cycle strain fatigue. During their normal residence time in a reactor, the fuel rods may be subjected to 1000 cycles with typical changes in power level from 50 to 100 percent of their steady state values.
The ZIRLO cladding fatigue analyses have been performed using the Zircaloy-4 fatigue model. Fatigue properties of materials can be generally correlated with tensile properties, and are frequently described in terms of the fatigue ratio, defined as the ratio of the fatigue limit to the tensile strength. While the ratio is considered to be an approximation, it can serve as a useful tool to predict fatigue properties within a given alloy system. At room temperature, ZIRLO cladding has an average tensile strength equivalent to Zircaloy-4 cladding. At elevated temperatures, ZIRLO cladding has a higher tensile strength than Zircaloy-4 cladding. On this basis, it is conservative to assume that the fatigue properties of ZIRLO cladding are equivalent to those of Zircaloy-4 cladding, and that on a best estimate basis, ZIRLO cladding has a higher fatigue strength than does Zircaloy-4 cladding at elevated temperatures.
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DCPP UNITS 1 & 2 FSAR UPDATE The evaluation of the fatigue life usage factor for extended burnup operation conservatively assumes daily load follow operation over the life of the fuel rod, plus ten (10) cold shutdowns per cycle. The Westinghouse fuel performance code, PAD 4.0, (Reference 34) is used to determine the strain range for the fatigue life usage analysis.
PAD 4.0, is used to determine the strain range for the fatigue usage analysis of the VANTAGE+ fuel design. The Langer-O'Donnell fatigue model is used with the calculated strains from PAD 4.0 to assure that the above criterion is met for VANTAGE+.
The Westinghouse analytical approach to strain fatigue results from evaluating several strain-fatigue models and the results of the Westinghouse experimental programs. In conclusion, the approach defined by Langer-O'Donnell (Reference 12) was retained, and the empirical factors of their correlation were modified to conservatively bound the results of the Westinghouse testing program.
The Langer-O'Donnell empirical correlation has the following form:
100 E
S ln S (4.2-1) a 4 100RA e N
f where:
Sa = 1/2 E =
pseudo-stress amplitude that causes failure in Nf cycles, lb/in2
= total strain range, in./in E = Young's Modulus, lb/in2 Nf = number of cycles to failure RA = reduction in area at fracture in a uniaxial tensile test, %
Se = endurance limit, lb/in2 Both RA and Se are empirical constants that depend on the type of material, the temperature, and the irradiation.
The results of the Westinghouse test programs provided information on different cladding conditions, including the effect of irradiation, hydrogen level, and temperature.
The Westinghouse design equations followed the concept for the fatigue design criterion according to the ASME BPVC Section III, namely:
(1) The calculated pseudo-stress amplitude (Sa) includes a safety factor of 2.
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DCPP UNITS 1 & 2 FSAR UPDATE (2) The allowable cycles for a given Sa are 5 percent of Nf or a safety factor of 20 on cycles.
The lesser of the two allowable numbers of cycles is selected. The cumulative fatigue life fraction is then computed as:
k nk N
1 1 (4.2-2) fk where:
nk = number of diurnal cycles of mode k 4.2.1.2.3 Fuel Rod Design Evaluation The fuel rod design ensures that the design bases are satisfied.
(1) Fuel Pellet Temperatures - The temperature of the VANTAGE+ fuel pellets is evaluated by the same methods as are used for all Westinghouse fuel designs. Rod geometries, thermal properties, heat fluxes, and temperature differences are modeled to calculate the temperature at the surface and centerline of the fuel pellets. Fuel centerline temperatures are calculated as a function of local power and rod burnup. To preclude fuel melting, the peak local power experienced during Condition I and Condition II events can be limited to a maximum value which is sufficient to ensure that the fuel centerline temperatures remain below the melting temperature at all burnups. Design scoping evaluations for Condition I and Condition II events show that fuel melting will not occur for achievable local powers and extended burnups (refer to Section 4.2.1.2.1).
(2) Internal Gas Pressure - The rod internal pressure of the VANTAGE+ fuel rod is evaluated in the same manner as is used for other Westinghouse fuel types. Gas inventories, gas temperature, and rod internal volumes are modeled and the resulting rod internal pressure is compared to the design limit. VANTAGE+ design evaluations to extended burnup levels verify that the fuel rod internal pressure as calculated will meet the design basis. This evaluation is discussed in detail in Appendix B, Section B.2.2 of Reference 29.
(3) Cladding Stress and Strain - Radial, tangential, and axial stress components due to pressure differential and fuel cladding contact pressure are combined into an effective stress using the maximum-distortion-energy theory. The von Mises criterion (Reference
- 22) is used to evaluate whether or not the yield strength has been exceeded. The criterion states that an isotropic material under multiaxial 4.2-12 Revision 24 September 2018
DCPP UNITS 1 & 2 FSAR UPDATE stress will begin to yield plastically when the effective stress (i.e.,
combined stress using maximum-distortion-energy theory) becomes equal to the material yield stress in simple tension, as determined by a uniaxial tensile test. Since general yielding is prohibited, the volume average effective stress determined by integrating across the cladding thickness is increased by an allowance for local nonuniformity effects before the stress is compared to the yield strength. The yield strength correlation is appropriate for irradiated cladding since the irradiated properties are attained at low exposure, whereas the fuel/cladding interaction conditions, which can lead to minimum margin to the design basis limit, always occur at much higher exposures. The clad stresses in the VANTAGE+ fuel rod clad caused by power transients are evaluated (refer to Section 4.2.1.2.2.5). These evaluations show that the clad stresses and clad strains for VANTAGE+ fuel rod designs (IFBA and non-IFBA) meet the design limits.
(4) Cladding Tensile Strain - Section 4.2.1.2.2.5 shows that the cladding tensile strain stresses for the VANTAGE+ fuel rod designs meet the design limits.
(5) Strain Fatigue - Clad fatigue for the VANTAGE+ fuel rod design is evaluated by the same methods as are used for other Westinghouse fuel designs. Computer modeling of the fuel rod simulates a daily load follow cycling scheme 100-15-100% power and 12-3-6-3 hour intervals for residence times of more than 60 months. Design evaluations have shown that the cumulative fatigue usage factor for the VANTAGE+ fuel rod will meet the design criterion. Details regarding this evaluation are presented in Reference 29, Section B.2.3.
(6) Fuel Clad Oxidation and Hydriding - The clad surface temperature and hydriding of the VANTAGE+ fuel rod is evaluated by the same methods as are used for other Westinghouse fuel designs. The coolant temperature rise over the length of the fuel rod and temperature rise through the film, crud and oxide layer are calculated and the temperature at the metal-oxide interface is determined. Calculations show that the clad surface temperature and hydriding of the VANTAGE+ fuel rod meet the design limits.
(7) Fuel Clad Wear - The design criteria with regard to clad fretting wear are met for VANTAGE+ to its design burnup. Details regarding this evaluation are presented in Reference 29, Section B.2.5.1.
(8) Fuel Clad Flattening - Calculations for VANTAGE+ fuel show that predicted clad flattening time exceeds residence times expected for extended burnup fuel management. The evaluation is discussed in detail in Reference 29, Section B.2.4.
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DCPP UNITS 1 & 2 FSAR UPDATE (9) Fuel Rod Axial Growth - The fuel assembly design is sized to provide sufficient fuel-rod-to-nozzle gaps, as discussed in Reference 29, Section 2.3.1.1. Sufficient top and bottom nozzle-to-fuel rod gaps are provided to assure that fuel rod growth is accommodated during the fuel design lifetime.
4.2.1.3 Fuel Assembly Structure 4.2.1.3.1 Fuel Assembly Structure Acceptance Criteria 4.2.1.3.1.1 Fuel Assembly Structural Integrity The VANTAGE+ fuel assembly must maintain its structural integrity in response to seismic and LOCA loads (refer to Sections 3.7.3.15.2 and 4.2.1.5.4). The stresses and deformations due to various loads on the fuel assemblies are discussed in Section 4.2.1.3.2.5, including the effects of Seismic/LOCA loads.
4.2.1.3.1.2 Fuel Assembly Shipping and Handling Loads The design acceleration limit for the fuel assembly handling and shipping loads is 4g's minimum.
4.2.1.3.1.3 Top Nozzle The top nozzle is required to transmit 4g shipping and handling loads without permanent deformation. The joints must transmit the same loads and be capable of disconnecting using a special disassembly tool.
4.2.1.3.1.4 Fuel Assembly Holddown Springs The fuel assembly holddown springs, Figure 4.2-2, are designed to keep the fuel assemblies resting on the lower core plate under transients associated with Condition I and Condition II events with the exception of the turbine overspeed transient associated with a loss of external load. The holddown springs are designed to tolerate the possibility of a deflection associated with fuel assembly liftoff for this case and provide contact between the fuel assembly and the lower core plate following this transient.
4.2.1.3.2 Fuel Assembly Structure Description The fuel assembly structure consists of a bottom nozzle, top nozzle, guide thimbles, and grids, as shown in Figure 4.2-2.
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DCPP UNITS 1 & 2 FSAR UPDATE 4.2.1.3.2.1 Bottom Nozzle The bottom nozzle is a box-like structure that serves as a bottom structural element of the fuel assembly and directs the coolant flow distribution to the assembly. The square nozzle is fabricated from Type 304 stainless steel. The legs form a plenum for the inlet coolant flow. The plate prevents a downward ejection of the fuel rods. The bottom nozzle is fastened to the fuel assembly guide tubes by locked screws that penetrate through the nozzle and mate with an inside fitting in each guide thimble tube.
The bottom nozzle design, known as the DFBN, reduces the possibility of fuel rod damage due to debris-induced fretting. The holes in the bottom nozzle plate are sized to minimize passage of debris particles large enough to cause damage while providing sufficient flow area, comparable pressure drop, and continued structural integrity of the nozzle. Tests to measure pressure drop and demonstrate structural integrity have been performed to verify that the DFBN meets the applicable mechanical design criteria for both Condition I and Condition II events.
The bottom nozzle design has a reconstitution feature that permits remote unlocking, removing, and relocking of the thimble screws. Coolant flow through the fuel assembly is directed from the plenum in the bottom nozzle upward through the penetrations in the plate to the channels between the fuel rods.
The weight and axial loads (holddown) imposed on the fuel assembly are transmitted through the bottom nozzle to the lower core plate. Indexing and positioning of the fuel assembly is controlled by alignment holes in two diagonally opposite bearing plates that mate with locating pins in the lower core plate. Any lateral loads on the fuel assembly are transmitted to the lower core plate through the locating pins.
Westinghouse has developed the standardized debris filter bottom nozzle (SDFBN) for use on its 17x17 fuel designs, including the 17x17 VANTAGE+ fuel design. The SDFBN improves the debris mitigation performance of the bottom nozzle by eliminating the side skirt communication flow holes. This nozzle has been evaluated and meets all of the applicable mechanical design criteria. In addition, there is no adverse effect on the thermal hydraulic performance of the SDFBN either with respect to the pressure drop or with respect to DNB. The SDFBN was implemented at Diablo Canyon beginning with Unit 2 Region 20 (Cycle 18 feed) and Unit 1 Region 21 (Cycle 19 feed).
4.2.1.3.2.2 Top Nozzle The RTN assembly functions as the upper structural element of the fuel assembly in addition to providing a partial protective housing for the RCCA. It consists of an adapter plate, enclosure, top plate, and pads. The integral welded assembly has holddown springs mounted on the assembly, as shown in Figure 4.2-2. The springs are made of Inconel 718. The other components are made of Type 304 stainless steel.
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DCPP UNITS 1 & 2 FSAR UPDATE The square adapter plate is provided with round penetrations and semicircular-ended slots that permit the flow of coolant upward through the top nozzle. Other round holes are provided to accept sleeves that are welded to the adapter plate and mechanically attached to the thimble tubes. The ligaments in the plate cover the tops of the fuel rods and prevent their upward ejection from the fuel assembly. The enclosure is a box-like structure that sets the distance between the adapter plate and the top plate. The top plate has a large square hole in the center to permit access for the control rods and the control rod spiders. Holddown springs are mounted on the top plate and are fastened in place by bolts and clamps located at two diagonally opposite corners. The Westinghouse integral nozzle (WIN) design provides a wedged joint for transfer of the fuel assembly hold-down forces into the top nozzle structure. On the other two corners, integral pads contain alignment holes to locate the upper end of the fuel assembly.
In the removable top nozzle design, a stainless steel nozzle insert is mechanically connected to the top nozzle adapter plate by means of a pre-formed circumferential bulge near the top of the insert. The insert engages a mating groove in the wall of the adapter plate thimble tube thru-hole. The insert has 4 equally spaced axial slots which allow the insert to deflect inwardly at the elevation of the bulge thus permitting the installation or removal of the nozzle. The insert bulge is positively held in the adapter plate mating groove by placing a lock tube, with a uniform ID identical to that of the thimble tube, into the insert. The lock tube is secured in place by locally deforming it into the concave side of the bulge in the insert.
The full complement of these joints comprises the structural connection between the top nozzle and the remainder of the fuel assembly. The nozzle insert-to-adapter plate bulge joints replace the uppermost grid sleeve-to-adapter plate welded joints found in current fuel assemblies. The nozzle insert-to-thimble tube multiple 4-lobe bulge joint located in the lower portion of the insert represents the structural connection between the insert and the remainder of the fuel assembly below the elevation of the insert.
4.2.1.3.2.3 Guide and Instrument Thimbles Guide thimbles are structural members that also provide channels for the neutron absorber rods, burnable poison rods, or neutron source assemblies. Each guide thimble is fabricated from zirconium alloy tubing having two different diameters. The larger tube diameter at the top provides the annular area necessary to permit rapid insertion of the control rods during a reactor trip. The lower portion of the guide thimble has a reduced diameter to produce a dashpot action near the end of the control rod travel during a reactor trip. Four holes are provided on the thimble tube above the dashpot to reduce the rod drop time. The dashpot is closed at the bottom by an end plug and thimble screw that is provided with a small flow port to avoid fluid stagnation in the dashpot volume during normal operation.
The central instrumentation tube, fabricated from zirconium alloy tubing, is constrained by its seating in counterbores of each nozzle. Incore neutron detectors pass through the bottom nozzle's large counterbore into the center instrumentation tube.
4.2-16 Revision 24 September 2018
DCPP UNITS 1 & 2 FSAR UPDATE The VANTAGE+ guide thimble tube ID provides an adequate nominal diametral clearance of 0.061 inches for the control rods. The VANTAGE+ thimble tube ID also provides sufficient diametral clearance for burnable absorber rods, source rods, and dually compatible thimble plugs.
The VANTAGE+ instrumentation tube allows sufficient diametral clearance for the flux thimble to traverse the tube without binding.
4.2.1.3.2.4 Grid Assemblies The fuel rods, as shown in Figure 4.2-2, are supported laterally at eight intervals along their length by grid assemblies that maintain the lateral spacing between the rods by the combination of support dimples and springs. The grid assembly consists of individual slotted straps interlocked and brazed or welded in an "egg-crate" arrangement to join the straps permanently at their points of intersection. The straps contain springs and support dimples. The outside straps on all grids contain mixing vanes which, in addition to their mixing function, help guide the grids and fuel assemblies past projecting surfaces during fuel handling or core loading and unloading.
Inconel 718 and zirconium alloys were chosen as the grid materials because of corrosion resistance, neutron economy, and high strength properties. The grids are connected to the thimble tube with 4-lobe bulge joints similar to the top nozzle (refer to Section 4.2.1.3.2.2). The magnitude of the grid restraining force on the fuel rod is set high enough to minimize possible fretting, without overstressing the cladding at the points of contact. The grid assemblies also allow axial thermal expansion of the fuel rods to prevent their buckling or distortion.
VANTAGE+ fuel assemblies have four types of grid assemblies: top and bottom grids, Mid-grids, IFM grids, and protective grids (P-Grid). Each fuel assembly has a total of twelve grids: one bottom and one top grid, six Mid-grids, three IFM grids, and one P-Grid.
The most critical grids from a standpoint of preventing fretting of fuel rods are the bottom and top grids (i.e., the end grids). The bottom and top grid assemblies do not include mixing vanes on the inner straps. The material of these grid assemblies is Inconel 718, chosen because of its corrosion resistance and high strength. These grid assemblies, being at the ends of the fuel assembly, are in the lower flux regions of the fuel assembly.
The Mid-grid assemblies consist of zirconium alloy straps arranged as described above and permanently joined by welding at their points of intersection. This material is primarily chosen for its low neutron absorption properties. These grids are used in the high heat flux region of the fuel assemblies. The inner straps include mixing vanes that project into the coolant stream and promote mixing of the coolant.
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DCPP UNITS 1 & 2 FSAR UPDATE The IFM grids, as shown in Figure 4.2-2, are located between the three uppermost spans between the Mid-grids and incorporate a similar mixing vane array. Their prime function is mid-span flow mixing in the hottest fuel assembly spans. Each IFM grid cell contains four dimples which are designed to prevent mid-span channel closure in the spans containing IFMs and fuel rod contact with the mixing vanes. This simplified cell arrangement allows short grid cells so that the IFM can accomplish its flow mixing objective with minimal pressure drop. The IFM grids are fabricated from Zircaloy. This material was selected to take advantage of the material's inherent low neutron capture cross-section.
The P-Grid, configured with the same egg crate support matrix and other internal structures as the other grid assemblies, is added to the bottom of the fuel assembly to provide an additional debris barrier thereby improving fuel reliability. This protective or P grid is thinner (i.e., shorter) than the normal grid to accommodate the gap between the bottom nozzle and the fuel rods. It has no mixing vanes and has shorter inner straps comprising the support matrix compared to the top and bottom grids. The P-Grid is composed of Inconel 718 material.
4.2.1.3.2.5 Stresses and Deflections Structural integrity of fuel assemblies is ensured by setting limits on stresses and deformations due to various loads, and by determining that the assemblies do not interfere with other components' functionality.
These stress and deformation limits are applied to the design and evaluation of the top and bottom nozzles, the guide thimbles, the grids, and the thimble joints.
The design bases for evaluating the structural integrity of the fuel assemblies are:
(1) Nonoperational - 4g minimum lateral loading with dimensional stability in both lateral and axial directions.
(2) Normal Operation (Condition I) and Incidents of Moderate Frequency (Condition II). The fuel assembly component structural design criteria are classified into two material categories: austenitic steels and zirconium alloys. Although not strictly fuel assembly components, reactor core elements that are made of stainless steel and are closely related to the fuel assembly design include the top and bottom nozzle, the RCCA cladding, and some burnable absorber rod's cladding.
The stress categories and strength theory presented in the ASME BPVC Section III, are used as a general guide.
Zirconium alloy structural components, which consist of guide thimbles and fuel tubes, are in turn subdivided into two categories because of material differences and functional requirements. The fuel tube design 4.2-18 Revision 24 September 2018
DCPP UNITS 1 & 2 FSAR UPDATE criteria are covered separately in Section 4.2.1.2.3. For the guide thimble design, the stress intensities, the design stress intensities and the stress intensity limits are calculated using the same methods as for the austenitic steel structural components. Un-irradiated zirconium alloy properties are used to define the stress limits.
The maximum shear stress theory (Tresca criterion {Reference 22}) for combined stresses is used to determine the stress intensities for the austenitic steel components. The stress intensity is defined as the numerically largest difference between the various principal stresses in a three-dimensional field. The allowable stress intensity value for austenitic stainless steel is given by the lowest of the following:
(a) One-third of the specified minimum tensile strength, or two-thirds of the minimum yield strength, at room temperature.
(b) One-third of the tensile strength or 90 percent of the yield strength, at temperature, but not to exceed two-thirds of the specified minimum yield strength at room temperature.
The stress intensity limits for the austenitic steel components are:
Stress Intensity Limits Categories Limit General Primary Membrane Stress Intensity 1.0 Sm Local Primary Membrane Stress Intensity 1.5 Sm Primary Membrane plus Bending Stress Intensity 1.5 Sm Total Primary plus Secondary Stress Intensity Range 3.0 Sm where Sm is the membrane stress.
(3) Abnormal Loads during Conditions III or IV - Worst cases are represented by combined seismic and blowdown loads. However, with NRC approval of the DCPP leak-before-break (LBB) analysis (Reference 30) and Amendments 221 and 223, to perform the fuel assembly structural analyses based on postulated pipe break locations that consider the application of LBB (Reference 32), the blowdown loads resulting from pipe rupture events in the main reactor coolant loop piping no longer have to be considered in the structural design basis analysis. Only the much smaller blowdown loads from RCS branch line breaks have to be considered (see Section 3.6.2.1.1.1).
(a) Deflections of components cannot interfere with reactor shutdown or emergency cooling of fuel rods.
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DCPP UNITS 1 & 2 FSAR UPDATE (b) The fuel assembly structural component stress under faulted conditions are evaluated using primarily the methods outlined in Appendix F of the ASME BPVC Section III. Since the current analytical methods utilize elastic analysis, the stress allowables are defined as the smaller value of 2.4 Sm or 0.70 Su (ultimate stress) for primary membrane and 3.6 Sm or 1.05 Su for primary membrane plus primary bending. For the austenitic steel fuel assembly components, the stress intensity is defined in accordance with the rules described in the previous section for normal operating conditions. For the Zircaloy components the stress intensity limits are set at two-thirds of the material yield strength, Sy, at reactor operating temperature. This results in Zircaloy stress limits being the smaller of 1.6 Sy or 0.70 Su for primary membrane and 2.4 Sy or 1.05 Su for primary membrane plus bending. For conservative purposes, the Zircaloy unirradiated properties are used to define the stress limits. The grid component strength criteria are based on experimental tests. For both Zircaloy and Inconel grids, the limit is the 95 percent confidence level on the true mean as taken from the distribution of measurements at operating temperature.
Stresses in the fuel rod due to thermal expansion and fuel cladding irradiation growth are limited by the relative motion of the rod as it slips over the grid spring and dimple surfaces. Clearances between the fuel rod ends and nozzles are provided so that fuel cladding irradiation growth does not produce interferences. Stresses due to hold-down springs opposing the hydraulic lift force are limited by the deflection characteristic of the springs. Stresses in the fuel assembly caused by tripping of the RCCA have little influence on fatigue because of the small number of events during the life of an assembly. Welded joints in the fuel assembly structure are considered in the structural analysis of the assembly. Appropriate material properties of welds ensure that the design bases are met. Assembly components and prototype fuel assemblies made from production parts were subjected to structural tests to verify that the design bases requirements were met.
Precautions are taken during fuel handling operations to minimize fuel assembly grid strap damage. These precautions include proper training of operators, confirmation of proper functioning and alignment of the fuel handling and transfer equipment, implementation of appropriate handling precautions, and Westinghouse recommendations.
The fuel assembly design loads for shipping have been established at 4g. Probes, permanently placed in the shipping cask, monitor and detect fuel assembly displacements that would result from loads in excess of the criteria. Experience indicates that loads which exceed the allowable limits rarely occur. Exceeding the limits requires re-inspection of the fuel assembly. Tests on various fuel assembly components such as the grid assembly, sleeves, inserts, and structure joints have been 4.2-20 Revision 24 September 2018
DCPP UNITS 1 & 2 FSAR UPDATE performed to ensure that the shipping design limits do not result in impairment of fuel assembly function.
The seismic/LOCA analysis methodology for fuel assembly analysis is presented in Reference 27. Specific seismic/LOCA analyses have been performed for Diablo Canyon and have demonstrated that all Condition II, III and IV load requirements are satisfied for the VANTAGE+ fuel design. The seismic analyses are performed for Design Earthquake (DE) (Condition II) and for both the Double Design Earthquake (DDE) and Hosgri Earthquake (HE) (Condition III and IV).The LOCA analysis is performed for the largest RCS branch lines (accumulator, pressure surge, residual heat removal). In all cases, the combined faulted condition seismic/LOCA grid impact forces were less than the grid strength and all stress values were found to be acceptable. Fuel assembly stresses and deflections meet the acceptance criteria for all loading conditions, as described in Section 4.2.1.3.2.5. Therefore, no fuel rod fragmentation will occur, long term core coolable geometry is maintained and RCCA insertability is maintained.
4.2.1.3.2.6 Dimensional Stability The dimensional stability of coolant flow channels is maintained by the grids and guide thimbles structure. The lateral spacing between fuel rods is controlled by the support dimples of adjacent grid cells plus the spring force and the internal moments generated between the spring and the support dimples.
No interference with control rod insertion into thimble tubes will occur during a postulated LOCA transient due to fuel rod swelling, thermal expansion, or bowing. In the early phase of the event, the high axial loads, which could be potentially generated by the difference in thermal expansion between fuel cladding and thimbles, are relieved by slippage of the fuel rods through the grids. The relatively low drag force restraint on the fuel rods will induce only minor thermal bowing, not enough to close the fuel rod-to-thimble tube gap. This rod-to-grid slip mechanism occurs simultaneously with control rod drop. Subsequent to the control rod insertion, the transient temperature increase of the fuel rod cladding can result in sufficient swelling to contact the thimbles.
4.2.1.3.2.7 Vibration and Wear The effect of a flow-induced vibration on the fuel assembly and individual fuel rods is minimal. Both fretting and vibration have been experimentally investigated. The cyclic stress range associated with deflections of such small magnitude is insignificant and has no effect on the structural integrity of the fuel rod.
The conclusion that the effect of flow-induced vibrations on the fuel assembly and fuel rod is minimal is based on test results and analysis documented in WCAP-8279 (Reference 13), which considered conditions normally encountered in reactor operation.
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DCPP UNITS 1 & 2 FSAR UPDATE The reaction on the grid support due to vibrational motions is correspondingly small and much less than the spring preload. Firm contact is therefore maintained. No significant cladding or grid support wear is expected during the life of the fuel assembly, as described in Section 4.2.1.4.
HISTORICAL INFORMATION IN ITALICS BELOW NOT REQUIRED TO BE REVISED.
During the mid-1970s, unexpected degradation of guide thimble tube walls was observed during examination of irradiated fuel assemblies taken from several operating PWRs. It was later determined that coolant up-flow through the guide thimble tubes and turbulent cross-flow above the fuel assemblies were responsible for inducing vibratory motion in normally fully withdrawn ("parked") control rods. When these vibrating rods were in contact with the inner surface of the thimble wall, a fretting wear of the thimble wall occurred. The extent of the observed wear is both time and nuclear steam supply system (NSSS) design-dependent and has been observed, in some non-Westinghouse cases, to extend through the guide tube walls, resulting in the formation of holes.
Guide thimble tubes function as the main structural members of the fuel assembly and as channels to guide and decelerate tripped control rods. Significant loss of mechanical integrity due to wear or hole formation could: (a) result in the inability of the guide thimble tubes to withstand their anticipated loadings for fuel handling accidents and transients, and (b) hinder RCCA trip.
The susceptibility and impact of guide thimble tube wear in Westinghouse plants of the DCPP design have been assessed in References 17 through 20. Included is a mechanistic wear model and the impact of the model's wear predictions on plant designs such as for DCPP.
Accordingly, the DCPP fuel design will experience less wear than that reported for other NSSS designs because the DCPP fuel design uses thinner, more flexible control rods that have relatively more lateral support in the guide tube assembly of the upper core structure. Such a design provides the housing and guide path for the RCCA above the core, and thus restricts control rod vibration due to lateral exit flow. The wear model is also believed to conservatively predict guide thimble tube wear and even with the worst anticipated wear conditions (both in the degree of wear and the location of wear), the guide thimble tubes will be able to fulfill their design functions.
Pacific Gas and Electric Company (PG&E) participated in a surveillance program to obtain data related to guide tube thimble wear (References 20 and 21). Data obtained from surveillance program examinations confirmed that guide thimble tubes used in DCPP meet design requirements.
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DCPP UNITS 1 & 2 FSAR UPDATE 4.2.1.3.3 Fuel Assembly Structure Design Evaluation 4.2.1.3.3.1 Structural Integrity Design Evaluation Structural integrity of fuel assemblies is ensured by setting limits on stresses and deformations due to various loads and by determining that the assemblies do not interfere with other components' operability (refer to Section 4.2.1.3.2.5). No interference with control rod insertion into thimble tubes will occur during a postulated LOCA transient due to fuel rod swelling, thermal expansion, or bowing (refer to Section 4.2.1.3.2.6). Deflections of components have been determined not to interfere with reactor shutdown or emergency cooling of fuel rods (refer to Section 4.2.1.3.2.5).
4.2.1.3.3.2 Fuel Assembly Shipping and Handling Loads Design Evaluation Fuel handling accelerations at both the manufacturing facility and reactor sites have been determined to be well below the 4g limit. Extensive over-the-road tests with shipping containers containing dummy (lead weighted) fuel assemblies were made during the current shipping container design development. Roads were specifically selected to ensure that the most realistic adverse conditions such as jumping curbs, railroad tracks, rough roads, etc. were encountered. Recording accelerometers confirmed that insignificant g loads were communicated to the fuel assembly carriage in the container.
4.2.1.3.3.3 Top Nozzle Design Evaluation Finite element analyses of the nozzle have been performed which show that the design is more than adequate to resist the 4g loading. Structural testing has verified this result.
The RTN feature was tested both as individual joints and functionally as part of the fuel assembly flow test assembly (refer to Appendix A, Sections A.1.0 and A.3.0 of Reference 26). The results of these tests demonstrate that the joint strength exceeds the structural (4g) and functional requirement.
4.2.1.3.3.4 Fuel Assembly Holddown Springs Design Evaluation The results of the flow testing serve as the basis for optimizing the spring force with due consideration of fuel assembly growth and holddown spring relaxation and holddown force requirements (refer to Appendix A, Section A.1.0 of Reference 26).
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DCPP UNITS 1 & 2 FSAR UPDATE 4.2.1.4 Operational Experience The operational experience of Westinghouse cores is presented in WCAP-8183 (Reference 8).
4.2.1.5 Safety Evaluation 4.2.1.5.1 General Design Criterion 2, 1967 - Performance Standards The seismic analyses for the fuel assemblies documents that the fuel assembly design can withstand the additional forces that might be imposed under DDE and HE conditions without fuel or clad damage that could interfere with continued effective core cooling (refer to Section 3.7.3.15.2).
4.2.1.5.2 General Design Criterion 10, 1971 - Reactor Design Section 4.2.1.2.1 provides the conservative design bases of the fuel rods adopted for Condition I and Condition II events. Section 4.2.1.2.3 describes how the fuel rods are designed to ensure that the design bases are satisfied for Condition I and Condition II events.
Section 4.2.1.3.1 provides the design bases to ensure the structural integrity of the fuel assembly adopted for Condition I and Condition II events. Section 4.2.1.3.3 describes how the fuel assembly meets its design basis.
4.2.1.5.3 Safety Function Requirements (1) Loads During Handling Section 4.2.1.3.3.2 documents that the fuel assembly shipping and handling load safety function requirements are met.
4.2.1.5.4 10 CFR 50.46(b)(4) - Coolable Geometry The fuel is designed with appropriate margin to assure that following a postulated LOCA, the calculated core geometry remains amenable to cooling by the ECCS. The calculated core geometry allows adequate ECCS cooling to assure the peak cladding temperature does not exceed 2200F and the maximum cladding oxidation nowhere exceeds 0.17 times the total cladding thickness prior to oxidation (refer to Sections 4.2.1.3.3.1 and 15.4.1).
Plant specific analyses were performed using the approved methodology in WCAP-9401 (Reference 27) to demonstrate structural integrity of fuel assemblies, in accordance with NUREG-0800, Standard Review Plan (SRP) for the Review of Safety Analysis Reports for Nuclear Power Plants, Section 4.2, Appendix A.
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DCPP UNITS 1 & 2 FSAR UPDATE 4.2.1.6 Tests and Inspections 4.2.1.6.1 Quality Assurance Program The Quality Assurance Program for Westinghouse nuclear fuel is summarized in the latest edition of the Westinghouse Nuclear Fuel Division Quality Assurance Program Plan, as listed in the PG&E Qualified Suppliers List.
4.2.1.6.2 Manufacturing The Westinghouse quality control philosophy during manufacturing is described in the Westinghouse Nuclear Fuel Division Quality Assurance Program Plan, as listed in the PG&E Qualified Suppliers List.
4.2.1.6.3 Onsite Inspection Onsite inspection of fuel assemblies, control rods, and reactor vessel internals is performed in accordance with the inspection program requirements discussed in Chapter 17.
Surveillance of fuel and reactor performance is routinely conducted on Westinghouse reactors. Power distribution is monitored using the excore fixed and incore movable detectors. Coolant activity and chemistry are followed, which permit early detection of any fuel cladding defects.
Visual examinations are routinely conducted during refueling outages. Additional fuel inspections are dependent on results of the operational monitoring and the visual examinations. Onsite examinations, if required, could include fuel integrity or other fuel performance evaluation examinations.
4.2.1.6.4 Removable Fuel Rod Assembly HISTORICAL INFORMATION IN ITALICS BELOW NOT REQUIRED TO BE REVISED.
As part of a continuing Westinghouse fuel performance evaluation program, one surveillance fuel assembly containing 88 removable fuel rods was included in Region 3 of the initial DCPP Unit 1 and Unit 2 core loading. The objective of this program was to facilitate interim and EOL fuel evaluation as a function of exposure. The rods could be removed, nondestructively examined, and reinserted at the end of intermediate fuel cycles. The rods could be removed easily and subjected to a destructive examination at EOL.
The overall dimensions, rod pitch, number of rods, and material are the same as for other Region 3 assemblies. These fuel rods were fabricated in parallel with the regular Region 3 rods using selected Region 3 cladding and pellets fabricated to the same 4.2-25 Revision 24 September 2018
DCPP UNITS 1 & 2 FSAR UPDATE manufacturing tolerance limits. Mechanically, the special assemblies differ from other Region 3 assemblies only in those features that facilitate removal and reinsertion.
Figure 4.2-6 compares the mechanical design of a removable fuel rod to a standard rod.
Figure 4.2-7 shows the removable rod fuel assembly, the modified upper nozzle adapter plate, and thimble plug assembly; it may be compared to the standard assembly shown in Figure 4.2-2. The location of the removable rods within the fuel assembly is shown in Figure 4.2-8. Fuel handling with removable fuel rods has been done routinely and without difficulty in many operating plants.
4.2.2 REACTOR VESSEL INTERNALS 4.2.2.1 Design Bases 4.2.2.1.1 General Design Criterion 2, 1967 - Performance Standards The reactor internals are designed to withstand the effects of, or are protected against, natural phenomena such as earthquakes.
4.2.2.1.2 General Design Criterion 4, 1987 - Environmental and Dynamic Effects Design Bases Consideration of the dynamic effects associated with main reactor coolant loop (RCL) piping postulated pipe ruptures are excluded from the DCPP design basis with the approval of leak-before-break (LBB) methodology by demonstrating that the probability of fluid system piping rupture is extremely low under conditions consistent with the design basis for the piping.
4.2.2.1.3 General Design Criterion 10, 1971 - Reactor Design The reactor vessel internals are designed with appropriate margin to assure that SAFDLs are not exceeded during any condition of normal operation, including the effects of anticipated operational occurrences.
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DCPP UNITS 1 & 2 FSAR UPDATE 4.2.2.1.4 Safety Function Requirements (1) Core Flow Distribution The reactor vessel internals, in conjunction with the fuel assemblies, direct reactor coolant through the core to achieve acceptable flow distribution and to restrict bypass flow so that the heat transfer performance requirements are met for all modes of operation. In addition, required cooling for the pressure vessel head is provided so that the temperature differences between the vessel flange and head do not result in leakage from the flange during reactor operation.
(2) Protection of the Reactor Pressure Vessel from Neutron Exposure In addition to neutron shielding provided by the reactor coolant, a thermal shield in Unit 1 and a neutron pad (Reference 7) assembly in Unit 2 limit the neutron exposure of the pressure vessel. Additionally, provisions are made to install the vessel material test specimens for the reactor pressure vessel material surveillance program (refer to Section 5.2).
(3) Incore Instrumentation Provisions are made to install incore instrumentation for plant operation.
4.2.2.2 Acceptance Criteria The acceptance criteria for the mechanical design of the reactor vessel internals components are:
(1) The core internals were designed to withstand mechanical loads arising from the DE, DDE, HE, and pipe ruptures (refer to Sections 3.7.3.15 and 3.9.2.1.3). The seismic and pipe rupture design of core internals is further discussed in Sections 3.7.3 and 3.9.3. This addresses GDC 2, 1967.
(2) The reactor has mechanical provisions to adequately support the core and internals and to ensure that the core is intact with acceptable heat transfer geometry following transients arising from abnormal operating conditions.
This addresses GDC 10, 1971.
(3) Following the design basis accident, the plant shall be capable of being shut down and cooled in an orderly fashion so that fuel cladding temperature is kept within 10 CFR 50.46 limits. This implies that the deformation of certain critical reactor internals must be kept sufficiently small to allow core cooling (refer to Section 3.9.2.3.1).
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DCPP UNITS 1 & 2 FSAR UPDATE 4.2.2.3 Reactor Vessel Internals Description The components of the reactor vessel internals consist of the lower core support structure (including the entire core barrel, the thermal shield on Unit 1, and the neutron shield pad assembly on Unit 2), the upper core support structure, and the incore instrumentation support structure. The reactor vessel internals support the core, maintain fuel alignment, limit fuel assembly movement, maintain alignment between fuel assemblies and CRDMs, direct coolant flow past the fuel elements, direct coolant flow to the pressure vessel head, and provide gamma and neutron shielding and guides for incore instrumentation.
In DCPP Unit 1, the coolant flows from the vessel inlet nozzles down the annulus between the core barrel and the vessel wall around the thermal shield. Most of the coolant then enters the plenum at the bottom of the vessel. The coolant then reverses and flows up through the core support and lower core plate. A small portion of the coolant passes through core barrel flow holes and flows downward between the core barrel and the baffle plate, providing additional cooling of the barrel, and enters the core between the bottom of the baffle plate and the top of the lower support plate. The coolant coming up through the lower core plate and the coolant coming out to the baffle region then passes through the core. After passing through the core, the coolant enters the upper support structure and flows radially to the core barrel outlet nozzles and directly through the vessel outlet nozzles. Additionally, a small amount of the entering flow is directed into the vessel head plenum and exits through the vessel outlet nozzles.
In DCPP Unit 2, all the flow entering the core barrel from the inlet nozzle flows downward around the thermal neutron pads into the lower plenum. The coolant then reverses and flows upwards through the lower core plate and into the core and the baffle barrel region. After passing through the core, the coolant enters the upper support structure and flows radially to the core barrel outlet nozzles and directly through the vessel outlet nozzles. A small portion passes up between the core barrel and the baffle plate, providing additional cooling of the barrel, and enters the upper support structure through flow holes in the baffle region top plate. Also for Unit 2, an additional modification has been made to reduce the upper head bulk fluid temperature to approximately T-cold. In this modification, reactor upper and lower internals were modified to provide additional flow in the upper head region.
The major material for the reactor vessel internals is Type 304 stainless steel. Parts not fabricated from Type 304 stainless steel include bolts and dowel pins, which are fabricated from Type 316 stainless steel, the radial support clevis inserts which are fabricated from alloy 600, and insert bolts, which are fabricated from alloy X-750. The Unit 1 reactor vessel internals hold-down springs are Type 304 stainless steel. The Unit 2 reactor vessel internals hold-down springs are made of Type 403 stainless steel treated in accordance with ASME BPVC-1971, Case Number 1337 to have a yield stress greater than 90,000 psi. Undue susceptibility to intergranular stress corrosion cracking is prevented by not using sensitized stainless steel, as recommended in Regulatory Guide 1.44, May 1973 (Reference 23).
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DCPP UNITS 1 & 2 FSAR UPDATE All reactor vessel internals are removable, thus permitting inspection of the vessel internal surface.
HISTORICAL INFORMATION IN ITALICS BELOW NOT REQUIRED TO BE REVISED.
With respect to previous plants, there is no change in the design configuration of reactor vessel internals and the reactor vessel internals core support structures. Moreover, since their mechanical properties (e.g., fuel assembly weight, beam stiffness) are virtually the same, the response of the reactor vessel internals core support structure will not change.
The qualification of identical plants by the first-of-a-kind analysis is further verified by the Internals Vibration Assurance Program discussed in Section 3.9.2.1.
4.2.2.3.1 Lower Core Support Structure The reactor vessel internals support member is the lower core support structure shown in Figures 4.2-9 and 4.2-10 for DCPP Unit 1 and Unit 2, respectively. This support structure assembly consists of the core barrel, the core baffle, the lower core plate and support columns, the thermal shield on Unit 1, and the neutron shield pad assembly on Unit 2 (the transition from a thermal shield to neutron shield pad assembly is explained in WCAP-7870 {Reference 7}), and the core support, which is welded to the core barrel.
All the major material for this structure is Type 304 stainless steel. The lower core support structure is supported at its upper flange from a ledge in the reactor vessel and its lower end is restrained from transverse motion by a radial support system attached to the vessel wall. Within the core barrel are an axial baffle and a lower core plate, both of which are attached to the core barrel wall and form the enclosure periphery of the core. The lower core support structure and core barrel provide passageways and direct the coolant flow. The lower core plate is positioned at the bottom level of the core below the baffle plates and provides support and orientation for the fuel assemblies.
The lower core plate contains the necessary flow distribution holes for each fuel assembly. On Unit 2, adequate coolant distribution is obtained through the use of the lower core plate and the flat core support. Unit 1 has a domed core support plate.
Adequate coolant distribution is obtained through the use of an intermediate flow diffuser plate and the lower core plate.
On Unit 1, the one-piece thermal shield is fixed to the core barrel at the top with rigid bolted connections. The bottom of the thermal shield is connected to the core barrel by means of axial flexures. Rectangular specimen guides in which material samples can be inserted, held by a preloaded spring device, and irradiated during reactor operation, are welded to the thermal shield.
On Unit 2, the neutron shield pad assembly, shown in Figure 4.2-11, consists of four panels, constructed of Type 304 stainless steel, that are bolted and pinned to the 4.2-29 Revision 24 September 2018
DCPP UNITS 1 & 2 FSAR UPDATE outside of the core barrel. Rectangular specimen guides in which material surveillance samples are inserted, held by a preloaded spring device, and irradiated during reactor operation, are bolted and pinned to the panels. Additional details of the neutron shielding pads and irradiation specimen holders are given in Reference 7.
Vertically downward loads from weight, fuel assembly preload, control rod dynamic loading, hydraulic loads, and earthquake acceleration are carried by the lower core plate into the lower core plate support flange on the core barrel shell, and through the lower support columns to the core support and then through the core barrel shell to the core barrel flange supported by the vessel flange. Transverse loads from earthquake acceleration, coolant cross flow, and vibration are carried by the core barrel shell and distributed between the lower radial support to the vessel wall and to the vessel flange.
Transverse loads of the fuel assemblies are transmitted to the core barrel shell by direct connection of the lower core plate to the barrel wall, and by upper core plate alignment pins that are welded into the core barrel.
The radial support system of the core barrel is accomplished by "key" and "keyway" joints to the reactor vessel wall. At six equally spaced points around the circumference, an Inconel clevis block is welded to the vessel inner diameter. An Inconel insert block is bolted to each of these clevis blocks, and has a keyway geometry. Opposite each of these is a key that is welded to the lower core support. During assembly, as the internals are lowered into the vessel, the keys engage the keyways in the axial direction.
Radial and axial expansions of the core barrel are accommodated, but this design restricts transverse movement of the core barrel. With this system, cyclic stresses in the internal structures are within the ASME BPVC Section III limits. In the event of an abnormal downward vertical displacement of the internals following a hypothetical failure, the load is transferred through energy absorbing devices of the lower internals to the vessel. The number and design of these absorbers are determined so as to limit the stresses imposed on all components (except the energy absorber) to less than yield stress (ASME BPVC Section III values).
To prevent fuel rod damage as a result of water jetting through lower internals baffle gaps in Unit 2, edge bolts have been added along the full length of the center injection baffle plate joints and the gaps have been peened after bolting. Unit 1 has edge bolts along the entire length of all corner and center injection baffle plate joints.
In addition, if baffle jetting is detected in Unit 2, anti-baffle jetting fuel clips may be used to dampen the amplitude of the fuel rod vibrations.
4.2.2.3.2 Upper Core Support Assembly The upper core support assembly, shown in Figures 4.2-12through 4.2-14, consists of the upper support assembly and the upper core plate between which are contained support columns and guide tube assemblies. The support columns establish the 4.2-30 Revision 24 September 2018
DCPP UNITS 1 & 2 FSAR UPDATE spacing between the upper support assembly and the upper core plate, and transmit the mechanical loadings between the upper support and upper core plate. The guide tube assemblies shield and guide the control rod drive shafts and control rods. Flow restrictors are installed in the guide tubes that formerly housed the part length CRDM drive shafts.
The upper core support assembly, which is removed as a unit during the refueling operation, is positioned in its proper orientation with respect to the lower support structure by slots in the upper core plate. Fuel assembly locating pins protrude from the bottom of the upper core plate and engage the fuel assemblies as the upper assembly is lowered into place, thus ensuring proper alignment of the lower core support structure, the upper core support assembly, the fuel assemblies, and control rods. The upper core support assembly is restrained from any axial movements by a large circumferential spring that rests between the upper barrel flange and the upper core support assembly. The spring is compressed when the reactor vessel head is installed on the pressure vessel.
Vertical loads from weight, earthquake acceleration, hydraulic loads, and fuel assembly preload are transmitted through the upper core plate, via the support columns, to the upper support assembly and then into the reactor vessel head. Transverse loads from coolant cross flow, earthquake acceleration, and possible vibrations are distributed by the support columns to the upper support and upper core plate. The upper support plate is particularly stiff to minimize deflection.
4.2.2.3.3 Incore Instrumentation Support Structures The incore instrumentation support structures consist of an upper system to convey and support thermocouples penetrating the vessel through the head, and a lower system to convey and support flux thimbles penetrating the vessel through the bottom (Figure 7.7-9 shows the basic flux-mapping system).
The upper system utilizes the reactor vessel head penetrations. Instrumentation port columns are slip-connected to in-line columns that are, in turn, fastened to the upper support plate. These port columns protrude through the head penetrations. The thermocouple conduits, made of Type 304 stainless steel, are supported from the columns of the upper core support system.
In addition to the upper incore instrumentation, there are reactor vessel bottom port columns that carry the retractable, cold-worked stainless steel flux thimbles that are pushed upward into the reactor core. Conduits extend from the bottom of the reactor vessel down through the concrete shield area and up to a thimble seal table. The thimbles are closed at the leading ends and serve as the pressure barrier between the reactor pressurized water and the containment atmosphere.
Mechanical seals between the retractable thimbles and conduits are provided at the seal table. During normal operation, the retractable thimbles are stationary and move 4.2-31 Revision 24 September 2018
DCPP UNITS 1 & 2 FSAR UPDATE only during refueling or for maintenance, at which time a space of approximately 15 feet above the seal table is cleared for the retraction operation.
The incore instrumentation support structure is designed for adequate support of instrumentation during reactor operation and is sturdy enough to resist damage or distortion under the conditions imposed by handling during the refueling sequence.
Reactor vessel surveillance specimen capsules are covered in Section 5.2.2.4.
4.2.2.4 Reactor Vessel Internals Design Evaluation The following show the acceptance criteria in Section 4.2.2.2 are satisfied.
4.2.2.4.1 Design Loading Conditions The design loading conditions for the reactor vessel internals are:
(1) Fuel assembly weight (2) Fuel assembly spring forces (3) Internals weight (4) Control rod scram (equivalent static load)
(5) Differential pressure (6) Spring preloads (7) Coolant flow forces (static)
(8) Temperature gradients (9) Differences in thermal expansion (a) Due to temperature differences (b) Due to expansion of different materials (10) Interference between components (11) Vibration (mechanically or hydraulically induced)
(12) One or more loops out of service (13) All operational transients listed in Table 5.2-4 4.2-32 Revision 24 September 2018
DCPP UNITS 1 & 2 FSAR UPDATE (14) Pump overspeed (15) Seismic loads (DE, DDE and HE)
(16) Blowdown forces (due to RCS branch line breaks)
Combined seismic and blowdown forces are included in the stress analysis by assuming the maximum amplitude of each force to act concurrently. In the original analyses, the blowdown forces were those resulting from breaks in the RCS cold and hot legs. However, with the acceptance of the DCPP LBB analysis by the NRC (Reference 30), the blowdown forces resulting from pipe rupture events in the main RCL piping no longer have to be considered in the design basis analyses. Only the much smaller loads from RCS branch line breaks have to be considered (refer to Section 3.6.2.1.1.1).
The design analysis ensures that allowable stress limits are not exceeded, that adequate design margin exists, and establishes deformation limits that are concerned primarily with components' operability. The stress limits are established not only to ensure that peak stresses do not reach unacceptable values, but also to limit the amplitude of the oscillatory stress component in consideration of material fatigue characteristics. Both low and high cycle fatigue stresses are considered when the allowable amplitude of oscillation is established. Dynamic analysis on the reactor internals is provided in Section 3.9.2.3.
As part of the evaluation of design loading conditions, extensive testing and inspections are performed from the initial selection of raw materials up to and including component installation and plant operation. Among these tests and inspections are those performed during component fabrication, plant construction, startup and checkout, and plant operation.
4.2.2.4.2 Design Loading Categories The combination of design loadings fits into either the normal, upset, or faulted conditions (refer to Table 5.2-4). The reactor vessel internals are designed to withstand stresses originating from RCS design transients, as summarized in Table 5.2-6.
Loads and deflections imposed on components due to shock and vibration are determined analytically and experimentally in both scaled models and operating reactors. The cyclic stresses due to these dynamic loads and deflections are combined with the stresses imposed by loads from component weights, hydraulic forces, and thermal gradients for the determination of the total stresses of the internals.
The scope and methodology of the stress analysis problem is discussed in Section 3.9.2.3.
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DCPP UNITS 1 & 2 FSAR UPDATE 4.2.2.4.3 Allowable Deflections For normal operating conditions, downward vertical deflection of the lower core support plate is negligible.
Limiting deflection values from the LOCA plus the earthquake (larger of the DDE or HE),
and for the deflection criteria of critical internal structures, are given in Section 3.9.2.3 and Table 3.9-10.
The criteria for the core drop accident are based on determining the total downward displacement of the internal structures following a hypothesized core drop resulting from loss of supports. The initial clearance between the secondary core support structures and the reactor vessel lower head in the hot condition is approximately 1/2 inch. An additional displacement of approximately 3/4 inch would occur due to strain of the energy absorbing devices of the secondary core support; thus the total drop distance is about 1-1/4 inches, which is insufficient to permit the tips of a fully withdrawn RCCA to come out of the guide thimble.
Specifically, the secondary core support is a device that will never be used, except during a hypothetical accident of the core support (core barrel, barrel flange, etc.).
There are four supports in each reactor. This device limits the fall of the core and absorbs the energy of the fall that otherwise would be imparted to the vessel. The energy of the fall is calculated assuming a complete and instantaneous failure of the primary core support and is absorbed during the plastic deformation of the controlled stainless steel volume loaded in tension.
4.2.2.4.4 Design Criteria Bases The structural adequacy of the reactor vessel internals is discussed in Section 3.9.2.3.5.1.
4.2.2.5 Safety Evaluation 4.2.2.5.1 General Design Criterion 2, 1967 - Performance Standards The design loading conditions for the reactor vessel internals include the additional forces that may be imposed by earthquakes. The design analysis ensures that allowable stress limits are not exceeded and establishes deformation limits that are concerned primarily with components' operability (refer to Section 4.2.2.4.1). Allowable deflections for normal operating conditions and the limiting deflection values from the LOCA plus the earthquake are discussed in Section 3.9.2.3.3.
Section 3.7.3.15.1 provides a detailed discussion of the seismic analyses for the reactor vessel internals.
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DCPP UNITS 1 & 2 FSAR UPDATE 4.2.2.5.2 General Design Criterion 4, 1987 - Environmental and Dynamic Effects Design Bases Protection from the dynamic effects of the most limiting breaks of auxiliary lines is considered. RCS branch line breaks and other high energy line breaks are provided.
Refer to Section 3.9.3.3 for additional discussion of LBB methodology and application.
The LOCA dynamic analyses are based on the accumulator, pressurizer surge, and residual heat removal system RCS branch line breaks credited LBB (Reference 32).
4.2.2.5.3 General Design Criterion 10, 1971 - Reactor Design The reactor vessel internals are designed with appropriate margin to assure that SAFDLs are not exceeded during any condition of normal operation, including the effects of anticipated operational occurrences. Section 4.2.2.4.2 documents that the reactor vessel internals are designed to withstand stresses originating from design transients. Section 3.9.2.3 discusses the design criteria used for normal operating conditions to evaluate calculated static and dynamic stresses. These calculated allowable stresses are considered appropriate and conservative.
4.2.2.5.4 Safety Function Requirements (1) Core Flow Distribution The components of the reactor vessel internals direct coolant flow past the fuel elements and direct coolant flow to the pressure vessel head (refer to Section 4.2.2.3).
The design loading conditions for the reactor vessel internals include temperature gradients and differences in thermal expansion due to temperature differences and the expansion of different materials (refer to Section 4.2.2.4.1).
(2) Protection of the Reactor Pressure Vessel from Neutron Exposure The Unit 1 one-piece thermal shield is fixed to the core barrel at the top with rigid bolted connections. The bottom of the thermal shield is connected to the core barrel by means of axial flexures. Rectangular specimen guides in which material samples can be inserted, held by a preloaded spring device, and irradiated during reactor operation, are welded to the thermal shield (refer to Section 4.2.2.3.1).
The Unit 2 neutron shield pad assembly consists of four panels, constructed of Type 304 stainless steel, that are bolted and pinned to the outside of the core barrel (refer to Figure 4.2-11). Rectangular specimen guides in which material surveillance samples are inserted, held by a preloaded spring device, and irradiated during reactor operation, are bolted and pinned to the panels. Additional details of the neutron shielding pads and irradiation specimen holders are given in Reference 7.
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DCPP UNITS 1 & 2 FSAR UPDATE The reactor vessel internals have provisions to install material test specimens for the reactor vessel material surveillance program (refer to Section 5.2.2.4.4).
(3) Incore Instrumentation The reactor vessel internals have provisions to install incore instrumentation. The incore instrumentation support structure is designed for adequate support of instrumentation during reactor operation (refer to Section 4.2.2.3.3).
4.2.3 REACTIVITY CONTROL SYSTEM 4.2.3.1 Design Bases 4.2.3.1.1 General Design Criterion 2, 1967 - Performance Standards The CRDMs are designed to withstand the effects of, or are protected against, natural phenomena, such as earthquakes.
4.2.3.1.2 General Design Criterion 4, 1987 - Environmental and Dynamic Effects Design Bases Consideration of the dynamic effects associated with main RCL piping postulated pipe ruptures are excluded from the DCPP design basis with the approval of LBB methodology by demonstrating that the probability of fluid system piping rupture is extremely low under conditions consistent with the design basis for the piping.
4.2.3.1.3 General Design Criterion 25, 1971 - Protection System Requirements for Reactivity Control Malfunctions The reactivity control system is designed to assure that no single malfunction (this does not include rod ejection) causes a violation of the acceptable fuel design limits.
4.2.3.1.4 General Design Criterion 26, 1971 - Reactivity Control System Redundancy And Capability The reactivity control system is provided with redundancy and capability such that two independent reactivity control systems of different design principles and capabilities of reliably controlling reactivity changes under conditions of normal operation, including anticipated operational occurrences, to assure acceptable fuel design limits are not exceeded.
In addition, one of the systems is capable of holding the reactor core subcritical under cold conditions. To meet this requirement the rod control system is designed to provide sufficient operational control and reliability during reactivity changes during normal and anticipated transients and the chemical and volume control system (CVCS) regulates the 4.2-36 Revision 24 September 2018
DCPP UNITS 1 & 2 FSAR UPDATE concentration of chemical neutron absorber in the reactor coolant to control reactivity changes.
4.2.3.1.5 General Design Criterion 29, 1971 - Protection Against Anticipated Operational Occurrences The reactivity control system is designed to ensure an extremely high probability of functioning during anticipated operational occurrences.
4.2.3.1.6 General Design Criterion 30, 1967 - Reactivity Holddown Capability At least one of the reactivity control systems provided is capable of making and holding the core subcritical under any conditions with appropriate margins for contingencies.
4.2.3.2 Reactivity Control System Acceptance Criteria 4.2.3.2.1 Design Stresses Acceptance Criteria The reactivity control system is designed to withstand stresses originating from the operating transients summarized in Table 5.2-4.
Allowable stresses for normal operating conditions are in accordance with ASME BPVC Section III. All components are analyzed as Class I components under Article NB-3000.
The cyclic stresses due to dynamic loads and deflections are combined with the stresses imposed by loads from component weights, hydraulic forces, and thermal gradients to determine the total stresses of the reactivity control system.
4.2.3.2.2 Material Compatibility Acceptance Criteria Materials are selected for compatibility in a PWR environment, adequate mechanical properties at room and operating temperature, resistance to adverse property changes in a radioactive environment, and compatibility with interfacing components.
4.2.3.2.3 Absorber Rods Acceptance Criteria The following design conditions, based on Article NB-3000 of the ASME BPVC Section III-1973, are considered.
(1) The external pressure equal to the RCS operating pressure (2) The wear allowance equivalent to 1000 reactor trips (3) Bending of the rod due to a misalignment in the guide tube (4) Forces imposed on the rods during rod drop 4.2-37 Revision 24 September 2018
DCPP UNITS 1 & 2 FSAR UPDATE (5) Loads caused by accelerations imposed by the CRDM (6) Radiation exposure for maximum core life.
The absorber material temperature shall not exceed its melting temperature (1470°F for silver-indium-cadmium absorber material {Reference 2}).
The Westinghouse RCCA and Enhanced Performance RCCA (EP-RCCA) model control rods that are cold-rolled Type 304 stainless steel is the only noncode material used in the control assembly. The stress intensity limit Sm for this material is defined as two-thirds of the 0.2 percent offset yield stress.
The Framatome RCCAs have an ion-nitrided 316L cladding material that improves wear resistance. The Framatome control rod noncode material stress intensity limit Sm is also two-thirds of the 0.2 percent offset yield stress.
4.2.3.2.4 Burnable Absorber Rods Acceptance Criteria The burnable absorber rod cladding (Zircaloy-4 for the wet annular burnable absorber
{WABA} design) is designed as a Class I component under Article NB-3000 of the ASME BPVC Section III-1973, for Conditions I and II. For Conditions III and IV loads, code stresses are not considered limiting. Failures of the burnable absorber rods during these conditions must not interfere with reactor shutdown or emergency cooling of the fuel rods.
The structural elements of the burnable absorber rod are designed to maintain absorber geometry. The rods are designed so that the Al2O3-B4C material is below 1200°F during normal operation or overpower transients.
4.2.3.2.5 Neutron Source Rods Acceptance Criteria The neutron source rods are designed to withstand:
(1) An external pressure equal to the RCS operating pressure (2) An internal pressure equal to the pressure generated by gases released over the neutron source rod life.
4.2.3.2.6 Thimble Plug Assembly Acceptance Criteria The thimble plug assemblies:
(1) Accommodate the differential thermal expansion between fuel assembly and core internals 4.2-38 Revision 24 September 2018
DCPP UNITS 1 & 2 FSAR UPDATE (2) Maintain positive contact with the fuel assembly and the core internals (3) Can be inserted into, or withdrawn from, the fuel assembly by a force not exceeding 65 pounds.
4.2.3.2.7 Control Rod Drive Mechanism Acceptance Criteria The CRDMs were designed to meet the following basic operational requirements:
(1) 5/8-inch per step (2) 142.5-inch travel (nominal)
(3) 360 pounds-force maximum load (4) Step in or out at 45 inches per minute (72 steps per minute) maximum (5) Power interruption shall initiate release of drive rod assembly (6) Trip delay of 150 milliseconds or less - Free fall of drive rod assembly shall begin less than 150 milliseconds after power interruption, no matter what holding or stepping action is being executed, with any load and coolant temperatures between 100°F and 650°F.
(7) 50-year design life with normal refurbishment (8) 13,200 complete travel excursions equaling 6 million steps with normal refurbishment 4.2.3.3 Reactivity Control System Description Reactivity control is provided by neutron absorbing rods and a soluble chemical neutron absorber (boric acid). The boric acid concentration is varied to control long-term reactivity changes such as:
(1) Fuel depletion and fission product buildup (2) Cold to hot, zero power reactivity change (3) Reactivity change produced by intermediate-term fission products such as xenon and samarium (4) Burnable poison depletion The concentration of boric acid in the reactor coolant is regulated by the CVCS, as described in Section 9.3.4.
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DCPP UNITS 1 & 2 FSAR UPDATE The RCCAs provide reactivity control for:
(1) Shutdown (2) Reactivity changes due to coolant temperature changes in the power range (3) Reactivity changes associated with the power coefficient of reactivity (4) Reactivity changes due to void formation The neutron source assemblies provide a means of verifying that the neutron instrumentation performs its function during periods of low neutron activity. They also provide the required count rate during startup.
The most effective reactivity control component is the RCCA and its corresponding drive rod assemblies. Figure 4.2-15 identifies the rod cluster control and drive rod assembly, in addition to the interfacing fuel assembly, guide tubes, and CRDM.
Guidance for the control rod cluster is provided by the guide tube, as shown in Figure 4.2-15. The guide tube provides two types of guidance:
(1) In the lower section, a continuous guidance system provides support immediately above the core. This system protects the rod against excessive deformation and wear due to hydraulic loading.
(2) The region above the continuous section provides support and guidance at uniformly spaced intervals.
The support envelope is determined by the RCCA pattern, as shown in Figure 4.2-16.
The guide tube ensures alignment and support of the control rods, spider body, and drive rod while maintaining trip times at or below required limits.
4.2.3.3.1 Reactivity Control Components Description The reactivity control components are subdivided into two categories:
(1) Permanent devices used to control or monitor the core (2) Optional burnable absorber assemblies The permanent type components are the RCCAs, control rod drive assemblies, and neutron source assemblies. Although the optional thimble plug assembly does not directly contribute to the reactivity control of the reactor, it is presented as a reactivity 4.2-40 Revision 24 September 2018
DCPP UNITS 1 & 2 FSAR UPDATE control system component in this document because it is used to restrict bypass flow through those thimbles not occupied by absorber, source or burnable poison rods.
The purpose of the optional burnable absorber assemblies is to control assembly power and ensure that the temperature coefficient of reactivity is less positive under normal operating conditions.
4.2.3.3.2 Rod Cluster Control Assembly Description The RCCA banks are divided into two categories: control and shutdown. Two criteria have been employed for selection of the control groups. First, the total reactivity worth must be adequate to meet the nuclear requirements. Second, because some of these rods may be partially inserted at power operation, the total power peaking factor should be low enough to ensure that the power capability is met. The control and shutdown groups provide adequate shutdown margin (SDM) which is defined as: the instantaneous amount of reactivity by which the reactor is subcritical, or would be subcritical from its present condition, assuming (1) all RCCAs are fully inserted, except for the single RCCA of highest reactivity worth which is assumed to be fully withdrawn (with any RCCA not capable of being fully inserted, the reactivity worth of the RCCA must be accounted for in the determination of SDM) and (2) when in MODE 1 or 2, the fuel and moderator temperatures are changed to the hot zero power temperatures.
An RCCA comprises a group of individual neutron absorber rods fastened at the top end to a common spider assembly, as illustrated in Figure 4.2-16.
The absorber material used in the control rods is a silver-indium-cadmium alloy that is essentially "black" to thermal neutrons and has sufficient additional resonance absorption to significantly increase its worth. The alloy is in the form of extruded rods that are sealed in stainless steel tubes to prevent the rods from coming in direct contact with the coolant. The silver-indium-cadmium rods are inserted into cold-worked stainless steel tubing. It is sealed at the bottom and top by welded end plugs, as shown in Figure 4.2-17. Sufficient diametral and end clearance is provided to accommodate relative thermal expansions.
The bullet-nosed bottom plugs reduce the hydraulic drag during reactor trip and guide the absorber rods smoothly into the dashpot section of the fuel assembly guide thimbles. The upper plug is threaded for assembly to the spider and has a reduced end section to make the joint more flexible.
The spider assembly is in the form of a central hub with radial vanes containing cylindrical fingers from which the absorber rods are suspended. Handling detents and detents for connection to the drive rod assembly are machined into the upper end of the 4.2-41 Revision 24 September 2018
DCPP UNITS 1 & 2 FSAR UPDATE hub. Coiled springs inside the spider body absorb the impact energy at the end of a trip insertion. All components of the spider assembly are made from Types 304 and 308 stainless steel, except for the retainer, which is made of 17-4 PH stainless steel material, and the springs, which are made of Inconel 718 alloy or, for the Westinghouse RCCA and EP-RCCA spiders only, an austenitic stainless steel where the springs do not contact the coolant. Other Framatome spider assembly components not made from 304 or 308 stainless steel are the spider itself, cast from Type 316L stainless steel, the cladding which is tempered and cold worked Type 316 stainless steel and the rod spring spacer which is Inconel 750. The absorber rods are fastened securely to the spider to ensure trouble-free service.
The overall length is such that when the assembly is withdrawn through its full travel, the tips of the absorber rods remain engaged in the guide thimbles so that alignment between rods and thimbles is always maintained. Because the rods are long and slender, they are relatively free to conform to any small misalignments with the guide thimble.
4.2.3.3.3 Burnable Absorber Assembly Description Each burnable absorber assembly consists of WABA burnable absorber rods attached to a hold down assembly.
A WABA rod (refer to Figure 4.2-18a) consists of annular pellets of alumina-boron carbide (Al2O3-B4C) burnable absorber material contained within two concentric Zircaloy tubes. These Zircaloy tubes, which form the inner and outer cladding for the WABA rod, are plugged and welded at each end to encapsulate the annular stack of absorber material. The assembled rod is then internally pressurized to 650 psig and seal welded.
The absorber stack lengths are positioned axially within the WABA rods by the use of Zircaloy bottom-end spacers. An annular plenum is provided within the rod to accommodate the helium gas released from absorber material depletion during irradiation. The reactor coolant flows inside the inner tube and outside the outer tube of the annular rod. Further design details are given in Section 3.0 of Reference 28.
The burnable absorber rods are statically suspended and positioned in selected guide thimbles within the fuel assemblies. The absorber rods in each assembly are attached together at the top end of the rods to a hold down assembly by a flat, perforated retaining plate which fits within the fuel assembly top nozzle and rests on the adapter plate. The absorber rod assembly is held down and restrained against vertical motion through a spring pack which is attached to the plate and is compressed by the upper core plate when the reactor upper internals assembly is lowered into the reactor. This arrangement ensures that the absorber rods cannot be ejected from the core by flow forces. Each rod is permanently attached to the base plate by a nut, which is locked into place.
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DCPP UNITS 1 & 2 FSAR UPDATE 4.2.3.3.4 Neutron Source Assembly Description The neutron source assembly provides a base neutron level to ensure that the detectors are operational and responding to core multiplication neutrons. Because there is very little neutron activity during core loading, refueling, hot and cold shutdown, and approach to criticality, neutron sources are placed in the reactor to help determine if source range detectors are properly responding.
During core loading, it is verified that active source assemblies provide the responding source range detectors with a sufficient count rate. For normal source range detectors (N-31 and N-32), and for alternate source range detectors (N-51 and N-52), the following count rate requirements must be met after an installed active source is neutronically coupled to a detector:
(1) For N-31 and N-32:
The count rate shall be equal to or greater than the maximum of the following count rates:
Twice the background radiation in counts per second (CPS), or 0.5 CPS + background radiation in CPS, or 1.0 CPS.
(2) For N-51 and N-52:
The count rate shall be equal to or greater than the maximum of the following count rates:
Twice the background radiation in CPS, or 0.05 CPS + background radiation in CPS, or 0.1 CPS.
The differences in required count rates are due to differences in detector sensitivity between the proportional counters (N-31 and N-32) and the fission chambers (N-51 and N-52).
The source assembly also permits detection of changes in the core multiplication factor during core loading, refueling, and approach to criticality. This can be done since the multiplication factor is related to an inverse function of the detector count rate.
Therefore, a change in the multiplication factor can be detected during addition of fuel assemblies while loading the core, a change in control rod positions, and changes in boron concentration.
HISTORICAL INFORMATION IN ITALICS BELOW NOT REQUIRED TO BE REVISED.
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DCPP UNITS 1 & 2 FSAR UPDATE The DCPP Unit 1 and Unit 2 reactor cores each employed two primary source assemblies and two secondary source assemblies in the first core. Each primary source assembly contained one primary source rod and between zero and twenty-three burnable absorber rods. The primary source rod, containing californium-252, spontaneously fissions and emits neutrons. After the primary source rod decays beyond the desired neutron flux level, neutrons are then supplied by the secondary source rod.
The secondary source rod contains a mixture of approximately half antimony and half beryllium by volume, which is activated by neutron bombardment during reactor operation. Activation of antimony results in the subsequent release of neutrons by the
(,n) reaction in beryllium. This becomes a source of neutrons during periods of low neutron flux, such as during refueling and subsequent startups.
Each of the two secondary source assemblies has six secondary source rods and no burnable poison rods (refer to Figure 4.2-21).
Neutron source assemblies are located at diametrically opposite sides of the core. The assemblies are inserted into the guide thimbles at selected unrodded locations.
The secondary source rods utilize slightly cold worked 304 SS material. The secondary source rods contain about 500 grams of stacked antimony-beryllium pellets, and the rod is internally pre-pressurized to 625 +/- 50 psig. The rods in each assembly are permanently fastened at the top end to a hold down assembly, which is identical to that of the burnable absorber assemblies.
The other structural members are fabricated from Type 304 and 308 stainless steel except for the springs exposed to the reactor coolant. They are wound from an age hardened nickel base alloy for corrosion resistance and high strength.
4.2.3.3.5 Thimble Plug Assembly Description Thimble plug assemblies are utilized, if desired, to further limit bypass flow through the guide thimbles in fuel assemblies that do not contain either control rods, source rods, or burnable absorber rods.
The thimble plug assemblies shown in Figure 4.2-22 consist of a flat base plate with short rods suspended from the bottom surface and a spring pack assembly. The 24 short rods, called thimble plugs, project into the upper ends of the guide thimbles to reduce the bypass flow area. Similar short rods may be also used on the source assemblies and burnable absorber assemblies to plug the ends of all vacant fuel assembly guide thimbles.
All components in the thimble plug assembly, except for the springs, are fabricated from Type 304 stainless steel. The springs are wound from an age-hardened nickel base alloy for corrosion resistance and high strength.
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DCPP UNITS 1 & 2 FSAR UPDATE 4.2.3.3.6 Control Rod Drive Mechanism Description All parts exposed to reactor coolant are made of metals that resist the corrosive action of the water. Three types of metals are used exclusively: stainless steels, nickel alloy, and cobalt-based alloys. Wherever magnetic flux is carried by parts exposed to the main coolant, 400 series stainless steel is used. Cobalt-based alloys are used for the pins and latch tips. Nickel alloy is used for the springs of the latch assemblies, and Type 304 stainless steel for all pressure-containing parts. Hard chrome plating provides wear surfaces on the sliding parts and prevents galling between mating parts.
A position indicator assembly slides over the CRDM rod travel housing. This position indicator assembly detects the drive rod assembly position by means of 42 discrete coils that magnetically sense the entry and presence of the rod drive line through its centerline over the normal length of the drive rod travel.
The CRDMs are located on the head of the reactor vessel. They are coupled to RCCAs. A CRDM schematic is shown in Figure 4.2-24.
The primary function of the CRDM is to insert or withdraw RCCAs into or from the core to control average core temperature and to shut down the reactor. The CRDM is a magnetically operated jack. A magnetic jack is an arrangement of three electromagnets that are energized in a controlled sequence by a power cycler to insert or withdraw the RCCAs of the reactor core in discrete steps. The CRDM consists of the pressure vessel, coil stack assembly, the latch assembly, and the drive rod assembly:
(1) The pressure vessel includes a one-piece integrated latch housing / head adaptor. A seal weld is located between the integrated latch housing and the rod travel housing. The integrated latch housing is butt welded to the CRDM nozzle; however, the butt weld is not part of the pressure housing assembly.
The latch housing is the lower portion of the vessel and contains the latch assembly. The rod travel housing is the upper portion of the vessel and provides space for the drive rod during its upward movement as the control rods are withdrawn from the core.
(2) The coil stack assembly includes the coil housings, an electrical conduit and connector, and three operating coils: (a) the stationary gripper coil, (b) the movable gripper coil, and (c) the lift coil.
Energizing the operation coils causes movement of the pole pieces and latches in the latch assembly.
(3) The latch assembly includes the guide tube, stationary pole pieces, movable pole pieces, and two sets of latches: (a) the movable gripper latch, and (b) the stationary gripper latch.
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DCPP UNITS 1 & 2 FSAR UPDATE The latches engage grooves in the drive rod assembly. The movable gripper latches are moved up or down in 5/8-inch steps by the lift pole to raise or lower the drive rod. The stationary gripper latches hold the drive rod assembly while the movable gripper latches are repositioned for the next 5/8-inch step.
(4) The drive rod assembly includes a flexible coupling, a drive rod, a disconnect button, a disconnect rod, and a locking button.
The drive rod has 5/8-inch grooves that receive the latches during holding or moving of the drive rod.
The disconnect button, disconnect rod, and locking button provide positive locking of the coupling to the RCCA and permit remote disconnection.
The CRDM has a trip design. Tripping can occur during any part of the power cycler sequencing if power to the coils is interrupted.
The mechanism can handle a 360-pound load, including the drive rod weight, at a rate of 45 inches per minute (72 steps per minute). Withdrawal of the RCCA is accomplished by magnetic forces while insertion is by gravity.
The mechanism internals are designed to operate in 650°F reactor coolant. The three operating coils are designed to operate at 392°F with forced air cooling required to maintain that temperature.
The CRDM, shown schematically in Figure 4.2-24, withdraws and inserts its control rod as electrical pulses are received by the operator coils. Position of the control rod is measured by 42 discrete coils mounted on the position indicator assembly surrounding the rod travel housing. Each coil magnetically senses the entry and presence of the top of the ferromagnetic drive rod assembly as it moves through the coil centerline.
During plant operation, the stationary gripper coil of the drive mechanism holds the control rod withdrawn from the core in a static position until the movable gripper coil is energized.
If power to the stationary gripper coil is cut off, the combined weight of the drive rod assembly and the RCCA is sufficient to move latches out of the drive rod assembly groove. The control rod falls by gravity into the core. The trip occurs as the magnetic field, holding the stationary gripper plunger half against the stationary gripper pole, collapses, and the stationary gripper plunger half is forced down by the weight acting upon the latches. After the drive rod assembly is released by the mechanism, it falls freely until the control rods enter the buffer section of their guide tubes.
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DCPP UNITS 1 & 2 FSAR UPDATE 4.2.3.4 Reactivity Control System Design Evaluation 4.2.3.4.1 Reactivity Control Components Design Evaluation The components are analyzed for loads corresponding to normal, upset, emergency, and faulted conditions. The analysis performed depends on the mode of operation under consideration.
The scope of the analysis requires many different techniques and methods, both static and dynamic.
Some of the loads that are considered on each component, where applicable, are:
(1) Control rod scram (equivalent static load)
(2) Differential pressure (3) Spring preloads (4) Coolant flow forces (static)
(5) Temperature gradients (6) Differences in thermal expansion (a) Due to temperature differences (b) Due to expansion of different materials (7) Interference between components (8) Vibration (mechanically or hydraulically induced)
(9) All operational transients listed in Table 5.2-4 (10) Pump overspeed (11) Seismic loads (refer to Section 3.7.3.15)
(12) Post-LOCA Rod Insertion (seismic plus LOCA loads)
The main objective of the analysis is to ensure that allowable stress limits are not exceeded, that an adequate design margin exists, and to establish deformation limits that are concerned primarily with the components' functioning. The stress limits are established not only to ensure that peak stresses will not reach unacceptable values, but also to limit the amplitude of the oscillatory stress component in consideration of 4.2-47 Revision 24 September 2018
DCPP UNITS 1 & 2 FSAR UPDATE fatigue characteristics of the materials. Standard methods of strength of materials are used to establish the stresses and deflections of these components. The dynamic behavior of the reactivity control components has been studied using experimental test data and experience from operating reactors.
The design of reactivity component rods provides sufficient cold void volume within the source rods to limit internal pressures to a value that satisfies the criteria in Section 4.2.1.2.3. The WABA rods have an annular plenum within the rod to accommodate the helium gas released from the absorber material during boron depletion (refer to Figure 4.2-18a). The internal pressure of source rods continues to increase from ambient until EOL; the internal pressure never exceeds that allowed by the criteria in Section 4.2.1.2.3. The stress analysis for the WABA rods assumed a maximum 30 percent gas release, consistent with Reference 28.
The WABA rod cladding and rod initial internal pressure have been designed so that the clad will not rely upon the pellets for support under all Condition I and Condition II events. Rodlet pre-pressurization will support the outer clad against irradiation induced creep collapse in the event of 0% gas release (worst case) from the Al2O3-B4C for the design life. Calculations also verify the clad integrity under Condition I and Condition II events with a conservative maximum gas release.
Thermal analyses have shown that the maximum absorber temperature is less than 1200°F for both normal operation (Condition I) and for Condition II upset events. This provides assurance that the He gas release will not exceed limits for the WABA rod's mechanical design life of 18,000 EFPH. This also assures that the Zircaloy clad strain limit is satisfied. A conservatively large heating rate was used in the thermal analyses.
The actual heating rate would be less for the reference design due to the lower B10 loading.
An evaluation of the WABA rod design is given in Reference 28.
Analysis of the RCCA spider indicates it is structurally adequate to withstand the various operating loads, including the higher loads that occur during the drive mechanism stepping action and rod drop. Verification of the spider structural capability has been experimentally demonstrated.
The material was selected on the basis of resistance to irradiation damage and compatibility with the reactor environment. No apparent degradation of construction material has occurred in operating plants with the DCPP reactivity control design.
Regarding material behavior in a radioactive environment, it should be noted that at high fluences, the austenitic material increases in strength with a corresponding decreased ductility (as measured by tensile tests), but energy absorption (as measured by impact tests) remains quite high. Corrosion of the material exposed to the coolant is quite low, and proper control of Cl- and O2 in the coolant prevents stress corrosion. All 4.2-48 Revision 24 September 2018
DCPP UNITS 1 & 2 FSAR UPDATE of the austenitic stainless steel base material used is processed and fabricated to preclude sensitization.
Analysis of the RCCA shows that if the drive mechanism housing ruptures, the RCCA will be ejected from the core by the pressure differential of the operating pressure and ambient pressure across the drive rod assembly. The ejection is also predicated on the failure of the drive mechanism to retain the drive rod/RCCA position. It should be pointed out that a drive mechanism housing rupture causes the ejection of only one RCCA with the other assemblies remaining in the core. For the Westinghouse RCCA only, analysis also showed that a pressure drop in excess of 4000 psi must occur across a two-fingered vane to break the vane/spider body joint, causing ejection of two neutron absorber rods from the core. Since the highest normal pressure of the primary system coolant is only 2250 psi, with the safety valves set to lift at 2485 psig, a pressure drop in excess of 4000 psi is not expected. Thus, ejection of the neutron absorber rods is not possible.
Ejection of a burnable absorber or thimble plug assembly is conceivable if one postulates that the hold-down bar fails and that the base plate and burnable absorber rods are severely deformed. In the unlikely event of hold-down bar failure, the upward displacement of the burnable absorber assembly only permits the base plate to contact the upper core plate. Since this displacement is small, the major portion of the absorber material remains positioned within the core. In the case of the thimble plug assembly, the thimble plugs will partially remain in the fuel assembly guide thimbles, thus maintaining a majority of the desired flow impedance. Further displacement or complete ejection would necessitate that the square base plate and burnable absorber rods be forced, thus plastically deformed, to fit up through a smaller diameter hole. As expected, this condition requires a substantially higher force or pressure drop than that of the hold-down bar failure.
To ensure subcriticality during hot leg recirculation following a LOCA (refer to Section 15.4.1.4.3), control rod insertion analyses and evaluations are performed that demonstrate that the control rods will insert under combined LOCA and seismic conditions for the applicable limiting design basis cold leg breaks.
Experience with control rods, burnable absorber rods, and source rods is discussed in Reference 8.
The mechanical design of the reactivity control components provides for the protection of the active elements to prevent the loss of control capability and functional failure of critical components. The components have been reviewed for potential failure and consequences of a functional failure of critical parts. The results of the review are summarized below.
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DCPP UNITS 1 & 2 FSAR UPDATE 4.2.3.4.2 Rod Cluster Control Assembly Design Evaluation (1) The basic absorbing material is sealed from contact with the primary coolant and the fuel assembly and guidance surfaces by a high quality stainless steel cladding. Potential loss of absorber mass or reduction in reactivity control material due to mechanical or chemical erosion or wear is therefore reliably prevented.
(2) A breach of the cladding for any postulated reason does not result in serious consequences. The silver-indium-cadmium absorber material is relatively inert and would still remain remote from high coolant velocity regions. Rapid loss of material resulting in significant loss of reactivity control material would not occur.
(3) The individually clad absorber rods are doubly secured to the retaining spider vane by a threaded joint and a welded lock pin. A failure of the joint would result in the insertion of the individual rod into the core. This results in reduced core reactivity which is a fail-safe condition.
(4) The spider finger braze joint that fastens the individual rods to the vanes on the Westinghouse RCCA and EP-RCCA models have also experienced many years of service, as described above, without failure. A failure of this joint would also result in insertion of the individual rod into the core. The Framatome RCCA spider is one-piece casting that includes vanes and fingers, a failure of which could also result in insertion of the individual rod into the core.
(5) The Westinghouse RCCA and EP-RCCA models radial vanes are brazed to the spider body and guidance of the RCCA is accomplished by the inner fingers of these vanes. They are therefore the most susceptible to mechanical damage. For the Framatome RCCA, the radial vanes are integral parts of the one-piece spider casting.
Failure of the vane-to-hub joint of a single rod vane could potentially result in failure of the separated vane and rod insertion. This could occur only at withdrawal elevation where the spider is above the continuous guidance section of the guide tube (in the upper internals). A rotation of the disconnected vane could cause it to hang on one of the guide cards in the intermediate guide tube. Such an occurrence would be evident from the failure of the RCCA to insert below a certain elevation, but with free motion above this point.
This possibility is considered extremely remote because the single rod vanes are subjected to only vertical loads and very light lateral reactions from the rods even during a seismic event (refer to Section 3.7.3.15). The consequences of such a failure are not considered critical since only one 4.2-50 Revision 24 September 2018
DCPP UNITS 1 & 2 FSAR UPDATE drive line of the reactivity control system would be involved. This condition is readily observed and can be cleared at shutdown.
(6) The spider hub, being of single unit cylindrical construction, is very rugged and has extremely low potential for damage. Should some unforeseen event cause fracture of the hub above the vanes, the lower portion with the vanes and rods attached would insert by gravity into the core causing reactivity decrease, again a fail-safe condition.
(7) The RCCA rods are provided a clear channel for insertion by the guide thimbles of the fuel assemblies. All fuel rod failures are protected against by providing this physical barrier between the fuel rod and the intended insertion channel. Distortion of the fuel rods by bending cannot apply sufficient force to damage or significantly distort the guide thimble. Fuel rod distortion by swelling, though precluded by design, would be terminated by fracture before contact with the guide thimble occurs. If such were not the case, a force reaction at the point of contact would cause a slight deflection of the guide thimble. The radius of curvature of the deflected shape of the guide thimbles would be sufficiently large to have a negligible influence on rod cluster control insertion.
4.2.3.4.3 Burnable Absorber Assemblies Design Evaluation The burnable absorber assemblies are static temporary reactivity control elements. The axial position is ensured by the holddown assembly that bears against the upper core plate. Their lateral position is maintained by the guide thimbles of the fuel assemblies.
The individual rods are shouldered against the underside of the retainer plate and securely fastened at the top by a threaded nut that is then locked in place. The square dimension of the retainer plate is larger than the diameter of the flow holes through the core plate. Therefore, failure of the holddown bar or spring pack does not result in ejection of the burnable absorber rods from the core.
The only incident that could result in ejection of the burnable absorber rods is a multiple fracture of the retainer plate. This is not considered credible because of the light loads borne by this component.
The burnable absorber rods are clad with either stainless steel or Zircaloy 4. The burnable absorber is Al2O3-B4C annular pellets contained within two concentric Zircaloy tubes. Burnable absorber rods are placed in static assemblies and are not subjected to motion which might damage the rods. Further, the guide thimble tubes of the fuel assembly afford additional protection from damage.
During the accumulated thousands of years of burnable absorber rodlet operating experience, only one instance of penetration of the stainless steel burnable absorber cladding has been observed. The consequences of cladding breach are also small. It 4.2-51 Revision 24 September 2018
DCPP UNITS 1 & 2 FSAR UPDATE is anticipated that upon cladding breach, the B4C would be leached by the coolant water and that localized power peaking of a few percent would occur; no design criteria would be violated. Additional information on the consequences of postulated WABA rod failures is presented in Reference 28.
4.2.3.4.4 Drive Rod Assemblies Design Evaluation All postulated failures of the drive rod assemblies, either by fracture or uncoupling, lead to the fail-safe condition. If the drive rod assembly fractures at any elevation, that portion remaining coupled falls with, and is guided by, the RCCA. This always results in a reactivity decrease.
4.2.3.4.5 Control Rod Drive Mechanism Design Evaluation The CRDMs are Code Class I components designed to meet the stress requirements for normal operating conditions of the ASME BPVC Section III, Division I, 2001 Edition through 2003 Addenda.
Structural analysis of the CRDMs was performed for normal and faulted conditions as described in Section 5.2.2.1.15.7.
4.2.3.4.5.1 Material Selection Design Evaluation Materials for all pressure-containing CRDM components comply with ASME BPVC Section III and were fabricated from austenitic (Type 304) stainless steel.
Magnetic pole pieces are fabricated from Type 410 stainless steel. All nonmagnetic parts, except pins and springs, are fabricated from Type 304 stainless steel. Cobalt alloy is used to fabricate link pins. Springs are made from nickel alloy. Latch arm tips are clad with Stellite 6 or ERCoCrA to provide improved wearability. Hard chrome plate and Stellite 6 or ERCoCrA are used selectively for bearing and wear surfaces.
The cast coil housings require a magnetic material. The choice was the ductile iron used in the CRDM. The finished housings are zinc-plated to provide corrosion resistance.
Coils are wound on bobbins of molded Dow Corning 302 material, with double glass-insulated copper wire. Coils are then vacuum-impregnated with silicon varnish. A wrapping of mica sheet is secured to the coil outer surface. The result is a well-insulated coil capable of sustained operation at 200°C.
The drive rod assembly uses a Type 410 stainless steel drive rod. The coupling is machined from Type 403 stainless steel. Other parts are Type 304 stainless steel with the exception of the springs, which are Inconel-X, and the locking button, which is Haynes 25.
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DCPP UNITS 1 & 2 FSAR UPDATE 4.2.3.4.5.2 Radiation Damage Design Evaluation As required by the equipment specification, the CRDMs are designed to accommodate a radiation dose rate of 10 rad/hr. The above radiation level, which amounts to 1.753 x 106 rads in 20 years, will not limit CRDM life.
4.2.3.4.5.3 Positioning Requirements Design Evaluation The mechanism has a step length of 5/8 inch that determines the positioning capabilities of the CRDM. (Note: Positioning requirements are determined by reactor physics.)
4.2.3.4.5.4 Evaluation of Materials' Adequacy Design Evaluation The ability of the pressure housing components to perform throughout the design lifetime as defined in the equipment specification is confirmed by the stress analysis report required by the ASME BPVC Section III.
The CRDM latch assembly may be significantly worn and should be replaced after 6 million individual steps.
4.2.3.4.5.5 Results of Dimensional and Tolerance Analysis With respect to the CRDM systems as a whole, critical clearances are present in the following areas:
(1) Latch assembly (diametral clearances)
(2) Latch arm-drive rod clearances (3) Coil stack assembly-thermal clearances (4) Coil fit in coil housing These clearances have been proven by life tests and actual field performance at operating plants:
(1) Latch Assembly - Thermal Clearances - The magnetic jack has several clearances where parts made of Type 410 stainless steel fit over parts made from Type 304 stainless steel. Differential thermal expansion is therefore important.
(2) Latch Arm - Drive Rod Clearances - The CRDM incorporates a load transfer action. The movable or stationary gripper latch is not under load during engagement due to load transfer action.
4.2-53 Revision 24 September 2018
DCPP UNITS 1 & 2 FSAR UPDATE Figure 4.2-25 shows latch clearance variation with the drive rod at minimum and maximum temperatures. Figure 4.2-26 shows clearance variations over the design temperature range.
(3) Coil Stack Assembly - Thermal Clearances - The assembly clearance of the coil stack assembly over the latch housing was selected so that the assembly could be removed under all anticipated conditions of thermal expansion.
(4) Coil Fit in Coil Housing - CRDM and coil housing clearances are selected so that coil heatup results in a close or tight fit. This facilitates thermal transfer and coil cooling in a hot CRDM.
4.2.3.5 Safety Evaluation 4.2.3.5.1 General Design Criterion 2, 1967 - Performance Standards A dynamic seismic analysis is performed on the CRDMs to confirm their ability to trip under a postulated seismic disturbance while maintaining resulting stresses under allowable values (refer to Section 5.2.2.1.15.7).
Additionally, seismic loads (DE, DDE, and HE) are considered on each component, where applicable (refer to Section 3.7.3.15.3).
The containment building, in which the CRDMs are located, is a PG&E Design Class I structure (refer to Section 3.8). As such it is designed to withstand the effects of winds (refer to Section 3.3), floods and tsunamis (refer to Section 3.4), external missiles (refer to Section 3.5), and earthquakes (refer to Section 3.7). This design protects the CRDMs, ensuring their safety function will be performed.
4.2.3.5.2 General Design Criterion 4, 1987 - Environmental and Dynamic Effects Design Bases The LBB methodology was applied to the primary loops of DCPP Unit 1 and Unit 2. The following postulated breaks were eliminated: the six terminal ends in the RCS cold, hot, and crossover legs; a split in the steam generator inlet elbow; and the loop closure weld in the crossover leg. Protection from the dynamic effects of the most limiting breaks of auxiliary lines is considered. RCS branch line breaks and other high energy line breaks are provided. Refer to Section 5.2 for additional discussion of LBB methodology and application.
The LOCA dynamic analyses are based on the accumulator, pressurizer surge, and residual heat removal system RCS branch line breaks credited LBB (Reference 32).
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DCPP UNITS 1 & 2 FSAR UPDATE 4.2.3.5.3 General Design Criterion 25, 1971 - Protection System Requirements for Reactivity Control Malfunctions The maximum reactivity insertion rate due to withdrawal of RCCAs, or by boron dilution, is limited as discussed in Section 4.3.4.4.
4.2.3.5.4 General Design Criterion 26, 1971 - Reactivity Control System Redundancy and Capability The rod control system relies on gravity to insert the rods into the core (refer to Section 4.2.3.3.6). The mechanical design of the reactivity control components provides for the protection of the active elements to prevent the loss of control capability and functional failure of critical components. The components have been reviewed for potential failure and consequences of a functional failure of critical parts (refer to Section 4.2.3.4).
The CVCS regulates the concentration of chemical neutron absorber in the reactor coolant to control reactivity changes resulting from the change in reactor coolant temperature between cold shutdown and hot full power (HFP) operation, burnup of fuel and burnable poisons, and xenon transients (refer to Section 9.3.4).
This documents that the reactivity control system is provided with redundancy and capability such that two independent reactivity control systems of different design principles and capable of reliably controlling reactivity changes under conditions of normal operation, including anticipated operational occurrences to assure acceptable fuel design limits are not exceeded and that one is capable of holding the reactor core subcritical under cold conditions.
4.2.3.5.5 General Design Criterion 29, 1971 - Protection Against Anticipated Operational Occurrences During plant operation, the stationary gripper coil of the drive mechanism holds the control rod withdrawn from the core in a static position until the movable gripper coil is energized. If power to the stationary gripper coil is cut off, the combined weight of the drive rod assembly and the RCCA is sufficient to move latches out of the drive rod assembly groove. The control rod falls by gravity into the core. After the drive rod assembly is released by the mechanism, it falls freely until the control rods enter the buffer section of their thimble tubes.
4.2.3.5.6 General Design Criterion 30, 1967 - Reactivity Holddown Capability The control and shutdown rod groups provide adequate SDM (refer to Section 4.2.3.3.2) which is defined as: the instantaneous amount of reactivity by which the reactor is subcritical, or would be subcritical from its present condition, assuming (1) All RCCAs are fully inserted, except for the single RCCA of highest reactivity worth which is assumed to be fully withdrawn (with any RCCA 4.2-55 Revision 24 September 2018
DCPP UNITS 1 & 2 FSAR UPDATE not capable of being fully inserted, the reactivity worth of the RCCA must be accounted for in the determination of SDM) and (2) When in MODE 1 or 2, the fuel and moderator temperatures are changed to the hot zero power temperatures.
4.2.3.6 Tests and Inspections 4.2.3.6.1 Reactivity Control Components Tests and inspections are performed on each reactivity control component to verify its mechanical characteristics. For the RCCA, prototype testing has been conducted and both manufacturing tests/inspections and functional testing are performed at the plant site.
During the component manufacturing phase, the following requirements apply to the reactivity control components to ensure proper functioning during reactor operation:
(1) To attain the desired standard of quality, all materials are procured to specifications.
(2) For the Westinghouse RCCA and EP-RCCA models only, all spiders are proof tested by an applied load to the spider body which is reacted on by the 16 peripheral, outermost fingers. This proof load subjects the spider assembly to a load greater than the acceleration loads caused by the CRDM stepping.
(3) All cladding/end plug welds are checked for integrity by visual inspection, X-ray, and helium leak tests. All the seal welds in the neutron absorber rods, burnable absorber rods, and source rods are checked in this manner.
(4) To ensure proper fitup with the fuel assembly, the rod cluster control, burnable absorber, and source assemblies are installed in the fuel assembly without restriction or binding in the dry condition.
HISTORICAL INFORMATION IN ITALICS BELOW NOT REQUIRED TO BE REVISED.
The RCCAs are functionally tested following initial core loading, but prior to criticality, to demonstrate reliable operation of the assemblies. Each assembly is operated (and tripped) once each at the following conditions:
no flow cold, full flow cold, no flow hot, and full flow hot. In addition, the slowest and fastest rods for each condition are tripped six more times.
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DCPP UNITS 1 & 2 FSAR UPDATE Rod drop tests following refueling outages are performed in accordance with the DCPP Technical Specifications (Reference 24) requirements.
4.2.3.6.2 Control Rod Drive Mechanisms The quality assurance program and testing for the CRDMs is discussed in the CRDM component design specification.
It is expected that all CRDMs will meet specified operating requirements for the duration of plant life with normal refurbishment. Technical Specifications define actions to be taken for rod group misalignment and rod inoperability.
To demonstrate continuous free movement of the RCCA and to ensure acceptable core power distributions during operation, partial-movement checks are performed in accordance with Technical Specifications. In addition, periodic drop tests of the RCCAs are performed at each refueling shutdown to demonstrate continued ability to meet trip time requirements. During these tests, the acceptable trip time of each assembly is not greater than the requirements listed in the Technical Specifications at full flow and operating temperature, from decay of the stationary gripper voltage to dashpot entry.
Periodic tests are also conducted during plant operation in accordance with the Technical Specifications.
4.2.3.7 Instrumentation Applications Instrumentation for determining reactor coolant average temperature is provided to create demand signals for moving groups of RCCAs to provide load follow (determined as a function of turbine impulse pressure) during normal operation, and to counteract operational transients. The hot and cold leg resistance temperature detectors (RTDs) are described in Section 7.2. The reactor control system, which controls the reactor coolant average temperature by regulation of control rod bank position, is described in Section 7.7.
Rod position indication instrumentation is provided to sense the actual position of each control rod so that it may be displayed to the operator. Signals are also supplied by this system as input to the rod deviation comparator. The rod position indication system is described in Chapter 7. The CVCS, one of whose functions is to permit adjustment of the reactor coolant boron concentration for reactivity control (as well as to maintain the desired operating fluid inventory in the volume control tank), consists of a group of instruments arranged to provide a manually preselected makeup composition that is borated or diluted, as required, to the charging pump suction header or the volume control tank. This system, as well as other systems, including boron sampling provisions that are part of the CVCS, is described in Section 9.3.4.
When the reactor is critical, the normal indication of reactivity status in the core is the position of the control bank in relation to reactor power (as indicated by the RCS loop 4.2-57 Revision 24 September 2018
DCPP UNITS 1 & 2 FSAR UPDATE T) and coolant average temperature. These parameters are used to calculate insertion limits for the control banks to warn the operator of excessive rod insertion.
Monitoring of the neutron flux for various phases of reactor power operation, as well as of core loading, shutdown, startup, and refueling is by means of the nuclear instrumentation system. The monitoring functions and readout and indication characteristics for the following reactivity monitoring systems are included in the discussion on the safety parameter display system (SPDS) in Section 7.5:
(1) Nuclear instrumentation system (2) Temperature indicators (a) T average (measured)
(b) T (measured)
(c) Auctioneered T average (3) Demand position of RCCA group (4) Actual rod position indicator.
4.
2.4 REFERENCES
- 1. J. A. Christensen, et al, Melting Point of Irradiated UO2, WCAP-6065, February 1965.
- 2. J. Cohen, Development and Properties of Silver-Based Alloys as Control Rod Materials for Pressurized Water Reactors, WAPD-214, December 1959.
- 3. Deleted in Revision 23.
- 4. Deleted in Revision 23.
- 5. F.T. Eggleston, Safety-Related Research and Development for Westinghouse Pressurized Water Reactors - Program Summaries, Winter 1977 - Summer 1978, WCAP-8768, Revision 2, October 1978.
- 6. Deleted.
- 7. S. Kraus, Neutron Shielding Pads, WCAP-7870, May 1972.
- 8. Westinghouse Electric Company, Operational Experience with Westinghouse Cores, WCAP-8183.
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- 9. J. M. Hellman (Ed.), Fuel Densification Experimental Results and Model For Reactor Application, WCAP-8218-P-A, March 1975 (Westinghouse Proprietary) and WCAP-8219-A, March 1975.
- 10. Deleted.
- 11. Deleted.
- 12. W. J. O'Donnell and B. F. Langer, "Fatigue Design Basis for Zircaloy Components," Nuclear Science and Engineering, 20, 1-12, 1964.
- 13. E. E. DeMario and S. Nakazato, Hydraulic Flow Test of the 17 x 17 Fuel Assembly, WCAP-8279, February 1974.
- 14. D. H. Risher, et al, Safety Analysis for the Revised Fuel Rod Internal Pressure Design Basis, WCAP-8963-P-A, August 1978.
- 15. R. A. George, et al, Revised Clad Flattening Model, WCAP-8377 (Westinghouse Proprietary) and WCAP-8381, July 1974.
- 16. Deleted.
- 17. Letter from T. M. Anderson (Westinghouse) to D. G. Eisenhut (NRC), Integrity of Control Rod Guide Thimble, NS-TMA-1936, September 1978.
- 18. Letter from T. M. Anderson (Westinghouse) to D. G. Eisenhut (NRC), Additional Information - Integrity of CRGT, NS-TMA-1992, December 1978.
- 19. Letter from T. M. Anderson (Westinghouse) to D. G. Eisenhut (NRC), Guide Thimble Tube Wear, NS-TMA-2102, June 1979.
- 20. Letter from P. A. Crane (PG&E) to J. F. Stolz (NRC), Response to NRC Questions on Guide Tube Wear, March 1980.
- 21. H. Kunishi and G.R. Schmidt, J. Skaritka (Ed.), Salem Unit 1 17 x 17 Fuel Assembly Guide Thimble Tube Wear Examination Report, Westinghouse Report, January 1982.
- 22. R. L. Cloud, et al. (Ed.), Pressure Vessels and Piping: Design and Analysis, Volume 1, Chapter 1, The American Society of Mechanical Engineers, 1972.
- 23. Regulatory Guide 1.44, Control of the Use of Sensitized Stainless Steel, USNRC, May 1973.
- 24. Technical Specifications, Diablo Canyon Power Plant Units 1 and 2, Appendix A to License Nos. DPR-80 and DPR-82, as amended.
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- 25. S. L. Davidson (Ed.), et al, Extended Burnup Evaluation of Westinghouse Fuel, WCAP-10125-P-A, December 1985.
- 26. S. L. Davidson (Ed.), Reference Core Report - VANTAGE 5 Fuel Assembly, WCAP-10444-P-A, September 1985 (Westinghouse Proprietary) and WCAP-10445-NP-A, September 1985.
- 27. S. L. Davidson (Ed.), et al, Verification Testing and Analysis of the 17 x 17 Optimized Fuel Assembly, WCAP-9401-P-A, August 1981.
- 28. J. Skaritka, Westinghouse Wet Annular Burnable Absorber Evaluation Report, WCAP-10021-P-A, Revision 1, October 1983.
- 29. S. L. Davidson (Ed.), et al, VANTAGE+ Fuel Assembly Reference Core Report, WCAP-12610-P-A, April 1995.
- 30. Letter from S.R. Peterson (NRC) to G.M. Rueger (PG&E), Leak-Before-Break Evaluation of Reactor Coolant System Piping for DCPP Units 1 and 2, (Docket Nos. 50-275 and 50-323), March 1993.
- 31. Deleted
- 32. D. Staub, et al., Diablo Canyon Replacement Reactor Vessel Closure Head and Integrated Head Assembly Project - Impact of IHA on Reactor Vessel, Internals, Fuel and Loop Piping, WCAP-16946-P, Revision 3, October 2011.
- 33. P. J. Kersting, et al., Assessment of Clad Flattening and Densification Power Spike Factor Elimination in Westinghouse Nuclear Fuel, WCAP-13589-A, March 1995 (Westinghouse Proprietary) and WCAP-14297-A, March 1995.
- 34. J.P. Foster, et al., Westinghouse Improved Performance Analysis and Design Model (PAD 4.0), WCAP-15063-P-A, Revision 1, with Errata, July 2000.
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DCPP UNITS 1 & 2 FSAR UPDATE 4.3 NUCLEAR DESIGN The nuclear design of the reactors for DCPP Unit 1 and Unit 2, including fuel and reactivity control systems, is described in this section; the analytical methods used in reactor design and evaluation are also discussed.
Before discussing the nuclear design bases, a brief review of the four major plant operation conditions, categorized in accordance with their anticipated frequency of occurrence and risk to the public, (refer to Section 4.2) are as follows:
(1) Condition I - Normal Operation (2) Condition II - Incidents of Moderate Frequency (3) Condition III - Infrequent Faults (4) Condition IV - Limiting Faults In general, Condition I occurrences are accommodated with margin between any plant parameter and the value of that parameter which would require either automatic or manual protective action. Condition II incidents are accommodated with, at most, a shutdown of the reactor with the plant capable of returning to operation after corrective action. Fuel damage (penetration of the fission product barrier; i.e., the fuel rod cladding) is not expected during Conditions I and II events. It is not possible, however, to preclude a very small number of rod failures. These are within the capability of the plant cleanup system and are consistent with the plant design bases.
Condition III incidents shall not cause more than a small fraction of the fuel elements in the reactor to be damaged, although sufficient fuel element damage might occur to preclude immediate resumption of operation. The release of radioactive material due to Condition III incidents should not be sufficient to interrupt or restrict public use of those areas beyond the exclusion radius. Furthermore, a Condition III incident shall not, by itself, generate a Condition IV fault or result in a consequential loss of function of the RCS or reactor containment barriers. Condition IV occurrences are faults that are not expected to occur, but are defined as limiting faults that must be considered in design.
Condition IV faults shall not cause a release of radioactive material that results in an undue risk to public health and safety.
The core design power distribution limits related to fuel integrity are met for Condition I occurrences through conservative design, and maintained by the action of the control system. The requirements for Condition II occurrences are met by providing an adequate protection system that monitors reactor parameters. The control and protection systems are described in Chapter 7 and the consequences of Conditions II, III, and IV occurrences are discussed in Chapter 15.
4.3-1 Revision 24 September 2018
DCPP UNITS 1 & 2 FSAR UPDATE 4.3.1 DESIGN BASES The design bases and functional requirements for the nuclear design of the fuel and reactivity control system, and the relationships of these design bases to the GDC of July 1971, are presented in this section. Where appropriate, supplemental criteria, such as 10 CFR 50.46, are addressed.
4.3.1.1 General Design Criterion 10, 1971 - Reactor Design The reactor core and associated coolant, control and protection systems are designed with appropriate margin to assure that SAFDLs are not exceeded during any condition of normal operation, including the effects of anticipated operational occurrences.
4.3.1.2 General Design Criterion 11, 1971 - Reactor Inherent Protection The reactor core is designed so that in the power operating range, the prompt inherent nuclear feedback characteristics tend to compensate for a rapid increase in reactivity.
4.3.1.3 General Design Criterion 12, 1971 - Suppression of Reactor Power Oscillations The reactor core is designed to assure that power oscillations that could result in conditions exceeding SAFDLs are not possible or can be reliably and readily detected and suppressed.
4.3.1.4 General Design Criterion 25, 1971 - Protection System Requirements for Reactivity Control Malfunctions The reactivity control system is designed to assure that no single malfunction (this does not include rod ejection) causes a violation of the acceptable fuel design limits.
4.3.1.5 General Design Criterion 26, 1971 - Reactivity Control System Redundancy and Capability Two independent reactivity control systems of different design principles are provided.
Each system has the capability to control the rate of reactivity changes resulting from planned, normal power changes. One of the systems is capable of reliably controlling anticipated operational occurrences. In addition, one of the systems is capable of holding the reactor core subcritical under cold conditions.
4.3.1.6 General Design Criterion 28, 1971 - Reactivity Limits The reactivity control systems are designed to assure that the effects of postulated reactivity accidents neither result in damage to the reactor coolant pressure boundary greater than limited local yielding nor cause sufficient damage to significantly impair the capability to cool the core.
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DCPP UNITS 1 & 2 FSAR UPDATE 4.3.2 NUCLEAR DESIGN ACCEPTANCE CRITERIA 4.3.2.1 Fuel Burnup Sufficient reactivity should be incorporated in the fuel to attain a desired region average discharge burnup.
4.3.2.2 Control of Power Distribution The nuclear design basis, with at least a 95 percent confidence level, is as follows:
(1) The fuel will not be operated at greater than 14.3 kW/ft under normal operating conditions, including an allowance of 2 percent for calorimetric error and not including the power spike factor due to densification effects (Reference 31).
(2) Under abnormal conditions, including the maximum overpower condition, the fuel peak power will not cause melting as defined in Section 4.4.2.2.
(3) The fuel will not operate with a power distribution that violates the DNB design basis (i.e., the departure from nucleate boiling ratio {DNBR} shall not be less than the design limit DNBR, as discussed in Section 4.4.2) under Conditions I and II events, including the maximum overpower condition.
(4) Fuel management will be such as to produce fuel rod powers and burnups consistent with the assumptions in the fuel rod mechanical integrity analysis of Section 4.2.
4.3.2.3 Negative Reactivity Feedbacks (Reactivity Coefficients)
The fuel temperature coefficient of reactivity will be negative, and the MTC of reactivity will be nonpositive for full power operating conditions, thus providing negative reactivity feedback characteristics over the operating range. Below 70 percent power, an MTC of up to +5 pcm (percent mille)/°F is allowed. From 70 percent to 100 percent the MTC limit decreases linearly from +5 to 0 pcm/°F.
4.3.2.4 Stability Spatial power oscillations within the core, with a constant core power output, should they occur, can be reliably and readily detected and suppressed.
4.3-3 Revision 24 September 2018
DCPP UNITS 1 & 2 FSAR UPDATE 4.3.2.5 Maximum Controlled Reactivity Insertion Rate The maximum reactivity insertion rate due to withdrawal of RCCAs, or by boron dilution, is limited. This limit, expressed as a maximum reactivity change rate (75 pcm/sec) is set such that the peak heat generation rate does not exceed the maximum allowable, and DNBR is not below the minimum allowable at overpower conditions (refer to Note (b) of Table 4.3-1).
4.3.2.6 Shutdown Margins Minimum SDM, as specified in the Core Operating Limits Reports, is required in all operational modes.
In all analyses involving reactor trip, the single, highest worth RCCA is postulated to remain untripped in its full-out position (stuck rod criterion).
4.
3.3 DESCRIPTION
4.3.3.1 Nuclear Design Description The reactor core consists of 193 fuel assemblies arranged in a pattern that approximates a right circular cylinder. Each fuel assembly contains a 17 x 17 rod array composed of 264 fuel rods, 24 RCCA guide tubes, and an incore instrumentation thimble. Each rod is held in place by spacer grids and top and bottom nozzles. The fuel rods are constructed of zirconium alloy tubing containing enriched UO2 fuel pellets.
A limited substitution of fuel rods by filler rods of zirconium alloy or stainless steel may be made for a particular design if justified by a cycle-specific reload analysis. Figure 4.2-1 shows a cross-sectional view of a fuel assembly and the related RCCA locations.
The fuel assembly design is discussed in Section 4.2.1.
All the fuel rods within a given assembly generally have the same nominal uranium enrichment. The exceptions are that the top and bottom portions of the rods may contain a low enriched or natural uranium blanket and that some assemblies may contain more than one enrichment as a result of reconstitution operations. Figure 4.3-1 shows a typical equilibrium 18-month cycle core loading of fresh and burned fuel assemblies. This "typical" loading pattern is modified for fuel cycles of longer length to accommodate the needed additional cycle energy.
A typical reload pattern employs low leakage fuel management in which more highly burned fuel is placed on the core periphery. Reload cores will operate approximately 12 months to 24 months between refuelings. The feed fuel enrichment is determined by the amount of fissionable material required to provide the desired core lifetime and energy production. Reactivity losses due to U-235 depletion and the buildup of fission products are partially offset by the buildup of plutonium produced by the capture of neutrons in U-238, as shown in Figure 4.3-2. At the beginning of any cycle, an excess reactivity to compensate for these losses over the specified cycle life must be "built" into 4.3-4 Revision 24 September 2018
DCPP UNITS 1 & 2 FSAR UPDATE the reactor. This excess reactivity is controlled by removable neutron absorbing material in the form of boron dissolved in the primary coolant and burnable absorber rods or boron coated fuel pellets.
Boric acid concentration in the primary coolant is varied to control and to compensate for long-term reactivity requirements, such as those due to fuel burnup, fission product poisoning, including xenon and samarium, burnable absorber material depletion, and the cold-to-operating moderator temperature change. Using its normal makeup path, the CVCS is capable of inserting negative reactivity at a rate of approximately 30 pcm/min when the reactor coolant boron concentration is 100 ppm. In an emergency, the CVCS can insert negative reactivity at approximately 65 pcm/min when the reactor coolant concentration is 1000 ppm, and 75 pcm/min when the reactor coolant boron concentration is 100 ppm. The peak xenon burnout rate is 25 pcm/min (Section 9.3.4 discusses the capability of the CVCS to counteract xenon decay). Rapid transient reactivity requirements and safe shutdown requirements are met with control rods.
As the boron concentration increases, the MTC becomes less negative. Using soluble poison alone would result in a positive MTC at BOL at full power operating conditions.
Therefore, burnable absorber and IFBA rods are used to reduce the soluble boron concentration sufficiently to ensure that the MTC is not positive for full power operating conditions. During operation, the absorber content in these rods is depleted, thus adding positive reactivity to offset some of the negative reactivity from fuel depletion and fission product buildup. The depletion rate of the burnable absorber material is not critical since chemical shim is always available and flexible enough to cover any possible deviations in the expected burnable absorber depletion rate. Figure 4.3-3 shows typical core depletion curves with burnable absorbers.
In addition to reactivity control, the burnable absorbers are strategically located to provide a favorable radial power distribution. Figures 4.3-4 and 4.3-5 show the typical burnable absorber distribution within a fuel assembly for the several burnable absorber patterns used for both discrete and IFBAs. The burnable absorber loading pattern for a typical equilibrium cycle reload core is shown in Figure 4.3-6.
Tables 4.1-1, and 4.3-1 through 4.3-3, summarize the reactor core design parameters for a typical reload fuel cycle, including reactivity coefficients, delayed neutron fraction, and neutron lifetimes.
4.3.3.2 Power Distribution DCPP employs two methods for performing core power distribution calculations: the power distribution monitoring system (PDMS) and the movable incore detector system (MIDS).
The PDMS generates a continuous measurement of the core power distribution using the methodology documented in References 32 and 33. The measured core power 4.3-5 Revision 24 September 2018
DCPP UNITS 1 & 2 FSAR UPDATE distribution is used to determine the most limiting core peaking factors, which are used to verify that the reactor is operating within the design limits.
The PDMS requires information on current plant and core conditions in order to determine the core power distribution using the core peaking factor measurement and measurement uncertainty methodology described in References 32 and 33. The core and plant condition information is used as input to the continuous core power distribution measurement software that continuously and automatically determines the current core peaking factor values. The core power distribution calculation software provides the measured peaking factor values at nominal one-minute intervals to allow operators to confirm that the core peaking factors are within design limits. In order for the PDMS to accurately determine the peaking factor values, the core power distribution measurement software requires accurate information about the current reactor power level average reactor vessel inlet temperature, control bank positions, the power range detector currents, and the core exit thermocouples.
Data obtained from the MIDS, described in Section 7.7.2.9.2, are used to calibrate the PDMS, and may also be used independent of the PDMS to generate a flux map of the core power distribution. The accuracy of these power distribution calculations has been confirmed under conditions similar to those expected for DCPP, as discussed in References 1 and 3 and in Section 4.3.3.2.7.
4.3.3.2.1 Definitions Power distributions are quantified in terms of hot channel factors. These factors are a measure of the peak pellet power within the reactor core and the total energy produced in a coolant channel and are expressed in terms of quantities related to the nuclear or thermal design; namely:
Power density is the thermal power produced per unit volume of the core (kW/liter).
Linear power density is the thermal power produced per unit length of active fuel (kW/ft).
Since fuel assembly geometry is standardized, this is the unit of power density most commonly used. For all practical purposes, it differs from kW/liter by a constant factor that includes geometry effects and the fraction of the total thermal power which is generated in the fuel rods.
Average linear power density is the total thermal power produced in the fuel rods divided by the total active fuel length of all rods in the core.
Local heat flux is the heat flux at the surface of the cladding (Btu ft-2hr-1). For nominal fuel rod parameters, this differs from linear power density by a constant factor.
Rod power or rod integral power is the linear power density in one rod integrated over its length (kW).
4.3-6 Revision 24 September 2018
DCPP UNITS 1 & 2 FSAR UPDATE Average rod power is the total thermal power produced in the fuel rods divided by the number of fuel rods.
The hot channel factors used in the discussion of power distributions in this section are defined as follows:
T F Q
, heat flux hot channel factor, is defined as the maximum local heat flux on the surface of a fuel rod divided by the average fuel rod heat flux, allowing for manufacturing tolerances on fuel pellets and rods.
N F Q
, nuclear heat flux hot channel factor, is defined as the maximum local fuel rod linear power density divided by the average fuel rod linear power density, assuming nominal fuel pellet and rod parameters. (No densification effects included.)
E F Q
, engineering heat flux hot channel factor, is the allowance on heat flux required for manufacturing tolerances. The engineering factor allows for local variations in enrichment, pellet density and diameter, surface area of the fuel rod, and eccentricity of the gap between pellet and clad.
Combined statistically, the net effect is a factor of 1.03 to be applied to the fuel rod surface heat flux.
N F H
, nuclear enthalpy rise hot channel factor, is defined as the ratio of the integral of linear power along the rod with the highest integrated power to the average rod power.
Manufacturing tolerances, hot channel power distribution, and surrounding channel power distributions are treated explicitly in the calculation of DNBR described in Section 4.4.
T For the purposes of discussion, it is convenient to define subfactors of F Q
- design limits are set, however, in terms of the total peaking factor
T FQ Total peaking factor or heat flux hot channel factor Maximum kW/ft (4.3-1)
Average kW/ft without densification effects.
T N E FQ F Q F Q
= max [F XYN (z) P(z)] F UN F QE (4.3-2) 4.3-7 Revision 24 September 2018
DCPP UNITS 1 & 2 FSAR UPDATE where:
N E F Q and F Q are defined above.
F UN = the measurement uncertainty associated with a full core flux map with movable detectors or PDMS N
F XY (z) = ratio of peak power density to average power density in the horizontal plane of peak local power P(z) = ratio of the power per unit core height in the horizontal plane at elevation Z to the average value of power per unit core height 4.3.3.2.2 Radial Power Distributions The power shape in horizontal sections of the core at full power is a function of the fuel and burnable absorber loading patterns, and the presence or absence of a single bank of control rods. Thus, at any time in the cycle, any horizontal section of the core can be characterized as unrodded, or with control banks inserted. These two situations, combined with burnup effects, determine the radial power shapes that can exist in the core at full power. The effects on radial power shapes of power level, xenon, samarium, and moderator density effects are also considered, but these are smaller. While radial power distributions in various planes of the core are often illustrated, the core radial enthalpy rise distribution, as determined by the power integral of each channel, is of greater interest. Figures 4.3-7 through 4.3-12 show representative radial power distributions for one-eighth of the core for representative operating conditions during the initial cycle, as follows:
Figure Conditions 4.3-7 HFP at BOL unrodded no xenon 4.3-8 HFP at BOL unrodded equilibrium xenon 4.3-9 HFP at BOL Bank D in equilibrium xenon - Unit 1 4.3-10 HFP at BOL Bank D in equilibrium xenon - Unit 2 4.3-11 HFP at middle of life (MOL) unrodded equilibrium xenon, and 4.3-12 HFP at EOL unrodded equilibrium xenon.
Since hot channel location varies from time to time, a single reference radial design power distribution is selected for DNB calculations. This reference power distribution, normalized to core average power, is chosen conservatively to concentrate power in one area of the core, minimizing the benefits of flow redistribution.
4.3-8 Revision 24 September 2018
DCPP UNITS 1 & 2 FSAR UPDATE 4.3.3.2.3 Assembly Power Distributions For the purpose of illustration, assembly power distributions for the BOL and EOL conditions corresponding to Figures 4.3-8 and 4.3-12 are given for the same assembly in Figures 4.3-13 and 4.3-14, respectively.
Since the detailed power distribution surrounding the hot channel varies from time to time, a conservatively flat assembly power distribution is assumed in the DNB analysis, described in Section 4.4, with the rod of maximum integrated power artificially raised to the design value of F NH . The nuclear design considers all fuel cycles and all operating conditions to ensure that a flatter assembly power distribution does not occur with limiting values of F NH .
4.3.3.2.4 Axial Power Distributions The shape of the power profile in the axial direction is largely under the control of the operator through the manual operation of the control rods or the automatic motion of the control rods responding to manual operation of the CVCS. Nuclear effects that cause variations in the axial power shape include moderator density, Doppler effect on resonance absorption, spatial xenon variations, fuel and burnable absorber material distribution and burnup. Automatically controlled variations in total power output and control rod motion are also important in determining the axial power shape at any time.
Signals are available to the operator from the excore ion chambers that run parallel to the axis of the core. Separate signals are taken from the top and bottom halves of the chambers. The difference between top and bottom signals for each of four pairs of detectors is called the flux difference, . If it deviates from the flux difference target band, an alarm is actuated.
Calculations of core average peaking factor for many plants and measurements from operating situations are associated with either or axial offset (AO) to place an upper bound on the peaking factor. For these correlations, AO is defined as:
t b AO = (4.3-4) t b where:
t and b are the top and bottom detector readings.
Representative axial power shapes for BOL, MOL, and EOL conditions covering a wide range, including power shape changes achieved by skewing xenon distributions, are shown in Figures 4.3-15 through 4.3-17.
4.3-9 Revision 24 September 2018
DCPP UNITS 1 & 2 FSAR UPDATE 4.3.3.2.5 Local Power Peaking Fuel densification causes fuel pellets to shrink both axially and radially. Pellet shrinkage combined with random hang-up of fuel pellets results in gaps in the fuel column when the pellets below the hung-up pellet settle in the fuel rod. The gaps vary in length and location in the fuel rod. Because of decreased neutron absorption in the vicinity of the gap, power peaking occurs in the adjacent fuel rods resulting in an increased power peaking factor. A quantitative measure of this local power peaking is given by the power spike factor S(z) where z is the axial location in the core.
In previous analyses of power peaking factors for DCPP Unit 1 and Unit 2, it was necessary to apply a penalty on calculated overpower transient FQ values to allow for interpellet gaps caused by pellet hang-ups and pellet shrinkage due to densification (Reference 22). This penalty is known as the densification spike factor. However, studies have shown (Reference 31) that this penalty can be eliminated for the fuel type present in the DCPP Unit 1 and Unit 2 cores.
4.3.3.2.6 Limiting Power Distributions As discussed in Section 4.3, Condition I occurrences are those expected frequently or regularly in the course of power operation, maintenance, or maneuvering of the plant.
Condition I occurrences are accommodated with margin between any plant parameter and the value of that parameter that would require either automatic or manual protective action. Since they occur frequently or regularly, Condition I occurrences affect the consequences of Conditions II, III, and IV events. Analysis of each fault condition is generally based on a conservative set of initial conditions corresponding to the most adverse set of conditions that can occur during a Condition I event.
The list of steady state and shutdown conditions, permissible deviations, and operational transients is given in Section 15.1. Implicit in the definition of normal operation is proper and timely action by the reactor operator. That is, the operator follows recommended operating procedures for maintaining appropriate power distributions and takes any necessary remedial actions when alerted to do so by plant instrumentation. Thus, as stated above, the worst or limiting power distribution that can occur during normal operation is considered as the starting point for analysis of Conditions II, III, and IV events.
Improper procedural actions or errors by the operator are assumed in the design as occurrences of moderate frequency (Condition II). The limiting power shapes that result from such Condition II events are, therefore, those power shapes, which deviate from the normal operating condition at the recommended AO band. Power shapes that fall in this category are used to determine reactor protection system setpoints in order to maintain margin to overpower or DNB limits.
Maintaining power distributions within the required hot channel factor limits is discussed in the Technical Specifications. A complete discussion of power distribution control in 4.3-10 Revision 24 September 2018
DCPP UNITS 1 & 2 FSAR UPDATE Westinghouse PWRs is included in References 2, 29, and 30. Detailed background information on the following design constraints on local power density in a Westinghouse PWR, on the defined operating procedures, and on the measures taken to preclude exceeding design limits is presented in References 23, 29, and 30.
The upper bound on peaking factors, F TQ and F NH , includes all of the nuclear effects that influence the radial and/or axial power distributions throughout core life for various modes of operation, including load follow, reduced power operation, and axial xenon transients.
Radial power distributions are calculated for full power, and fuel and moderator temperature feedback effects are included for the average enthalpy plane of the reactor.
Steady state nuclear design calculations are done for normal flow with the same mass flow in each channel and flow redistribution is calculated explicitly where it is important to the DNB analysis of accidents. The effect of xenon on radial power distribution is small (compare Figures 4.3-7 and 4.3-8), but is included as part of the normal design process.
The core average axial profile can experience significant changes that can occur rapidly as a result of rod motion and load changes, and more slowly due to xenon distribution.
To study points of closest approach to axial power distribution limits, several thousand cases are examined. Since the nuclear design properties dictate what axial shapes can occur, the limits of interest can be set in terms of parameters that are readily observed.
Specifically, the following nuclear design parameters are significant to the axial power distribution analysis:
(1) Core power level (2) Core height (3) Coolant temperature and flow (4) Coolant temperature program as a function of reactor power (5) Fuel cycle lifetimes (6) Rod bank worths (7) Rod bank overlaps Normal plant operation assumes compliance with the following conditions:
(1) Control rods in a single bank move together with no individual rod insertion differing by more than 12 steps (indicated) from the bank demand position 4.3-11 Revision 24 September 2018
DCPP UNITS 1 & 2 FSAR UPDATE (2) Control banks are sequenced with overlapping banks (3) Control bank insertion limits are not violated (4) Axial power distribution procedures, which are given in terms of flux difference control and control bank position, are observed.
The above axial power distribution procedures are part of the normal plant operating procedures. Briefly, they require control of the AO (refer to Equation 4.3-4) at all power levels, within a permissible operating band. This minimizes xenon transient effects on the axial power distribution, since the procedures essentially keep the xenon distribution in phase with the power distribution.
Calculations are performed for normal reactor operation at beginning, middle, and end of cycle conditions. Different operation histories are implicitly included in the methodology. These different histories cover both base loaded operation and extensive load following.
These cases represent many possible reactor states in the life of one fuel cycle. They are considered to be necessary and sufficient to generate a local power density limit which, when increased by 5 percent for conservatism, will not be exceeded with a 95 percent confidence level. Many of the points do not approach the limiting envelope.
However, they are generated as part of the process that leads to the shapes, which do define the envelope.
Thus, it is not possible to single out any transient or steady state condition that defines the most limiting case. It is not even possible to separate out a small number, which form an adequate analysis. The process of generating a myriad of shapes is essential to the philosophy that leads to the required level of confidence. A set of parameters that produces a limiting case for one reactor fuel cycle (defined as approaching the line of Figure 4.3-23) is not necessarily a limiting case for another reactor or fuel cycle with different control bank worths or insertion limits, enrichments, burnup, reactivity coefficient, etc. The shape of the axial power distribution calculated for a particular time depends on the operating history of the core up to that time, and on the manner in which the operator conditioned xenon in the days immediately before that time.
The calculated points are synthesized from axial calculations combined with the radial factors appropriate for rodded and unrodded planes. In these calculations, the effects on the radial peak of xenon redistribution that occur, following the withdrawal of a control bank (or banks) from a rodded region, are obtained from three-dimensional calculations. The factor to be applied to the radial peak is obtained from calculations in which the xenon distribution is preconditioned by the presence of control rods and then allowed to redistribute for several hours. A detailed discussion of this effect may be found in References 23 and 29. In addition to the 1.05 conservatism factor, the calculated values are increased by a factor of 1.03 for the engineering factor F QE .
4.3-12 Revision 24 September 2018
DCPP UNITS 1 & 2 FSAR UPDATE The envelope drawn over the calculated (max F TQ x Power) points, as shown in Figure 4.3-23 is an example of an upper bound envelope on local power density versus elevation in the core. Cycle-specific values are calculated each cycle.
Finally, this upper bound envelope is based on operation within an allowed range of axial flux difference limits. These limits are detailed in the Core Operating Limits Reports and rely only on excore surveillance supplemented by the required normal monthly power distribution measurement. If the axial flux difference exceeds the allowable range, an alarm is actuated.
Allowing for fuel densification, the average linear power is 5.445 kW/ft for both units at 3,411 MWt. The conservative upper bound value of normalized local power density, including uncertainty allowances, is 2.58, corresponding to a peak linear power of 14.3 kW/ft at 102 percent power.
To determine reactor protection system setpoints, with respect to power distributions, three categories of events are considered: rod control equipment malfunctions, operator errors of commission, and operator errors of omission. In evaluating these three categories, the core is assumed to be operating within the four constraints described above.
The first category is uncontrolled rod withdrawal (with rods moving in the normal bank sequence). Also included are motions of the banks below their insertion limits, which could be caused, for example, by uncontrolled dilution or primary coolant cooldown.
Power distributions were calculated, assuming short-term corrective action. That is, no transient xenon effects were considered to result from the malfunction. The event was assumed to occur from typical normal operating situations, which include normal xenon transients. It was also assumed that the total power level would be limited by the reactor trip to below 118 percent. Results are given in Figure 4.3-21 in units of kW/ft.
The peak power density, which can occur in such events, assuming reactor trip at or below 118 percent, is less than that required for fuel centerline melt, including uncertainties.
The second category, also appearing in Figure 4.3-21, assumes that the operator mispositions the rod bank in violation of insertion limits and creates short-term conditions not included in normal operating conditions.
The third category assumes that the operator fails to take action to correct a flux difference violation. The results shown in Figure 4.3-22 are F TQ multiplied by 102 percent power, including an allowance for calorimetric error. The peak linear power does not exceed 22.0 kW/ft, provided the operator's error does not continue for a period which is long compared to the xenon time constant. It should be noted that a reactor overpower accident is not assumed to occur coincident with an independent operator error. Additional detailed discussion of these analyses is presented in Reference 23.
4.3-13 Revision 24 September 2018
DCPP UNITS 1 & 2 FSAR UPDATE Analyses of possible operating power shapes for the DCPP reactor show that the appropriate hot channel factors F TQ and F NH for peak local power density, and for DNB analysis at full power, are the values given in Table 4.3-1 and addressed in the Technical Specifications.
The maximum allowable F TQ can increase with decreasing power, as shown in the Technical Specifications. Increasing F NH with decreasing power is permitted by the DNB protection setpoints and allows radial power shape changes with rod insertion to the insertion limits, as described in Section 4.4.3.13. The allowances for increased F NH permitted is:
F NH = 1.65 [1 + 0.3 (1-P)] for VANTAGE+ fuel (4.3-5)
This becomes a design basis criterion, which is used for establishing acceptable control rod patterns and control bank sequencing. Likewise, fuel loading patterns for each cycle are selected with consideration of this design criterion. The worst values of F NH for possible rod configurations occurring in normal operation are used in verifying that this criterion is met. Typical radial factors and radial power distributions are shown in Figures 4.3-7 through 4.3-12. The worst values generally occur when the rods are assumed to be at their insertion limits. As discussed in Reference 3, it has been determined that the Technical Specification limits are met, provided the above conditions (1) through (4) are observed. These limits are taken as input to the thermal-hydraulic design basis, as described in Section 4.4.3.13.1.
If the possibility exists during normal operation of local power densities exceeding those assumed as the precondition for a subsequent hypothetical accident, but which would not itself cause fuel failure, administrative controls and alarms are provided to return the core to a safe condition. These alarms are described in Chapter 7 and in the Technical Specifications.
4.3.3.2.7 Experimental Verification of Power Distribution Analysis This subject, which is discussed in depth in Reference 1, is summarized here.
To measure the peak local power density, F TQ , with the movable detector system described in Sections 4.4.7 and 7.7.2.9.2, the following uncertainties are considered:
(1) Reproducibility of the measured signal (2) Errors in the calculated relationship between detector current and local flux (3) Errors in the calculated relationship between detector flux and peak rod power some distance from the measurement thimble 4.3-14 Revision 24 September 2018
DCPP UNITS 1 & 2 FSAR UPDATE Allowance for (1) has been quantified by repetitive measurements made with several intercalibrated detectors using the common thimble features of the incore detector system. This system allows more than one detector to access any thimble. Item (2) above is quantified to the extent possible by using the fluxes measured at one thimble location to predict fluxes at another location, which is also measured. Local power distribution predictions are verified in critical experiments on arrays of rods with simulated guide thimbles, control rods, burnable poisons, etc.
Reference 1 concludes that the uncertainty associated with the peak nuclear heat flux factor, F TQ , is 4.58 percent at the 95 percent confidence level with only 5 percent of the measurements greater than the inferred value.
In comparing measured power distributions (or detector currents) against the calculations for the same situations, it is not possible to subtract out the detector reproducibility. Thus, a comparison between measured and predicted power distributions must consider measurement error. Such a comparison is illustrated in Figure 4.3-25 for one of the maps of Reference 1, which is similar to hundreds of maps taken since then on various reactors, confirming the adequacy of the 5 percent uncertainty allowance on F TQ .
A similar analysis for the uncertainty in F NH (rod integral power) measurements results in an allowance of 3.68 percent at the equivalent of a 2 confidence level. For historical reasons, an 8 percent uncertainty factor is allowed in the nuclear design basis; that is, the predicted rod integrals at full power must not exceed the design F NH less 8 percent.
This 8 percent may be reduced in final design to 4 percent to allow a wider range of acceptable axial power distributions in the DNB analysis and still meet the acceptance criteria of Section 4.3.2.2.
A measurement in the second cycle of a 121-assembly, 12-foot core, was compared with a simplified one-dimensional core average axial calculation in Figure 4.3-26. This calculation does not give explicit representation to the fuel grids.
The accumulated data on power distributions in actual operation is basically of three types:
(1) Much of the data is obtained in steady state operation at constant power in the normal operating configuration.
(2) Data with unusual values of AO are obtained as part of the excore detector calibration exercise which is performed monthly.
(3) Special tests have been performed in load follow and other transient xenon conditions which have yielded useful information on power distributions.
4.3-15 Revision 24 September 2018
DCPP UNITS 1 & 2 FSAR UPDATE These data are presented in detail in Reference 3. Figure 4.3-27 contains a summary of measured values of F TQ as a function of AO for five plants from that report.
4.3.3.2.8 Testing An extensive series of physics tests is performed on first cores. These tests and the criteria for satisfactory results are described in detail in Chapter 14. Since not all limiting situations can be created at BOL, the main purpose of the tests is to provide a check on the calculation methods used in the predictions for the conditions of the test.
Physics testing is also performed at the beginning of each reload cycle to ensure that the operating characteristics of the core are consistent with design predictions.
4.3.3.2.9 Monitoring Instrumentation The adequacy of instrument numbers, spatial deployment, and required correlations between readings and peaking factors, calibration, and errors is described in References 1, 2, and 3. The relevant conclusions are summarized in Sections 4.3.3.2.7 and 4.4.7.
References 32 and 33 describe the instrumentation requirements and calibration of the PDMS, and the uncertainties applied to the calculated peaking factors.
If the limitations given in Section 4.3.3.2.6 on rod insertion and flux difference are observed, the excore detector system provides adequate monitoring of power distributions.
Further details of specific limits on the observed rod positions and flux difference are given in the Core Operating Limits Reports, together with a discussion of their bases.
Limits for alarms, reactor trip, etc., are given in the Technical Specifications. System descriptions are provided in Section 7.7.
4.3.3.3 Reactivity Coefficients Reactor core kinetic characteristics determine the response of the core to changing plant conditions, or to operator adjustments made during normal operation, as well as the core response during abnormal or accidental transients. These kinetic characteristics are quantified in reactivity coefficients. The reactivity coefficients reflect changes in the neutron multiplication due to varying plant conditions such as power, moderator or fuel temperatures, or, less significantly, due to a change in pressure or void conditions. Since reactivity coefficients change during the life of the core, ranges of coefficients are employed in transient analysis to determine the response of the plant throughout life. The analytical methods and calculational models used in calculating the reactivity coefficients are given in Section 4.3.3.10.
4.3-16 Revision 24 September 2018
DCPP UNITS 1 & 2 FSAR UPDATE 4.3.3.3.1 Fuel Temperature (Doppler) Coefficient The fuel temperature (Doppler) coefficient is defined as the change in reactivity per degree change in effective fuel temperature and is primarily a measure of the Doppler broadening of U-238 and Pu-240 resonance absorption peaks. Doppler broadening of other isotopes such as U-236, Np-237, etc., are also considered, but their contributions to the Doppler effect is small. An increase in fuel temperature increases the effective resonance absorption cross-sections of the fuel and produces a corresponding reduction in reactivity.
The fuel temperature coefficient is calculated by two-group two or three-dimensional calculations. Moderator temperature is held constant and the power level is varied.
Spatial variation of fuel temperature is taken into account by calculating the effective fuel temperature as a function of power density, as discussed in Section 4.3.3.10.1.
A typical Doppler temperature coefficient is shown in Figure 4.3-28 as a function of the effective fuel temperature (at BOL and EOL conditions). The effective fuel temperature is lower than the volume averaged fuel temperature since the neutron flux distribution is nonuniform through the pellet and gives preferential weight to the surface temperature.
A typical Doppler-only contribution to the power coefficient (defined later) is shown in Figure 4.3-29 as a function of percent core power. The integral of the differential curve in Figure 4.3-29 is the Doppler contribution to the power defect and is shown in Figure 4.3-30 as a function of percent power. The Doppler coefficient changes as a function of core life, representing the combined effects of the fuel temperature reduction with burnup and the buildup of Pu-240 (refer to Section 4.3.3.10.1). The upper and lower limits of Doppler coefficient used in accident analyses are given in Chapter 15.
4.3.3.3.2 Moderator Coefficients The moderator coefficient is a measure of the change in reactivity due to a change in specific coolant parameters such as density, temperature, pressure, or void.
4.3.3.3.2.1 Moderator Density and Temperature Coefficients The MTC (density) is defined as the change in reactivity per degree change in the moderator temperature. Generally, the effect of the changes in moderator density, as well as the temperature, are considered together. A decrease in moderator density means less moderation which results in a negative MTC. An increase in coolant temperature, keeping the density constant, leads to a hardened neutron spectrum resulting in greater resonance absorption in U-238, Pu-240, and other isotopes. The hardened spectrum also causes a decrease in the fission to capture ratio in U-235 and Pu-239. Both of these effects make the MTC more negative. Since water density decreases as temperature increases, the MTC (density) becomes more negative with increasing temperature.
4.3-17 Revision 24 September 2018
DCPP UNITS 1 & 2 FSAR UPDATE The soluble boron also affects the MTC (density) since its density, like that of water, also decreases when the coolant temperature rises. Therefore, an increase in the soluble poison concentration will result in an increase to the moderator coefficient.
Indeed, if the concentration of soluble poison is large enough, the net value of the coefficient may be positive. With the burnable poison rods present, however, the initial hot boron concentration is sufficiently low, making the MTC negative at full power operating temperatures. The effect of control rods is to make the moderator coefficient more negative by reducing the required soluble boron concentration and by increasing "leakage" from the core.
With burnup, the MTC normally becomes more negative primarily as a result of boric acid dilution, but also, to a significant extent, from the effects of plutonium and fission products buildup.
The MTC is calculated for various plant conditions by performing two-group two or three dimensional calculations, varying the moderator temperature (and density) by about 5°F about each of the mean temperatures. The MTC is shown in Figures 4.3-31 through 4.3-33 as a function of core temperature and boron concentration for a typical reload unrodded and rodded core. The temperature range covered is from cold (68°F) to about 600°F. The contribution due to Doppler coefficient (because of change in moderator temperature) has been subtracted from these results. Figure 4.3-34 shows the hot, full power MTC as a function of cycle lifetime for the critical boron concentration condition based on the design boron letdown condition (refer to Figure 4.3-3) for a typical reload cycle.
4.3.3.3.2.2 Moderator Pressure Coefficient The moderator pressure coefficient relates the change in moderator density, resulting from a reactor coolant pressure change, to the corresponding effect on neutron production. This coefficient is of much less significance than the MTC. A change of 50 psi in pressure has approximately the same effect on reactivity as a half-degree change in moderator temperature. This coefficient can be determined from the MTC by relating change in pressure to the corresponding change in density. The moderator pressure coefficient is negative over a portion of the moderator temperature range at BOL (-0.004 pcm/psi, BOL) but is always positive at operating conditions and becomes more positive during life (+0.3 pcm/psi, EOL).
4.3.3.3.2.3 Moderator Void Coefficient The moderator void coefficient relates the change in neutron multiplication to the presence of voids in the moderator. In a PWR, this coefficient is not very significant because of the low void content in the coolant. The core void content is less than one-half of 1 percent and is due to local or statistical boiling. The void coefficient varies from 50 pcm/% void at BOL and low temperatures to -250 pcm/% void at EOL and at operating temperatures. The negative void coefficient at operating temperature becomes more negative with fuel burnup.
4.3-18 Revision 24 September 2018
DCPP UNITS 1 & 2 FSAR UPDATE 4.3.3.3.3 Power Coefficient The combined effect of moderator temperature and fuel temperature change as the core power level changes is called the total power coefficient, and is expressed in terms of reactivity change per percent power change. A typical power coefficient at BOL, MOL, and EOL conditions is given in Figure 4.3-35. It becomes more negative with burnup, reflecting the combined effect on moderator and fuel temperature coefficients of burnup.
A typical power defect (integral reactivity effect) at BOL, MOL, and EOL is given in Figure 4.3-36.
4.3.3.3.4 Comparison of Calculated and Experimental Reactivity Coefficients The accuracy of the current analytical model is discussed in Section 4.3.3.10.3.
Experimental verification of the calculated coefficients was performed during the physics startup tests described in Chapter 14.
4.3.3.3.5 Reactivity Coefficients Used in Transient Analysis Table 4.3-1 gives representative ranges for the reactivity coefficients used in the transient analysis. The exact values of the coefficient used in the analysis depend on whether the transient of interest is examined at the BOL or EOL, whether the most negative or the most positive (least negative) coefficients are appropriate, and whether spatial nonuniformity must be considered in the analysis. Conservative values of coefficients are always used in the transient analysis, as described in Chapter 15.
The values listed in Table 4.3-1, and illustrated in Figures 4.3-29 through 4.3-36, apply to the core shown in Figure 4.3-1. Appropriate coefficients for use in other cycles depend on the core's operating history, the number and enrichment of fresh fuel assemblies, the loading pattern of burned and fresh fuel, and the number and location of any burnable poison rods. The need for a reevaluation of any accident in a subsequent cycle is contingent on whether or not the coefficients for that cycle fall within the range used in the analysis presented in Chapter 15. For information only, control rod requirements are given in Tables 4.3-2 and 4.3-3 for a hypothetical equilibrium cycle.
4.3.3.4 Control Requirements To ensure SDM availability under cooldown to ambient temperature conditions, concentrated soluble boron is added to the coolant. Boron concentrations for several core conditions are listed in Table 4.3-1. They are all well below the solubility limit. The RCCAs are employed to bring the reactor to the hot shutdown condition. The minimum SDM required is given in the Core Operating Limits Reports.
4.3-19 Revision 24 September 2018
DCPP UNITS 1 & 2 FSAR UPDATE The ability to shut down from hot conditions is demonstrated in Tables 4.3-2 and 4.3-3 by comparing the difference between the reactivity available in the RCCA, allowing for the rod with the highest worth being stuck, with that required for control and protection.
The SDM allows 10 percent for analytic uncertainties (refer to Section 4.3.3.4.9). The largest reactivity control requirement appears at EOL when the MTC reaches its peak negative value as reflected in the larger power defect.
Control rods are required to provide sufficient reactivity to compensate for the power defect from full power to zero power and the required SDM. The reactivity addition resulting from power reduction consists of contributions from Doppler, variable average moderator temperature, flux redistribution, and reduction in void content.
4.3.3.4.1 Doppler Control requirements to compensate for the Doppler effect are listed in Tables 4.3-2 and 4.3-3, for DCPP Unit 1 and Unit 2, respectively.
4.3.3.4.2 Variable Average Moderator Temperature When the core is shut down to the hot zero power condition, the average moderator temperature changes from the equilibrium full load value, determined by the steam generator and turbine characteristics (such as steam pressure, heat transfer, and tube fouling), to the equilibrium no-load value, which is based on the steam generator shell side design pressure. The design change in temperature is conservatively increased by 4°F to account for control dead band measurement errors.
Since the moderator coefficient is negative, there is a reactivity addition with power reduction. The MTC becomes more negative as the fuel depletes because the boron concentration decreases. This effect is the major contribution to the increased requirement at EOL.
4.3.3.4.3 Redistribution During full power operation, the coolant density decreases with core height and this, together with partial insertion of control rods, results in less fuel depletion near the top of the core. Under steady state conditions, the relative power distribution will be slightly asymmetric towards the bottom of the core. On the other hand, at hot zero power conditions, the coolant density is uniform and there is no flattening due to Doppler. The result is a flux distribution that at zero power can be skewed toward the top of the core.
The reactivity insertion due to the skewed distribution is calculated with an allowance for the most adverse effects of xenon distribution.
4.3-20 Revision 24 September 2018
DCPP UNITS 1 & 2 FSAR UPDATE 4.3.3.4.4 Void Content A small void content in the core is due to nucleate boiling at full power. The void collapse that results from a power reduction makes a small reactivity contribution.
4.3.3.4.5 Rod Insertion Allowance At full power, the control bank is operated within a prescribed travel band to compensate for small periodic changes in boron concentration, temperature, and very small changes in the xenon concentration not compensated for by a change in boron concentration. When the control bank reaches either limit of this band, a change in boron concentration is required to compensate for additional reactivity changes. Since the insertion limit is set by a rod travel limit, a conservatively high calculation of the inserted worth is made which exceeds the normally inserted reactivity.
4.3.3.4.6 Burnup Excess reactivity of 10 percent to 25 percent (hot) is installed at the beginning of each cycle to provide sufficient reactivity to compensate for fuel depletion and fission products buildup throughout the cycle. This reactivity is controlled by the addition of soluble boron to the coolant and by burnable absorber. Representative soluble boron concentrations for several core configurations and the boron coefficient in the primary coolant are given in Table 4.3-1. Since the excess reactivity for burnup is controlled by soluble boron and/or burnable absorbers, it is not included in control rod requirements.
4.3.3.4.7 Xenon and Samarium Poisoning Changes in xenon and samarium concentrations in the core occur at a sufficiently slow rate, even following rapid power level changes, so that the resulting reactivity change is controlled by changing the soluble boron concentration.
4.3.3.4.8 pH Effects Changes in reactivity due to a change in coolant pH, if any, are sufficiently small in magnitude and occur slowly enough to be controlled by the boron system. Further details are available in Reference 4.
4.3.3.4.9 Experimental Confirmation The most appropriate assessment of the experimental confirmation of the analytical methods calls for comparison of key physics parameter predictions against directly measured plant data. A total of seven reactor cores covering a diversified range of advanced fuel product features, modern fuel management schemes, and different reactor loop types formulate the basis for the qualification assessment. Three-dimensional Advanced Nodal Code (ANC) models employing PHOENIX-P based cross-sections were developed to predict all relevant physics parameters. To encompass the 4.3-21 Revision 24 September 2018
DCPP UNITS 1 & 2 FSAR UPDATE entire range of available measured quantities, hot zero power physics predictions are compared against measured data in addition to full power core analyses. Refer to Reference 27 for additional details of the experimental confirmation.
These and other measurements demonstrate the ability of the methods described in Section 4.3.3.10 to accurately predict the total shutdown reactivity of the core.
4.3.3.5 Control Core reactivity is controlled by means of a chemical neutron absorber (chemical shim) dissolved in the coolant, RCCAs, and burnable poison rods as described below.
4.3.3.5.1 Chemical Shim Boron in solution as boric acid is used to control relatively slow reactivity changes associated with:
(1) The moderator temperature defect in going from cold shutdown at ambient temperature to the hot operating temperature at zero power (2) Transient xenon and samarium poisoning, such as that following power changes or changes in RCCA position (3) The excess reactivity required to compensate for the effects of fissile inventory depletion and buildup of long-life fission products (4) The burnable absorber depletion The boron concentrations for various core conditions are presented in Table 4.3-1.
4.3.3.5.2 Rod Cluster Control Assemblies As shown in Table 4.1-1, 53 RCCAs are used in these reactors. The RCCAs are used for shutdown and control purposes to offset fast reactivity changes associated with:
(1) The required SDM in the hot zero power, stuck rods condition (2) The increase in power above hot zero power (power defect including Doppler and moderator reactivity changes)
(3) Unprogrammed fluctuations in boron concentration, coolant temperature, or xenon concentration (with rods not exceeding the allowable rod insertion limits)
(4) Reactivity ramp rates resulting from load changes 4.3-22 Revision 24 September 2018
DCPP UNITS 1 & 2 FSAR UPDATE Control bank reactivity insertion at full power is limited to maintain shutdown capability.
As the power level is reduced, control rod reactivity requirements are reduced and more rod insertion is allowed. The control bank position is monitored and the operator is notified by an alarm if the limit is approached. The determination of the insertion limit uses conservative xenon distributions and axial power shapes. In addition, the RCCA withdrawal pattern obtained from these analyses is used in determining power distribution factors, and in determining the maximum reactivity worth during an ejection accident of an inserted RCCA. The Technical Specifications discuss rod insertion limits.
Power distribution, rod ejection, and rod misalignment analyses are based on the arrangement of the shutdown and control RCCA groups shown in Figures 4.3-37 and 4.3-38, for Unit 1 and Unit 2, respectively. All shutdown RCCAs are withdrawn before control banks withdrawal is initiated. In going from zero to 100 percent power, control banks A, B, C, and D are withdrawn sequentially. Rod position limits and the basis for rod insertion limits are provided in the Core Operating Limits Reports.
4.3.3.5.3 Burnable Absorber Rods Burnable absorber rods (either discrete or integral type) provide partial control of excess reactivity during the fuel cycle. These rods prevent the MTC from being positive at normal operating conditions. They perform this function by reducing the requirement for soluble boron in the moderator at the beginning of the fuel cycle, as described above.
The burnable absorber patterns, used together with a typical number of rods per assembly, are shown in Figure 4.3-6. The arrangements within an assembly for discrete and integral absorber types are displayed in Figures 4.3-4 and 4.3-5 respectively. The critical concentration of soluble boron resulting from the slow burnup of boron in the rods is such that the MTC remains negative at all times for full power operating conditions.
4.3.3.5.4 Peak Xenon Startup Peak xenon buildup is compensated by the boron control system. Startup from the peak xenon condition is accomplished with a combination of rod motion and boron dilution. Boron dilution may be made at any time, including the shutdown period, provided the SDM is maintained.
4.3.3.5.5 Load Follow Control and Xenon Control The DCPP units are usually base loaded; however, it is expected that during certain times of certain years some load following may be required.
Should load following become a desired mode of operation, then, during load follow maneuvers, power changes would be accomplished using control rod motion, dilution or boration by the boron systems as required, and reductions in coolant average temperature. Control rod motion limitations are discussed in Section 4.3.3.5.2 and the 4.3-23 Revision 24 September 2018
DCPP UNITS 1 & 2 FSAR UPDATE Technical Specifications. Reactivity changes due to the changing xenon concentration can be controlled by rod motion and/or soluble boron concentration changes.
4.3.3.5.6 Burnup The excess reactivity available for burnup is controlled with soluble boron and/or burnable absorbers. The boron concentration must be limited during operating conditions to ensure the MTC is negative at full power. Sufficient burnable absorbers are installed at the beginning of a cycle to give the desired cycle lifetime without exceeding the boron concentration limit. The practical minimum boron concentration is 10 ppm.
4.3.3.6 Control Rod Patterns and Reactivity Worths The RCCAs are designated by function as the control groups and the shutdown groups.
The terms "group" and "bank" are used synonymously throughout this chapter to describe a particular grouping of control assemblies. The RCCA patterns are displayed in Figures 4.3-37 and 4.3-38 for Unit 1 and Unit 2, respectively. These patterns are not expected to change during the life of the units. The control banks are labeled A, B, C, and D, and the shutdown banks are labeled SA, SB, SC and SD.
The two criteria used to select the control groups are: (a) the total reactivity worth must be adequate to meet the requirements specified in Tables 4.3-2 and 4.3-3, and (b) because these rods may be partially inserted at power operation, the total power peaking factor should be low enough to ensure that power capability requirements are met. Analyses indicate that the first requirement can be met by one or more banks whose total worth equals at least the required amount. Since the shape of the axial power distribution would be more peaked following movement of a single group of rods worth 3 to 4 percent , four banks, each worth approximately 1 percent , were selected.
The position of control banks for criticality under any reactor condition is determined by the boron concentration in the coolant. On an approach to criticality, boron is adjusted to ensure criticality will be achieved with control rods above the insertion limit set by shutdown and other considerations (refer to the Technical Specifications). Early in the cycle there may also be a withdrawal limit at low power to maintain an MTC more negative than the Technical Specification limit. Usual practice is to adjust boron to ensure that the rod position lies within the so-called maneuvering band so that an escalation from zero power to full power does not require further adjustment of boron concentration.
Ejected rod worths are given in Section 15.4.6 for several different conditions.
Experimental confirmation of ejected rod worths can be found by reference to startup test reports such as Reference 5.
4.3-24 Revision 24 September 2018
DCPP UNITS 1 & 2 FSAR UPDATE Allowable deviations due to misaligned control rods are discussed in the Technical Specifications.
A representative calculation for two banks of control rods withdrawn simultaneously (rod withdrawal accident) at EOL is shown in Figure 4.3-39. Calculation of control rod reactivity worth versus time following reactor trip involves both control rod velocity and differential reactivity worth. Rod position versus time of travel after rod release is shown in Figure 4.3-40. The reactivity worth versus rod position is calculated by a series of steady state calculations at various control rod positions assuming all rods out of the core as the initial position to minimize the initial reactivity insertion rate. To be conservative, the rod of highest worth is assumed stuck out of the core and the flux distribution (and thus reactivity importance) is assumed to be skewed to the bottom of the core. The result of these calculations is shown in Figure 4.3-41.
The shutdown groups provide additional negative reactivity to ensure an adequate SDM. SDM is defined as the instantaneous amount of reactivity by which the reactor is subcritical, or would be subcritical from its present condition, assuming:
(1) all RCCAs are fully inserted except for the single RCCA of highest reactivity worth, which is assumed to be fully withdrawn (with any RCCA not capable of being fully inserted, the reactivity worth of the RCCA must be accounted for in the determination of SDM) and (2) when in MODE 1 or 2, the fuel and moderator temperatures are changed to the hot zero power temperatures.
The loss of control rod worth due to material irradiation is negligible, since only bank D rods may be in the core under full power operating conditions.
Tables 4.3-2 and 4.3-3 show that the available reactivity in withdrawn RCCAs provides the design bases minimum SDM allowing for the highest worth cluster to be at its fully withdrawn position in DCPP Unit 1 and Unit 2, respectively. An allowance for uncertainty in the calculated worth of N-1 rods is made before determination of the SDM.
4.3.3.7 Criticality of Fuel Assemblies Criticality of fuel assemblies outside the reactor is precluded by adequate design of fuel transfer and fuel storage facilities, and by administrative control procedures (refer to Section 9.1.1).
Verification that appropriate shutdown criteria, including uncertainties, are met during refueling is achieved using standard Westinghouse reactor design methods. Core subcriticality during refueling is continuously monitored as described in the Technical Specifications.
4.3-25 Revision 24 September 2018
DCPP UNITS 1 & 2 FSAR UPDATE 4.3.3.8 Stability 4.3.3.8.1 Introduction The stability of PWR cores against xenon-induced spatial oscillations, and the control of such transients, is discussed extensively in References 2, 6, 7, and 8.
Due to the negative power coefficient of reactivity, PWR cores are inherently stable to oscillations in total power. In a large reactor core, however, xenon-induced oscillations can take place with no corresponding change in total core power. The oscillation may be caused by a power shift in the core that occurs rapidly in comparison with the xenon-iodine time constants. Such a power shift occurs in the axial direction when a plant load change is made by control rod motion, and results in a change in the moderator density and fuel temperature distributions. Such a power shift in the diametral plane of the core could result from abnormal control action.
4.3.3.8.2 Stability Index Power distributions, either in the axial direction or in the X-Y plane, can undergo oscillations due to perturbations introduced in the equilibrium distributions without changing total core power. The xenon-induced oscillations are essentially limited to the first flux overtones in the current PWRs, and the stability of the core against xenon-induced oscillations can be determined in terms of the eigenvalues of the first flux harmonics. Writing the eigenvalue of the first flux harmonic, either in the axial direction or in the X-Y plane, as:
b ic, i2 1 ; (4.3-6) where b is defined as the stability index and T = 2/c as the oscillation period of the first harmonic. The time-dependence of the first harmonic in the power distribution can now be represented as:
t A et ae bt cos ct (4.3-7) where A and a are constants. The stability index can also be obtained approximately by:
1 A n 1 b= ln (4.3-8)
T An where An, An+1 are the successive peak amplitudes of the oscillation, and T is the time period between the successive peaks.
4.3-26 Revision 24 September 2018
DCPP UNITS 1 & 2 FSAR UPDATE 4.3.3.8.3 Prediction of the Core Stability The stability of the DCPP cores in relation to xenon-induced spatial oscillations is expected to be equal to that of earlier designs because: (a) the overall core size is unchanged and spatial power distributions are similar, (b) the MTC is expected to be similar, and (c) the Doppler coefficient of reactivity is expected to be similar at full power.
HISTORICAL INFORMATION IN ITALICS BELOW NOT REQUIRED TO BE REVISED.
4.3.3.8.4 Stability Measurements (1) Axial Measurements Two axial xenon transient tests conducted in a PWR with a core height of 12 feet and 121 fuel assemblies, at approximately 10 and 50 percent of cycle life, are reported in Reference 9.
The AO of power was obtained as a function of time for both tests as shown in Figure 4.3-42. The total core power was maintained constant during these spatial xenon tests, and the stability index and the oscillation period were obtained from a least-square fit of the AO data to Equation 4.3-8. The conclusions of the tests are as follows:
(a) The core was stable against induced axial xenon transients both at the core average burnups of 1550 MWD/MTU and 7700 MWD/MTU.
(b) The reactor core becomes less stable as fuel burnup progresses, and the axial stability index was essentially zero at 12,000 MWD/MTU.
(2) Measurements in the X-Y Plane Two X-Y xenon oscillation tests were performed at a PWR plant with a core height of 12 feet and 157 fuel assemblies. This plant had the highest power output of any Westinghouse PWR operating in 1972. The first test was conducted at a core average burnup of 1540 MWD/MTU and the second at a core average burnup of 12900 MWD/MTU. Both of the X-Y xenon tests show that the core was stable in the X-Y plane at both burnups. The second test shows that the core became more stable as the fuel burnup increased and all Westinghouse PWRs with 121 and 157 assemblies are expected to be stable throughout their burnup cycles.
In each of the two X-Y tests, a perturbation was introduced to the equilibrium power distribution through an impulse motion of one RCCA located along the diagonal axis. Following the perturbation, the 4.3-27 Revision 24 September 2018
DCPP UNITS 1 & 2 FSAR UPDATE uncontrolled oscillation was monitored using the movable detector and thermocouple system and the excore power range detectors. The quadrant tilt difference (QTD) is the quantity that properly represents the diametral oscillation in the X-Y plane of the reactor core in that the difference of the quadrant average powers over two symmetrically opposite quadrants essentially eliminates the contribution to the oscillation from the azimuthal mode. The QTD data were least-square fitted to the form of Equation 4.3-8. A stability index of -0.076 hr-1 with a period of 29.6 hours6.944444e-5 days <br />0.00167 hours <br />9.920635e-6 weeks <br />2.283e-6 months <br /> was obtained from the thermocouple data shown in Figure 4.3-43.
In the second X-Y xenon test, the PWR core with 157 fuel assemblies became more stable due to increased fuel depletion.
4.3.3.8.5 Comparison of Calculations with Measurements Axial xenon transient tests were analyzed in an axial slab geometry using a flux synthesis technique. The PANDA code (Reference 11) was used for direct simulation of the AO data. X-Y xenon transient tests analyses were performed with the modified TURTLE code (Reference 12). Both the PANDA and TURTLE codes solve the two-group time-dependent neutron diffusion equation with time-dependent xenon and iodine concentrations. The fuel temperature and moderator density feedback is limited to a steady state model. All the X-Y calculations were performed in an average enthalpy plane.
The basic nuclear cross-sections used in this study were generated from a unit cell depletion program that evolved from the codes LEOPARD (Reference 13) and CINDER (Reference 14). The detailed experimental data during the tests, including the reactor power level, enthalpy rise, and the impulse motion of the control rod assembly, as well as the plant follow burnup data, were closely simulated in the study.
The results of the stability calculation for the axial tests are compared with the experimental data in Table 4.3-4. The calculations show conservative results for both of the axial tests with a margin of approximately 0.01 hr-1 in the stability index.
An analytical simulation of the first X-Y xenon oscillation test shows a calculated stability index of -0.081 hr-1, in good agreement with the measured value of -0.076 hr-1. As indicated earlier, the second X-Y xenon test showed that the core had become more stable compared to the first test. The increase in the core stability in the X-Y plane due to increased fuel burnup is due mainly to the increased magnitude of the negative MTC.
Previous studies of the physics of xenon oscillations, including three-dimensional analysis, are reported in References 6, 7, 8, 9, and Section 1 of Reference 10.
4.3-28 Revision 24 September 2018
DCPP UNITS 1 & 2 FSAR UPDATE 4.3.3.8.6 Stability Control and Protection The excore detector system provides indications of xenon-induced spatial oscillations.
The readings from the excore detectors are available to the operator and also form part of the protection system.
(1) Axial Power Distribution To maintain proper axial power distributions, the operator is instructed to maintain an AO within a prescribed operating band, based on the excore detector readings. Should the AO move far enough outside this band, the protection limit will be reached and the power will be automatically cut back.
(2) Radial Power Distribution The DCPP cores are calculated to be stable with respect to xenon-induced oscillations in the X-Y plane during the plant's lifetime.
The X-Y stability of large PWRs has been further verified as part of the startup physics test program at a PWR core with 193 fuel assemblies.
The measured X-Y stability of the PWR core with 157 assemblies, and the good agreement between the calculated and measured stability index for this core, as discussed in Sections 4.3.3.8.4 and 4.3.3.8.5, make it very unlikely that a sustained X-Y oscillation can occur in a core with 193 assemblies. In the unlikely event that X-Y oscillations occur, backup actions are possible and would be implemented, if necessary, to increase the natural stability of the core until tests demonstrate a suitable stability, by making the MTC more negative.
A more detailed discussion of the power distribution control in PWR cores is presented in Reference 2.
4.3.3.9 Vessel Irradiation Pressure vessel irradiation and the corresponding material surveillance program are discussed in Sections 5.2.2.4 and 5.4.1. A brief review of the methodology used to determine neutron and gamma flux attenuation between the core and pressure vessel follows.
The primary shielding material used to attenuate high energy neutron and gamma flux originating in the core consists primarily of the core baffle, core barrel, the thermal shield for Unit 1 and the neutron pads for Unit 2, and associated water annuli, all of which are within the region between the core and the pressure vessel.
4.3-29 Revision 24 September 2018
DCPP UNITS 1 & 2 FSAR UPDATE In general, few group neutron diffusion theory and nodal analysis codes are used to determine flux and fission power density distributions within the active core, and the accuracy of these analyses is verified by incore measurements on operating reactors.
Refer to Section 5.2.2.4 for methods used outside the active core.
The neutron flux distribution and spectrum in the various structural components varies significantly from the core to the pressure vessel. Representative values of the neutron flux distribution and spectrum are presented in Table 4.3-5. The values listed are based on equilibrium cycle reactor core parameters and power distributions and are thus suitable for long-term neutron fluence projections and for correlation with radiation damage estimates.
4.3.3.10 Analytical Methods Calculations required in nuclear design consist of the following three distinct types, which are performed in sequence:
(1) Determination of effective fuel temperatures (2) Generation of macroscopic few-group parameters (3) Space-dependent, few-group diffusion calculations 4.3.3.10.1 Fuel Temperature (Doppler) Calculations Temperatures vary radially within the fuel rod, depending on heat generation rate in the pellet, the conductivity of the materials in the pellet, gap and cladding, and coolant temperature.
The FIGHTH code (Reference 34) performs a simplified calculation of effective temperatures in low enrichment, sintered UO2 fuel rods for use in nuclear design. The model includes radial variations of heat generation rate, thermal conductivity, and thermal expansion in the fuel pellet, elastic deflection of the cladding, and a pellet-clad gap conductance which depends on the kind of initial fill gas, the hot open gap dimension, and the fraction of the pellet circumference over which the gap is effectively closed due to pellet cracking. The steady-state radial temperature distribution in the fuel rod is calculated at a specified burnup, given the local value of the linear heat generation rate in the pellet and the moderator temperature and flow rate. The effective resonance temperatures of U-238 and Pu-240 are obtained by appropriate radial weighting of the temperature distribution. An effective flat pellet temperature for expansion which reproduces the hot pellet outer radius is also determined.
The observed burnup dependence of the Doppler defects and Doppler coefficients in operating plants was used to construct an empirical model of progressive pellet cracking which, primarily through the increased gap conductance due to gap closure, produces effective temperatures which lead to adequate predictions of the measurements. The 4.3-30 Revision 24 September 2018
DCPP UNITS 1 & 2 FSAR UPDATE effects of fission gas release, clad creep, and pellet swelling are not treated explicitly in this simplified model, described in Section 4.2.1.2.2.5.
Fuel temperatures for use in some past nuclear design Doppler calculations were obtained from a simplified version of the Westinghouse fuel rod design model described in Section 4.2.1.2.2, which considers the effect of radial variation of pellet conductivity, expansion-coefficient and heat generation rate, elastic deflection of the cladding, and a gap conductance which depends on the initial fill gas, the hot open gap dimension, and the fraction of the pellet over which the gap is closed. The fraction of the gap assumed closed represents an empirical adjustment to produce good agreement with observed reactivity data at BOL. Further gap closure occurs with burnup and accounts for the decrease in Doppler defect with burnup which has been observed in operating plants.
For detailed calculations of the Doppler coefficient, such as for use in xenon stability calculations, a more sophisticated temperature model is used which accounts for the effects of fuel swelling, fission gas release, and plastic cladding deformation.
Radial power distributions in the pellet as a function of burnup were obtained from LASER (Reference 15) calculations.
The effective U-238 temperature for resonance absorption was obtained from the radial temperature distribution by applying a radially dependent weighting function. The weighting function was determined from REPAD (Reference 16) Monte Carlo calculations of resonance escape probabilities in several steady state and transient temperature distributions. In each case, a flat pellet temperature was determined which produced the same resonance escape probability as the actual distribution. The weighting function was empirically determined from these results.
The effective Pu-240 temperature for resonance absorption was determined by a convolution of the radial distribution of Pu-240 number densities from LASER burnup calculations and the radial weighting function. The resulting temperature is burnup dependent, but the difference between U-238 and Pu-240 temperatures, in terms of reactivity effects, is small.
The effective pellet temperature for pellet dimensional change is that value which produces the same outer pellet radius in a virgin pellet as that obtained from the temperature model. The effective cladding temperature for dimensional change is its average value.
The temperature calculational model has been validated by plant Doppler defect data as shown in Table 4.3-6 and Doppler coefficient data as shown in Figure 4.3-44. Stability index measurements also provide a sensitive measure of the Doppler coefficient near full power (refer to Section 4.3.3.8). It can be seen that Doppler defect data are typically within 0.2 percent of prediction.
4.3-31 Revision 24 September 2018
DCPP UNITS 1 & 2 FSAR UPDATE 4.3.3.10.2 Macroscopic Group Constants There are two lattice codes used for the generation of macroscopic group constants for use in the spatial few group diffusion codes. They are a version of the LEOPARD and CINDER codes and PHOENIX-P. A detailed description of each follows. Macroscopic few-group constants and analogous microscopic cross-sections (needed for feedback and microscopic depletion calculations) can be generated for fuel cells by a Westinghouse version of the LEOPARD and CINDER codes, which are linked internally and provide burnup-dependent cross-sections. Normally, a simplified approximation of the main fuel chains is used; however, where needed, a complete solution for all the significant isotopes in the fuel chains from Th-232 to Cm-244 is available (Reference 17). Cross-section library tapes contain microscopic cross-sections from the ENDF/B (Reference 18) library, with a few exceptions, where other data provide better agreement with critical experiments, isotopic measurements, and plant critical boron values.
The effect on the unit fuel cell of nonlattice components in the fuel assembly is obtained by supplying an appropriate volume fraction of these materials in an extra region which is homogenized with the unit cell in the fast (MUFT) and thermal (SOFOCATE) flux calculations. In the thermal calculation, the fuel rod, cladding, and moderator are homogenized by energy-dependent disadvantage factors derived from an analytical fit to integral transport theory results.
Group constants for burnable absorber cells, guide thimbles, instrument thimbles, and interassembly gaps are generated in a manner analogous to the fuel cell calculation.
Reflector group constants are taken from infinite medium LEOPARD calculations.
Baffle group constants are calculated from an average of core and radial reflector microscopic group constants for stainless steel.
Group constants for control rods are calculated in a linked version of the HAMMER (Reference 19) and AIM (Reference 20) codes to provide an improved treatment of self-shielding in the broad resonance structure of the appropriate isotopes at epithermal energies than is available using LEOPARD. The Doppler broadened cross-sections of the control rod material are represented as smooth cross-sections in the 54-group LEOPARD fast group structure and in 30 thermal groups. The four-group constants in the rod cell and appropriate extra region are generated in the coupled space-energy transport HAMMER calculation. A corresponding AIM calculation of the homogenized rod cell with extra region is used to adjust the absorption cross-sections of the rod cell to match the reaction rates in HAMMER. These transport-equivalent group constants are reduced to two-group constants for use in space-dependent diffusion calculations.
In discrete X-Y calculations only one mesh interval per cell is used, and the rod group constants are further adjusted for use in this standard mesh by reaction rate matching the standard mesh unit assembly to a fine-mesh unit assembly calculation.
Validation of the cross-section method is based on analysis of critical experiments (refer to Table 4.3-7), isotopic data (refer to Table 4.3-8), plant critical boron (CB) values at hot 4.3-32 Revision 24 September 2018
DCPP UNITS 1 & 2 FSAR UPDATE zero power (HZP), BOL (refer to Table 4.3-9), and at HFP as a function of burnup (refer to Figures 4.3-45 through 4.3-47). Control rod worth measurements are shown in Table 4.3-10. Confirmatory critical experiments on burnable absorbers are described in Reference 21.
The PHOENIX-P computer code is a two-dimensional, multi-group, transport based lattice code capable of providing all necessary data for PWR analysis. Being a dimensional lattice code, PHOENIX-P does not rely on pre-determined spatial/spectral interaction assumptions for a heterogeneous fuel lattice, hence, will provide a more accurate multi-group flux solution than versions of LEOPARD/CINDER. The PHOENIX-P computer code is approved by the USNRC as the lattice code for generating macroscopic and microscopic few group cross-sections for PWR analysis (Reference 27).
The solution for the detailed spatial flux and energy distribution is divided into two major steps in PHOENIX-P (Reference 27). In the first step, a two-dimensional fine energy group nodal solution is obtained which couples individual subcell regions (pellet, cladding and moderator) as well as surrounding pins. PHOENIX-P uses a method based on the Carlvik's collision probability approach and heterogeneous response fluxes which preserves the heterogeneity of the pin cells and their surroundings. The nodal solution provides accurate and detailed local flux distribution, which is then used to spatially homogenize the pin cells to fewer groups.
The second step in the solution process solves for the angular flux distribution using a standard S4 discrete ordinates calculation. This step is based on the group-collapsed and homogenized cross-sections obtained from the first step of the solution. The S4 fluxes are then used to normalize the detailed spatial and energy nodal fluxes. The normalized nodal fluxes are used to compute reaction rates, power distribution and to deplete the fuel and burnable absorbers. A standard B1 calculation is employed to evaluate the fundamental mode critical spectrum and to provide an improved fast diffusion coefficient for the core spatial codes.
The PHOENIX-P code employs an energy group library, which has been derived mainly from ENDF/B (Reference 18) files. The PHOENIX-P cross-sections library was designed to properly capture integral properties of the multi-group data during group collapse, and enabling proper modeling of important resonance parameters. The library contains all neutronic data necessary for modeling fuel, fission products, cladding and structural data, coolant, and control/burnable absorber materials present in Light Water Reactor cores.
Group constants for burnable absorber cells, guide thimbles, instrument thimbles, control rod cells and other non-fuel cells can be obtained directly from PHOENIX-P without any adjustments such as those required in the cell or 1D lattice codes.
4.3-33 Revision 24 September 2018
DCPP UNITS 1 & 2 FSAR UPDATE 4.3.3.10.3 Spatial Few-Group Diffusion Calculations Spatial few-group diffusion calculations have primarily consisted of two group X-Y calculations using an updated version of the TURTLE code, and two-group axial calculations using an updated version of the PANDA code. However, with the advent of VANTAGE+ and hence axial features such as axial blankets and part length burnable absorbers, there is a greater reliance on three-dimensional nodal codes such as 3D PALADON (Reference 25) and APOLLO code (Reference 35) which performs one-dimensional, two group calculations using typical finite difference techniques, and 3D ANC (References 26 and 27) which provides both the radial and axial power distributions.
Nodal three-dimensional calculations are carried out to determine the critical boron concentrations and power distributions. The moderator coefficient is evaluated by varying the inlet temperature in the same calculations used for power distribution and reactivity predictions.
Axial calculations are used to determine differential control rod worth curves (reactivity versus rod insertion) and axial power shapes during steady state and transient xenon conditions. Group constants are obtained from three-dimensional nodal calculations homogenized by flux volume weighting.
Validation of the spatial codes for calculating power distributions involves the use of incore and excore detectors, and is discussed in Section 4.3.3.2.7.
The agreement in the PHOENIX-P methodology and PHOENIX-P/ANC core model qualification demonstrates the high accuracy of the PHOENIX-P/ANC advanced design system for multidimensional nuclear analysis of PWR cores. This qualification data base demonstrates the performance of this system for a wide range of applications performed in the design, safety, licensing and operational follow of PWR cores. Refer to Reference 27 for additional details on experimental verification.
The accuracy of APOLLO predictions for reactivity and axial power shapes is comparable to that for 3D ANC-based calculations (References 26 and 27).
4.3.4 SAFETY EVALUATION 4.3.4.1 General Design Criterion 10, 1971 - Reactor Design The reactor core and associated coolant, control, and protection systems are designed with appropriate margin to assure that SAFDLs are not exceeded during any condition of normal operation, including the effects of anticipated operational occurrences.
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DCPP UNITS 1 & 2 FSAR UPDATE 4.3.4.1.1 Fuel Burnup Fuel burnup is a measure of fuel depletion that represents the integrated energy output of the fuel (MWD/MTU) and is a convenient means for quantifying fuel exposure.
The core design lifetime or design discharge burnup is achieved by installing sufficient initial excess reactivity in each fuel region, and by following a fuel replacement program (such as that described in Section 4.3.3) that meets all safety-related criteria in each cycle of operation.
Initial excess reactivity in the fuel, although not a design basis, must be sufficient to maintain core criticality at full power operating conditions throughout cycle life with equilibrium xenon, samarium, and other fission products present. The end of design cycle life is defined to occur when the chemical shim concentration is essentially zero, with control rods present to the degree necessary for operational requirements (e.g., the controlling bank at the "bite" position). In terms of chemical shim boron concentration, this represents approximately 10 ppm with no control rod insertion.
4.3.4.1.2 Control of Power Distribution Calculation of the extreme power shapes that affect fuel design limits is performed with proven methods as described in Section 4.3.3.10 and verified frequently with results from measurements in operating reactors. The conditions under which limiting power shapes are assumed to occur are chosen conservatively with regard to any permissible operating state.
Even though there is good agreement between peak power calculations and measurements, a nuclear uncertainty margin is applied to calculated peak local power.
Such a margin is provided both for the analysis of normal operating states and for anticipated transients.
4.3.4.2 General Design Criterion 11, 1971 - Reactor Inherent Protection The reactor core and associated coolant systems is designed so that in the power operating range the net effect of the prompt inherent nuclear feedback characteristics tend to compensate for a rapid increase in reactivity.
When compensation for a rapid increase in reactivity is considered, there are two major effects. These are the resonance absorption effects (Doppler) associated with changing fuel temperature, and the spectrum effect resulting from changing moderator density.
These basic physics characteristics are often identified by reactivity coefficients. The use of slightly enriched uranium ensures that the Doppler coefficient of reactivity, which provides the most rapid reactivity compensation, is negative. The core is also designed to have an overall negative MTC of reactivity at full power so that average coolant temperature or void content provides another, slower, compensatory effect. A small positive MTC is allowed at low power. The negative MTC at full power can be achieved 4.3-35 Revision 24 September 2018
DCPP UNITS 1 & 2 FSAR UPDATE through use of fixed burnable absorbers and/or boron coated fuel pellets and/or control rods by limiting the reactivity held down by soluble boron.
Burnable absorber content (quantity and distribution) is not stated as a design basis other than as it relates to achieving a nonpositive MTC at power operating conditions, as discussed above.
4.3.4.3 General Design Criterion 12, 1971 - Suppression of Reactor Power Oscillations The reactor core and associated coolant, control, and protection systems are designed to assure that power oscillations which can result in conditions exceeding SAFDLs are not possible or can be reliably and readily detected and suppressed.
Oscillations in total core power output, from whatever cause, are readily detected by loop temperature sensors and by nuclear instrumentation. If power increased unacceptably, a reactor trip would occur, thus preserving margins to fuel design limits.
The stability of the turbine/steam generator/core systems and the reactor control system ensure that core power oscillations do not normally occur. Protection circuits' redundancy ensures an extremely low probability of exceeding design power levels.
The core is designed so that diametral and azimuthal oscillations due to spatial xenon effects are self-damping, and no operator action or control action is required to suppress them. Stability against diametral oscillations is so great that this excitation is highly improbable. Convergent azimuthal oscillations can be excited by prohibited motion of individual RCCAs. Such oscillations are readily observable and alarmed, using the excore long ion chambers. Indications are also continuously available from incore thermocouples and loop temperature measurements. The MIDS can be activated to provide more detailed information. In all presently proposed cores, these horizontal plane oscillations are self-damping by virtue of reactivity feedback effects designed into the core.
Axial xenon spatial power oscillations can be excited by power level changes or by control rod motion/misalignments. The oscillations are inherently convergent at the beginning of core life, but become divergent as the core ages. The time in core life when oscillations may become divergent depends on core characteristics. Xenon oscillations studies performed for plants similar to DCPP concluded that oscillations can diverge as early as 50 EFPD. The magnitude of oscillations increases with increasing core burnup, although the period is unaffected. The type of oscillation (convergence or divergence) does not depend on the amplitude of the initial oscillation but is a function of initial conditions at the start of the transient.
The excore detectors provide monitoring of axial power distribution. The operator actions (control rod movement or power level changes) are expected to suppress and control axial xenon transients.
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DCPP UNITS 1 & 2 FSAR UPDATE The limits on measured axial flux difference assure that the fuel design limits (Fq) are not exceeded during either normal operation or a xenon transient. The measured axial flux difference is also used as an input to the OTT trip function so that the DNB design bases are not exceeded.
4.3.4.4 General Design Criterion 25, 1971 - Protection System Requirements for Reactivity Control Malfunctions The protection system is designed to assure that SAFDLs are not exceeded for any single malfunction of the reactivity control systems, such as accidental withdrawal (not ejection or dropout) of control rods.
Reactivity addition associated with an accidental withdrawal of a control bank (or banks) is limited by the maximum rod speed (or travel rate) and by the worth of the bank(s).
For this reactor the maximum control rod speed is 45 inches per minute and the maximum rate of reactivity change considering two control banks moving is less than 75 pcm/sec. The reactivity rate used in the boron dilution analysis at power is discussed in Section 15.2.4.3.4.
4.3.4.5 General Design Criterion 26, 1971 - Reactivity Control System Redundancy and Capability Two independent reactivity control systems are provided: control rods and soluble boron in the coolant.
The control rod system can compensate for the reactivity effects of the fuel and water temperature changes accompanying power level changes over the range from full load to no load. In addition, the control rod system provides the minimum SDM under Condition I events and is capable of making the core subcritical rapidly enough to prevent exceeding acceptable fuel damage limits, assuming that the highest worth control rod is stuck out upon trip.
The boron system can compensate for all xenon burnout reactivity changes and will maintain the reactor in cold shutdown. Thus, backup and emergency shutdown provisions are provided by a mechanical and a chemical shim control system.
When fuel assemblies are in the pressure vessel and the vessel head is not in place, keff will be maintained at or below 0.95 with control rods and soluble boron. Further, the fuel will be maintained sufficiently subcritical that removal of all RCCAs will not result in criticality.
4.3.4.6 General Design Criterion 28, 1971 - Reactivity Limits The reactivity control system is designed with appropriate limits on the potential amount and rate of reactivity increase to assure that the effects of postulated reactivity accidents can neither (1) result in damage to the reactor coolant pressure boundary 4.3-37 Revision 24 September 2018
DCPP UNITS 1 & 2 FSAR UPDATE greater than limited local yielding nor (2) sufficiently disturb the core, its support structures or other reactor pressure vessel internals to impair significantly the capability to cool the core. These postulated reactivity accidents shall include consideration of rod ejection (unless prevented by positive means), rod dropout, steam line rupture, changes in reactor coolant temperature and pressure, and cold water addition.
The maximum control rod reactivity worth and the maximum rates of reactivity insertion using control rods are limited to preclude either rupture of the coolant pressure boundary or disruption of the core internals to a degree that would impair core cooling capacity in the event of a rod withdrawal or ejection accident (refer to Chapter 15).
Following any Condition IV event (such as rod ejection and steam line break), the reactor can be brought to the shutdown condition and the core will maintain acceptable heat transfer geometry.
4.
3.5 REFERENCES
- 1. F. L. Langford and R. J. Nath, Jr., Evaluation of Nuclear Hot Channel Factor Uncertainties, WCAP-7308-L, April 1969 (Westinghouse Proprietary) and WCAP-7810, December 1971.
- 2. J. S. Moore, Power Distribution Control of Westinghouse Pressurized Water Reactors, WCAP-7208, September 1968 (Westinghouse Proprietary) and WCAP-7811, December 1971.
- 3. A. F. McFarlane, Power Peaking Factors, WCAP-7912-P-A, January 1975 (Westinghouse Proprietary) and WCAP-7912-A, January 1975.
- 4. J. O. Cermak et al, Pressurized Water Reactor pH - Reactivity Effect, Final Report, WCAP-3696-8 (EURAEC-2074), October 1968.
- 5. J. E. Outzs, Plant Startup Test Report, H. B. Robinson Unit No. 2, WCAP-7844, January 1972.
- 6. C. G. Poncelet and A. M. Christie, Xenon-Induced Spatial Instabilities in Large PWRs, WCAP-3680-20, (EURAEC-1974), March 1968.
- 7. F. B. Skogen and A. F. McFarlane, Control Procedures for Xenon-Induced X-Y Instabilities in Large PWRs, WCAP-3680-21, (EURAEC-2111), February 1969.
- 8. F. B. Skogen and A. F. McFarlane, Xenon-Induced Spatial Instabilities in Three-Dimensions, WCAP-3680-22 (EURAEC-2116), September 1969.
- 9. J. C. Lee, et al, Axial Xenon Transient Tests at the Rochester Gas and Electric Reactor, WCAP-7964, June 1971.
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- 10. C. J. Kubit, Safety Related Research and Development for Westinghouse Pressurized Water Reactors, Program Summaries, Fall 1972, WCAP-8004, January 1973..
- 11. S. Altomare, et al., The PANDA Code, WCAP-7048-P-A, February 1975 (Westinghouse Proprietary) and WCAP-7757-A, February 1975.
- 12. S. Altomare and R. F. Barry, The TURTLE 24.0 Diffusion Depletion Code, WCAP-7213-P-A, February 1975 (Westinghouse Proprietary) and WCAP-7758-A, February 1975.
- 13. R. F. Barry, LEOPARD - A Spectrum Dependent Non-Spatial Depletion Code for the IBM-7094, WCAP-3269-26, September 1963.
- 14. T. R. England, CINDER - A One-Point Depletion and Fission Product Program.
WAPD-TM-334, August 1962.
- 15. C. G. Poncelet, LASER - A Depletion Program for Lattice Calculations Based on MUFT and THERMOS, WCAP-6073, April 1966.
- 16. J. E. Olhoeft, The Doppler Effect for a Non-Uniform Temperature Distribution in Reactor Fuel Elements, WCAP-2048, July 1962.
- 17. R. J. Nodvik, et al, Supplementary Report on Evaluation of Mass Spectrometric and Radiochemical Analyses of Yankee Core I Spent Fuel, Including Isotopes of Elements Thorium Through Curium, WCAP-6086, August 1969.
- 18. M. K. Drake, (Ed), Data Formats and Procedure for the ENDF Neutron Cross Section Library, BNL-50274, ENDF-102, Vol. I, 1970.
- 19. J. E. Suich and H. C. Honeck, The HAMMER System, Heterogeneous Analysis by Multigroup Methods of Exponentials and Reactors, DP-1064, January 1967.
- 20. H. P. Flatt and D. C. Baller, AIM-5, A Multigroup, One Dimensional Diffusion Equation Code, NAA-SR-4694, March 1960.
- 21. J. S. Moore, Nuclear Design of Westinghouse Pressurized Water Reactors with Burnable Poison Rods, WCAP-7806, December 1971.
- 22. J. M. Hellman, (Ed), Fuel Densification Experimental Results and Model for Reactor Application, WCAP-8218-P-A, March 1975 (Westinghouse Proprietary) and WCAP-8219-A, March 1975.
- 23. T. Morita, et al., Power Distribution Control and Load Following Procedures, WCAP-8385, September 1974 (Westinghouse Proprietary) and WCAP-8403, September 1974.
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- 24. Deleted.
- 25. T. M. Camden. et al., PALADON - Westinghouse Nodal Computer Code, WCAP-9485-P-A. December 1979 and Supplement 1, September 1981.
- 26. S. L. Davidson, (Ed), et al., ANC: A Westinghouse Advanced Nodal Computer Code, WCAP-10965-P-A, September 1986 (Westinghouse Proprietary) and WCAP-10966-A, September 1986.
- 27. T. Q. Nguyen, et al, Qualification of the PHOENIX-P/ANC Nuclear Design System for Pressurized Water Reactor Cores, WCAP-11596-P-A, June 1988.
- 28. Deleted.
- 29. S. L. Davidson, et al., Relaxation of Constant Axial Offset Control FQ Surveillance Technical Specification, WCAP-10216-P-A, Revision 1A, February 1994.
- 30. S. L. Davidson, et al., Westinghouse Reload Safety Evaluation Methodology, WCAP-9272-P-A, July 1985.
- 31. P.J. Kersting, et al., Assessment of Clad Flattening and Densification Power Spike Factor Elimination in Westinghouse Nuclear Fuel, WCAP-13589-A, March 1995 (Westinghouse Proprietary) and WCAP-14297-A, March 1995.
- 32. C. L. Beard, et al., BEACON Core Monitoring and Operations Support System, WCAP-12472-P-A, August 1994.
- 33. W.A. Boyd, et al., BEACON Core Monitoring and Operation Support System, WCAP-12472-P-A, Addendum 4, Revision 0, September 2012.
- 34. W.B. Henderson, et al., FIGHTH - A Simplified Calculation of Effective Temperatures in PWR Fuel Rods for Use in Nuclear Design, WCAP-9522, Revision 1, October 1985.
- 35. M. B. Yarbrough, et al., APOLLO - A One Dimensional Neutron Diffusion Theory Program, WCAP-13524-P-A, Revision 1-A, September 1997.
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DCPP UNITS 1 & 2 FSAR UPDATE 4.4 THERMAL AND HYDRAULIC DESIGN This section discusses the thermal and hydraulic design of the DCPP reactors.
The objective of the thermal and hydraulic design of the reactor core is to provide adequate heat transfer that is compatible with the heat generation distribution in the core, so that heat removal by the RCS or the ECCS (when applicable) meets the following performance and safety criteria:
(1) Fuel damage is not expected during normal operation and operational transients (Condition I) or any transient conditions arising from faults of moderate frequency (Condition II). It is not possible, however, to preclude a very small number of rod failures. These will be within the capability of the plant cleanup system and are consistent with the plant design bases.
(2) The reactor can be brought to a safe state following a Condition III event with only a small fraction of fuel rods damaged although sufficient fuel damage might occur to preclude resumption of operation without considerable outage time.
(3) The reactor can be brought to a safe state and the core can be kept subcritical with acceptable heat transfer geometry following transients arising from Condition IV events.
Note that fuel damage as used here is defined as penetration of the fission product barrier (i.e., the fuel rod cladding).
4.4.1 DESIGN BASES 4.4.1.1 General Design Criterion 10, 1971 - Reactor Design The reactor core is designed with appropriate margin to assure that SAFDLs are not exceeded during any condition of normal operation or anticipated operational occurrences.
4.4.1.2 General Design Criterion 12, 1971 - Suppression of Reactor Power Oscillations The reactor core is designed to assure that power oscillations that could result in conditions exceeding SAFDLs are not possible or can be reliably and readily detected and suppressed.
4.4-1 Revision 24 September 2018
DCPP UNITS 1 & 2 FSAR UPDATE 4.4.2 THERMAL AND HYDRAULIC DESIGN ACCEPTANCE CRITERIA 4.4.2.1 Departure from Nucleate Boiling Acceptance Criteria There will be at least a 95 percent probability that DNB will not occur on the limiting fuel rods during normal operation and operational transients and any transient conditions arising from faults of moderate frequency (Conditions I and II events) at a 95 percent confidence level.
4.4.2.2 Fuel Temperature Acceptance Criteria During Condition I and Condition II events, the maximum fuel temperature shall be less than the melting temperature of UO2. The UO2 melting temperature for at least 95 percent of the peak kW/ft fuel rods will not be exceeded at the 95 percent confidence level.
4.4.2.3 Core Flow Acceptance Criteria A minimum of 92.5 percent (Unit 1) and 91 percent (Unit 2) of the thermal flowrate (refer to Section 5.1) will pass through the fuel rod region of the core and be effective for fuel rod cooling. Coolant flow through the thimble tubes, as well as leakage from the core barrel-baffle region into the core, is not effective for heat removal.
4.4.2.4 Hydrodynamic Stability Acceptance Criteria Modes of operation associated with Condition I and Condition II events shall not lead to hydrodynamic instability.
4.4.3 SYSTEM DESCRIPTION 4.4.3.1 Summary Comparison The core design parameters of the DCPP Unit 1 and Unit 2 reactors are presented in Table 4.1-1.
The reactor core is designed to a minimum DNBR greater than or equal to the design limit DNBR as well as no fuel centerline melting during normal operation, operational transients, and faults of moderate frequency.
4.4.3.2 Fuel Cladding Temperatures A discussion of fuel cladding integrity is presented in Section 4.2.1.2.2.
The thermal-hydraulic design ensures that the maximum fuel pellet temperature is below the melting point of UO2 (refer to Section 4.4.2.2). To preclude center melting and establish overpower protection system setpoints, a calculated centerline fuel 4.4-2 Revision 24 September 2018
DCPP UNITS 1 & 2 FSAR UPDATE temperature of 4700°F has been selected as the overpower limit. The temperature distribution within the fuel pellet is predominantly a function of the local power density and UO2 thermal conductivity. However, the computation of radial fuel temperature distributions combines crud, oxide, cladding, gap, and pellet conductances. The factors that influence these conductances, such as gap size (or contact pressure), internal gas pressure, gas composition, pellet density, and radial power distribution within the pellet, etc., have been combined into a semiempirical thermal model (refer to Section 4.4.3.2.4) with modifications for time-dependent fuel densification (Reference 68). The temperature predictions have been compared to incore fuel temperature measurements (References 3 through 9) and melt radius data (References 10 and 11) with good results.
4.4.3.2.1 Effect of Fuel Densification on Fuel Rod Temperatures Fuel densification results in fuel pellet shrinkage. This affects the fuel temperatures in the following ways:
(1) Pellet radial shrinkage increases the pellet diametral gap that results in increased thermal resistance of the gap and thus higher fuel temperatures (refer to Section 4.2.1.2.2).
(2) Pellet axial shrinkage may produce pellet-to-pellet gaps that result in local power spikes, described in Section 4.3.3.2.1, and thus higher total heat flux hot channel factor, F QT and local fuel temperatures. However, studies have shown that this penalty can be eliminated for the fuel type present in the DCPP Unit 1 and Unit 2 cores (refer to Section 4.3.3.2.5).
(3) Pellet axial shrinkage results in a fuel stack height reduction and an increase in the linear power generation rate (kW/ft) for a constant core power level. Using the methods of Reference 68, the increase in linear power for the fuel rod specifications listed in Table 4.1-1 is 0.2 percent.
Fuel rod thermal parameters (fuel centerline, average, and surface temperatures) are determined throughout its lifetime considering time-dependent densification. Maximum fuel average and surface temperatures, shown in Figure 4.4-1 as a function of linear power density (kW/ft), are peak values attained during the fuel lifetime. Similarly, Figure 4.4-2 presents the peak value of fuel centerline temperature versus linear power density, attained during its lifetime. The maximum linear power density throughout the lifetime of the fuel is maintained below the value that would cause UO2 fuel centerline melt.
The maximum pellet temperature at the hot spot during full power steady state and at the maximum overpower T trip point is shown in Table 4.1-1 for Unit 1 and Unit 2. The principal factors employed in fuel temperature determinations are discussed below.
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DCPP UNITS 1 & 2 FSAR UPDATE 4.4.3.2.2 UO2 Thermal Conductivity The thermal conductivity of UO2 was evaluated from data reported in References 7 and 12 through 24.
At the higher temperatures, thermal conductivity is best obtained by utilizing the integral conductivity to melt, which can be determined with more certainty. From an examination of the data, it has been concluded that the best estimate for the value of 2800 C kdT is 93 watts/cm. This conclusion is based on the integral values reported in 0
References 10 and 24 through 28.
The design curve for the thermal conductivity is shown in Figure 4.4-3. The section of the curve at temperatures between 0 and 1300°C is in excellent agreement with the recommendation of the International Atomic Energy Agency (IAEA) panel (Reference 29). The section of the curve above 1300°C is derived for an integral value of 93 watts/cm (Reference 89).
Thermal conductivity for UO2 at 95 percent theoretical density can be represented best by the following equation:
1 k 8.775 10 13 T 3 (4.4-1) 11.80.0238T where:
k is in watts/cm-°C, and T is in °C 4.4.3.2.3 Radial Power Distribution in UO2 Fuel Rods An accurate radial power distribution as a function of burnup is needed to determine the power level for incipient fuel melting and other important performance parameters, e.g., pellet thermal expansion, fuel swelling, and fission gas release rates.
This UO2 fuel rods radial power distribution as a function of core burnup is determined using a 3D ANC calculation (refer to Sections 4.3.3.10.3 and 15.1.2.4) or with the neutron transport theory LASER (Reference 81) code that has been validated by comparing code predictions on radial burnup and isotopic distributions with measured radial microdrill data (References 30 and 31). "Microdrill data" are data obtained from the physical examination of irradiated pellets in a hot cell. Small core samples are removed from different radial positions in a pellet (using a "microdrill"). Isotopic measurements of the fuel samples determine actual UO2 burnups at the sample points.
A "radial power depression factor," f, is determined using radial power distribution predicted by LASER (Reference 81).
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DCPP UNITS 1 & 2 FSAR UPDATE 4.4.3.2.4 Gap Conductance The temperature drop across the pellet-cladding gap is a function of the gap size and the thermal conductivity of the gas in the gap. The gap conductance model is selected such that when combined with the UO2 thermal conductivity model, the calculated fuel centerline temperatures reflect the inpile temperature measurements. A more detailed discussion of the gap conductance model is presented in Reference 101.
4.4.3.2.5 Surface Heat Transfer Coefficients The fuel rod surface heat transfer coefficients during subcooled forced convection and the outer cladding wall temperature for the onset of nucleate boiling is presented in Section 4.4.3.8.1.
4.4.3.2.6 Fuel Cladding Temperatures The fuel rod outer surface at the hot spot operates at a temperature of approximately 660°F for steady state operation at rated power throughout core life, due to the onset of nucleate boiling. At BOL, this temperature is that of the cladding metal outer surface.
During operation over the life of the core, the buildup of oxides and crud on the fuel rod cladding outer surface causes the cladding surface temperature to increase. Allowance is made in the fuel center melt evaluation for this temperature rise. The thermal-hydraulic DNB limits ensure that adequate heat transfer is provided between the fuel cladding and the reactor coolant so that cladding temperature does not limit core thermal output. Figure 4.4-4 shows the axial variation of average cladding temperature for a representative average power rod both at BOL and EOL.
4.4.3.2.7 Treatment of Peaking Factors The total heat flux hot channel factor, F QT , is defined by the ratio of the maximum to core average heat flux. The design value of F QT for normal operation is 2.58, allowing for fuel densification effects, as shown in Table 4.3-1. This results in a peak local linear power density of 14.3 kW/ft at full power. The corresponding peak local power at the maximum overpower trip point (118 percent total power) is 16.6 kW/ft. Centerline temperature at this kW/ft must be below the UO2 melt temperature over the lifetime of the rod including allowances for uncertainties. From Figure 4.4-2, the centerline temperature at the maximum overpower trip point is well below that required to produce melting. The maximum linear power density throughout the lifetime of the fuel is maintained below the value that would cause UO2 fuel centerline melt. Fuel centerline and average temperature at rated (100 percent) power and at the maximum overpower trip point for Unit 1 and Unit 2 are presented in Table 4.1-1.
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DCPP UNITS 1 & 2 FSAR UPDATE 4.4.3.3 Departure from Nucleate Boiling Ratio The minimum DNBRs for the rated power, and anticipated transient conditions are given in Table 4.1-1 for Unit 1 and Unit 2. The minimum DNBR in the limiting flow channel will occur downstream of the peak heat flux location (hot spot) due to the increased downstream enthalpy rise.
DNBRs are calculated by using the correlation and definitions described in Section 4.4.3.3.1. The THINC-IV (Reference 47) computer code (discussed in Section 4.4.3.15.1) determines the flow distribution in the core and the local conditions in the hot channel for use in the DNB correlation. The use of hot channel factors is discussed in Section 4.4.3.13.1 (nuclear hot channel factors) and in Section 4.4.3.3.4 (engineering hot channel factors).
4.4.3.3.1 Departure from Nucleate Boiling Technology The WRB-2 DNB correlation (Reference 85) was developed to take credit for the VANTAGE+ fuel assembly mixing vane design. A DNBR limit of 1.17 is also applicable for the WRB-2 correlation. Figure 4.4-20 shows measured critical heat flux (CHF) plotted against predicted CHF using the WRB-2 correlation.
In several cases, the W-3 DNB correlation is used where the WRB-2 is not applicable.
For example, Section 15.2.14.1 uses the W-3 DNB correlation since the system pressure for the limiting statepoint is below 1,000 psia.
The W-3 DNB correlation, and several modifications, have been used in Westinghouse CHF calculations. The W-3 DNB was originally developed from single tube data (Reference 34), but was subsequently modified to apply to the 0.422 inch, OD rod low parasitic reinforced grid (R-grid) (References 35 and 92) and low parasitic grid (L-grid)
(Reference 36), as well as the 0.374 inch OD (References 37 and 38) rod bundle data.
These modifications to the W-3 DNB correlation have been demonstrated to be adequate for reactor rod bundle design.
A description of the 17 x 17 fuel assembly test program and a summary of the results are described in detail in Reference 37.
Figure 4.4-5 shows the data obtained in this test program. The test results indicate that a reactor core using this geometry may operate with a minimum DNBR of 1.28 and satisfy the design criterion.
HISTORICAL INFORMATION IN ITALICS BELOW NOT REQUIRED TO BE REVISED.
The WRB-1 correlation (Reference 84) was developed based exclusively on the large bank of mixing vane grid rod bundle CHF data (over 1100 points) that Westinghouse has collected. The WRB-1 correlation, based on local fluid conditions, represents the rod bundle data with better accuracy over the wide range of variables than the previous 4.4-6 Revision 24 September 2018
DCPP UNITS 1 & 2 FSAR UPDATE correlation used in design. This correlation accounts directly for both typical and thimble cold wall cell effects, uniform and non-uniform heat flux profiles, and variations in rod heated length and in grid spacing.
Figure 4.4-19 shows measured CHF plotted against predicted CHF using the WRB-1 correlation.
CHF tests which model the 17x17 optimized fuel assembly have been performed with the results described in detail in Reference 87. It was concluded that the CHF characteristics of the 17x17 optimized fuel assembly design are not significantly different from those of 17x17 LOPAR design, and can be adequately described by the "R" grid form of the WRB-1 CHF correlation. Furthermore, the new data can be incorporated into the "R" grid data base such that the WRB-1 correlation can be applied to 17x17 LOPAR fuel design without changing the DNBR design criterion of 1.17.
4.4.3.3.2 Definition of Departure from Nucleate Boiling Ratio The DNBR, as applied to this design for both typical and thimble cold wall cells is:
q "DNB, Predicted DNBR (4.4-5) q "actual For the W-3 DNB (R-Grid) correlation,
" qEU, W 3 FS qDNB,Predicted (4.4-6)
F when all flow cell walls are heated and q" EU, W-3 is the uniform DNB heat flux as predicted by W-3 DNB correlation and F is the flux shape factor which accounts for non-uniform axial heat flux distributions (Reference 39) with the "C" term modified as in Reference 34.
F S' is the modified spacer factor described in Reference 37 using an axial grid spacing coefficient, KS = 0.046, and a thermal diffusion coefficient (TDC) of 0.038, based on the 26-inch grid spacing data. Since the actual grid spacing is approximately 20 inches, these values are conservative since the DNB performance was found to improve and TDC increase as axial grid spacing is decreased (References 35 and 40).
When a cold wall is present for the W-3 DNB correlation, qDNB,Predicted qEU,W 3,CW FS (4.4-7) 4.4-7 Revision 24 September 2018
DCPP UNITS 1 & 2 FSAR UPDATE where:
q "EU, W 3,Dh q"EU, W 3,CW CWF (4.4-7A)
F qEU, W 3,Dh is the uniform DNB heat flux as predicted by the W-3 DNB cold wall correlation (Reference 34) when not all flow cell walls are heated (thimble cold wall cell).
The cold wall factor (CWF) is provided in References 34 and 39. For the WRB-2 correlation,
" q "WRB 2 q DNB,Predicted (4.4-8)
F HISTORICAL INFORMATION IN ITALICS BELOW NOT REQUIRED TO BE REVISED.
" q "WRB 1 qDNB, Predicted for WRB1 correlation F
where:
F is the same flux shape factor that is used with the W-3 DNB correlation.
4.4.3.3.3 Mixing Technology The rate of heat exchange by mixing between flow channels is proportional to the difference in the local mean fluid enthalpy of the respective channels, the local fluid density, and the flow velocity. The proportionality is expressed by the dimensionless TDC, which is defined as:
w' TDC (4.4-9)
Va where:
w' = flow exchange rate per unit length, lbm/ft-sec
= fluid density, lbm/ft3 V = fluid velocity, ft/sec a = lateral flow area between channels per unit length, ft2/ft The application of the TDC in the THINC analysis for determining the overall mixing effect or heat exchange rate is presented in Reference 41.
The TDC is determined by comparing the THINC code predictions with the measured subchannel exit temperatures. Data for 26-inch axial grid spacing are presented in Figure 4.4-6 where the TDC is plotted versus the Reynolds number. The TDC is found 4.4-8 Revision 24 September 2018
DCPP UNITS 1 & 2 FSAR UPDATE to be independent of the Reynolds number, mass velocity, pressure, and quality over the ranges tested.
The two-phase data (local and subcooled boiling) fell within the scatter of the single-phase data. The effect of two-phase flow on the value of TDC has been demonstrated by Cadek (Reference 40), Rowe and Angle (References 42 and 43), and Gonzalez-Santalo and Griffith (Reference 44). In the subcooled boiling region, the values of TDC were indistinguishable from the single-phase values. In the quality region, Rowe and Angle show that in the case with rod spacing similar to that in PWR reactor core geometry, the value of TDC increased with quality to a point and then decreased but never below the single-phase value. Gonzalez-Santalo and Griffith showed that the mixing coefficient increased as the void fraction increased.
The data from these tests on the R-grid showed that a design TDC value of 0.038 (for 26 inch grid spacing) can be used in determining the effect of coolant mixing in the THINC analysis. A mixing test program similar to the one described above was conducted at Columbia University for the 17 x 17 geometry and mixing vane grids on 26-inch spacing (Reference 45). The mean value of TDC obtained from these tests was 0.059, and all data were well above the current design value of 0.038.
Because the reactor grid spacing is approximately 20 inches, additional margin is available for this design, as the value of TDC increases as grid spacing decreases (Reference 40).
The inclusion of three IFM grids in the upper span of the VANTAGE+ fuel assembly results in a grid spacing of approximately 10 inches. Therefore, the design value of 0.038 for TDC is a conservatively low value for use in VANTAGE+ to determine the effect of coolant mixing in the core thermal performance analysis.
4.4.3.3.4 Hot Channel Factors The total hot channel factors for heat flux and enthalpy rise are defined as the maximum-to-core average ratios of these quantities. The heat flux hot channel factor considers the local maximum linear heat generation rate at a point (the "hot spot"), and the enthalpy rise hot channel factor involves the maximum integrated value along a channel (the "hot channel").
Each of the total hot channel factors considers a nuclear hot channel factor (refer to Section 4.4.3.13) describing the neutron power distribution and an engineering hot channel factor, which allows for variations in flow conditions and fabrication tolerances.
The engineering hot channel factors are made up of subfactors that account for the influence of the variations of fuel pellet diameter, density, enrichment and eccentricity; fuel rod diameter pitch and bowing; inlet flow distribution; flow redistribution; and flow mixing.
4.4-9 Revision 24 September 2018
DCPP UNITS 1 & 2 FSAR UPDATE 4.4.3.3.4.1 Heat Flux Engineering Hot Channel Factor, F EQ The heat flux engineering hot channel factor is used to evaluate the maximum heat flux.
This subfactor is determined by statistically combining the tolerances for the fuel pellet diameter, density, enrichment, eccentricity, and the fuel rod diameter, and has a value of 1.03. Measured manufacturing data on Westinghouse fuel verify that this value was not exceeded for 95 percent of the limiting fuel rods at a 95 percent confidence level.
As shown in Reference 99, no DNB penalty need be taken for the short, relatively low intensity heat flux spikes caused by variations in the above parameters.
4.4.3.3.4.2 Enthalpy Rise Engineering Hot Channel Factor, F EH The effect of variations in flow conditions and fabrication tolerances on the hot channel enthalpy rise is directly considered in the THINC core thermal subchannel analysis (refer to Section 4.4.3.15.1) under any reactor operating condition. The following items contribute to the enthalpy rise engineering hot channel factor:
(1) Pellet Diameter, Density and Enrichment, Fuel Rod Diameter, Pitch, and Bowing Design values employed in the THINC analysis are based on applicable limiting tolerances such that design values are met for 95 percent of the limiting channels at a 95 percent confidence level. The effect of variations in pellet diameter and enrichment is employed in the THINC analysis as a direct multiplier on the hot channel enthalpy rise, while the fuel rod diameter, pitch, and bowing variation, including incore effects, enter in the preparation of the THINC input values.
(2) Inlet Flow Maldistribution Inlet flow maldistribution in the core thermal performances is discussed in Section 4.4.3.12.2. A design basis of 5 percent reduction in coolant flow to the hot assembly is used in the THINC-IV analysis.
(3) Flow Redistribution The flow redistribution accounts for the flow reduction in the hot channel resulting from the high flow resistance in the channel due to the local or bulk boiling. The effect of the non-uniform power distribution is inherently considered in the THINC analysis.
(4) Flow Mixing The subchannel mixing model incorporated in the THINC code and used in reactor design is based on experimental data (Reference 46), as discussed in Section 4.4.3.15.1. The mixing vanes incorporated in the 4.4-10 Revision 24 September 2018
DCPP UNITS 1 & 2 FSAR UPDATE spacer grid design induce additional flow mixing between the various flow channels in a fuel assembly, as well as between adjacent assemblies.
This mixing reduces the enthalpy rise in the hot channel resulting from local power peaking or unfavorable mechanical tolerances.
4.4.3.3.5 Effects of Rod Bow on Departure from Nucleate Boiling Ratio The phenomenon of fuel rod bowing, as described in Reference 79, must be accounted for in the DNBR safety analysis of Condition I and Condition II events for each plant application. Applicable generic credits for margin resulting from retained conservatism in the evaluation of DNBR and/or margin obtained from measured plant operating parameters (such as F NH or core flow), which are less limiting than those required by the plant safety analysis, can be used to offset the effect of rod bow.
The safety analysis for DCPP cores maintains sufficient margin between the safety analysis DNBR limits and the design DNBR limits. The design DNBR limits are shown below to accommodate full flow and low flow DNBR penalties identified in Reference 80, which are applicable to 17x17 VANTAGE+ fuel assembly analysis utilizing the WRB-2 correlation.
However, for the upper assembly span of VANTAGE+ fuel where additional restraint is provided with the IFM grids, the grid-to-grid spacing in DNB limiting span is approximately 10 inches . Using the rod bow topical report methods (Reference 79),
and scaling with the NRC approved factor results in predicted channel closure in the limiting spans of less than 50 percent closure; no rod bow DNBR penalty is required in the 10 inch spans in the VANTAGE+ safety analyses.
VANTAGE+
Design Limit Typical Cell 1.34 Thimble Cell 1.32 Safety Limit Typical Cell 1.71 Thimble Cell 1.68 The maximum rod bow penalties accounted for in the design safety analysis are based on an assembly average burnup of 24,000 MWD/MTU based on Reference 100. At burnups greater than 24,000 MWD/MTU, credit is taken for the effect of F NH burndown.
Due to the decrease in fissionable isotopes and the buildup of fission product inventory, no additional rod bow penalty is required.
4.4-11 Revision 24 September 2018
DCPP UNITS 1 & 2 FSAR UPDATE HISTORICAL INFORMATION IN ITALICS BELOW NOT REQUIRED TO BE REVISED.
4.4.3.3.6 Transition Core The Westinghouse transition core DNB methodology is given in References 89 and 90 and has been approved by the NRC via Reference 91. Using this methodology, transition cores are analyzed as if they were full cores of one assembly type (full LOPAR or full VANTAGE 5), applying the applicable transition core penalties. This penalty was included in the safety analysis limit DNBRs such that sufficient margin over the design limit DNBR existed to accommodate the transition core penalty and the appropriate rod bow DNBR penalty. However, since the transition to a full VANTAGE 5 core has been completed, various analyses, such as large break and small LOCA analysis, have assumed a full VANTAGE 5 core and no longer assume a transition core penalty.
The LOPAR and VANTAGE 5 designs have been shown to be hydraulically compatible in Reference 85.
4.4.3.4 Flux Tilt Considerations Significant quadrant power tilts are not anticipated during normal operation since this phenomenon is caused by asymmetric perturbations. A dropped or misaligned RCCA could cause changes in hot channel factors. These events are analyzed separately in Chapter 15.
Other possible causes for quadrant power tilts include X-Y xenon transients, inlet temperature mismatches, enrichment variations within tolerances, and so forth.
In addition to unanticipated quadrant power tilts, other readily explainable asymmetries may be observed during calibration of the excore detector quadrant power tilt alarm.
During operation, at least one power distribution measurement is taken per effective-full-power month. Each of these power distribution measurements is reviewed for deviations from the expected power distributions. The acceptability of an observed asymmetry, planned or otherwise, depends solely on meeting the required accident analyses assumptions. In practice, once acceptability has been established by review of the power distribution measurements, the quadrant power tilt alarms and related instrumentation are adjusted to indicate zero quadrant tilt, 1.00 quadrant power tilt ratio, as the final step in the calibration process. Proper functioning of the quadrant power tilt alarm is significant because no allowances are made in the design for increased hot channel factors due to unexpected developing flux tilts since all likely causes are prevented by design or procedures or specifically analyzed. Finally, in the event that unexplained flux tilts do occur, the Technical Specification (Reference 82) stipulates appropriate corrective actions to ensure continued safe operation of the reactor.
4.4-12 Revision 24 September 2018
DCPP UNITS 1 & 2 FSAR UPDATE 4.4.3.5 Void Fraction Distribution The calculated core average and the hot subchannel maximum and average void fractions are presented in Tables 4.4-1 and 4.4-2 for operation at full power with design hot channel factors for Unit 1 and Unit 2, respectively. The void fraction distribution in the core is presented in Reference 47. The void fraction as a function of thermodynamic quality is shown in Figure 4.4-10. The void models used in the THINC-IV computer code are described in Section 4.4.3.8.3.
4.4.3.6 Core Coolant Flow Distribution Coolant enthalpy rise and flow distributions are shown for the 4-foot elevation (1/3 of core height) in Figure 4.4-7, 8-foot elevation (2/3 of core height) in Figure 4.4-8, and at the core exit in Figure 4.4-9. These distributions correspond to a representative Westinghouse 4-loop plant. The THINC code analysis for this case utilized a uniform core inlet enthalpy and inlet flow distribution.
4.4.3.7 Core Pressure Drops and Hydraulic Loads 4.4.3.7.1 Core Pressure Drops The analytical model and experimental data used to calculate the pressure drops, for the full power conditions given in Table 4.1-1, are described in Section 4.4.3.8.2. The core pressure drop consists of the fuel assembly, lower core plate, and upper core plate pressure drops. These pressure drops are based on the best estimate flow, as described in Section 5.1.6. Section 5.1.6 also defines the thermal design flow (minimum flow), which is the basis for reactor core thermal performance, and the mechanical design flow (maximum flow), which is used in the mechanical design of the reactor vessel internals and fuel assemblies. Since the best estimate flow is that which is most likely to exist in an operating plant, the calculated core pressure drops in Table 4.1-1 are greater than pressure drops previously quoted using the thermal design flow. The relation between best estimate flow, thermal design flow, and mechanical design flow is illustrated in Figure 5.1-2.
4.4.3.7.2 Hydraulic Loads Maximum flow conditions are limiting because hydraulic loads are a maximum. The most adverse flow conditions occur during a LOCA, as discussed in Section 15.4.1.
Hydraulic loads at normal operating conditions are calculated considering the best estimate flow and accounting for the best estimate core bypass flow. Core hydraulic loads at cold plant startup conditions are based on the cold best estimate flow, but are adjusted to account for the coolant density difference. Conservative core hydraulic loads for a pump overspeed transient, which are based on flowrates 20 percent greater than the mechanical design flow (refer to Section 5.1.6), are evaluated to be greater than twice the fuel assembly weight.
4.4-13 Revision 24 September 2018
DCPP UNITS 1 & 2 FSAR UPDATE The hydraulic verification tests are discussed in References 48 and 87.
4.4.3.8 Correlation and Physical Data 4.4.3.8.1 Surface Heat Transfer Coefficients Forced convection heat transfer coefficients are obtained from the familiar Dittus-Boelter correlation (Reference 49), with the properties evaluated at bulk fluid conditions:
0.8 C 0.4 hD e D G p 0.023 e (4.4-10) k k where:
h = heat transfer coefficient, Btu/hr-ft2-°F De = equivalent diameter, ft k = thermal conductivity, Btu/hr-ft-°F G = mass velocity, lbm/hr-ft2
= dynamic viscosity, lbm/ft-hr Cp = heat capacity, Btu/lbm-°F This correlation has been shown to be conservative (Reference 50) for rod bundle geometries with pitch-to-diameter ratios in the range used by PWRs.
The onset of nucleate boiling occurs when the cladding wall temperature reaches the amount of superheat predicted by Thom's (Reference 51) correlation. After this occurrence, the outer cladding wall temperature is determined by:
Tsat = [0.072 exp (-P/1260)] (q")0.5 (4.4-11) where:
TSAT = wall superheat, Tw - Tsat, °F q" = wall heat flux, Btu/hr-ft2 P = pressure, psia Tw = outer cladding wall temperature, °F TSAT = saturation temperature of coolant at P, °F 4.4.3.8.2 Total Core and Vessel Pressure Drop Pressure losses occur as a result of viscous drag (friction) and/or geometry changes (form) in the fluid flowpath. The flow field is assumed to be incompressible, turbulent, single-phase water. Two-phase considerations are neglected in the vessel pressure 4.4-14 Revision 24 September 2018
DCPP UNITS 1 & 2 FSAR UPDATE drop evaluation because the core average void is negligible (refer to Section 4.4.3.5 and Tables 4.4-1 and 4.4-2).
Two-phase flow considerations in the core thermal subchannel analyses are considered and the models are discussed in Section 4.4.3.12.3. Core and vessel pressure losses are calculated by equations of the form:
FL V 2 PL K (4.4-12)
D e 2gc (144) where:
PL = pressure drop, lbf/in2
= fluid density, lbm/ft3 L = length, ft De = equivalent diameter, ft V = fluid velocity, ft/sec 1b m ft gc = 32.174 1b f sec 2 K = form loss coefficient, dimensionless F = friction loss coefficient, dimensionless Fluid density is assumed to be constant at an appropriate value for each component in the core and vessel. Because of the complex core and vessel flow geometry, precise analytical values for the form and friction loss coefficients are not available. Therefore, experimental values for these coefficients are obtained from geometrically similar models.
The results of full-scale tests of core components and fuel assemblies were utilized in developing the core pressure loss characteristics. The pressure drop for the vessel was obtained by combining the core pressure loss with correlation of 1/7th scale model hydraulic test data on a number of vessels (References 52 and 53) and form loss relationships (Reference 54). Moody (Reference 55) curves were used to obtain the single-phase friction factors.
Tests of the primary coolant loop flowrates are made (refer to Section 4.4.6.1) prior to initial criticality to verify that the flowrates used in the design are conservative.
4.4.3.8.3 Void Fraction Correlation Three separate void regions are considered in flow boiling in a PWR as illustrated in Figure 4.4-10. They are the wall void region (no bubble detachment), the subcooled boiling region (bubble detachment), and the bulk boiling region.
4.4-15 Revision 24 September 2018
DCPP UNITS 1 & 2 FSAR UPDATE In the wall void region, local boiling begins at the point where the cladding temperature reaches the amount of superheat predicted by Thom's (Reference 51) correlation (refer to Section 4.4.3.8.1). The void fraction in this region is calculated using Maurer's (Reference 56) relationship. The bubble detachment point, where the superheated bubbles break away from the wall, is determined by using Griffith's (Reference 57) relationship.
The void fraction in the subcooled boiling region (i.e., after the detachment point) is calculated from the Bowring (Reference 58) correlation. This correlation predicts the void fraction from the detachment point to the bulk boiling region.
The void fraction in the bulk boiling region is predicted by using homogeneous flow theory and assuming no slip. The void fraction in this region is, therefore, a function of steam quality only.
4.4.3.9 Thermal Effects of Operational Transients DNB core safety limits are generated as a function of coolant temperature, pressure, core power, and axial power imbalance. Steady state operation within these safety limits ensures that the minimum DNBR is not less than the safety limit DNBR.
Figure 15.1-1 shows lines at the safety limit DNBR and the resulting overtemperature T trip lines (which are part of the Technical Specifications), plotted as T versus T-average for various pressures. This system provides adequate protection against anticipated operational transients that are slow with respect to fluid transport delays in the primary system. In addition, for fast transients (e.g., uncontrolled rod bank withdrawal at power incident) (refer to Section 15.2.2), specific protection functions are provided as described in Section 7.2; their use is described in Chapter 15 (refer to Table 15.1-2). Fuel rod thermal response is discussed in Section 4.4.3.18.
4.4.3.10 Uncertainties in Estimates 4.4.3.10.1 Uncertainties in Fuel and Cladding Temperatures As discussed in Section 4.4.3.2, the fuel temperature is a function of crud, oxide, cladding, gap, and pellet conductance. Uncertainties in the fuel temperature calculation are essentially of two types: fabrication uncertainties, such as variations in the pellet and cladding dimensions and the pellet density; and model uncertainties, such as variations in the pellet conductivity and the gap conductance. These uncertainties have been quantified by comparison of the thermal model to the incore thermocouple measurements (References 3 through 9), by out-of-pile measurements of the fuel and cladding properties (References 7 and 12 through 23), and by measurements of the fuel and cladding dimensions during fabrication. The effect of densification on fuel temperature uncertainties is presented in Reference 68.
4.4-16 Revision 24 September 2018
DCPP UNITS 1 & 2 FSAR UPDATE In addition, the measurement uncertainty in determining the local power, and the effect of density and enrichment variations on local power, are considered in establishing the heat flux hot channel factor.
Uncertainty in determining cladding temperature results from uncertainties in the crud and oxide thickness. Because of the excellent heat transfer between the surface of the rod and the coolant, the film temperature drop does not appreciably contribute to the uncertainty.
Reactor trip setpoints, as specified in the Technical Specifications, include allowance for instrument and measurement uncertainties such as calorimetric error, instrument drift, and channel reproducibility.
4.4.3.10.2 Uncertainties in Pressure Drops Core and vessel pressure drops based on the best estimate flow, as described in Section 5.1, are quoted in Table 4.1-1. The uncertainties quoted are based on the uncertainties in both the test results and the analytical extension of these values to the reactor application.
4.4.3.10.3 Uncertainties Due to Inlet Flow Maldistribution Uncertainties in the inlet flow maldistribution criteria used in the core thermal analyses are discussed in Section 4.4.3.12.2.
4.4.3.10.4 Uncertainty in Departure from Nucleate Boiling Correlation The uncertainty in the DNB correlation (refer to Section 4.4.3.3) can be written as a statement on the probability of not being in DNB based on the DNB data statistics. This is discussed in Section 4.4.3.3.2.
4.4.3.10.5 Uncertainties in Departure from Nucleate Boiling Ratio Calculations The uncertainties in the DNBRs calculated by THINC analysis (refer to Section 4.4.3.15.1) due to uncertainties in the nuclear peaking factors are accounted for by applying conservatively high values of the nuclear peaking factors and including measurement error allowances in the statistical evaluation of the limit DNBR (refer to Section 4.4.4.1) using the Improved Thermal Design Procedure (ITDP) (Reference 86).
In addition, conservative values for the engineering hot channel factors are used (refer to Section 4.4.3.3.4). The results of a sensitivity study with THINC-IV show that the minimum DNBR in the hot channel is relatively insensitive to variations in the core-wide radial power distribution (for the same value of F NH ).
4.4-17 Revision 24 September 2018
DCPP UNITS 1 & 2 FSAR UPDATE The ability of the THINC-IV computer code to accurately predict flow and enthalpy distributions in rod bundles is discussed in Section 4.4.3.15.1 and in Reference 59. The sensitivity of the minimum DNBR in the hot channel to the void fraction correlation (refer to Section 4.4.3.8.3), the inlet velocity and exit pressure distributions, and the grid pressure loss coefficients have been studied (Reference 47). The results show that the minimum DNBR in the hot channel is relatively insensitive to variations in these parameters.
4.4.3.10.6 Uncertainties in Flowrates The uncertainties associated with loop flowrates are discussed in Section 5.1. For core thermal performance evaluations, a thermal design loop flow is used which is less than the best estimate loop flow (by approximately 4 percent). In addition, another 7.5 percent (Unit 1) and 9.0 percent (Unit 2) of the thermal design flow is assumed to be ineffective for core heat removal capability because it bypasses the core through the various available flowpaths described in Section 4.4.3.12.1.
4.4.3.10.7 Uncertainties in Hydraulic Loads As discussed in Section 4.4.3.7.2, hydraulic loads on the fuel assembly are evaluated for a pump overspeed transient that creates flowrates 20 percent greater than the mechanical design flow.
4.4.3.10.8 Uncertainty in Mixing Coefficients The value of the mixing coefficient, TDC, used in THINC analyses for this application is 0.038. The mean value of TDC obtained in the R-grid mixing tests described in Section 4.4.3.3.2 was 0.042 (for 26-inch grid spacing). The value of 0.038 is one standard deviation below the mean value and 90 percent of the data gives values of TDC greater than 0.038 (Reference 40).
The results of the mixing tests discussed in Section 4.4.3.3.3, had a mean TDC value of 0.059 and standard deviation of = 0.007. Hence, the current design TDC value is almost three standard deviations below the mean for 26-inch grid spacing.
Since the value of TDC increases as grid spacing decreases, the design value of 0.038 for TDC is a conservatively low value for use in VANTAGE+ to determine the effect of coolant mixing in the core thermal performance analysis. Refer to Section 4.4.3.3.3 for a discussion of the current reactor grid spacing of approximately 20 inches and IFM grid spacing of approximately 10 inches.
4.4.3.11 Plant Configuration Data Plant configuration data for the thermal-hydraulic and fluid systems external to the core are provided in Chapters 5, 6, and 9. Implementation of the ECCS is discussed in Chapters 6 and 15. Some specific areas of interest are:
4.4-18 Revision 24 September 2018
DCPP UNITS 1 & 2 FSAR UPDATE (1) Total coolant flow rate for the RCS is provided in Table 5.1-1.
(2) Total RCS volume is given in Table 5.1-1.
(3) The flowpath length through each volume is calculated from physical data and the component pressure drops as provided in Table 5.1-1.
(4) The height of fluid in each component of the RCS may be determined from the physical data and data provided in Table 5.1-1. The RCS components are water-filled during power operation with the pressurizer being approximately 60 percent water-filled.
(5) ECCS components are located to meet the criteria for net positive suction head described in Section 6.3.
(6) Line lengths and sizes for the safety injection system are determined so as to guarantee a total system resistance that will provide, as a minimum, the fluid delivery rates assumed in the safety analyses described in Chapter 15.
(7) The design data for RCS components, including volumes (diameter),
flows, and pressure and temperature, are presented in Section 5.5.
(8) RCS steady state pressure and temperature distribution are presented in Table 5.1-1.
4.4.3.12 Core Hydraulics 4.4.3.12.1 Flowpaths Considered in Core Pressure Drop and Thermal Design The following flowpaths are considered:
(1) Flow through the spray nozzles into the upper head for cooling purposes (2) Flow entering the RCCA guide thimbles to cool core components (3) Leakage flow from the vessel inlet nozzle directly to the vessel outlet nozzle through the gap between vessel and barrel (4) Flow entering the core from the baffle-barrel region through the gaps between the baffle plates (5) Flow introduced between baffle and barrel to cool these components (6) Flow through the empty guide thimble tubes 4.4-19 Revision 24 September 2018
DCPP UNITS 1 & 2 FSAR UPDATE The above contributions are evaluated to confirm that the design basis value of 7.5 percent (Unit 1) and 9.0 percent (Unit 2) core bypass flow is met.
4.4.3.12.2 Inlet Flow Distribution Data from several 1/7 scale hydraulic reactor model tests (References 52, 53, and 60) have been considered in arriving at the core inlet flow maldistribution criteria to be used in the THINC analyses (refer to Section 4.4.3.15.1). THINC (Reference 41) analyses have indicated that a conservative design basis is to consider a 5 percent reduction in the flow to the hot assembly (Reference 61).
The experimental error in the inlet velocity distribution has been estimated in Reference 47. The sensitivity of changes in inlet velocity distributions to hot channel thermal performance is shown to be small.
The effect of the total flowrate on the inlet velocity distribution was studied in the experiments of Reference 52. As expected, no significant variation could be found in inlet velocity distribution with reduced flowrate.
4.4.3.12.3 Empirical Friction Factor Correlations, F EQ Two empirical friction factor correlations are used in the THINC-IV computer code (refer to Section 4.4.3.15.1).
The friction factor in the axial direction, parallel to the fuel rod axis, uses the Novendstern-Sandberg (Reference 62) correlation. This correlation consists of the following:
(1) For isothermal conditions, this correlation uses the Moody (Reference 55) friction factor, including surface roughness effects.
(2) Under single-phase heating conditions, a factor is applied based on the values of the coolant density and viscosity at the temperature of the heated surface and at the bulk coolant temperature.
(3) Under two-phase flow conditions, the homogeneous flow model proposed by Owens (Reference 63) is used with a modification to account for a mass velocity and heat flux effect.
The flow in the lateral directions, normal to the fuel rod axis, views the reactor core as a large tube bank. Thus, the lateral friction factor proposed by Idel'chik (Reference 54) is applicable. This correlation is of the form:
FL A Re L 0.2 (4.4-13) 4.4-20 Revision 24 September 2018
DCPP UNITS 1 & 2 FSAR UPDATE where:
A is a function of the rod pitch and diameter as given in Reference 54 ReL is the lateral Reynolds number based on rod diameter Extensive comparisons of THINC-IV predictions using these correlations to experimental data are given in Reference 59, and verify the applicability of these correlations in PWR design.
4.4.3.13 Influence of Power Distribution The core power distribution, which at BOL is largely established by fuel enrichment, loading pattern, and core power level, is also a function of variables such as control rod worth and position and fuel depletion throughout lifetime. Although radial power distributions in various planes of the core are often illustrated for general interest, the core radial enthalpy rise distribution, as determined by the integral of power over each channel, is of greater importance for DNB analyses. These radial power distributions, N
characterized by FH (defined in Section 4.3.3.2.2), as well as axial heat flux profiles, are discussed in the following two sections.
4.4.3.13.1 Nuclear Enthalpy Rise Hot Channel Factor, F NH Given the local power density q' (kW/ft) at a point x, y, z in a core with N fuel rods and height H:
Max o q'x o , y o , z dz H
hot rod power N
FH (4.4-14) average rod power l H o q'(x,y,z) dz N all rods The location of minimum DNBR depends on the axial profile and its magnitude depends on the enthalpy rise up to that point. The maximum value of the rod integral is used to identify the most likely rod for minimum DNBR. An axial power profile is obtained which, when normalized to the design value of F NH , recreates the axial heat flux along the limiting rod. The surrounding rods are assumed to have the same axial profile with rod average powers that are typical of distributions found in hot assemblies. In this manner, worst case axial profiles can be combined with worst case radial distributions for reference DNB calculations.
Local heat fluxes are obtained by using hot channel and adjacent channel explicit power shapes which take into account variations in horizontal power shapes throughout the core. The sensitivity of the THINC-IV analysis to radial power shapes is discussed in Reference 47.
For operation at a fraction P of full power, the design F NH is given by:
4.4-21 Revision 24 September 2018
DCPP UNITS 1 & 2 FSAR UPDATE F NH = 1.59 [1 + 0.3 (1-P)] (VANTAGE+) (4.4-15)
HISTORICAL INFORMATION IN ITALICS BELOW NOT REQUIRED TO BE REVISED.
F NH = 1.56 [1 + 0.3 (1-P)] (LOPAR)
The permitted relaxation of F NH is included in the DNB protection setpoints and allows radial power shape changes with rod insertion to the insertion limits (Reference 64),
thus allowing greater flexibility in the nuclear design.
4.4.3.13.2 Axial Heat Flux Distributions As discussed in Section 4.3.3.2, the axial heat flux distribution can vary as a result of rod motion, power change, or due to spatial xenon transients that may occur in the axial direction. Consequently, it is necessary to measure the axial power imbalance by means of the excore nuclear detectors (refer to Section 4.3.3.2.7) and protect the core from excessive axial power imbalance. The reactor trip system provides automatic reduction of the trip setpoint in the overtemperature T channels on excessive axial power imbalance, i.e., when an extremely large AO corresponds to an axial shape that could lead to a DNBR, which is less than that calculated for the reference DNB design axial shape.
The reference DNB design axial shape is a chopped cosine with a peak-to-average ratio of 1.55.
4.4.3.14 Core Thermal Response A general summary of the steady state thermal-hydraulic design parameters including thermal output, flowrates, etc., is provided in Table 4.1-1. As stated in Section 4.4.2, the acceptance criteria are to prevent DNB and to prevent fuel melting for Conditions I and II events. The protective systems described in Chapter 7 (Instrumentation and Controls) are designed to meet these bases. The response of the core to Condition II transients is given in Chapter 15.
4.4.3.15 Analytical Techniques 4.4.3.15.1 Core Analysis The objective of reactor core thermal analysis is to determine the maximum heat removal capability in all flow subchannels, and to show that the core safety limits, as presented in the Technical Specifications, are not exceeded. The thermal design considers local variations in dimensions, power generation, flow redistribution, and mixing. THINC-IV is a realistic three-dimensional matrix model developed to account for hydraulic and nuclear effects on the enthalpy rise in the core (References 47 and 4.4-22 Revision 24 September 2018
DCPP UNITS 1 & 2 FSAR UPDATE 59). The behavior of the hot assembly is determined by superimposing the power distribution among the assemblies upon the inlet flow distribution, while allowing for flow mixing and distribution between assemblies. The average flow and enthalpy in the hottest assembly is obtained from the core-wide assembly-by-assembly analysis. The local variations in power, fuel rod and pellet fabrication, and mixing within the hottest assembly are then superimposed on the average conditions of the hottest assembly to determine conditions in the hot channel.
The following sections describe the use of the THINC code in the thermal-hydraulic design evaluation.
4.4.3.15.1.1 Steady State Analysis The THINC-IV computer program determines coolant density, mass velocity, enthalpy, vapor void, static pressure, and DNBR distributions along parallel flow channels within a reactor core under all expected operating conditions. The core region being studied is made up of a number of contiguous elements in a rectangular array extending the full length of the core. An element may represent any region of the core from a single assembly to a subchannel.
The momentum and energy exchange between elements in the array are described by the conservation of energy and mass equations, the axial momentum equation, and two lateral momentum equations that couple each element with its neighbors. The momentum equations used in THINC-IV incorporate frictional loss terms that represent the combined effects of frictional and form drag due to the presence of the grids and fuel assembly nozzles in the core. The cross flow resistance model used in the lateral momentum equations was developed from experimental data for flow normal to tube banks (Reference 54). The energy equation for each element also contains additional terms that represent the energy gain or loss due to the cross flow between elements.
The unique feature in THINC-IV is that lateral momentum equations, which include both inertial and cross flow resistance terms, are incorporated into the calculation scheme.
Another important consideration in THINC-IV is that the entire velocity field is solved, en masse, by a field equation. The solution method is complex and some simplifying techniques must be employed. Because the reactor flow is chiefly in the axial direction, the core flow field is primarily one-dimensional, and it is reasonable to assume that the lateral velocities and the parameter gradients are larger in the axial direction than the lateral direction. Thus, a perturbation technique is used to represent separately the axial and lateral parameters in the conservation equations.
Three THINC-IV computer runs constitute one design run: a core-wide analysis, a hot-assembly analysis, and a hot subchannel analysis.
The first computation is a core-wide assembly-by-assembly analysis that uses an inlet velocity distribution modeled from experimental reactor models (References 52, 53, and
- 60) (refer to Section 4.4.3.12.2). The core is made up of a number of contiguous fuel 4.4-23 Revision 24 September 2018
DCPP UNITS 1 & 2 FSAR UPDATE assemblies divided axially into increments of equal length. The system of perturbed and unperturbed equations is solved for this array giving the flow, enthalpy, pressure drop, temperature, and void fraction in each assembly. This computation determines the inter-assembly energy and flow exchange at each elevation for the hot assembly.
THINC-IV stores this information, then uses it for the subsequent hot assembly analysis.
In the second computation, each computational element represents one-fourth of the hot assembly. The inlet flow and the amount of momentum and energy interchange at each elevation are known from the previous core-wide calculation. The same solution technique is used to solve for the local parameters in the hot one-quarter assembly.
The third computation further divides the hot assembly into channels consisting of individual fuel rods to form flow channels. The local variations in power, fuel rod and pellet fabrication, fuel rod spacing and mixing (engineering hot channel factors) within the hottest assembly are imposed on the average conditions of the hottest fuel assembly to determine the conditions in the hot channel. Engineering hot channel factors are described in Section 4.4.3.3.4.
4.4.3.15.1.2 Experimental Verification An experimental verification (Reference 59) of the THINC-IV analysis for core-wide assembly-by-assembly enthalpy rises, as well as enthalpy rises in a non-uniformly heated rod bundle, have been obtained. In these tests, system pressure, inlet temperature, mass flowrate, and heat fluxes were typical of present PWR core designs.
During reactor operation, various incore monitoring systems obtain measured data indicating core performance. Assembly power distributions and assembly mixed mean temperature are measured and can be converted into the proper three-dimensional power input needed for the THINC programs. These data can then be used to verify the Westinghouse thermal-hydraulic design codes.
One standard startup test is the natural circulation test in which the core is held at a very low power (2 percent) and the pumps are turned off. The core will then be cooled by the natural circulation currents created by the power differences in the core. During natural circulation, a thermal siphoning effect occurs, resulting in the hotter assemblies gaining flow, thereby creating significant inter-assembly cross flow. Tests with significant cross flow are of more value in code verification.
Inter-assembly cross flow is caused by radial variations in pressure that are caused in turn by variations in the axial pressure drops in different assemblies. Under normal operating conditions (subcooled forced convection), the axial pressure drop is due mainly to friction losses. Because all assemblies have the same geometry, all assemblies have nearly the same axial pressure drops, and cross flow velocities are small. However, under natural circulation conditions (low flow) the axial pressure drop is due primarily to the difference in elevation head (or coolant density) between assemblies. This phenomenon can result in relatively large radial pressure gradients 4.4-24 Revision 24 September 2018
DCPP UNITS 1 & 2 FSAR UPDATE and, therefore, in higher cross flow velocities than at normal reactor operating conditions.
Incore instrumentation was used to obtain the assembly-by-assembly core power distribution during a natural circulation test. Assembly exit temperatures during the natural circulation test on a 157-assembly three-loop plant were predicted using THINC-IV. The predicted data points were plotted as assembly temperature rise versus assembly power, and a least squares fitting program was used to generate an equation that best fits the data. The result is the straight line presented in Figure 4.4-11 and is predicted closely by THINC-IV. This agreement verifies the lateral momentum equations and the cross flow resistance model used in THINC-IV.
Data have been obtained for Westinghouse plants operating from 67 to 101 percent of full power. A representative cross-section of the data obtained from a two-loop and a three-loop reactor was analyzed to verify the THINC-IV predictions that are compared with the experimental data in Figures 4.4-12 and 4.4-13.
The predicted assembly exit temperatures were compared with the measured exit temperatures for each data run. The standard deviations of the measured and predicted assembly exit temperatures are compared for both THINC-IV and THINC-I, and are given in Table 4.4-3. THINC-IV generally fits the data somewhat more accurately than THINC-I. Both codes are conservative and predict exit temperatures higher than measured values for the high-powered assemblies. Experimental verification of the THINC-IV subchannel calculation has been obtained from exit temperature measurements in a non-uniformly heated rod bundle (Reference 66).
Figure 4.4-14 compares, for a typical run, the measured and predicted temperature rises as a function of the power density in the channel. The THINC-IV results correctly predict the temperature gradient across the bundle.
In Figure 4.4-15, the measured and predicted temperature rises are compared for a series of runs at different pressures, flows, and power levels. Again, the measured points represent the average of the measurements taken in the various quadrants. The THINC-IV predictions provide a good representation of the data.
Thus, the THINC-IV analysis provides a realistic evaluation of the core performance and is used in the thermal analyses as described above.
4.4.3.15.1.3 Transient Analysis The THINC-III thermal-hydraulic computer code (Reference 41) is the third section of the THINC program and has transient DNB analysis capability.
The conservation equations needed for the transient analysis are included in THINC-III by adding the necessary accumulation terms to the conservation equations used in the 4.4-25 Revision 24 September 2018
DCPP UNITS 1 & 2 FSAR UPDATE steady state analysis. The input description must now include one or more of the following time arrays:
(1) Inlet flow variation (2) Heat flux distribution (3) Inlet pressure history At the beginning of the transient, the calculation procedure is carried out as in the steady state analysis. The THINC-III code is first run in the steady state mode to ensure conservatism with respect to THINC-IV and to provide the steady state initial conditions at the start of the transient. The time is incremented by an amount determined either by the user or by the program. At each new time step, the accumulation terms are evaluated using the information from the previous time step.
This procedure is continued until a preset maximum time is reached.
At various times during the transient, steady state THINC-IV is applied to show that the application of THINC-III is conservative. The THINC-III code does not have the capability for evaluating fuel rod thermal response. This is treated by the methods described in Section 15.1.9.
4.4.3.15.2 Fuel Temperatures As discussed in Section 4.4.3.2, fuel rod behavior is evaluated with a semiempirical thermal model that considers, in addition to the thermal aspects, such items as cladding creep, fuel swelling, fission gas release, release of absorbed gases, cladding corrosion, elastic deflection, and helium solubility.
A detailed description of the thermal model can be found in References 67 and 83 with the modifications for the time-dependent densification given in Reference 68.
4.4.3.16 Hydrodynamic and Flow-Power Coupled Instability Boiling flow may be susceptible to thermohydrodynamic instabilities (Reference 69).
These instabilities may cause a change in thermo-hydraulic conditions that may lead to a reduction in the DNB heat flux relative to that observed during a steady flow condition, or to undesired forced vibrations of core components. Thus, the thermo-hydraulic design criterion states that operation during Condition I and Condition II events shall not lead to thermohydrodynamic instabilities.
Two specific types of flow instabilities are considered by Westinghouse for PWR operation. These are the Ledinegg, or flow excursion, type of static instability and the density wave type of dynamic instability.
4.4-26 Revision 24 September 2018
DCPP UNITS 1 & 2 FSAR UPDATE A Ledinegg instability involves a sudden change in flowrate from one steady state to another. This instability occurs (Reference 69) when the slope of the RCS pressure drop-flowrate curve becomes algebraically smaller than the slope of the loop supply (pump head) pressure drop-flowrate curve. The Westinghouse pump head curve has a negative slope whereas the RCS pressure drop-flow curve has a positive slope over the Conditions I and II operational ranges. Thus, Ledinegg instability will not occur.
The mechanism of density wave oscillations in a heated channel has been described by Lahey and Moody (Reference 70).
The method developed by Ishii (Reference 71) for parallel closed channel systems evaluates if a given condition is stable with respect to the density wave type of dynamic instability. This method had been used to assess the stability of typical Westinghouse reactor designs (References 72, 73, and 74) under operating Conditions I and II. The results indicate that a large margin to density wave instability exists (e.g., an increase in the order of 200 percent of rated reactor power would be required) for the inception of this type of instability.
Flow instabilities that have been observed have occurred almost exclusively in closed channel systems operating at low pressures relative to the Westinghouse PWR operating pressures. Kao, Morgan, and Parker (Reference 75) analyzed parallel closed channel stability experiments simulating a reactor core flow. These experiments were conducted at pressures up to 2200 psia. The results showed that for flow and power levels typical of power reactor conditions, no flow oscillations could be induced above 1200 psia. Additional evidence that flow instabilities do not adversely affect thermal margin is provided by data from rod bundle DNB tests.
In summary, it is concluded that thermo-hydrodynamic instabilities will not occur under Conditions I and II events for Westinghouse PWR designs. A large power margin exists to predicted inception of such instabilities. Analysis has been performed and shows that minor plant-to-plant differences in Westinghouse reactor designs such as fuel assembly arrays, core power flow ratios, fuel assembly length, etc., will not result in gross deterioration of the above power margins.
4.4.3.17 Temperature Transient Effects Analysis Waterlogging damage of a fuel rod could occur as a consequence of a power increase on a rod after water has entered the fuel rod through a cladding defect and will continue until the fuel rod internal pressure equals the reactor coolant pressure. A subsequent power increase raises the temperature and, hence, could raise the pressure of the water contained within the fuel rod. Zirconium alloy-clad fuel rods, which have failed due to waterlogging (References 76 and 77) indicate that very rapid power transients are required for fuel failure. Normal operational transients are limited to about 40 cal/gm-min. (peak rod) while the Spert tests (Reference 76) indicate that 120 to 150 cal/gm is required to rupture the clad even with very short transients (5.5 msec. period).
Release of the internal fuel rod pressure is expected to have a minimal effect on the 4.4-27 Revision 24 September 2018
DCPP UNITS 1 & 2 FSAR UPDATE RCS (Reference 76) and is not expected to result in failure of additional fuel rods (Reference 77). Ejection of fuel pellet fragments into the coolant stream is not expected (References 76 and 77). A cladding breach due to waterlogging is thus expected to be similar to any fuel rod failure mechanism that exposes fuel pellets to the reactor coolant stream. Waterlogging has not been identified as the mechanism for cladding distortion or perforation of any Westinghouse Zirconium alloy-clad rods.
An excessively high fuel rod internal gas pressure could cause cladding failure. During operational transients, fuel rod cladding rupture due to high internal gas pressure is precluded by adopting a design basis that the fuel rod internal gas pressure remains below the value that causes the fuel-cladding diametral gap to increase due to outward cladding creep.
4.4.3.18 Potentially Damaging Temperature Effects During Transients A fuel rod experiences many operational transients (intentional maneuvers) while in the core. Several thermal effects must be considered when designing and analyzing fuel rod performance.
The cladding can be in contact with the fuel pellet at some time in the fuel lifetime.
Cladding-pellet interaction occurs if fuel pellet temperature is increased after the cladding is in contact with the pellet. Cladding-pellet interaction is discussed in Section 4.2.1.2.3.
Increasing fuel temperature results in an increased fuel rod internal pressure. One of the fuel rod acceptance criteria is that the fuel rod internal pressures remain below values that can cause the fuel-cladding diametral gap to increase due to outward cladding creep (refer to Section 4.2.1.2.1).
The potential effects of operation with waterlogged fuel were discussed in Section 4.4.3.17, which concluded that waterlogging is not a concern during operational transients.
Clad flattening, as noted in Section 4.2.1.2.3, has been observed in some operating power reactors. Thermal expansion (axial) of the fuel rod stack against a flattened section of cladding could cause cladding failure. This is no longer a concern because clad flattening is precluded by pre-pressurization.
A differential thermal expansion between the fuel rods and the guide thimbles can occur during a transient. Excessive bowing of fuel rods can occur if the grid assemblies do not allow axial movement of the fuel rods relative to the grids. Thermal expansion of fuel rods is considered in the grid design so that axial loads imposed on the fuel rods during a thermal transient will not result in excessively bowed fuel rods (refer to Section 4.2.1.3.2).
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DCPP UNITS 1 & 2 FSAR UPDATE 4.4.3.19 Energy Release During Fuel Element Burnout As discussed in Section 4.4.3.14, the core is protected from going through DNB over the full range of possible operating conditions. At full power operation, the typical minimum DNBR was calculated for VANTAGE+ fuel for Unit 1 and for Unit 2 and is listed in Table 4.1-1. This means that, for these conditions, the probability of a rod going through DNB is less than 0.1 percent at 95 percent confidence level based on the statistics of the WRB-2 correlations (References 84 and 85). In the extremely unlikely event that DNB should occur, cladding temperature will rise due to steam blanketing the rod surface and the consequent degradation in heat transfer. During this time a potential for a chemical reaction between the cladding and the coolant exists. Because of the relatively good film boiling heat transfer following DNB, the energy release from this reaction is insignificant compared to the power produced by the fuel.
These results have been confirmed in DNB tests conducted by Westinghouse (References 66 and 78).
4.4.3.20 Energy Release During Rupture of Waterlogged Fuel Elements A full discussion of waterlogging including energy release is contained in Section 4.4.3.17.
4.4.3.21 Fuel Rod Behavior Effects from Coolant Flow Blockage Coolant flow blockage can occur within the coolant channels of a fuel assembly or external to the reactor core. The effect of coolant flow blockage within the fuel assembly on fuel rod behavior is more pronounced than external blockages of the same magnitude. In both cases, the flow blockages cause local reductions in coolant flow.
The amount of local flow reduction, its location in the reactor, and how far downstream reduction persists, are considerations that influence fuel rod behavior. Coolant flow blockage effects in terms of maintaining rated core performance are determined both by analytical and experimental methods. The experimental data are usually used to augment analytical tools such as the THINC-IV program. Inspection of the DNB correlation (refer to Section 4.4.3.3) shows that the predicted DNBR depends on local values of quality and mass velocity.
The THINC-IV code can predict the effects of local flow blockages on DNBR within the fuel assembly on a subchannel basis, regardless of where the flow blockage occurs.
THINC-IV accurately predicts the flow distribution within the fuel assembly when the inlet nozzle is completely blocked (Reference 59). For the DCPP reactors operating at nominal full power conditions as specified in Table 4.1-1, the effects of an increase in enthalpy and decrease in mass velocity in the lower portion of the fuel assembly would not result in the reactor reaching the safety limit DNBR.
The analyses, which assume fully developed flow along the full channel length, show that a reduction in local mass velocity greater than approximately 53 percent would be 4.4-29 Revision 24 September 2018
DCPP UNITS 1 & 2 FSAR UPDATE required before the safety limit DNBRs are reached. In reality, a local flow blockage is expected to promote turbulence and thus would likely not affect DNBR.
Coolant flow blockages induce local cross flows as well as promoting turbulence. Fuel rod vibration could occur, caused by this cross flow component, through vortex shedding or turbulent mechanisms. If the cross flow velocity exceeds the limit established for fluid elastic stability, large amplitude whirling will result in, and can lead to, mechanical wear of the fuel rods at the grid support locations. The limits for a controlled vibration mechanism are established from studies of vortex shedding and turbulent pressure fluctuations. Fuel rod wear due to flow-induced vibration is considered in the fuel rod fretting evaluation (refer to Section 4.2.1.3.2.7).
4.4.4 THERMAL AND HYDRAULIC DESIGN EVALUATION 4.4.4.1 Departure from Nucleate Boiling DNB will not occur on at least 95 percent of the limiting fuel rods during normal operation and operational transients and any transient conditions arising from faults of moderate frequency (Condition I and Condition II events) at a 95 percent confidence level.
This criterion has been conservatively met by adhering to the following thermal design basis: there must be at least a 95 percent probability that the minimum DNBR of the limiting power rod during Condition I and II events is greater than or equal to the DNBR limit of the DNB correlation being used. The DNBR limit for the correlation is established based on the variance of the correlation such that there is a 95 percent probability with 95 percent confidence that DNB will not occur when the calculated DNBR is at the DNBR limit.
Historically, this DNBR limit has been 1.30 for Westinghouse applications. In this application the WRB-1 correlation (Reference 84) for LOPAR fuel and the WRB-2 correlation (Reference 85) for VANTAGE 5 fuel are employed (refer to Section 15.1.4.2). With the significant improvement in the accuracy of the CHF prediction by using these correlations instead of previous DNB correlations, a DNBR limit of 1.17 is applicable in this application. For 17 x 17 VANTAGE+ fuel, a DNBR limit of 1.17 is applicable to thermal-hydraulic analyses performed with the WRB-2 correlation.
The design method employed to meet the DNB design basis is the ITDP (Reference 86). Uncertainties in plant operating parameters, nuclear and thermal parameters, and fuel fabrication parameters are considered statistically such that there is at least a 95 percent probability that the minimum DNBR will be greater than or equal to 1.17 for the limiting power rod. Plant parameter uncertainties are used to determine the plant DNBR uncertainties. These DNBR uncertainties, combined with the DNBR limit, establish a design DNBR value, which must be met in plant safety analyses. Since the parameter uncertainties are considered in determining the design DNBR value, the plant safety analyses are performed using values of input parameters without 4.4-30 Revision 24 September 2018
DCPP UNITS 1 & 2 FSAR UPDATE uncertainties. Table 4.4-4 lists the Chapter 15 non-LOCA accident analyses and the applicable DNB design method.
This design procedure is illustrated in Figure 4.4-18. For this application, the minimum required DNBR design limits for the VANTAGE+ fuel analysis are 1.32 for thimble cells (three fuel rods and a thimble tube) and 1.34 for typical cells (four fuel rods).
In addition to the above considerations, a plant-specific DNBR margin has been considered in the analyses. In particular, safety analysis DNBR limits of 1.68 for thimble cells and 1.71 for typical cells for the VANTAGE+ fuel, were employed in the safety analyses. The plant allowance available between the DNBRs used in the safety analyses and the design DNBR values is not required to meet the design basis discussed earlier. This allowance is used for the flexibility in the design, operation, and analyses of DCPP.
By preventing DNB, adequate heat transfer is ensured between the fuel cladding and the reactor coolant, thereby preventing cladding damage. Maximum fuel rod surface temperature is not a design basis because it will be within a few degrees of coolant temperature during operation in the nucleate boiling region. Limits provided by the nuclear control and protection systems are such that this design basis will be met for transients associated with Condition II events including overpower transients. The DNBR margin at rated power operation and during normal operating transients is substantially larger (refer to Table 4.1-1).
4.4.4.2 Fuel Temperature During Conditions I and II events, the maximum fuel temperature shall be less than the melting temperature of UO2. The UO2 melting temperature for at least 95 percent of the peak kW/ft fuel rods will not be exceeded at the 95 percent confidence level.
The melting temperature of UO2 is taken as 5080°F (Reference 1) unirradiated, and decreasing 58°F per 10,000 megawatt days per metric ton of uranium (MWD/MTU). By precluding UO2 melting, the fuel geometry is preserved and possible adverse effects of molten UO2 on the cladding are eliminated. To preclude center melting and establish overpower protection system setpoints, a calculated centerline fuel temperature of 4700°F has been selected as the overpower limit, thus providing sufficient margin for uncertainties. The peak linear power value used in the design evaluation is 22.0 kW/ft.
This value corresponds to a peak centerline temperature which is less than 4700°F.
Fuel rod thermal evaluations are performed at rated power, maximum overpower, and during transients at various burnups. These analyses ensure that this design basis, as well as the fuel integrity design bases given in Section 4.2, are met. They also provide input for the evaluation of Conditions III and IV faults given in Chapter 15.
4.4-31 Revision 24 September 2018
DCPP UNITS 1 & 2 FSAR UPDATE 4.4.4.3 Core Flow A minimum of 92.5 percent (Unit 1) and 91 percent (Unit 2) of the thermal flowrate (refer to Section 5.1) will pass through the fuel rod region of the core and be effective for fuel rod cooling. Coolant flow through the thimble tubes, as well as leakage from the core barrel-baffle region into the core, is not effective for heat removal.
Core cooling evaluations are based on the thermal flowrate (minimum flow) entering the reactor vessel. On Unit 1 the design core bypass flow is 7.5%, which accounts for thimble plugs removed, intermediate flow mixing vanes, and allowance for possible future upflow conversion. On Unit 2 the design core bypass flow is 9.0%, which accounts for thimble plugs removed, intermediate flow mixing vanes, upflow conversion and upper head temperature reduction.
4.4.4.4 Hydrodynamic Stability Operation during Condition I and Condition II events shall not lead to hydrodynamic instability.
4.4.5 SAFETY EVALUATION 4.4.5.1 General Design Criterion 10, 1971 - Reactor Design The preceding evaluations show that the SAFDLs are not exceeded during any condition of normal operation or anticipated operational occurrence. Specifically:
Section 4.4.4.1 shows that the DNB limits are not exceeded; Section 4.4.4.2 shows that the maximum fuel temperature limits are not exceeded; and Section 4.4.4.3 shows that the minimum core flow is met.
The above design bases, together with the fuel cladding and fuel assembly design bases given in Section 4.2.1.1, are sufficient. Fuel cladding integrity criteria cover possible effects of cladding temperature limitations. As noted in Section 4.2.1.2.3, the fuel rod conditions change with time. A single cladding temperature limit for Conditions I or II events is not appropriate since of necessity it would be overly conservative. A cladding temperature limit is applied to the LOCA (refer to Section 15.4.1), control rod ejection accident (Reference 2), and locked rotor accident (Reference 67).
4.4.5.2 General Design Criterion 12, 1971 - Suppression of Reactor Power Oscillations Thermo-hydrodynamic instabilities (i.e., oscillations) will not occur under Condition I and Condition II events for DCPP. A large power margin exists to predicted inception of 4.4-32 Revision 24 September 2018
DCPP UNITS 1 & 2 FSAR UPDATE such instabilities. Analyses have been performed and show that in Westinghouse reactor designs, parameters such as fuel assembly arrays, core power flow ratios, fuel assembly length, etc., do not result in gross deterioration of the above power margins (refer to Section 4.4.3.16).
4.4.6 TESTS AND INSPECTIONS 4.4.6.1 Testing Prior to Initial Criticality Reactor coolant flow tests, as noted in Tests 3.9 and 3.10 of Table 14.1-2, are performed following fuel loading, but prior to initial criticality. Coolant loop pressure drop data are obtained in this test. These data, in conjunction with coolant pump performance information, allow determination of the coolant flowrates at reactor operating conditions. This test verifies that proper coolant flowrates have been used in the core thermal and hydraulic analysis.
4.4.6.2 Initial Power Plant Operation Core power distribution measurements are made at several core power levels (refer to Section 4.3.3.2.7) during startup and initial power operation. These tests are used to verify that conservative peaking factors were used in the core thermal and hydraulic design and analysis.
4.4.6.3 Component and Fuel Inspections Inspections performed on the manufactured fuel are delineated in Section 4.2.1.6.
Fabrication measurements critical to thermal and hydraulic analysis are obtained to verify that the engineering hot channel factors employed in the design analyses (refer to Section 4.4.3.3.4) are met.
4.4.7 INSTRUMENTATION APPLICATIONS 4.4.7.1 Incore Instrumentation Instrumentation is located in the core so that by correlating movable neutron detector information with fixed thermocouple information the radial core characteristics may be obtained for all core quadrants.
The incore instrumentation system is comprised of thermocouples, positioned to measure fuel assembly coolant outlet temperatures at preselected positions, and movable fission chamber detectors positioned in guide thimbles that run the length of selected fuel assemblies to measure the neutron flux distribution. Figures 4.4-16 and 4.4-17 show the number and location of instrumented assemblies in the core for Unit 1 and Unit 2, respectively. In the Unit 1 reactor, four of these thermocouples have been moved to the upper head for monitoring conditions in the upper head.
4.4-33 Revision 24 September 2018
DCPP UNITS 1 & 2 FSAR UPDATE The core exit thermocouples provide a backup for the flux monitoring instrumentation to monitor power distribution. The routine, systematic collection of thermocouple readings provides a data base. From this data base, abnormally high or abnormally low readings, quadrant temperature tilts, or systematic departures from a prior reference map can be deduced.
The movable incore neutron detector system is used for more detailed mapping should the thermocouple system indicate an abnormality. These two complementary systems are more useful when taken together than taken alone. The incore instrumentation system is described in more detail in Section 7.7.2.9.
Incore instrumentation is provided to obtain data from which fission power density distribution in the core, coolant enthalpy distribution in the core, and fuel burnup distribution may be determined.
4.4.7.2 Overtemperature and Overpower T Instrumentation The overtemperature T trip protects the core against low DNBR. The overpower T trip protects against excessive power (fuel rod rating protection).
As discussed in Section 7.2.2.1.2, factors included in establishing the overtemperature T and overpower T trip setpoints include the reactor coolant temperature in each loop. The axial distribution of core power, as determined by the two-section (upper and lower) excore neutron detectors, is also a factor in establishing the overtemperature T trip.
4.4.7.3 Instrumentation to Limit Maximum Power Output The output of the three ranges (source, intermediate, and power) of detectors, with the associated nuclear instrumentation electronics, is used to limit the maximum power output of the reactor.
Eight instrument wells are located around the reactor periphery in the primary shield, 45° apart from each other, at an equal distance from the reactor vessel.
Two of the positions, on opposite flat portions of the core, directly across from the secondary neutron source positions, each contain a BF3 proportional counter to cover the source range, and a compensated ionization chamber for the intermediate range.
The source range detector is located at an elevation of approximately one-fourth of the core height; the compensated ionization chambers are positioned at an elevation corresponding to one-half of the core height. The two positions opposite the other two flat portions of the core house the post-accident neutron flux monitor detectors.
Four dual-section uncompensated ionization chamber assemblies are installed vertically in the instrumentation wells directly across from the four corners of the core.
They are used as power range detectors. To minimize neutron flux pattern distortions, vessel. Each dual-section they are placed within 1 foot of the reactor4.4-34 uncompensated Revision 24 September 2018 ionization
DCPP UNITS 1 & 2 FSAR UPDATE chamber assembly provides two signals that correspond to the neutron flux in the upper and in the lower positions of a core quadrant, thus permitting the determination of relative axial power production.
Signals from the detectors in the three ranges (source, intermediate, and power) provide inputs which, when combined, monitor neutron flux from a completely shutdown condition to 120 percent of full power, with the capability of recording overpower excursions up to 200 percent of full power.
The difference in neutron flux readings between the upper and lower sections of the power range detectors is used to limit the overtemperature T and overpower T trip setpoints and to provide the operator with an indication of the core power AO. In addition, the output of the power range channels is used as follows:
(1) For the rod speed control function (2) To alert the operator to an excessive power unbalance between the quadrants (3) To protect the core against rod ejection accidents (4) To protect the core against adverse power distributions resulting from dropped rods Details of the neutron detectors and nuclear instrumentation design and the control and trip logic are given in Chapter 7. The limits on neutron flux operation and trip setpoints are given in the Technical Specifications.
4.
4.8 REFERENCES
- 2. D. H. Risher, Jr., An Evaluation of the Rod Ejection Accident in Westinghouse Pressurized Water Reactors Using Spatial Kinetics Methods, WCAP-7588, Revision 1, December 1971.
- 3. G. Kjaerheim and E. Rolstad, In Pile Determination of UO2 Thermal Conductivity, Density Effects and Gap Conductance, HPR-80, December 1967.
- 4. G. Kjaerheim, In-Pile Measurements of Centre Fuel Temperatures and Thermal Conductivity Determination of Oxide Fuels, paper IFA-175 presented at the European Atomic Energy Society Symposium on Performance Experience of Water-Cooled Power Reactor Fuel, Stockholm, Sweden, October 1969.
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- 5. I. Cohen, et al, Measurement of the Thermal Conductivity of Metal-Clad Uranium Oxide Rods during Irradiation, WAPD-228, 1960.
- 6. D. J. Clough and J. B. Sayers, The Measurement of the Thermal Conductivity of UO2 under Irradiation in the Temperature Range 150°F-1600°C, AERE-R-4690, UKAEA Research Group, Harwell, December 1964.
- 7. J. P. Stora, et al, Thermal Conductivity of Sintered Uranium Oxide Under In-Pile Conditions, EURAEC-1095, August 1964.
- 8. I. Devold, A Study of the Temperature Distribution in UO2 Reactor Fuel Elements, AE-318, Aktiebolaget Atomenergi, Stockholm, Sweden, 1968.
- 9. M. G. Balfour, et al, In-Pile Measurement of UO2 Thermal Conductivity, WCAP-2923, March 1966.
- 10. R. N. Duncan, Rabbit Capsule Irradiation of UO2, CVTR Project, CVNA-142, June 1962.
- 11. R. C. Nelson, et al, Fission Gas Release from UO2 Fuel Rods with Gross Central Melting, GEAP-4572, July 1964.
- 12. V. C. Howard and T. G. Gulvin, Thermal Conductivity Determinations on Uranium Dioxide by a Radial Flow Method, UKAEA IG-Report 51, November 1960.
- 13. C. F. Lucks and H. W. Deem, "Thermal Conductivity and Electrical Conductivity of UO2," in Progress Reports Relating to Civilian Applications, BMI-1448 (Revision) for June 1960, BMI-1489 (Revision) for December 1960 and BMI-1518 (Revision) for May 1961.
- 14. J. L. Daniel, et al, Thermal Conductivity of UO2, HW-69945, September 1962.
- 15. A. D. Feith, Thermal Conductivity of UO2 by a Radial Heat Flow Method, TID-21668, 1962.
- 16. J. Vogt, et al, Determination of the Thermal Conductivity of Unirradiated Uranium Dioxide, AB Atomenergi Report RMB-527, 1964, Quoted by IAEA Technical Report Series No. 59, "Thermal Conductivity of Uranium Dioxide."
- 17. T. Nishijima, et al, "Thermal Conductivity of Sintered UO2 and Al2O3 at High Temperatures," J. American Ceramic Society, 48, 1965, pp. 31-34.
- 18. J. B. Ainscough and M. J. Wheeler, "The Thermal Diffusivity and Thermal Conductivity of Sintered Uranium Dioxide," in Proceedings of the Seventh Conference on Thermal Conductivity, National Bureau of Standards, Washington, D.C., 1968, p. 467.
4.4-36 Revision 24 September 2018
- 19. T. G. Godfrey, et al, Thermal Conductivity of Uranium Dioxide and Armco Iron by an Improved Radial Heat Flow Technique, ORNL-3556, June 1964.
- 20. Deleted.
- 21. A. J. Bush, Apparatus for Measuring Thermal Conductivity to 2500°C, Westinghouse Research Laboratories Report 64-1P6-401-R3, February 1965.
- 22. R. R. Asamoto, et al, The Effect of Density on the Thermal Conductivity of Uranium Dioxide, GEAP-5493, April 1968.
- 23. O. L. Kruger, Heat Transfer Properties of Uranium and Plutonium Dioxide, Paper 11-N-68F presented at the Fall meeting of Nuclear Division of the American Ceramic Society, September 1968, Pittsburgh.
- 24. J. A. Gyllander, In-Pile Determination of the Thermal Conductivity of UO2 in the Range 500-2500°C, AE-411, January 1971.
- 25. M. F. Lyons, et al, UO2 Powder and Pellet Thermal Conductivity During Irradiation, GEAP-5100-1, March 1966.
- 26. D. H. Coplin, et al, The Thermal Conductivity of UO2 by Direct In-reactor Measurements, GEAP-5100-6, March 1968.
- 27. A. S. Bain, "The Heat Rating Required to Produce Center Melting in Various UO2 Fuels," ASTM Special Technical Publication, No. 306, Philadelphia, 1962, pp. 30-46.
- 28. J. P. Stora, "In-Reactor Measurements of the Integrated Thermal Conductivity of UO2 - Effects of Porosity," Trans. ANS, 13, 1970, pp. 137-138.
- 29. International Atomic Energy Agency, "Thermal Conductivity of Uranium Dioxide,"
Report of the Panel held in Vienna, April 1965, IAEA Technical Reports Series, No. 59, Vienna, The Agency, 1966.
- 30. C. G. Poncelet, Burnup Physics of Heterogeneous Reactor Lattices, WCAP-6069, June 1965.
- 31. R. J. Nodvick, Saxton Core II Fuel Performance Evaluation, WCAP-3385-56, Part II, Evaluation of Mass Spectrometric and Radiochemical Analyses of Irradiated Saxton Plutonium Fuel, July 1970.
- 32. R. A. Dean, Thermal Contact Conductance Between UO2 and Zircaloy-2, CVNA-127, May 1962.
4.4-37 Revision 24 September 2018
- 33. A. M. Ross and R. L. Stoute, Heat Transfer Coefficient Between UO2 and Zircacloy-2, AECL-1552, June 1962.
- 34. L. S. Tong, Boiling Crisis and Critical Heat Flux, AEC Critical Review Series, TID-25887, 1972.
- 35. F. E. Motley and F. F. Cadek, DNB Tests Results for New Mixing Vane Grids (R),
WCAP-7695 -P-A, January 1975 (Westinghouse Proprietary) and WCAP-7958-A, January 1975.
- 36. F. E. Motley, and F. F. Cadek, Application of Modified Spacer Factor to L. grid Typical and Cold Wall Cell DNB, WCAP-7988-P-A, January 1975 (Westinghouse Proprietary) and WCAP-8030-A, January 1975.
- 37. F. E. Motley, et al, Critical Heat Flux Testing of 17 x 17 Fuel Assembly Geometry with 22-Inch Grid Spacing, WCAP-8536, May 1975 (Westinghouse Proprietary) and WCAP-8537, May 1975.
- 38. K.W. Hill, et al, Effect of 17 x 17 Fuel Assembly Geometry on DNB, WCAP-8296-P-A, February 1975 (Westinghouse Proprietary) and WCAP-8297-A, February 1975.
- 39. L. S. Tong, "Prediction of Departure from Nucleate Boiling for an Axially Non-Uniform Heat Flux Distribution," J. Nucl. Energy, 21, 1967, pp. 241-248.
- 40. F. F. Cadek, et al, Effect of Axial Spacing on Interchannel Thermal Mixing with The R Mixing Vane Grid, WCAP-7941-P-A, January 1975 (Westinghouse Proprietary) and WCAP-7959-A, January 1975.
- 41. H. Chelemer, et al, Subchannel Thermal Analysis of Rod Bundle Core, WCAP-7015, Revision 1, January 1969.
- 42. D. S. Rowe and C. W. Angle, Crossflow Mixing Between Parallel Flow Channels During Boiling, Part II Measurement of Flow and Enthalpy in Two Parallel Channels, BNWL-371, Part 2, December 1967.
- 43. D. S. Rowe and C. W. Angle, Crossflow Mixing Between Parallel Flow Channels During Boiling, Part III Effect of Spacers on Mixing Between Two Channels, BNWL-371, Part 3, January 1969.
- 44. J. M. Gonzalez-Santalo and P. Griffith, Two-Phase Flow Mixing in Rod Bundle Subchannels, ASME Paper 72-WA/NE-19, 1972.
- 45. F. E. Motley, et al, The Effect of 17 x 17 Fuel Assembly Geometry on Interchannel Thermal Mixing, WCAP-8298-P-A, January 1975 (Westinghouse Proprietary) and WCAP-8299-A, January 1975.
4.4-38 Revision 24 September 2018
- 46. F. F. Cadek, Interchannel Thermal Mixing with Mixing Vane Grids, WCAP-7667-P-A, January 1975 (Westinghouse Proprietary), and WCAP-7755-A, January 1975.
- 47. L. E. Hochreiter, Application of the THINC-IV Program to PWR Design, WCAP-8054-P-A, February 1989 (Westinghouse Proprietary), and WCAP-8195-A, February 1989.
- 48. S. Nakazato and E. E. DeMario, Hydraulic Flow Test of the 17 x 17 Fuel Assembly, WCAP-8279, February 1974.
- 49. F. W. Dittus and L. M. K. Boelter, "Heat Transfer in Automobile Radiators of the Tubular Type," Calif. Univ. Publication in Eng., 2, No. 13, 1930, pp. 443-461.
- 50. J. Weisman, "Heat Transfer to Water Flowing Parallel to Tube Bundles,"
Nucl. Sci. Eng., 6, 1959, pp. 78-79.
- 51. J. R. S. Thom, et al, "Boiling in Sub-cooled Water During Flowup Heated Tubes or Annuli," Proc. Instn. Mech. Engrs., 180, Pt. C, 1965-66, pp. 226-46.
- 52. G. Hetsroni, Hydraulic Tests of the San Onofre Reactor Model, WCAP-3269-8, June 1964.
- 53. G. Hetsroni, Studies of the Connecticut-Yankee Hydraulic Model, NYO-3250-2, June 1965.
- 54. I. E. Idel'chik, Handbook of Hydraulic Resistance, AEC-TR-6630, 1960.
- 55. L. F. Moody, "Friction Factors for Pipe Flow," Transaction of the American Society of Mechanical Engineers, 66, 1944, pp. 671-684.
- 56. G. W. Maurer, A Method of Predicting Steady State Boiling Vapor Fractions in Reactor Coolant Channels, WAPD-BT-19, June 1960, pp. 59-70.
- 57. P. Griffith, et al, Void Volumes in Subcooled Boiling Systems, ASME Paper No. 58-HT-19, 1958.
- 58. R. W. Bowring, Physical Model, Based on Bubble Detachment, and Calculation of Steam Voidage in the Subcooled Region of a Heated Channel, HPR-10, December 1962.
- 59. L. E. Hochreiter, et al, THINC-IV An Improved Program for Thermal-Hydraulic Analysis of Rod Bundle Cores, WCAP-7956-A, February 1989.
4.4-39 Revision 24 September 2018
- 60. F. D. Carter, Inlet Orificing of Open PWR Cores, WCAP-9004, January 1969, (Westinghouse Proprietary), and WCAP-7836, January 1972.
- 61. J. Shefcheck, Application of the THINC Program to PWR Design, WCAP-7359-L, August 1969 (Westinghouse Proprietary), and WCAP-7838, January 1972.
- 62. E. H. Novendstern and R. O. Sandberg, Single Phase Local Boiling and Bulk Boiling Pressure Drop Correlations, WCAP-2850, April 1966 (Westinghouse Proprietary), and WCAP-7916, April 1966.
- 63. W. L. Owens, Jr., "Two-Phase Pressure Gradient," International Developments in Heat Transfer, Part II, ASME, New York, 1961, pp. 363-368.
- 64. A. F. McFarlane, Power Peaking Factors, WCAP-7912-P-A, January 1975 (Westinghouse Proprietary) and WCAP-7912-A, January 1975.
- 65. Deleted.
- 66. J. Weisman, et al, "Experimental Determination of the Departure from Nucleate Boiling in Large Rod Bundles at High Pressures," Chem. Eng. Prog. Symp.
Ser. 64, No. 82, 1968, pp. 114-125.
- 67. Supplemental information on fuel design transmitted from R. Salvatori, Westinghouse NES, to D. Knuth, AEC, as attachments to Letters NS-SL-518 (12/22/72), NS-SL-521 (1/4/73), NS-SL-524 (1/4/73), and NS-SL-543 (1/12/73) (Westinghouse Proprietary), and supplemental information on fuel design transmitted from R. Salvatori, Westinghouse NES, to D. Knuth, AEC, as attachments to Letters NS-SL-527 (1/4/73), and NS-SL-544 (1/12/73).
- 68. J. M. Hellman (ed), Fuel Densification Experimental Results and Model for Reactor Application, WCAP-8218-P-A, March 1975 (Westinghouse Proprietary) and WCAP-8219-A, March 1975.
- 69. J. A. Boure, et al, Review of Two-Phase Flow Instability, ASME Paper 71-HT-42, August 1971.
- 70. R. T. Lahey and F.J. Moody, The Thermal Hydraulics of a Boiling Water Reactor, American Nuclear Society, 1977.
- 71. M. Ishii, et al, "An Experimental Investigation of the Thermally Induced Flow Oscillations in Two-Phase Systems," J. of Heat Transfer, November 1976, pp. 616-622.
- 72. Virgil C. Summer FSAR, Docket #50-395.
- 73. Byron/Braidwood FSAR, Docket #50-456.
4.4-40 Revision 24 September 2018
- 75. H. S. Kao, et al., "Prediction of Flow Oscillation in Reactor Core Channel,"
Trans. ANS. Vol. 16, 1973, pp. 212-213.
- 76. L. A. Stephan, The Effects of Cladding Material and Heat Treatment on the Response of Waterlogged UO2 Fuel Rods to Power Bursts, IN-ITR-111, January 1970.
- 77. Western New York Nuclear Research Center Correspondence with the AEC on February 11 and August 27, 1971, Docket 50-57.
- 78. L. S. Tong, et al., Critical Heat Flux (DNB) in Square and Triangular Array Rod Bundles, presented at the Japan Society of Mechanical Engineers Semi-International Symposium held at Tokyo, Japan, September 1967, pp. 25-34.
- 79. J. Skaritka, (Ed.), Fuel Rod Bow Evaluation, WCAP-8691, Revision 1 July 1979.
- 80. "Partial Response to Request Number 1 for Additional Information on WCAP-8691, Revision 1" Letter, E. P. Rahe, Jr., (Westinghouse) to J. R. Miller (NRC), NS-EPR-2515, dated October 1981; "Remaining Response to Request Number 1 for Additional Information on WCAP-8691, Revision 1" Letter, E. P. Rahe, Jr., (Westinghouse) to J.R. Miller (NRC), NS-EPR-2572, dated March 1982.
- 81. C. G. Poncelet, LASER - A Depletion Program for Lattice Calculations Based on MUFT and THERMOS, WCAP-6073, April 1966.
- 82. Technical Specifications, Diablo Canyon Power Plant Units 1 and 2, Appendix A to License Nos. DPR-80 and DPR-82, as amended.
- 83. Deleted in Revision 23.
- 84. F. E. Motley, K. W. Hill, F. F. Cadek and J. Shefcheck, New Westinghouse Correlation WRB-1 for Predicting Critical Heat Flux in Rod Bundles with Mixing Vane Grids, WCAP-8762-P-A, July 1984 (Westinghouse Proprietary) and WCAP-8763-A, July 1984.
- 85. S. L. Davidson and W. R. Kramer, (Ed.), Reference Core Report VANTAGE 5 Fuel Assembly, WCAP-10444-P-A, September 1985 (Westinghouse Proprietary) and WCAP-10445-NP-A, September 1985.
4.4-41 Revision 24 September 2018
- 86. H. Chelemer, L. H. Boman and D. R. Sharp, Improved Thermal Design Procedure, WCAP-8567-P-A, February 1989 Westinghouse Proprietary) and WCAP-8568-NP-A (Non-Proprietary), February 1989.
- 87. S. L. Davidson, F. E. Motley, Y. C. Lee, T. Bogard and W. J. Bryan, Verification Testing and Analyses of the 17x17 Optimized Fuel Assembly, WCAP-9401-P-A, August 1981 (Westinghouse Proprietary) and WCAP-9402-A, August 1981.
- 88. Deleted.
- 89. S. L. Davidson, J. A. Iorii, (Ed.) Reference Core Report - 17x17 Optimized Fuel Assembly, WCAP-9500-A, May 1982.
- 90. Letter from E. P. Rahe (Westinghouse) to Miller (NRC), NS-EPR-2573, WCAP-9500 and WCAPs 9401/9402 NRC SER Mixed Core Compatibility Items, March 1982.
- 91. Letter from C. O. Thomas (NRC) to Rahe (Westinghouse) - Supplemental Acceptance No. 2 for Referencing Topical Report, WCAP-9500, January 1983.
- 92. F. E. Motley and F. F. Cadek, DNB Test Results for R Grid Thimble Cold Wall Cells, WCAP-7695-P-A, Addendum 1, January 1975 (Westinghouse Proprietary) and WCAP-7958-A1-A, January 1975.
- 93. Deleted.
- 94. Deleted.
- 95. Deleted.
- 96. Deleted.
- 97. Deleted.
- 98. Deleted.
- 99. K. W. Hill, et al., Effect of Local Heat Flux Spikes on DNB in Non Uniformly Heated Rod Bundles, WCAP-8174-P-A, February 1975 (Westinghouse Proprietary) and WCAP-8202-A, February 1975.
100. S. L. Davidson (Ed.), et al., VANTAGE+ Fuel Assembly Reference Core Report, WCAP-12610-P-A, April 1995.
101. J.P. Foster, et al., Westinghouse Improved Performance Analysis and Design Model (PAD 4.0), WCAP-15063-P-A, Revision 1, with Errata, July 2000.
4.4-42 Revision 24 September 2018
DCPP UNITS 1 & 2 FSAR UPDATE TABLE 4.1-1 Sheet 1 of 7 REACTOR DESIGN COMPARISON Thermal and Hydraulic Design Parameters Unit 1 Unit 2 (Using ITDP) (a)
Reactor Core Heat Output, MWt 3,411 3,411 Reactor Core Heat Output, 10 Btu/hr 6
11,639 11,639 Heat Generated in Fuel, % 97.4 97.4 Reactor Coolant Pressure, psia 2,250 2,250 Fuel Type Vantage+ Vantage+
Design DNBR Limit Typical Cell 1.34 1.34 Thimble Cell 1.32 1.32 Safety Analysis DNBR Limit Typical Cell 1.71 1.71 Thimble Cell 1.68 1.68 DNB Correlation WRB-2 WRB-2 Revision 23 December 2016
DCPP UNITS 1 & 2 FSAR UPDATE TABLE 4.1-1 Sheet 2 of 7 Unit 1 Unit 2 Vessel Minimum Measured Flow Rate (including Bypass) gpm 359,200 362,500 Vessel Thermal Design Flow Rate (e)
(including Bypass) gpm 350,800 354,000 Thermal and Hydraulic Design Parameters (Based on Thermal Design Flow)
Core Flow Rate (excluding Bypass) 10 lbm/hr 6
122.3 123.4 gpm 324,490 327,450 Effective Core Flow Area for Heat Transfer, 54.13 54.13 ft2 Average Velocity along Fuel (k)
Rods, ft/sec 14.0 14.2 Core Inlet Mass Velocity, 10 lbm/hr-ft 6
(V-5) 2.26 2.28 Revision 23 December 2016
DCPP UNITS 1 & 2 FSAR UPDATE TABLE 4.1-1 Sheet 3 of 7 Unit 1 Unit 2 Nominal Vessel/Core Inlet Temperature, °F 531.7 - 544.5 531.9 -545.1 Vessel Average Temperature, °F 565.0 - 577.3 565.0 - 577.6 Core Average Temperature, °F 569.1 - 581.5 569.6 - 582.3 Vessel Outlet Temperature, °F 598.3 - 610.1 598.1 - 610.1 Average Temperature Rise in Vessel, °F 66.6 - 65.6 66.2 - 65.0 Average Temperature Rise in Core, °F 71.6 - 70.4 72.1 - 70.7 Heat Transfer Active Heat Transfer Surface Area, ft 2 57,505 57,505 Average Heat Flux, Btu/hr-ft 2 197,180 197,180 Maximum Heat Flux for Normal (h)
Operation, Btu/hr-ft 2 508,720 508,720 Average Linear Power, kW/ft 5.445 5.445 Peak Linear Power for Normal Operation, kW/ft (h) 14.3 14.3 Peak Linear Power for Prevention of Centerline Melt, kW/ft 22.0 (i) 22.0 (i)
Pressure Drop (j)
Across Core, psi
- 25.5 + 2.6 27.2 + 2.7 Across Vessel, (n) including nozzle, psi
- 52.8 + 5.3 48.2 + 4.8 Thermal and Hydraulic Design Parameters (Fuel Rod Design)
Heat Flux Hot Channel Factor, FQ T 2.58 2.58 Temperature at Peak Linear Power for 4,700 4,700 Prevention of Centerline Melt, °F Fuel Centerline Temperature, °F Peak at 102% power <3,360 <3,360 Peak at maximum thermal output for maximum overpower DT trip point <3,770 <3,770
- Pressure drop values for mechanical design flow and low inlet temperatures of 531.7°F and 531.9°F for Unit 1 and Unit 2.
Revision 23 December 2016
DCPP UNITS 1 & 2 FSAR UPDATE TABLE 4.1-1 Sheet 4 of 7 Core Mechanical Design Parameters Unit 1 Unit 2 Fuel Assemblies Design RCC Canless RCC Canless Number of fuel assemblies 193 193 Rod array 17 X 17 17 X 17 U02 rods per assembly 264 264 Rod pitch, in 0.496 0.496 Overall dimensions, in 8.424 x 8.424 8.424 x 8.424 Fuel weight (as UO2) lb 200,720/205,159 200,720/205,159 Zirconium alloy weight, lb 51,917 51,917 Number of grids per assembly 12 12 2 non-mixing vane type 2 non-mixing vane type 6 mixing vane type 6 mixing vane type 3 IFM; 1 P-Grid 3 IFM; 1 P-Grid Composition of grids Inconel Alloy 718 or Inconel Alloy 718 or ZIRLO ZIRLO Weight of grids, lb 1841/2820 1841/2820 Number of guide thimbles per assembly 24 24 Composition of guide thimbles ZIRLO ZIRLO Diameter of guide thimbles (ID x OD), in.
Upper part 0.442 x 0.474 0.442 x 0.474 Lower part 0.397 x 0.430 0.397 x 0.430 Diameter of instrument guide thimbles, in. 0.442 x 0.474 0.442 x 0.474 Revision 23 December 2016
DCPP UNITS 1 & 2 FSAR UPDATE TABLE 4.1-1 Sheet 5 of 7 Core Mechanical Design Parameters (Cont'd) Unit 1 Unit 2 Fuel Rods Number 50,952 50,952 Outside diameter, in 0.360 0.360 Diametral gap, in 0.0062 0.0062 Cladding thickness, in 0.0225 0.0225 Cladding material ZIRLO ZIRLO Gap material Helium Helium Fuel Pellets Material UO2 sintered UO sintered 2
Density, % of theoretical 95 95 Diameter, in 0.3088 +/- 0.0005 0.3088 +/-
0.0005 Length, in Enriched or IFBA Pellets 0.370 +/- 0.025 0.370 +/- 0.025 Axial Blanket Pellets (refer to Section 0.500 +/- 0.025 0.500 +/- 0.025 4.2.1.2.2)
Mass of UO , lb/ft of fuel rod 2 0.334 0.334 Rod Cluster Control Assemblies Neutron absorber, Ag-In-Cd Ag-In-Cd Composition 80%, 15%, 5% 80%, 15%, 5%
Diameter, in 0.341 +/- 0.001* 0.341 +/- 0.001*
Nominal length of absorber material, in. 142 142 Density, lb/in3 0.367 0.367 Cladding material Framatome RCCAs Ion-nitrided, cold Ion-nitrided, worked, Type 316L cold worked, SS Type 316L SS Westinghouse RCCAs Chrome Plated, cold Chrome worked, Type 304 SS Plated, cold worked, Type 304 SS Cladding thickness, in 0.0185 0.0185 Number of RCCAs 53 53 Number of absorber rods per cluster 24 24 Core Structure Core barrel, ID/OD, in 148.0/152.5 148.0/152.5 Thermal shield, ID/OD, in 158.5/164.0 Neutron pad Design
- Diameter reduced to 0.336 +/- 0.001 over bottom 12 inches.
Revision 23 December 2016
DCPP UNITS 1 & 2 FSAR UPDATE TABLE 4.1-1 Sheet 6 of 7 Nuclear Design Parameters Unit 1 Unit 2 Structure Characteristics Core diameter, in (equivalent) 132.7 132.7 Core average active fuel height, in. 144 144 Reflector Thickness and Composition Top - water plus steel, in. ~10 ~10 Bottom - water plus steel, in. ~10 ~10 Side - water plus steel, in ~15 ~15 H2O/U, cold molecular ratio lattice 2.73 2.73 Fuel Enrichment, Wt%
Maximum feed enrichment 5.0 5.0 Burnable Absorbers Type IFBA IFBA Number (typical range) 2000 - 15000 2000 - 15000 Material ZrB2 ZrB2 B10 Loading, mg/inch (typical) 2.25 2.25 (a) Includes the effect of fuel densification (b) Deleted (c) Based on T = 545.1°F (Unit 1) and T = 545.7°F (Unit 2) corresponding to Minimum in in Measured Flow of each unit (d) Deleted (e) Includes 0 to 10 percent steam generator tube plugging (f) Deleted (g) Deleted (h) This limit is associated with the value of T = 2.58 FQ (i) Refer to Section 4.3.3.2.6 (j) Based on best estimate reactor flow rate(refer to Section 5.1)
(k) At core average temperature Revision 23 December 2016
DCPP UNITS 1 & 2 FSAR UPDATE TABLE 4.1-1 Sheet 7 of 7 (l) Deleted (m) Deleted (n) Deleted Revision 23 December 2016
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DCPP UNITS 1 & 2 FSAR UPDATE TABLE 4.1-3 DESIGN LOADING CONDITIONS FOR REACTOR CORE COMPONENTS (1) Fuel assembly weight (2) Fuel assembly spring forces (3) Internals weight (4) Control rod scram (equivalent static load)
(5) Differential pressure (6) Spring preloads (7) Coolant flow forces (static)
(8) Temperature gradients (9) Differences in thermal expansions (a) Due to temperature differences (b) Due to expansion of different materials (10) Interference between components (11) Vibration (mechanically or hydraulically induced)
(12) One or more loops out of service (13) All operational transients listed in Table 5.2-4 (14) Pump overspeed (15) Seismic loads (DE and DDE)
(16) Blowdown forces (due to RCS branch line breaks)(a)
(a) In the original analysis, the blowdown forces used were those resulting from breaks in the RCS cold and hot legs. However, with the acceptance of the DCPP leak-before-break analysis by the NRC, the blowdown forces resulting from pipe rupture events in the main reactor coolant loop piping no longer have to be considered in the design basis structural analyses and included in the loading combinations. Only the much smaller forces from RCS branch line breaks have to be considered (see Section 3.6.2.1.1.1).
Revision 12 September 1998
DCPP UNITS 1 & 2 FSAR UPDATE TABLE 4.3-1 Sheet 1 of 2 NUCLEAR DESIGN PARAMETERS (Typical)
Core Average Linear Power, kW/ft, including densification effects 5.445 (a)
Total Heat Flux Hot Channel Factor, FQT 2.58 Nuclear Enthalpy Rise Hot Channel Factor, FN 1.65 VANTAGE+
H Reactivity Coefficients Doppler coefficient Refer to 4.3-29 Moderator temperature coefficient at operating conditions, pcm/°F(b)
+5 to -40 Boron coefficient in primary coolant, pcm/ppm -16 to -8 Delayed Neutron Fraction and Lifetime eff BOL, (EOL) 0.0069, (0.0051) l BOL, (EOL), sec 19.2, (18.6)
Control Rod Worths Rod requirements Refer to Table 4.3-2 Maximum bank worth, pcm < 2000 Maximum ejected rod worth Refer to Chapter 15 Revision 23 December 2016
DCPP UNITS 1 & 2 FSAR UPDATE TABLE 4.3-1 Sheet 2 of 2 Boron Concentrations (ppm)
Refueling 2000 k = 0.95, cold, rod cluster eff control assemblies in 2000 Full power, no xenon, k = 1.0, hot, eff rod cluster control assemblies out 1876 (d)
Full power, equilibrium xenon, k = 1.0, eff hot, rod cluster control assemblies out 1536 (d)
Reduction with fuel burnup Typical reload cycle, ppm/GWD/MTU (c)
Refer to Figure 4.3-3 (a) Data in table based on Unit 1 and Unit 2 (b) 1 pcm = percent mille
= 10 where is calculated from two state point values of k by ln (k /k )
-5 eff 2 1 (c) Gigawatt day (GWD) = 1000 megawatt days (1000 MWD)
(d) These values are representative values used for analytical purposes only.
Revision 23 December 2016
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DCPP UNITS 1 & 2 FSAR UPDATE HISTORICAL INFORMATION IN ITALICS BELOW NOT REQUIRED TO BE REVISED.
TABLE 4.3-4 AXIAL STABILITY INDEX PWR CORE WITH A 12-FT HEIGHT Burnup Stability Index, hr -1 (MWD/MTU) F Z C (ppm)
B Exp. Calc.
1550 1.34 1065 -0.041 -0.032 7700 1.27 700 -0.014 -0.006 Difference: +0.027 +0.026 Revision 23 December 2016
DCPP UNITS 1 & 2 FSAR UPDATE TABLE 4.3-5 TYPICAL NEUTRON FLUX LEVELS (n/cm2-sec) AT FULL POWER E > 1 MeV 5.53 keV E 0.625 eV E E < 0.625 eV 1 MeV < 5.53 keV (nv)0 13 Core center 6.51 x 1013 1.12 x 1014 8.50 x 1013 3.00 x 10 Core outer radius 3.23 x 1013 5.74 x 1013 4.63 x 1013 8.60 x 1012 at midheight Core top, on axis 1.53 x 1013 2.42 x 1013 2.10 x 1013 1.63 x 1013 Core bottom, on axis 2.36 x 1013 3.94 x 1013 3.50 x 1013 1.46 x 1013 Pressure vessel inner wall, 2.77 x 1010 5.75 x 1010 6.03 x 1010 8.38 x 1010 azimuthal peak, core midheight Revision 11 November 1996
DCPP UNITS 1 & 2 FSAR UPDATE TABLE 4.3-6 COMPARISON OF MEASURED AND CALCULATED DOPPLER DEFECTS Core Burnup Measured Calculated Plant Fuel Type (MWD/MTU) (pcm)(a) (pcm) 1 Air filled 1800 1700 1710 2 Air filled 7700 1300 1440 3 Air and 8460 1200 1210 helium filled (a) pcm = 10-5 . See footnote in Table 4.3-1 Revision 11 November 1996
DCPP UNITS 1 & 2 FSAR UPDATE TABLE 4.3-7 BENCHMARK CRITICAL EXPERIMENTS Description of No. of LEOPARD keff Using Experiments Experiments Experimental Bucklings UO2 Al clad 14 1.0012 SS clad 19 0.9963 Borated H2O 7 0.9989 Total 40 0.9985 U-metal Al clad 41 0.9995 Unclad 20 0.9990 Total 61 0.9993 Grand Total 101 0.9990 Revision 11 November 1996
DCPP UNITS 1 & 2 FSAR UPDATE TABLE 4.3-8 SAXTON CORE II ISOTOPICS ROD MY, AXIAL ZONE 6 Atom Ratio Measured 2 Precision, % LEOPARD Calculation U-234/U 4.65 x 10-5 +/- 29 4.60 x 10-5 U-235/U 5.74 x 10-3 +/- 0.9 5.73 x 10-3 U-236/U 3.55 x 10-4 +/- 5.6 3.74 x 10-4 U-238/U 0.99386 +/- 0.01 0.99385 Pu-238/Pu 1.32 x 10-3 +/- 2.3 1.222 x 10-3 Pu-239/Pu 0.73971 +/- 0.03 0.74497 Pu-240/Pu 0.19302 +/- 0.2 0.19102 Pu-241/Pu 6.014 x 10-2 +/- 0.3 5.74 x 10-2 Pu-242/Pu 5.81 x 10-3 +/- 0.9 5.38 x 10-3 Pu/U(a) 5.938 x 10-2 +/- 0.7 5.970 x 10-2 Np-237/U-238 1.14 x 10-4 +/- 15 0.86 x 10-4 Am-241/Pu-239 1.23 x 10-2 +/- 15 1.08 x 10-2 Cm-242/Pu-239 1.05 x 10-4 +/- 10 1.11 x 10-4 Cm-244/Pu-239 1.09 x 10-4 +/- 20 0.98 x 10-4 (a) Weight ratio Revision 11 November 1996
DCPP UNITS 1 & 2 FSAR UPDATE TABLE 4.3-9 CRITICAL BORON CONCENTRATIONS, AT HZP, BOL Plant Type Measured Calculated 2-loop, 121 assemblies, 10-foot core 1583 1589 2-loop, 121 assemblies, 12-foot core 1625 1624 2-loop, 121 assemblies, 12-foot core 1517 1517 3-loop, 157 assemblies, 12-foot core 1169 1161 3-loop, 157 assemblies, 12-foot core 1344 1319 4-loop, 193 assemblies, 12-foot core 1370 1355 4-loop, 193 assemblies, 12-foot core 1321 1306 Revision 11 November 1996
DCPP UNITS 1 & 2 FSAR UPDATE TABLE 4.3-10 COMPARISON OF MEASURED AND CALCULATED ROD WORTH 2-Loop Plant, 121 Assemblies, 10-foot core Measured, pcm Calculated, pcm Group B 1885 1893 Group A 1530 1649 Shutdown group 3050 2917 ESADA Critical, 0.69-in pitch, 2 wt.% PuO2, 8% Pu240, 9 Control Rods 6.21-in rod separation 2250 2250 2.07-in rod separation 4220 4160 1.38-in rod separation 4100 4010 Revision 11 November 1996
CPP UNITS 1 & 2 FSAR UPDATE HISTORICAL INFORMATION IN ITALICS BELOW NOT REQUIRED TO BE REVISED.
TABLE 4.3-11 COMPARISON OF MEASURED AND CALCULATED MODERATOR TEMPERATURE COEFFICIENTS AT HZP, BOL Measured iso (a)
Calculated iso Plant Type/Control Bank Configuration pcm/°F pcm/°F 3-Loop, 157 Assemblies, 12-foot core D at 160 steps -0.50 -0.50 D in, C at 190 steps -3.01 -2.75 D in, C at 28 steps -7.67 -7.02 B, C, and D in -5.16 -4.45 2-Loop, 121 Assemblies, 12-foot core D at 180 steps +0.85 +1.02 D in, C at 180 steps -2.40 -1.90 C and D in, B at 165 steps -4.40 -5.58 B, C, and D in, A at 174 steps -8.70 -8.12 (a) Isothermal coefficients, which include the Doppler effect in the fuel k2
= 105 ln( k )/T(°F) 1 iso Revision 23 December 2016
CPP UNITS 1 & 2 FSAR UPDATE TABLE 4.4-1 UNIT 1 VOID FRACTIONS AT NOMINAL REACTOR CONDITIONS WITH DESIGN HOT CHANNEL FACTORS Average Maximum Core (VANTAGE+) 0.17% --
Hot subchannel (VANTAGE+) 0.89% 2.11%
Revision 23 December 2016
CPP UNITS 1 & 2 FSAR UPDATE TABLE 4.4-2 UNIT 2 VOID FRACTIONS AT NOMINAL REACTOR CONDITIONS WITH DESIGN HOT CHANNEL FACTORS Average Maximum Core (VANTAGE+) 0.19% --
Hot subchannel (VANTAGE+) 3.92% 14.51%
Revision 23 December 2016
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DCPP UNITS 1 & 2 FSAR UPDATE TABLE 4.4-4 NON-LOCA DNB ANALYSIS METHOD UFSAR Event DNB Analysis Section(s) Method (a) 15.2.1 RCCA Bank Withdrawal from Subcritical Non-ITDP 15.2.2 RCCA Bank Withdrawal at Power ITDP 15.2.3 RCCA Misoperation (Dropped Rod) ITDP 15.2.4 Uncontrolled Boron Dilution Not Applicable (b) 15.2.5 Partial Loss of Flow ITDP 15.2.6 Startup of an Inactive Loop (c)
Non-ITDP 15.2.7 Loss of Electrical Load / Turbine Trip ITDP 15.2.8, 15.2.9 Loss of Normal Feedwater / Loss of Offsite Power Not Applicable (Non-DNB) 15.2.10 Feedwater Malfunction (excess flow) ITDP 15.2.11 Sudden Feedwater Temperature Reduction ITDP 15.2.12 Excessive Load Increase ITDP 15.2.13 Accidental RCS Depressurization ITDP 15.2.14 Accidental MSS Depressurization (d)
Non-ITDP 15.2.15 Spurious SI Actuation at Power ITDP 15.3.4 Complete Loss of Flow ITDP 15.3.5 Single RCCA Withdrawal at Power ITDP 15.4.2.1 Rupture of a Main Steam Line (HZP) Non-ITDP 15.4.2.2 Rupture of a Main Feedwater Pipe Not Applicable (Non-DNB) 15.4.2.3 Steam Break at Full Power ITDP 15.4.4 Locked Rotor ITDP 15.4.6 RCCA Ejection Not Applicable (Non-DNB)
(a) The DNB analysis method specified (ITDP or non-ITDP) refers to the method used to analyze for the DNB criterion. This is not applicable for cases analyzed for other acceptance criteria (e.g., peak RCS pressure) or for events that are not explicitly analyzed for DNB.
(b) For the time period from the start of dilution until reactor trip the Boron Dilution at Power case with the reactor in manual control is bounded by the DNB analysis for RCCA Bank Withdrawal at Power (refer to Section 15.2.2).
(c) Precluded by Technical Specifications; no longer analyzed.
(d) Bounded by Section 15.4.2.1; no longer analyzed.
Revision 23 December 2016
Revision 24 September 2018 Revision 24 September 2018 HISTORICAL FSAR UPDATE UNITS 1 AND 2 DIABLO CANYON SITE FIGURE 4.2-6 REMOVABLE ROD COMPARED TO STANDARD ROD Revision 11 November 1996 Revision 23 December 2016
HISTORICAL FSAR UPDATE UNITS 1 AND 2 DIABLO CANYON SITE FIGURE 4.2-7 REMOVABLE FUEL ROD ASSEMBLY OUTLINE Revision 1111November Revision November 1996 1996 Revision 23 December 2016
FSAR UPDATE UNITS 1 AND 2 DIABLO CANYON SITE FIGURE 4.2-8 LOCATION OF REMOVABLE RODS WITHIN AN ASSEMBLY Revision 1111November Revision November 1996 1996 Revision 23 December 2016
FSAR UPDATE UNIT 1 DIABLO CANYON SITE FIGURE 4.2-9 LOWER CORE SUPPORT ASSEMBLY Revision 11 November 1996
FSAR UPDATE UNIT 2 DIABLO CANYON SITE FIGURE 4.2-10 LOWER CORE SUPPORT ASSEMBLY Revision 11 November 1996
FSAR UPDATE UNIT 2 DIABLO CANYON SITE FIGURE 4.2-11 NEUTRON SHIELD PAD LOWER CORE SUPPORT STRUCTURE Revision 11 November 1996
FSAR UPDATE UNIT 1 DIABLO CANYON SITE FIGURE 4.2-12 UPPER CORE SUPPORT STRUCTURE Revision 11 November 1996
FSAR UPDATE UNIT 2 DIABLO CANYON SITE FIGURE 4.2-13 UPPER CORE SUPPORT STRUCTURE Revision 11 November Revision 11 November 1996 1996
FSAR UPDATE UNITS 1 AND 2 DIABLO CANYON SITE FIGURE 4.2-14 PLAN VIEW OF UPPER CORE SUPPORT STRUCTURE Revision 11 November 1996
Revision 24 September 2018 Revision 24 September 2018 Revision 24 September 2018 FSAR UPDATE UNITS 1 AND 2 DIABLO CANYON SITE FIGURE 4.3-1 FUEL LOADING ARRANGEMENT Revision 11 November 1996
Revision 24 September 2018 FSAR UPDATE UNITS 1 AND 2 DIABLO CANYON SITE FIGURE 4.3-3 BORON CONCENTRATION VS CYCLE BURNUP WITH BURNABLE ABSORBER RODS Revision 11 November 1996
Revision 24 September 2018 Revision 24 September 2018 Revision 24 September 2018 Revision 24 September 2018 Revision 24 September 2018 Revision 24 September 2018 FSAR UPDATE UNITS 1 AND 2 DIABLO CANYON SITE FIGURE 4.3-6 BURNABLE ABSORBER LOADING PATTERN Revision 11 November 1996
FSAR UPDATE UNITS 1 AND 2 DIABLO CANYON SITE FIGURE 4.3-7 NORMALIZED POWER DENSITY DISTRIBUTION NEAR BEGINNING OF LIFE (BOL), UNRODDED CORE, HOT FULL POWER, NO XENON Revision 11 November 1996
FSAR UPDATE UNITS 1 AND 2 DIABLO CANYON SITE FIGURE 4.3-8 NORMALIZED POWER DENSITY DISTRIBUTION NEAR BOL UNRODDED CORE, HOT FULL POWER, EQUILIBRIUM XENON Revision 11 November 1996
FSAR UPDATE UNIT 1 DIABLO CANYON SITE FIGURE 4.3-9 NORMALIZED POWER DENSITY DISTRIBUTION NEAR BOL GROUP D AT INSERTION LIMIT, HOT FULL POWER, EQUILIBRIUM XENON Revision 11 November 1996
FSAR UPDATE UNIT 2 DIABLO CANYON SITE FIGURE 4.3-10 NORMALIZED POWER DENSITY DISTRIBUTION NEAR BOL GROUP D AT INSERTION LIMIT, HOT FULL POWER, EQUILIBRIUM XENON Revision 11 November 1996
FSAR UPDATE UNITS 1 AND 2 DIABLO CANYON SITE FIGURE 4.3-11 NORMALIZED POWER DENSITY DISTRIBUTION NEAR MIDDLE OF LIFE (MOL)
UNRODDED CORE, HOT FULL POWER, EQUILIBRIUM XENON Revision 11 November 1996
FSAR UPDATE UNITS 1 AND 2 DIABLO CANYON SITE FIGURE 4.3-12 NORMALIZED POWER DENSITY DISTRIBUTION NEAR END OF LIFE (EOL)
UNRODDED CORE, HOT FULL POWER, EQUILIBRIUM XENON Revision 11 November 1996
FSAR UPDATE UNITS 1 AND 2 DIABLO CANYON SITE FIGURE 4.3-13 RODWISE POWER DISTRIBUTION IN A TYPICAL ASSEMBLY (G-10)
NEAR BOL, HOT FULL POWER, EQUILIBRIUM XENON, UNRODDED CORE Revision 11 November 1996
FSAR UPDATE UNITS 1 AND 2 DIABLO CANYON SITE FIGURE 4.3-14 RODWISE POWER DISTRIBUTION IN A TYPICAL ASSEMBLY (G-10)
NEAR EOL, HOT FULL POWER, EQUILIBRIUM XENON, UNRODDED CORE Revision 11 November 1996
Revision 24 September 2018 Revision 24 September 2018 Revision 24 September 2018 Revision 24 September 2018 Revision 24 September 2018 FSAR UPDATE UNITS 1 AND 2 DIABLO CANYON SITE FIGURE 4.3-23 MAXIMUM F QT x POWER vs AXIAL HEIGHT DURING NORMAL OPERATIONS Revision 11 November 1996
Revision 24 September 2018 Revision 24 September 2018 Revision 24 September 2018 FSAR UPDATE UNITS 1 AND 2 DIABLO CANYON SITE FIGURE 4.3-28 DOPPLER TEMPERATURE COEFFICIENT AT BOL AND EOL Revision 11 November 1996
FSAR UPDATE UNITS 1 AND 2 DIABLO CANYON SITE FIGURE 4.3-29 DOPPLER ONLY POWER COEFFICIENT AT BOL AND EOL Revision 11 November 1996
FSAR UPDATE UNITS 1 AND 2 DIABLO CANYON SITE FIGURE 4.3-30 DOPPLER ONLY POWER DEFECT AT BOL AND EOL Revision 11 November 1996
FSAR UPDATE UNITS 1 AND 2 DIABLO CANYON SITE FIGURE 4.3-31 MODERATOR TEMPERATURE COEFFICIENT AT BOL, NO RODS Revision 11 November 1996
FSAR UPDATE UNITS 1 AND 2 DIABLO CANYON SITE FIGURE 4.3-32 MODERATOR TEMPERATURE COEFFICIENT AT EOL Revision 11 November 1996
FSAR UPDATE UNITS 1 AND 2 DIABLO CANYON SITE FIGURE 4.3-33 MODERATOR TEMPERATURE COEFFICIENT AS A FUNCTION OF BORON CONCENTRATION AT BOL, NO RODS Revision 11 November 1996
FSAR UPDATE UNITS 1 AND 2 DIABLO CANYON SITE FIGURE 4.3-34 HOT FULL POWER MODERATOR TEMPERATURE COEFFICIENT FOR CRITICAL BORON CONCENTRATION Revision 11 November 1996
FSAR UPDATE UNITS 1 AND 2 DIABLO CANYON SITE FIGURE 4.3-35 TOTAL POWER COEFFICIENT AT BOL AND EOL Revision 11 November 1996
FSAR UPDATE UNITS 1 AND 2 DIABLO CANYON SITE FIGURE 4.3-36 TOTAL POWER DEFECT AT BOL AND EOL Revision 11 November 1996
Revision 24 September 2018 Revision 24 September 2018 Revision 24 September 2018 Revision 24 September 2018 Revision 24 September 2018 Revision 24 September 2018 Revision 24 September 2018 FSAR UPDATE UNITS 1 AND 2 DIABLO CANYON SITE FIGURE 4.3-44 CALCULATED AND MEASURED DOPPLER DEFECT AND COEFFICIENTS AT BOL, FOR A TWO-LOOP PLANT WITH A 12-FT CORE HEIGHT AND 121 ASSEMBLIES Revision 11 November 1996
FSAR UPDATE UNITS 1 AND 2 DIABLO CANYON SITE FIGURE 4.3-45 COMPARISON OF CALCULATED AND MEASURED BORON CONCENTRATION FOR A TWO-LOOP PLANT WITH A 12-FT CORE HEIGHT AND 121 ASSEMBLIES Revision 11 November 1996
FSAR UPDATE UNITS 1 AND 2 DIABLO CANYON SITE FIGURE 4.3-46 COMPARISON OF CALCULATED AND MEASURED BORON FOR A TWO LOOP PLANT WITH A 12-FT CORE HEIGHT AND 121 ASSEMBLIES Revision 11 November 1996
FSAR UPDATE UNITS 1 AND 2 DIABLO CANYON SITE FIGURE 4.3-47 COMPARISON OF CALCULATED AND MEASURED BORON IN A 3-LOOP PLANT WITH A 12-FT CORE HEIGHT AND 157 ASSEMBLIES Revision 11 November 1996
Revision 24 September 2018 Revision 24 September 2018 Revision 24 September 2018 Revision 24 September 2018 Revision 24 September 2018 Revision 24 September 2018 Revision 24 September 2018 Revision 24 September 2018 Revision 24 September 2018 Revision 24 September 2018 Revision 24 September 2018 Revision 24 September 2018 Revision 24 September 2018 Revision 24 September 2018 Revision 24 September 2018 FSAR UPDATE UNIT 1 DIABLO CANYON SITE FIGURE 4.4-16 DISTRIBUTION OF INCORE INSTRUMENTATION Revision 11 November 1996
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FSAR UPDATE UNIT 2 DIABLO CANYON SITE FIGURE 4.4-17 DISTRIBUTION OF INCORE INSTRUMENTATION Revision 11 November 1996
Revision 24 September 2018 Revision 24 September 2018 Revision 24 September 2018