L-2016-198, Redacted Turkey Point, Units 3 & 4, Updated Final Safety Analysis Report, Chapter 3, Reactor
Text
TABLE OF CONTENTS Section Title Page 3 REACTOR 3.1.1-1 3.1Design Bases 3.1.1-1 3.1.1 Performance Objectives 3.1.1-1 References Section 3.1.1 3.1.1-3 3.1.2 Principal Design Criteria 3.1.2-1 Reactor Core Design 3.1.2-1 Suppression of Power Oscillations 3.1.2-2 Redundancy of Reactivity Control 3.1.2-3 Reactivity Hot Shutdown Capability 3.1.2-3 Reactivity Shutdown Capability 3.1.2-4 Reactivity Holddown Capability 3.1.2-5 Reactivity Control Systems Malfunction 3.1.2-6 Maximum Reactivity Worth of Control Rods 3.1.2-7 3.1.3 Design Objectives 3.1.3-1 Nuclear 3.1.3-1 Reactivity Control 3.1.3-2 Thermal and Hydraulic 3.1.3-2 Mechanical 3.1.3-3 Reactor Internals 3.1.3-3 Fuel Assemblies 3.1.3-5 Rod Cluster Control Assemblies 3.1.3-7 Control Rod Drive Assembly 3.1.3-7 3.2Reactor Design 3.2.1-1 3.2.1 Nuclear Design and Evaluation 3.2.1-1 Nuclear Characteristics of the Design 3.2.1-1 Reactivity Control Aspects 3.2.1-1 Chemical Shim Control 3.2.1-2 Control Rod Requirements 3.2.1-2 Total Power Reactivity Defect 3.2.1-3 Control Rod Bite 3.2.1-3 Xenon Stability Control 3.2.1-5 Excess Reactivity Insertion Upon 3.2.1-5 Reactor Trip Calculated Rod Worths 3.2.1-5 Reactor Core Power Distribution 3.2.1-6 Power Distribution Control 3.2.1-7 Reactivity Coefficients 3.2.1-8 Moderator Temperature Coefficient 3.2.1-8 Moderator Pressure Coefficient 3.2.1-10 Moderator Density Coefficient 3.2.1-10 Doppler and Power Coefficients 3.2.1-10 Nuclear Evaluation 3.2.1-12 Reactivity Analysis 3.2.1-12 Depletion Analysis 3.2.1-12 Power Peaking Analysis 3.2.1-13 Gross Power Distribution Analysis 3.2.1-14 RCC Assembly Worth Analysis 3.2.1-15 Moderator Coefficient Analysis 3.2.1-16 Doppler and Power Coefficient Analysis 3.2.1-17 Comparison of Predicted and Measured Boron Concentrations 3.2.1-20 References, Section 3.2.1 3.2.1-21 3-i Revised 04/17/2013 C26C26C26C26C26C26C26C26 TABLE OF CONTENTS (Continued)
Section Title Page 3.2.2 Thermal and Hydraulic Design and Evaluation 3.2.2-1 Thermal and Hydraulic Characteristics of the Design 3.2.2-1 Thermal Data 3.2.2-1 Fuel and Cladding Temperature 3.2.2-1 U0 2 Thermal Conductivity 3.2.2-2 Operational Experience with Westinghouse Cores 3.2.2-3 Heat Flux Ratio and Data Correlation 3.2.2-4 Definition of Departure from Nucleate Boiling Ratio 3.2.2-4 W-3 Equivalent Uniform Flux DNB Correlation 3.2.2-5 W-3 DNB Correlation 3.2.2-6 WRB-1 DNB Correlation 3.2.2-7 Surface Heat Transfer Coefficients 3.2.2-9 Hot Channel Factors 3.2.2-10 Definition of Engineering Hot Channel Factor 3.2.2-10 Heat Flux Engineering Subfactor 3.2.2-10 Enthalpy Rise Engineering Subfactor 3.2.2-11 Pellet diameter, density and enrichment 3.2.2-11 Inlet Flow Maldistribution 3.2.2-11 Flow Distribution 3.2.2-12 Flow Mixing 3.2.2-12 Nuclear Enthalpy Rise Hot Channel Factor 3.2.2-12 Pressure Drop and Hydraulic Forces 3.2.2-11 DNBR Design Methodology 3.2.2-12 Steady State Analysis 3.2.2-15 Experimental Verification 3.2.2-15 Transient Analysis 3.2.2-16 Effects of Rod Bow on DNBR 3.2.2-16 Transition Core DNB Methodology 3.2.2-17 Effects of DNB on Neighboring Rods 3.2.2-17 DNB With Return to Nucleate Boiling 3.2.2-17 Hydrodynamic and Flow Power Coupled Instability 3.2.2-18 Effect of Fuel Densification 3.2.2.20 Thermal-Hydraulic Effect of Reactor Vessel Upper Internals During Refueling 3.2.2-21 References, Section 3.2.2 3.2.2-23 3.2.3 Mechanical Design and Evaluation 3.2.3-1 Reactor Internals 3.2.3-3 Design Description 3.2.3-3 Lower Core Support Structure 3.2.3-5 Upper Core Support Assembly 3.2.3-8 In-core Instrumentation Support Structures 3.2.3-10 Evaluation of Core Barrel and Thermal Shield 3.2.3.11
3-ii Revised 04/17/2013 C26C26 TABLE OF CONTENTS (Continued)
Section Title Page 3.2.3 Core Components 3.2.3-12 (Cont'd) Design Description 3.2.3-12 Fuel Assembly 3.2.3-12 Bottom Nozzle 3.2.3-13 Top Nozzle 3.2.3-14 Guide Thimbles 3.2.3-16 Grids 3.2.3-16 Fuel Rods 3.2.3-18 Process Control 3.2.3-19 Rod Cluster Control Assemblies 3.2.3-20 Neutron Source Assemblies 3.2.3-21 Thimble Plug Assemblies 3.2.3-22 Burnable Poison Rods 3.2.3-22 Removable Rod Assemblies 3.2.3-24 Evaluation of Core Components 3.2.3-26 Fuel Rod Evaluation 3.2.3-26 Evaluation of Burnable Poison Rods 3.2.3-28 Effects of Vibration and Thermal Cycling on 3.2.3-28 Fuel Assemblies Control Rod Drive Mechanism 3.2.3-29 Full Length Rods 3.2.3-29 Design Description 3.2.3-29 Latch Assembly 3.2.3-31 Rod Drive Mechanism Housing 3.2.3-31 Operating Coil Stack 3.2.3-31 Drive Shaft Assembly 3.2.3-32 Position Indicator Coil Stack 3.2.3-32 Drive Mechanism Materials 3.2.3-32 Principles of Operation 3.2.3-33 Control Rod Withdrawal 3.2.3-33 Control Rod Insertion 3.2.3-34 Control Rod Tripping 3.2.3-35 Reactor Vessel Level Measuring Probes 3.2.3-35 Fuel Assembly and RCC Mechanical Evaluation 3.2.3-35 Reactor Evaluation Center (WREC) Tests 3.2.3-36 Loading and Handling Tests 3.2.3-36 Axial and Lateral Bending Tests 3.2.3-36 CRDM Housing Mechanical Failure Evaluation 3.2.3-37 Effect of Rod Travel Housing Longitudinal 3.2.3-37 Failures Effect of Rod Travel Housing Circumferential 3.2.3-37 Failures Summary 3.2.3-38 References 3.2.3-39
3-iii Revised 04/17/2013 C26 LIST OF TABLES
Table Title 3.2.1-1 Nuclear Design Data (FIRST CYCLE)
3.2.1-2 Reactivity Requirements for Control Rods
3.2.1-3 Calculated Rod Worths, For First Cycle With Burnable Poison Rods 3.2.1-4 Results of Calculation as a Function of Laboratory Providing Experimental Data
3.2.1-5 Calculated and Measured Reactivity Effects of Void Tubes
3.2.1-6 Core Startup Critical Boron Concentration
3.2.2-1 Thermal and Hydraulic Design Parameters
3.2.2-2 Engineering Hot Channel Factors (First Cycle)
3.2.3-1 Core Mechanical Design Parameters
3-iv Revised 04/17/2013
LIST OF FIGURES Figure Title
3.2.1-1 Pattern of Control Rod Cluster Banks
3.2.1-2 Deleted
3.2.1-3 Average Power Density (BOL), Control Bank D In
3.2.1-4 Deleted
3.2.1-5 Average Power Density (BOL), No Control Rods In
3.2.1-6 Schematic Demonstration of Typical kW/ft. Limits
3.2.1-7 Burnable Poison Cluster Locations
3.2.1-8 Pattern of Poison Rod Locations
3.2.1-9 Moderator Temperature Coefficient Vs. Moderator Temperature
3.2.1-10 Doppler Coefficient Vs. Effective Fuel Temperature (BOL)
3.2.1-11 Power Coefficient
3.2.1-12 Power Coefficient (Closed Gap Model)
3.2.1-13 Plutonium/Uranium Mass Ratio as a Function of Uranium-235 Depletion
3.2.1-14 Fraction of Plutonium-239 in Plutonium as a Function of Uranium-235 Depletion
3.2.1-15 Composition of Plutonium as a Function of Uranium-235 Depletion
3.2.1-16 Plan of Critical Experiment (Unborated Case)
3.2.1-17 Plan of Critical Experiment (Borated Case)
3.2.1-18 Borated Power Distribution Comparison
3.2.1-19 Unborated Power Distribution Comparison
3.2.1-20 Comparison of Experimental and Calculated Power Distribution Using One Mesh Spacing Per Fuel Rod
3.2.1-21 Comparison of Calculated Power Distribution With Experimental Power Scans - Unborated Core
3.2.1-22 Comparison of Calculated Power Distribution With Experimental Power Scans - Borated Core
3.2.1-23 Radial Fuel Rod Scan
3-v Revised 04/17/2013
LIST OF FIGURES (cont'd)
Figure Title
3.2.1-24 Yankee Core I Power Distribution Comparison
3.2.1-25 Yankee Core I Burnup Distribution Comparison
3.2.1-26 Fast Absorption of Cluttered Absorbers
3.2.1-27 Fast Absorption of Uniformly Distributed Absorbers
3.2.1-28 Selni Temperature Coefficient Vs. Moderator Temperature (1600 ppm Boron)
3.2.1-29 Moderator Temperature Coefficient Vs. Boron Concentration
3.2.1-30 Comparison of Calculated and Measured Moderator Temperature Coefficient Vs. Burnup
3.2.1-31 Comparison of Resonance Integral Correlations
3.2.1-32 Fuel Temperature Changes Vs. Power Density
3.2.1-33 Alpha Vs. Heat Flux
3.2.1-34 Comparison of Effective Fuel Temperature With Changing Heat Flux
3.2.2-1 Thermal Conductivity of Uranium Dioxide
3.2.2-2 Comparison of W-3 Prediction and Uniform Flux Data
3.2.2-3 W-3 Correlation Probability Distribution Curve
3.2.2-4 Comparison of W-3 Correlation With Rod Bundle DNB Data (Simple Grid Without Mixing Vanes)
3.2.2-5 Comparison of W-3 Correlation With Rod Bundle DNB Data (Simple Grid With Mixing Vanes)
3.2.2-6 Measured Vs. Predicted Critical Heat Flux WRB-1 Correlation
3.2.2-7 Comparison of W-3 Prediction and Non-Uniform Flux Data
3.2.2-8 Comparison of W-3 Prediction With Measured DNB Location
3.2.2-9 Measured Vs predicted Critical Heat Flux -ABB-NV Correlation 3.2.2-10 Measured Vs predicted Critical Heat Flux -WLOP Correlation
3.2.3-1 Reactor Core Cross Section
3.2.3-2 Reactor Vessel Internals
3.2.3-3 Three Region Core Loading - First Cycle
3-vi Revised 04/17/2013 C26 LIST OF FIGURES (cont'd)
Figure Title
3.2.3-4 Typical Rod Cluster Assembly
3.2.3-5 Lower Core Support Assembly
3.2.3-6 Upper Core Support Assembly
3.2.3-7 Guide Tube Assembly
3.2.3-8 Fuel Assembly and Control Cluster Cross Section
3.2.3-9 LOPAR Fuel Assembly Outline
3.2.3-9A OFA-LOPAR Fuel Assembly Outlines
3.2.3-9B Bottom Nozzle to Thimble Tube Connection
3.2.3-9C 15x15 OFA/DRFA Fuel Assembly Comparison
3.2.3-9D 15x15 DRFA/Upgrade Fuel Assembly Comparison
3.2.3-10 Spring Clip Grid Assembly
3.2.3-10A Comparison of OFA and LOPAR Plugging Device
3.2.3-11 Detail of Burnable Poison Rod
3.2.3-11A Wet Burnable Poison Absorber Rod
3.2.3-12 Control Rod Drive Mechanism Assembly
3.2.3-13 Control Rod Drive Mechanism Schematic
3.2.3-14 Reduced Length HVFD Absorber Rod
3.2.3-15 Head Adapter Plug Design
3.2.3-16 Pipe Cap Design
3.2.3-17 CRGT Flow Restrictor Assembly
3.2.3-18 CRDM Plug
3-vii Revised 04/17/2013
3 REACTOR 3.1 DESIGN BASIS
3.1.1 PERFORMANCE OBJECTIVES.
The reactor core is a three-region cycled core. The fuel rods are cold worked partially annealed Zircaloy-4, ZIRLO or Optimized ZIRLOŽ tubes containing slightly enriched uranium dioxide fuel. All fuel rods are pressurized with helium during fabrication to reduce stresses and strains and to increase fatigue life. A more detailed discussion is given in Reference 1.
The fuel assembly consists of the rod cluster control (RCC) guide thimbles fastened to the grids and the top and bottom nozzles. The fuel rods are held in this assembly at seven points along their length by spring-clip grids which provide a very stiff support for the fuel rods.
The Turkey Point Units are loaded with Westinghouse seven grid 15 x 15 Upgrade Assemblies (Upgrade) starting with Unit 3 Cycle 25 and Unit 4 Cycle
- 26. Previously, Westinghouse 15 x 15 Debris Resistant Fuel Assembles (DRFA) were used. Beginning with Unit 3 Cycle 12 and Unit 4 Cycle 13 DRFA assemblies were loaded. Additional details and analysis of the DRFA design are provided in Reference 8. Prior to DRFA, 15 x 15 Optimized Fuel Assemblies (OFA) and 15 x 15 Low Parastic Fuel (LOPAR) was used. See Figures 3.2.3-9A, 3.2.3-9C, and 3.2.3-9D for the overall configurations of these fuel designs. These figures were included for historical purposes. The top and bottom grids (non-mixing) will continue to be manufactured using Inconel.
Differences between the Upgrade and DRFA are the addition of Intermediate Flow Mixing Grids (IMFs) and a Protective Grid (P-grid), balanced vaned mid grids, shorter end plug, a Debris Filter Bottom Nozzle (DFBN), and a Tube-in-Tube dashpot.
Full length rod cluster control assemblies (RCCA), secondary sources, thimble plug devices and burnable poison rods are inserted into the guide thimbles of the fuel assemblies. The absorber sections of the control rods are fabricated of silver-indium-cadmium alloy sealed in stainless steel tubes.
The absorber material in the fixed burnable poison rods is in the form of borosilicate glass sealed in stainless steel tubes.
Three other types of burnable poison rods and absorbers are employed:
a) Wet Annular Burnable Absorbers (WABA)
(2), each consisting of an aluminum oxide-boron carbide annulus sealed in Zircaloy, and
3.1.1-1 Revised 03/11/2016 C28 b) Reduced length Annular Hafnium Vessel Flux Depression (HVFD) absorbers which may be placed in peripheral assemblies as part of the flux reduction program.
(3)* c) Integral Fuel Burnable Absorbers (IFBA)
(4), consisting of a Zirconium diboride coating on the surface of the fuel pellets.
Evaluations (5) have been performed to support the complete or partial removal of thimble plugs from Turkey Point Units 3 & 4. These evaluations have
addressed the effect of thimble plug removal on core design, core thermal
hydraulics, reactor pressure vessel system thermal hydraulics and the non-
LOCA and LOCA safety analyses. Based on these evaluations, it has been
determined that it is acceptable to remove all or any combination of thimble
plugs from Turkey Point Units 3 & 4. In addition, the secondary sources have
been removed from Turkey Point Units 3 & 4 (References 6 and 7, respectively)
The Control Rod Drive Mechanisms (CRDM) for the RCCA are of the magnetic
latch type. The latches are controlled by three magnetic coils. They are so
designed that upon a loss of power to the coils, the RCC assembly is released
and falls by gravity to shut down the reactor.
The mechanisms for the former part length rod cluster control assemblies have
been immobilized in a fully withdrawn position. Two of these positions have
been modified to accommodate the installation of heated junction thermocouple
probes for the Reactor Vessel Level Monitoring System (RVLMS).
The replacement reactor vessel closure heads (RVCHs) do not have nozzles and
CRDM adapters for the mechanisms for the former part length rod cluster
control assemblies. The mechanisms for the former part length rod cluster
control assemblies that were formerly immobilized in the fully withdrawn
position have been removed (Reference 9 & 11). The two core positions that
were formerly modified to accommodate the installation of heated junction
thermocouple probes for the RVLMS have been retained as RVLMS locations on
the Unit 3 and Unit 4 replacement RVCHs. When the former part length rod
cluster control assemblies were immobilized in the fully withdrawn position, the drive shaft was immobilized in such a position that it was inserted into
the upper portion of the reactor internals and provided flow restriction to
flow up through the reactor upper internals. The flow restriction attributed
to the removed drive shaft has been maintained by installing flow restricting
devices into the upper internals in those positions from which the drive
shafts were removed. Evaluations have been performed to support the design
of the flow restrictor device to insure there is no effect on core thermal
hydraulics during all design bases operating conditions (References 9, 10, and 11). 3.1.1-2 Revised 03/01/2011 C25C25 Turkey Point Units 3 & 4 will make a transition from all DRFA fuel cores to all Upgrade fuel cores with the exception that DRFA fuel assemblies, once
removed from the reactor core, may be re-inserted to meet cycle specific
energy requirements and to enhance fuel efficiency.
REFERENCES, SECTION 3.1.1
- 1. WCAP-9002 (2/69), "Use of Internally Pressurized Fuel Rods in Westinghouse Pressurized Water Reactors", PROPRIETARY. A
NON-PROPRIETARY version of this report is WCAP-7855.
- 2. Letter from Thomas, C.D., NRC, to Rahe, E.P., Westinghouse,
Subject:
Acceptance for Referencing of Licensing Topical Report WCAP-10021 (P),
Revision 1, and WCAP-10377 (NP), "Westinghouse Wet Annular Burnable
Absorber Evaluation Report", August 9, 1983.
- 3. Letter from Uhrig, R.E., FP&L to Varga, S.A., NRC,
Subject:
Pressurized Thermal Shock, Letter No L-83-180, March 25, 1983.
- 4. WCAP-10444 Addendum 1 (NP) "Reference Core Report Vantage 5 Fuel Assembly," via Westinghouse transmittal letter NS-NRC-85-3090, December
23, 1985, to Standardization and Special Projects Directorate, U.S.N.R.C.
- 5. Westinghouse Letter #87FP*-G-0052, "Evaluation Report for Thimble Plug Removal for Turkey Point Units 3 & 4", December 18, 1987.
- 6. PCM-91-072, "Turkey Point Unit 3 Removal of Startup Sources," July 11.
1991.
- 7. PCM-91-043, "Turkey Point Unit 4 Removal of Startup Sources," July 11, 1991.
- 8. WCAP-12346, Turkey Point Units 3 and 4, 15x15 Debris Resistant Fuel Assembly Design Report, July 1989.
- 9. PC/M 03-057, Rev.01, "Reactor Vessel Closure Head Replacement".
- 10. Framatome ANP Document 32-5017014, Rev.3, Dated 3/9/2004, "Turkey Point 3 and 4 Part-Length CRDM Nozzle Repair Hydraulic Evaluation".
- 11. PC/M 03-058, "Reactor Vessel Closure Head Replacement- Unit 4"
- 12. Letter from Jason Paige, NRC, to Mano Nazar, FPL, "Turkey Point Units 3 and 4 - Exemption from the Requirements of 10 CFR part 50, Appendix G and 10 CFR Part 50, Section 50.61 (TAC Nos. ME 1007 and ME 1008), March 11, 2010.
- The HVFD absorbers were removed from both vessels in 2009. The NRC approved (12) the exemption request for alternate material properties bases with use of Framatome Topical Report BAW-2308, Revisions 1A and 2A. The alternate material properties bases allows enough PTS margin that the HVFD absorbers can be eliminated. HVFD absorbers could be used again in the future, if desired.
3.1.1-3 Revised 03/01/2011 C25C25 3.1.2 PRINCIPAL DESIGN CRITERIA
Reactor Core Design Criterion: The reactor core with its related controls and protection systems shall be designed to function throughout its design lifetime without exceeding acceptable fuel damage limits which have been stipulated and justified. The core and related auxiliary system designs shall provide this integrity under all expected conditions of normal operation with appropriate margins for uncertainties and for specified transient situations which can be anticipated (1967 Proposed GDC 6)
The reactor core, with its related control and protection system, is designed to function throughout its design lifetime without exceeding fuel limits specified to preclude damage. The core design, together with reliable process and decay heat removal systems, provides for this capability under all expected conditions of normal operation with appropriate margins for uncertainties and anticipated transient situations, including, as examples, the effects of the loss of reactor coolant flow (Section 14.1.9), loss of external electrical load (Section 14.1.10), loss of normal feedwater flow (Section 14.1.11) and loss of non-emergency A.C. power to the plant auxiliaries (Section 14.1.12).
The Reactor Control and Protection system is designed to actuate a reactor trip for any anticipated combination of plant conditions, when necessary, to ensure a minimum Departure from Nucleate Boiling (DNB) ratio greater than or equal to the DNBR limit of the applicable DNB correlation. The integrity of the fuel cladding is ensured by preventing excessive fuel swelling, excessive clad heating, and excessive cladding stress and strain. This is achieved by designing the fuel rods so that the following conservative limits are not exceeded during normal operation or any anticipated transient condition:
a) Minimum DNB ratio equal to or greater than the DNBR limit of the applicable DNB correlation. b) Fuel Center temperature below melting point of U0
- 2. c) The combined maximum stress intensity meets the criteria based on the American Society of Mechanical Engineers (ASME) code in Addendum 1-A of Reference 21 or clad stresses less than the Zircaloy-4, ZIRLO or Optimized ZIRLO clad yield strength (as irradiated). d) Clad strain less than 1%. e) Cumulative fatigue life shall not exceed the design failure life design strain (as irradiated).
3.1.2-1 Revised 03/11/2016 C28 f) The internal gas pressure of the lead rod in the reactor is maintained below a value that could cause (1) the diametrical gap
to increase due to outward clad creep during steady-state
operations and (2) extensive DNB propagation to occur.
The ability of fuel designed and operated to these criteria to withstand
postulated normal service conditions is described in this chapter. Abnormal
service conditions as shown by analyses are presented in Chapter 14 to
substantiate that the demands of plant operation are well within applicable
regulatory limits.
The reactor coolant pumps have sufficient rotational inertia to maintain an
adequate flow coastdown in the event of a simultaneous loss of power to all
pumps. The flow coastdown inertia is sufficient such that the reduction in
heat flux obtained with a low flow reactor trip prevents core damage.
In the unlikely event of a turbine trip from full power without immediate
reactor trip, the subsequent reactor coolant temperature increase and volume
insurge to the pressurizer results in a high pressurizer pressure trip and
thereby prevents fuel damage from this transient. A loss of external
electrical load is controlled by RCCA insertion together with a controlled
steam dump to the condenser and atmosphere to prevent a large temperature
and pressure increase in the Reactor Coolant System. In this case, trip
signals due to overpower delta-T, overtemperature delta-T, high pressurizer
pressure or water lever, and low-low steam generator water level would guard
against any combination of pressure, temperature and power which could result
in a DNB ratio less than the DNBR limit of the applicable DNB correlation
during the course of the transient. Details of this event are described in
FSAR Section 14.1.10.
In neither the turbine trip nor the loss-of-flow events do the changes in
coolant conditions cause a nuclear power excursion because of the large
system thermal inertia and relatively small void fraction. Protection
circuits actuated directly by the coolant conditions identified with core
limits will prevent core damage.
Suppression of Power Oscillations
Criterion: The design of the reactor core with its related controls
and protection systems shall ensure that power oscillations, the magnitude of
which could cause damage in excess of acceptable fuel damage limits, are not
possible or can be readily suppressed. (1967 Proposed GDC 7)
3.1.2-2 Revised 04/17/2013 The potential for possible spatial oscillations of power distribution for
this core has been reviewed. It is concluded that low frequency xenon
oscillations, which may occur in the axial direction, can be controlled by
control rod movement. The core is expected to be stable to xenon
oscillations in the X-Y dimension. Excore instrumentation is provided to
obtain necessary information concerning power distribution. This
instrumentation is adequate to enable the operator to monitor and control
xenon induced oscillations. (Incore instrumentation is used to periodically
calibrate and verify the information provided by the excore instrumentation).
The analysis, detection and control of these oscillations is discussed in
reference (3) of Section 3.2.1.
Redundancy of Reactivity Control
Criterion: Two independent reactivity control systems, preferably of different principles, shall be provided. (1967 Proposed GDC 27)
Two independent reactivity control systems are provided, one involving rod
cluster control assemblies (RCCA) and the other involving chemical shimming.
he control rod system provides the minimum shutdown margin under Condition I
events and is capable of making the core subcritical rapidly enough to
prevent exceeding acceptable fuel damage limits assuming that the highest
worth control rod is stuck in the fully withdrawn position.
The boron system can compensate for all xenon burnout reactivity changes and
will maintain the reactor in cold shutdown.
Reactivity Hot Shutdown Capability Criterion: The reactivity control systems provided shall be capable of making and holding the core subcritical from any hot
standby or hot operating condition. (1967 Proposed GDC 28)
The two reactivity control systems provided are capable of making and holding
the core subcritical from any hot standby or hot operating condition, including those resulting from power changes. The maximum excess reactivity
expected for the core occurs for the cold, clean condition at the beginning
of life of the initial core.
3.1.2-3 Revised 04/17/2013 The Rod Cluster Control Assemblies (RCCA) are divided into two categories
comprising control and shutdown groups. The control banks used in
combination with chemical shim control provide control of the reactivity
changes of the core throughout the life of the core at power conditions.
This group of RCCA's is used to compensate for short term reactivity changes
at power that might be produced due to variations in reactor power level or
in coolant temperature. The chemical shim control is used to compensate for
the more slowly occurring changes in reactivity throughout core life such as
those due to fuel depletion and fission product buildup.
Reactivity Shutdown Capability Criterion: One of the reactivity control systems provided shall be capable of making the core subcritical under any
anticipated operating condition (including anticipated
operational transients) sufficiently fast to prevent
exceeding acceptable fuel damage limits. Shutdown margin
should assure subcriticality with the most reactive control
rod fully withdrawn. (1967 Proposed GDC 29)
The reactor core, together with the reactor control and protection system is designed so that the minimum allowable DNBR is greater than or equal to the
DNBR limit of the DNB correlation being used, and there is no fuel melting
during normal operation including anticipated transients.
The shutdown groups are provided to supplement the control groups of RCCA's to make the reactor at least one percent subcritical at the hot, zero power
condition (k eff = 0.99) following a trip from the most credible operating condition assuming the most reactive RCCA remains in the fully withdrawn
position.
The negative reactivity worth of all the RCCAs, assuming the most reactive
RCCA remains in the fully withdrawn position, is not sufficient to maintain
the core subcritical for the most severe anticipated cooldown transient
associated with a single active failure e.g., accidental opening of a steam
bypass, or a safety valve stuck open. In this transient the core may become
critical and return to power; however, the core is ultimately shutdown with a
combination of RCCA's and automatic boron addition via the emergency core
cooling system with the most reactive RCCA's assumed to be fully withdrawn.
Manually controlled boric acid addition is used to maintain the shutdown
margin for the long term conditions of xenon decay and plant cooldown.
3.1.2-4 Revised 04/17/2013 Reactivity Holddown Capability Criterion: The reactivity control systems provided shall be capable of making the core subcritical under credible accident conditions with appropriate margins for contingencies and limiting any
subsequent return to power such that there will be no undue risk
to the health and safety of the public. (1967 Proposed GDC 30)
Initial reactivity shutdown capability is provided within 2.4 seconds following a trip signal by control rods with boric acid injection used to
limit any subsequent return to power and to compensate for the long term
xenon decay transient and for plant cooldown. As discussed in the previous
paragraph the combined shutdown capability of the RCCAs and boron addition
limits the return to power and results in no undue risk to the health and
safety of the public as a result of the cooldown associated with a safety
valve stuck fully open.
Any time that the reactor is at power, the quantity of boric acid retained in
the boric acid tanks and ready for injection always exceeds that required to
support a cooldown to cold shutdown conditions without letdown. Under these
conditions, adequate boration can be achieved simply by providing makeup for
coolant contraction from a boric acid tank and the refueling water storage
tank. The minimum volume maintained in the boric acid tanks, therefore, is
that volume necessary to increase the RCS boron concentration during the
early phase of the cooldown of each unit such that subsequent use of the
refueling water storage tank for contraction makeup will maintain the
required shutdown margin throughout the remaining cooldown. In addition, the
boric acid tanks have sufficient boric acid solution to achieve cold shutdown
for each unit if the most reactive RCCA is not inserted. This quantity also
exceeds that required to bring the reactor to hot standby and to compensate
for subsequent xenon decay.
Boric acid is pumped from the boric acid tanks by one of two boric acid
transfer pumps to the suction of one of three charging pumps which inject
boric acid into the reactor coolant. Any charging pump and either boric acid
transfer pump can be operated from diesel generator power on loss of offsite
power. Boric acid can be injected by one pump at a rate which takes the
reactor to hot standby with no rods inserted in less than forty minutes when
a feed and bleed process is utilized (less than 30 minutes when the available
pressurizer volume is utilized). In forty additional minutes, enough boric
acid can be injected to compensate for xenon decay although xenon decay below
the equilibrium operating level does not begin until approximately 15 hours1.736111e-4 days <br />0.00417 hours <br />2.480159e-5 weeks <br />5.7075e-6 months <br />
after shutdown.
3.1.2-5 Revised 04/17/2013 If two boric acid pumps are available, these time periods are reduced.
Additional boric acid injection is employed if it is desired to bring the
reactor to cold shutdown conditions.
On the basis of the above, the injection of boric acid is shown to afford
backup reactivity shutdown capability, independent of control rod clusters
which normally serve this function in the short term situation. Shutdown for
long term and reduced temperature conditions can be accomplished with boric
acid injection using redundant components, thus achieving the measure of
reliability implied by the criterion.
Alternately, boric acid solution at lower concentration can be supplied from
the refueling water tank. This solution can be transferred directly by the
charging pumps. The reduced boric acid concentration lengthens the time
required to achieve equivalent shutdown.
If pressure is reduced in the primary system, a second alternative method
comprises the injection of boric acid solution by operation of the safety
injection pumps, taking suction from the refueling water storage tank.
Reactivity Control Systems Malfunction
Criterion: The reactor protection systems shall be capable of protecting against any single malfunction of the reactivity control system, such as unplanned continuous withdrawal (not ejection or dropout)
of a control rod, by limiting reactivity transients to avoid
exceeding acceptable fuel damage limits. (1967 Proposed GDC 31)
The reactor protection systems are capable of protecting against any single
anticipated malfunction of the reactivity control system, by limiting
reactivity transients to avoid exceeding fuel limits specified to preclude
damage. (1967 Proposed GDC 6).
Reactor shutdown with rods is completely independent of the normal rod
control functions since the trip breakers completely interrupt the power to
the latch type rod mechanisms regardless of existing control signals.
Details of the effects of continuous withdrawal of a control rod and
continuous deboration are described in Sections 14.1.1, 14.1.2 and 14.1.5, respectively.
3.1.2-6 Revised 04/17/2013 Maximum Reactivity Worth of Control Rods
Criterion: Limits, which include reasonable margin, shall be placed on the maximum reactivity worth of control rods or elements and on rates at which reactivity can be increased to ensure that the potential
effects of a sudden or large change of reactivity cannot (a)
rupture the reactor coolant pressure boundary or (b) disrupt the
core, its support structures, or other vessel internals
sufficiently to lose capability of cooling the core. (1967
Proposed GDC 32)
Limits, which include margin, are placed on the maximum reactivity worth of
control rods or elements and on rates at which reactivity can be increased to
ensure that the potential effects of a sudden or large reactivity change
cannot (a) rupture the reactor coolant pressure boundary, or (b) disrupt the
core, its support structures, or other vessel internals so as to lose
capability to cool the core.
The reactor control system employs RCCA's approximately half of which are
fully withdrawn during power operation, serving as shutdown rods. The
remaining rods comprise the control groups which are used to control load and
reactor coolant temperature. The full length rod cluster drive mechanisms
are wired into preselected groups, and are therefore prevented from being
withdrawn in other than their respective groups. The rod drive mechanism is
of the magnetic latch type and the coil actuation is sequenced to provide
variable speed rod travel.
The maximum reactivity insertion rate assumed in the RCCA Withdrawal safety
analysis bounds the rate corresponding to the maximum differential rod worth
for two overlapping groups moving together in the highest worth region of the
core. The assumed maximum reactivity insertion rate is well within the
capability of the overpower-overtemperature protection circuits to prevent
core damage.
No single mechanical or electrical control system malfunction can cause a rod
cluster to be withdrawn at a speed greater than 77 steps per minute (48
inches per minute). This represents the maximum theoretical limit. However, the nominal maximum design limit of 72 steps per minute is used in the design
calculation.
3.1.2-7 Revised 04/17/2013 3.1.3 DESIGN OBJECTIVES
The reactor is capable of meeting the performance objectives throughout core life under both steady state and transient conditions without violating the
integrity of the fuel cladding. Thus the release of unacceptable amounts of
fission products to the coolant is prevented.
The limiting conditions discussed below are the highest functional capacity or performance levels for the nuclear, control, thermal and hydraulic, and
mechanical aspects of design permitted to assure safe operation of the
facility.
Nuclear At a full power level (license application power) the nuclear heat flux hot
channel factor, F N q , specified in Table 3.2.1-1, Line 18, is not exceeded.
The nuclear axial peaking factor F N Z , and the nuclear enthalpy rise hot channel factor F NH are limited in their combined relationship so as not to exceed the F q or DNBR limits.
The limiting nuclear hot channel factors are higher than those calculated at full power for the range from all control rods fully withdrawn to maximum allowable control rod insertion. Control rod insertion limits as a function of power are delineated in the Core Operating Limits Report in Appendices 14A and 14B for Units 3 and 4 respectively to ensure that hot channel factors do
not exceed those specified in Table 3.2.1-1 at lower power levels due to control rod insertion and that the DNB ratio is always greater at part power
than at full power.
The protection system ensures that the nuclear core limits are not exceeded.
3.1.3-1 Revised 02/18/2011 C25C25 Reactivity Control The control system and the operational procedures provide adequate control of
the core reactivity and power distribution. The following criteria are met:
a) Sufficient control is available to produce a hot shutdown margin of at least 1% k/k. The minimum hot shutdown margin required by Technical Specifications is available assuming at least a 7% uncertainty in
control rod worth.
b) The shutdown margin is maintained with the most reactive RCCA stuck in the fully withdrawn position.
c) The shutdown margin is maintained at ambient temperature by the use of soluble poison.
Thermal and Hydraulic
The reactor core is designed to meet the following limiting thermal and
hydraulic criteria:
a) At least a 95 percent probability with 95% confidence that DNB will not occur on the limiting fuel rods during normal operations and operational transients and during transient conditions arising from faults of
moderate frequency (condition I and II events).
b) No fuel melts during normal operation, including Condition I and II events.
To maintain fuel rod integrity and prevent fission product release, it is necessary to prevent overheating of the fuel and possible cladding perforation which would result in the release of fission products to the reactor coolant.
Overheating of the fuel cladding is prevented by restricting fuel operation to within the nucleate boiling regime where the heat transfer coefficient is large and the cladding surface temperature is slightly above the coolant
saturation temperature.
3.1.3-2 Revised 02/18/2011 C25C25 Operation above the upper boundary of the nucleate boiling regime could result in excessive cladding temperatures because of the onset of departure from nucleate boiling (DNB) and the resultant sharp reduction in heat transfer coefficient. DNB is not a directly measurable parameter during operation and therefore THERMAL POWER and Reactor Coolant Temperature and Pressure have been related to DNB. This relation has been developed to predict the DNB flux and the location of DNB for axially uniform and non-uniform heat flux distributions. The local DNB heat flux ratio, DNBR, defined as the ratio of the heat flux that would cause DNB at a particular core location to the local heat flux, is indicative of the margin to DNB. The correlation DNBR limit is established based on the entire applicable experimental data set such that
there is a 95 percent probability with 95 percent confidence that DNB will not
occur when the minimum DNBR is at the DNBR limit.
Mechanical
Reactor Internals
The reactor internal components are designed to withstand the stresses resulting from startup, steady state operation with any number of pumps running, and shutdown conditions. No damage to the reactor internals occurs
as a result of a loss of pumping power.
Lateral deflection and torsional rotation of the lower end of the core barrel is limited to prevent excessive movements resulting from seismic disturbances and thus prevent interference with rod cluster control assemblies (See Appendix 5A). Core drop in the event of failure of the normal supports is
limited so that the RCCA's do not disengage from the fuel assembly guide
thimbles.
3.1.3-3 Rev. 16 10/99 The structural internals are designed to maintain their functional integrity in the event of a major loss-of-coolant accident or a maximum hypothetical earthquake. The dynamic loading resulting from the pressure oscillations because of a loss-of-coolant accident does not prevent rod cluster control
assembly insertion.
The following components of the reactor internals were checked for buckling under the combined effect of design earthquake and a double ended pipe break:
upper barrel during hot leg break, upper and lower support columns, and fuel
assemblies thimbles, for both cold and hot leg break.
The internals were analyzed by applying to each component the excitation forces due to the transient postulated condition. Maximum stresses and deflections were obtained from the structural response and compared with allowable values. The stress analysis was performed obtaining the maximum dynamic response for each component and computing the corresponding stress
intensity using standard strength of materials formulas.
Resulting stresses were then combined in the most unfavorable manner with the seismic stresses and the maximum stress intensities were obtained for each component. The dynamic analysis has been performed using the following
conservative assumptions:
- 1. The mechanical and hydraulic analyses were performed separately without including the effect of the water-solid interaction. Peak pressures
obtained from the hydraulic analysis will be attenuated by the
deformation of the structures.
- 2. When applying the hydraulic forces no credit was taken for the stiffening effect of the fluid environment which will reduce the
deflections and stresses in the structure.
- 3. The multi-mass model was considered to have enough degrees-of-freedom to represent the most important modes of vibration in the vertical
direction. This model is conservative in the sense that further
mass-spring resolution of the system would lead to further attenuation
of the shock effects obtained with the present model.
3.1.3-4 Rev. 16 10/99 To assure that the components will not fail, an allowable stress criterion was established as explained in Section 14.3.3. This criterion limits the maximum strain to percentages of the material uniform strain which is an indication of
the adopted margin.
Uncertainties in the dynamic loads have increased the margin contained in the maximum stresses. The major factor of conservatism is the assumption that the double ended break will occur in 0.001 sec.; larger breaking times will reduce all the stresses. The hydraulic analysis, with which the loads were computed, is performed by the MULTIFLEX Code which solves hydraulic equations by the
method of characteristics.
Uncertainties in the geometric modeling technique, when applied to a full scale reactor, have been also determined analytically by studying the same reactor using models of different complexity. Results indicate that the loads
obtained with the present model are conservative.
Fuel Assemblies
The fuel assemblies are designed to perform satisfactorily throughout their lifetime. The loads, stresses, and strains resulting from the combined effects of flow induced vibrations, earthquakes, reactor pressure, fission gas pressure, fuel growth, thermal strain, and differential expansion during both steady state and transient reactor operating conditions have been considered in the design of the fuel rods and fuel assembly. The assembly is also structurally designed to withstand handling and shipping loads prior to irradiation, and to maintain sufficient integrity at the completion of design burnup to permit safe removal from the core and subsequent handling during
cooldown, storage and shipment.
3.1.3-5 Rev. 16 10/99 The fuel rods are supported at several locations along their length within the fuel assemblies by brazed grid assemblies which are designed to maintain control of the lateral spacing between the fuel rods throughout the design life of the assemblies. The magnitude of the support loads provided by the grids are established to minimize possible fretting without overstressing the cladding at the points of contact between the grids and fuel rods. The grid
assemblies also allow axial thermal expansion of the fuel rods without imposing restraint of sufficient magnitude to result in buckling or distortion
of the rods.
The fuel rod cladding is designed to withstand operating pressure loads without collapse or rupture and to maintain encapsulation of the fuel
throughout the design life.
3.1.3-6 Rev. 16 10/99 Rod Cluster Control Assemblies The criteria used for the design of the cladding on the individual absorber rods in the RCCA's are similar to those used for the fuel rod cladding. The cladding is designed to be free standing under all operating conditions and will maintain encapsulation of the absorber material throughout the absorber rod design life. Allowance for wear during operation is included in the RCCA
cladding thickness.
Adequate clearance is provided between the absorber rods and the fuel assembly guide thimbles which position the rods within the fuel assemblies so that
coolant flow along the length of the absorber rods is sufficient to remove the heat generated without overheating of the absorber cladding. The clearance is also sufficient to compensate for any misalignment between the absorber rods and fuel assembly guide thimbles and to prevent mechanical interference between the absorber rods and fuel assembly guide thimbles under any operating
and accident conditions.
Control Rod Drive Assembly
Each control rod drive assembly is designed as a hermetically sealed unit to
prevent leakage of reactor coolant. All pressure-containing components are designed to meet the requirements of the ASME Code,Section III, Nuclear
Vessels for Class A vessels.
The control rod drive assemblies provide RCCA insertion and withdrawal rates
consistent with the required reactivity changes for reactor operational load
changes. The maximum reactivity addition rate is specified to limit the
magnitude of a possible nuclear excursion resulting from a control system
malfunction or operator error.
3.1.3-7 Rev 16 10/99 Also, the control rod drive assemblies for the full length rods provide a fast insertion rate during a "trip" of the RCCA's which results in a rapid shutdown
of the reactor for conditions that cannot be handled by the reactor control system. This rate is based on the results of various reactor emergency analyses, including instrument and control delay times and the amount of reactivity that must be inserted before deceleration of the RCCA's
occurs.
3.1.3-8 Rev. 16 10/99 3.2 REACTOR DESIGN 3.2.1 NUCLEAR DESIGN AND EVALUATION
This section presents the nuclear characteristics of the core and an evaluation
of the characteristics and design parameters which are significant to design
objectives. The capability of the reactor to achieve these objectives while
performing safely under normal operational modes, including both transient and
steady state, is described.
Nuclear Characteristics of the Design
For historical purposes, a summary of the reactor nuclear design characteristics
is presented in Table 3.2.1-1. This table includes design parameters for Cycle
1 only. Subsequent cycle specific values are calculated and their impact on
plant operation and safety analyses is evaluated prior to each cycle. The
results of these evaluations for the current cycles are presented in Appendices
14A and 14B for Units 3 and 4, respectively. The cycle specific design
parameters for the current reloads are documented in nuclear design reports that
are used for design verification and operational guidance.
Reactivity Control Aspects
Reactivity control is provided by neutron absorbing RCCA and by a soluble
chemical neutron absorber (boric acid) in the reactor coolant. The
concentration of boric acid is varied as necessary during the life of the core
to compensate for: (1) changes in reactivity which occur with change in
temperature of the reactor coolant from cold shutdown to hot operating, zero
power conditions; (2) changes in reactivity associated with changes in the
fission product poisons xenon and samarium; (3) reactivity losses associated
with the depletion of fissile inventory and buildup of long-lived fission
product poisons (other than xenon and samarium); and (4) changes in reactivity
due to burnable poison burnup.
The RCCAs provide reactivity control for: (1) fast shutdown; (2) reactivity
changes associated with changes in the average coolant temperature above hot
zero power (core average coolant temperature is increased with power level);
(3) reactivity associated with any void formation; and (4) reactivity changes
associated with the power coefficient of reactivity.
3.2.1-1 Revised 04/17/2013 Chemical Shim Control Control to render the reactor subcritical at temperatures below the operating
range is provided by a chemical neutron absorber (boron). The boron
concentration during refueling has been established for Cycle 1 as shown in
Table 3.2.1-1, line 29. This concentration together with the RCCA provides
approximately 10 percent shutdown for these operations. In Reference 17 of
Section 3.2.1, the refueling shutdown margin has been revised to 5 percent K/k. The concentration is also sufficient to maintain the core shutdown without any
RCCA during refueling. For historical purposes, for cold shutdown in Cycle 1, at
the beginning of core life, a concentration (shown in Table 3.2.1-1, line 37) is
sufficient for one percent shutdown with all but one stuck rod inserted. The
boron concentration (Table 3.2.1-1, line 29) for refueling is equivalent to less
than two percent by weight of boric acid (H 3 BO 3) and is well within solubility limits at ambient temperature. This concentration is also maintained in the
spent fuel pit since it is directly connected with the refueling canal during
refueling operations.
For historical purposes, the initial full power boron concentration for Cycle 1
without equilibrium xenon and samarium is specified in Table 3.2.1-1, line 34.
As these fission product poisons are built up, the boron concentration is
reduced to that specified in Table 3.2.1-1, line 36.
This initial boron concentration is that which permits the withdrawal of the
control banks to their operational limits. The xenon-free hot zero power
shutdown (k = 0.99) with all but one stuck rod inserted, can be maintained with
the boron concentration specified in Table 3.2.1-1, line 38 (Cycle 1). This
concentration is less than the full power operating value with equilibrium
Control Rod Requirements
Neutron-absorbing RCCA provide control to compensate for more rapid variations
in reactivity. The rods are divided into two categories according to their
function. Some rods compensate for changes in reactivity due to variations in
operating conditions of the reactor such as power or temperature. These rods
comprise the control group of rods. The remaining rods, which provide shutdown
reactivity, are termed shutdown rods. The total shutdown worth of all the rods
is also specified to provide adequate shutdown with the most reactive rod stuck
out of the core.
3.2.1-2 Revised 04/17/2013 For historical purposes, RCCA reactivity requirements at beginning and end of
life are summarized in Table 3.2.1-2. The installed worth of the RCCA is shown
in Table 3.2.1-3. The difference is available for excess shutdown upon reactor
trip. Summaries of the calculated reactivity requirements and shutdown margins
for the current cycles are presented in appendices 14A and 14B for Units 3 and
4, respectively.
Total Power Reactivity Defect
RCCA must be available to compensate for the reactivity change incurred with a
change in power level due to the Doppler effect. The magnitude of this change
has been established by correlating the experimental results of numerous
operating cores.
The average temperature of the reactor coolant is increased with power level in
the reactor. Since this change is actually a part of the power dependent
reactivity change, along with the Doppler effect and void formation, the
associated reactivity change must be controlled by rods. The largest amount of
reactivity that must be controlled is at the end of life when the moderator
temperature coefficient has its most negative value. For historical purposes, the moderator temperature coefficient range for Cycle 1 is given in Table
3.2.1-1, line 42, while the cumulative reactivity change is shown in the first
line of Table 3.2.1-2. By the end of the fuel cycle, the nonuniform axial
depletion causes a severe power peak at low power. The reactivity associated
with this peak is part of the power defect. The moderator temperature
coefficient range for the current cycles is given in Appendices 14A and 14B for
Units 3 and 4, respectively.
Control Rod Bite
Control rod bite is the reactivity worth of the control rods inserted into the
core during power operation to enable a more rapid reactivity control response
with control rod motion. Operation with control rod bite is typically used for
a load follow operation strategy when rapid load variations require quick
response with higher reactivity ramp rates with the rods in automatic control.
However, consistent with a base load full power operation strategy, the Turkey
Point units currently have the auto withdrawal feature of the rod control system
disabled and employ an all rods out (ARO) full power operation strategy that
eliminates the need for and the disadvantages of long term operation with a
control rod bite. Nevertheless, the text below provides a discussion of
operation with control rod bite.
3.2.1-3 Revised 04/17/2013 To facilitate load follow operations and partial load rejection transients, the
reactor control system is designed to accommodate the changes in power and
temperature from a 10% step load increase or decrease, a ramp load increase or
decrease of 5%/minute and a 50% load rejection without safeguards actuation or
reactor trip. The ability of the unit to accept major load variations is
distinct from safety considerations, since the reactor would be tripped and
shutdown safely if the rods could not follow the imposed load variations.
During times in core life when the at power moderator temperature coefficient (MTC) is sufficiently negative, a quick response with higher reactivity ramp
rates with the rods in automatic control is not as critical to managing these
load variations. The disabling of control rod auto withdrawal only impacts the
10% step load increase from 90% power transient. In this case, provided that the MTC is sufficiently negative (more negative than -5.0 pcm/
ºF), the operators would have sufficient time to respond and bring the plant back to equilibrium without safeguards actuation or reactor trip. Analysis of all the design load
variation transients has shown that for the typical 18 month fuel cycle core
designs the reactor can be controlled with rod withdrawal in manual such that
safeguards actuation or reactor trip would not occur.
To more easily accommodate load follow operation with rods in automatic control (insertion and withdrawal) during times when the MTC is less negative, one
control bank of rods can remain inserted about 10 percent into the core at the
beginning of life. The control rod bite position for D bank in steady state
operation is about 210 steps withdrawn to about 220 steps withdrawn.
Temperature control is still adequate at the bite position. The reactivity
associated with this bite is less than 0.1 percent.
For base load operation the apparent advantages of leaving rods inserted at the
bite position are outweighed by the disadvantages. The ability to add
reactivity to the core quickly by withdrawing rods is not required during long
term steady state operation at full power. The ability to change flux
difference in the positive direction is similarly not required; indeed, the flux
difference is more stable with rods fully withdrawn.
The major disadvantage of operation with rods inserted to the bite position is
the shadowing of fuel burnup in the top of the core. This leads to relatively
worse axial power distributions in the subsequent fuel cycles and will restrict
the permissible flux difference operating band. A small effect is a loss of
reactivity and, therefore, a reduction in cycle lifetime in the current cycle
due to less than optimum axial burnup distribution.
3.2.1-4 Revised 04/17/2013 Further, the consequences of many accidents are actually worse starting from
deep rod insertion, even though these worst cases have already been assumed in
the accident analysis for short term operation with rods inserted.
Xenon Stability Control
Excore instrumentation is provided to obtain necessary information concerning
power distribution. This instrumentation is adequate to enable the operator to
monitor xenon induced power oscillations. Extensive analysis, with confirmation
of methods by spatial transient experiments at Haddam Neck, has shown that any
induced radial or diametrical xenon transients would die away naturally. A full
discussion of xenon stability control can be found in Reference 1.
Excess Reactivity Insertion Upon Reactor Trip
The control rod requirements have been based on providing one percent shutdown
at hot zero power (HZP) conditions with the highest worth rod stuck in its fully
withdrawn position. The condition where excess reactivity insertion is most
critical is at the end of a cycle when the steam break accident is considered.
Calculated Rod Worths
The complement of 45 full length rods arranged in the pattern shown in Figure
3.2.1-1 meets the shutdown requirements. Table 3.2.1-3 lists the calculated
worths of this rod configuration for beginning and end of the first cycle. In
order to be sure of maintaining a conservative margin between calculated and
required rod worths the calculated reactivity worths listed are decreased in the
design by at least 7 percent to account for any errors or uncertainties in the
calculation. This worth is established for the condition that the highest worth
rod is stuck in the fully withdrawn position in the core. The reactivity
requirements and shutdown margins evaluation presented in Appendices 14A and 14B
lists the calculated worths for beginning and end of the current cycles for
Units 3 and 4, respectively.
A comparison between calculated and measured rod worths in the operating reactor
shows the calculation to be well within the allowed uncertainty of at least 7
percent. (Reference 19)
3.2.1-5 Revised 04/17/2013 Reactor Core Power Distribution
In order to meet the performance objectives, without violating safety limits, the peak to average power density must be within the limits set by the nuclear
hot channel factors. For the peak power point in the core, the nuclear heat
flux hot channel factor, F N q , was established as specified for Cycle 1 in Table 3.2.1-1, Line 18. For the hottest channel the nuclear enthalpy rise hot channel
factor, F NH , was established as specified in Table 3.2.1-1, Line 19. Variation in hot channel factors is illustrated for typical rod configurations for the
first cycle in Figures 3.2.1-3 and 3.2.1-5. The calculations shown in these
figures do not include the power flattening effect of equilibrium xenon and
non-uniform Doppler broadening. Reactivity feedback effects associated with
non-uniform xenon, water boron density and Doppler broadening are discussed in
detail in Reference 1. Nuclear hot channel factors for the current cycles are
provided in Appendices 14A and 14B for Units 3 and 4, respectively.
Incore instrumentation is employed to check the power distributions throughout
core lifetime.
The excore nuclear instrumentation system supplies the information on core power
distribution. This information is derived from four independently operating
channels. Each channel employs a dual section long ion chamber for monitoring
the upper and lower section of the core. Current signals from these detectors
are summed, conditioned and calibrated against incore power distribution
obtained from the movable incore detector system so that the eight individual
signals are directly related to the power generated in the adjacent section of
the core. This essentially divides the core into eight sections, four in the
upper half, and four in the lower half.
The relationship between core power distribution and excore nuclear
instrumentation reading is established during the startup testing program.
Incore flux measurements are made for reactor power in the range of 25 to 100
percent. These measurements, together with long ion chamber currents, are
processed to yield the relationships between power distribution parameters
calculated with an incore flux map and those calculated with excore nuclear
instrumentation. These relationships can be checked during operation to assess
the effect of core burnup on the sensitivity between incore power distribution
and excore readings.
3.2.1-6 Revised 04/17/2013 Power Distribution Control Limits placed on the axial flux difference are designed to assure that the heat
flux hot channel factor Fq is maintained within acceptable limits. The Constant
Axial Offset Control (CAOC) operating procedure described in Reference 1
requires control of the axial flux difference at all power levels within a
permissible operational band about a target value corresponding to the
equilibrium full power value. The Relaxed Axial Offset Control (RAOC)
procedures which were implemented in Unit 3 Cycle 13, Unit 4 Cycle 14, and
beyond (described in Reference 18) were developed to provide wider control bands
and, consequently, more operating flexibility. These wider operating limits, particularly at lower power levels, can increase plant availability by allowing
quicker plant startups and increased maneuvering flexibility without trip or
reportable occurrences.
In the standard CAOC analysis described in Reference 1, the generation of the
normal operation power distribution is constrained by the Rod Insertion Limit (RIL) and the axial flux difference (AFD) band limits. The purpose of the RAOC
is to find the widest possible AFD-power operating space by analyzing a wide
range of axial flux differences. Therefore, the generation of normal operation
power distribution is constrained only by the RIL.
For a CAOC analysis, load follow simulations were performed covering the allowed
CAOC operating space to generate a typical range of allowed axial xenon
distributions, which in turn were used to calculate axial power distribution in
both normal operation and Condition II accident conditions. For an RAOC
analysis, however, as described in Reference 18 a more practical method is used
to create an axial xenon distribution covering the wider AFD-power operating
space allowed with RAOC operation. Each resulting power shape is analyzed to
determine if LOCA constraints are met or exceeded. The total peaking factor, F T q, is determined using standard synthesis methods as described in Reference 1.
Following the guidance of Generic Letter 88-16, the RAOC AFD limits were removed
from the Technical Specifications and placed in the Core Operating Limits Report (COLR). These reports are presented in Appendix 14A and 14B for Units 3 and 4, respectively.
3.2.1-7 Revised 04/17/2013 Developments with regard to Emergency Core Cooling System (ECCS) criteria for LOCA have imposed new requirements on allowable k W/ft conditions. A schematic demonstration of the various limits and their effect on allowable local power densities (k W/ft), and hence operational flexibility, is presented in Figure 3.2.1-6. This figure shows that there are many limits that must be met, but also that these limits are, in general, a function of the elevation in the core.
Allowable local k W/ft limits are lower near the top of the core because of, for example, the axial power shape dependence on DNB and the reduced heat transfer upon re-flood for a large break LOCA near the top of the core.
Reactivity Coefficients
The response of the reactor core to unit conditions or operator adjustments
during normal operation, as well as the response during abnormal or accidental
transients, is evaluated by means of a detailed plant simulation. In these
calculations, reactivity coefficients are required to couple the response of the
core neutron multiplication to the variables which are set by conditions
external to the core. Since the reactivity coefficients change during the life
of the core, a range of coefficients is established to determine the response of
the unit throughout life.
Moderator Temperature Coefficient
The moderator temperature coefficient relates a change in neutron multiplication
to the change in reactor coolant temperature. Reactors employing soluble boron
as a means of reactivity control possess less negative moderator temperature
coefficients than cores controlled solely by RCCA's. There are two reasons for
this:
a) Soluble poison density is decreased with the water density when the coolant temperature rises; and
b) In a chemical shim core the control rods are only partially inserted.
A deep insertion tends to increase effective length of the core and
thus causing moderator coefficient to become more negative.
In order to reduce the dissolved poison requirement for control of excess
reactivity, burnable poison rods can be incorporated in the core design. The
result is that changes in the coolant density will have less effect on the
density of poison and the moderator temperature coefficient will become less
positive
3.2.1-8 Revised 04/17/2013 C26 For historical purposes, in Cycle 1 there were 816 of the borosilicate glass
rods in the form of clusters distributed throughout the core in vacant rod
cluster control guide tubes as illustrated in Figures 3.2.1-7 and 3.2.1-8.
Information regarding research, development and nuclear evaluation of the
burnable poison rods can be found in Reference 3 and 4. These rods initially
controlled the installed excess reactivity shown on lines 40 and 41 of Table
3.2.1-1, and their addition resulted in a reduction of the initial hot zero
power boron concentration in the coolant to the value shown on line 34. The
moderator temperature coefficient was negative at the operating coolant
temperature with this boron concentration and with burnable rods installed.
In a typical reload cycle, several hundreds of Wet Annular Burnable Absorber (WABA) rods in the form of clusters are distributed throughout the core in
vacant rod cluster control guide tubes. Additionally, recent cores utilize
several thousands of Integral Fuel Burnable Absorbers (IFBAs) in the form of a
zirconium diboride coating on the surface of the fuel palette. The number and
distribution of these burnable absorbers for the current cycles are presented in
Appendices 14A and 14B for Units 3 and 4, respectively.
The effect of burnup on the moderator temperature coefficient is calculated and
the coefficient becomes more negative with increasing burnup. This is due to
the buildup of fission products with burnup and dilution of the boron
concentration with burnup. The latter effect is considerably more important.
However, the buildup of equilibrium xenon contributes a positive increment to
the coefficient for a constant boron concentration. For historical purposes, the calculated net effect and the predicted unrodded moderator temperature
coefficient with equilibrium xenon at BOL for Cycle 1 is shown in table 3.2.1-1, Line 42. With core burnup, the coefficient will become more negative as boron
is removed because of a shift in the neutron energy spectrum due to the buildup
of plutonium and fission products. For Cycle 1 at end of life with no boron or
rods in the core, the moderator coefficient is specified in Table 3.2.1-1, Line
- 42.
The current Technical Specifications allows a +5 pcm/ F (+5 x 10
-5 K/K/ F) MTC below 70 percent of rated power, ramping to a 0 pcm/ F MTC at 100 percent power and above. A power-level dependent MTC was chosen to minimize the effect of the specification on postulated accidents at high power levels.
Moreover, as the power level is raised, the average core water temperature
becomes higher as allowed by programmed average temperature for the plant, tending to bring the moderator coefficient more negative.
3.2.1-9 Revised 04/17/2013 Also, the boron concentration can be reduced as xenon builds into the core.
Thus, there is less need to allow a positive coefficient as full power is
approached. As fuel burnup is achieved, boron is further reduced and the
moderator coefficient will become negative over the entire operating power
range.
The control rods provide a negative contribution to the moderator temperature
coefficient as can be seen in Figure 3.2.1-9 for Cycle 1. Moderator temperature
coefficients for current cycles are given in Appendices 14A and 14B for Units 3
and 4, respectively.
Moderator Pressure Coefficient
The moderator pressure coefficient has an opposite sign to the moderator
temperature coefficient. The effect on the total coefficient is small because
the pressure coefficient is 100 times smaller. For historical purposes the
calculated beginning and end of life pressure coefficients for Cycle 1 are
specified in Table 3.2.1-1, Line 43.
Moderator Density Coefficient
A uniform moderator density coefficient is defined as a change in the neutron
multiplication per unit change in moderator density. The range of the moderator
density coefficient from BOL to EOL for Cycle 1 is specified in Table 3.2.1-1, Line 44. The most positive moderator density coefficient is calculated as part
of the reload safety analyses evaluation for each cycle and compared to the
limit included in Appendices 14A and 14B for Unit 3 and 4, respectively.
Doppler and Power Coefficients
The Doppler coefficient is defined as the change in neutron multiplication per
degree change in fuel temperature. The coefficient is obtained by calculating
neutron multiplication as a function of effective fuel temperature. The results
are shown in Figure 3.2.1-10 for BOL conditions for the first cycle. The
coefficient becomes slightly more negative with increasing fuel burnup. Doppler
coefficients for current cycles are given in Appendices 14A and 14B for Units 3
and 4, respectively.
In order to know the change in reactivity with power, it is necessary to know
the change in the effective fuel temperature with power as well as the Doppler
coefficient. It is very difficult to predict the effective temperature of the
fuel using a conventional heat transfer model because of uncertainties in
predicting the behavior of the fuel pellets.
3.2.1-10 Revised 04/17/2013 Therefore, an empirical approach is taken to calculate the power coefficient, based on operating experience of existing Westinghouse cores. For historical
purposes, Figure 3.2.1-11 shows the power coefficient as a function of power
obtained by this method for non-collapsed helium filled rods for Cycle 1. The
results presented for BOL do not include any moderator coefficient even though
the moderator temperature changes with power level, and the coefficient becomes
slightly more negative with increasing fuel burnup.
As the fuel pellet temperature increases with power, the absorption in U-238
increases due to Doppler broadening of the resonances. A large temperature drop
occurs across the fuel pellet-clad gap. Under certain conditions, this gap may
be closed, thus resulting in lower pellet temperature. The net effect is a
lower effective fuel temperature, a higher Doppler coefficient, and a lower
power coefficient than that which exists with a pellet-clad gap. The power
coefficient, which is determined using a closed gap model, is shown in Figure
3.2.1-12 for Cycle 1.
Calculations indicate the stability of the reactor to Xenon oscillations is
relatively insensitive to the thermal model used to obtain the power
coefficient. The damping factor associated with the fuel Doppler effect is
f = K eff T T P where T = fuel temperature P = power
K eff = reactivity T The quantity P is larger for the gap model than for the no gap case but since the Doppler coefficient varies as T
-1/2 the term K eff is smaller. The net T effect is that f is relatively insensitive to the thermal model in the range of power 0.5 to 1.5 of core average which is the range of interest for stability.
3.2.1-11 Revised 04/17/2013 Nuclear Evaluation The basis for confidence in the procedures and design methods comes from the
comparison of these methods with many experimental results and actual measured
data from Turkey Point over numerous cycles of operation (Reference 19). These
experiments include criticals performed at the Westinghouse Reactor Evaluation
Center (WREC) and other facilities, and also measured data from operating power
reactors. A summary of the results and discussion of the agreement between
calculated and measured values is given in the following paragraphs and also
documented in References 19 and 20.
Reactivity Analysis
Data from 55 oxide and 56 metal lattice critical and exponential experiments
have been evaluated (5). The results of these studies are summarized in Table 3.2.1-4. The values of neutron multiplication k are computed using
experimentally measured material bucklings, and should equal unity. Table
3.2.1-4 demonstrates that much of the scatter can be attributed to variations in
results from one experimental laboratory to another, whereas the evaluation
demonstrates that errors do not develop with variations of certain significant
parameters. As the calculational accuracy is independent of variations in
hydrogen to uranium ratio, uranium enrichment, pellet diameter and buckling,
extrapolation from experiments to operating cores or extrapolation from one
operating core to another does not lead to any significant error.
It can be seen from Table 3.2.1-4 that if only W APD experimental results are considered, the computational method predicts k to a standard deviation of 0.36
percent which is a better estimate of the accuracy of the method because of the
more detailed information available. Much of the additional scatter in the
standard deviation for the other cases can be attributed to insufficient
information on the dimensions and results of many of the cases published.
Depletion Analysis Data from the Yankee Core Evaluation Program have been compared with calculated data using the design techniques. The results are summarized in Figures
3.2.1-13 through 3.2.1-15. Uranium depletion and net plutonium production have
a direct bearing on the core lifetime. The figures show the comparison between
calculations (solid lines) and measured concentrations of the various isotopes.
Although some small deviations can be observed between analysis and experiment, they are considered negligible.
3.2.1-12 Revised 04/17/2013 Power Peaking Analysis
A series of critical experiments were carried out at the Westinghouse Reactor
Evaluation Center (WREC) to determine the power peaking in fuel rods adjacent to
water holes and to determine the effects of voids on power distribution.
The power peaking experiment was performed in a 30 x 30 array of 2.72 percent
enriched fuel with a water-to-uranium ratio of 3.5 with and without boron in the
moderator. The pattern of 16 water holes was symmetrical about the center of
the core. The core arrangement and pattern of fuel rods scanned are shown in
Figure 3.2.1-16 for the unborated core and Figure 3.2.1-17 for the same core
with 479 ppm boron in the water.
The analysis consists of PDQ calculations using two-group constants obtained
from LEOPARD. Mixed Number Density thermal constants are used, and "soft
spectrum" microscopic constants are used in the reflector and water holes. In
the PDQ analysis, two mesh spacings per fuel rod are used. Also, in the
unborated core a calculation is performed for one mesh space per fuel rod. The
experimental data are normalized to the PDQ results using the average of the
four central rods. The experimental and calculated results for the borated and
unborated cores with two mesh spacers per fuel rod are shown in Figures 3.2.1-18
and 3.2.1-19, respectively, and in Figure 3.2.1-20 for the unborated core
calculated with one mesh spacer per fuel rod. Each block in the figures
represents a fuel rod. The experimental values correspond to the average values
of counts taken at five positions on the fuel rod. The agreement between
analysis and experiment is within 2 to 3 percent and is of the same order as the
scatter in the experimental data. There is no consistent difference in
over-estimating or underestimating peaking using the one mesh per fuel rod or
two mesh per fuel rod representation.
The void experiments were performed for two different core configurations. The
first series of experiments was carried out in a 47 x 47 square core of 2.7%
enriched fuel with a W/U of 2.9, with no boron. The second series was performed
using a 53 x 53 square core of 3.7% enriched fuel with a W/U of 2.9, and with
1046 ppm boron in the water. In both cores voids were simulated by empty 0.1875
inch O.D. 0.022 inch wall aluminum tubes inserted between fuel rods. The
moderator in the voided region consisted of 11.52% aluminum, 16.29% void and
72.19% water. Data were taken for the following cases:
- 1. No void tubes
- 2. Four void tubes (2x2) located around the central fuel rod
- 3. Sixteen void tubes (4x4) at core center
- 4. One hundred ninety-six void tubes (14x14) are core center
3.2.1-13 Revised 04/17/2013 The analysis again consisted of PDQ using two-group constants from LEOPARD, with
MND thermal constants and "soft spectrum" water hole and reflector constants.
The calculated power distribution is compared with the experimental power scans
in Figure 3.2.1-21 and 3.2.1-22 for the unborated and borated cores for the four
cases examined. The agreement between experiment and calculation is good except
at the transition region between voided and non-voided regions. Here the
calculated peaks are higher than those obtained by experimental measurements.
The reactivity effects of the void tubes were calculated assuming a constant
axial reflector savings. Calculation and experiment for each case examined are
compared in Table 3.2.1-5. Calculations overestimate the reactivity effect of
the voids by approximately 10%, which is good agreement in view of the small
magnitude of the effects being studied.
The adequacy of the current methods for peaking factor analysis is demonstrated
in the Turkey Point specific evaluations of Reference 19 by comparison of
prediction to actual measured values obtained using the flux map analysis code, INCORE. The nuclear enthalpy rise hot channel factor (FH<see FSAR Table 3.2.1-5> ) and the heat flux hot channel factor (FQ) were measured using the INCORE
code. Predicted peaking factors were obtained from three-dimensional ANC
calculations performed for core conditions similar to those at the time of the
measurements. Power peaking factors measured during Unit 4 Cycles 12, 13, and
14 are compared to predicted values in Reference 19. For FH the mean difference between the measured and predicted values for the three cycles is
2.02% with a standard deviation of 1.27%; for FQ the mean difference is 3.33%
with a standard deviation of 1.86%.
Gross Power Distribution Analysis The ability to evaluate power distributions in multiregion critical cores with no burnup has been evaluated in detail.(6) Agreement for all situations, including those with large enrichment variation and small regions, is found to
be good as is illustrated in Figure 3.2.1-23. The ability to evaluate power
distributions in depleted cores at power has been demonstrated by core
evaluation programs using in-core instrumentation data from Yankee, Saxton and
SELNI.
As an example of such a comparison, a power distribution is shown in Figure
3.2.1-24 for the end of life in Yankee Core I, which was not controlled by
chemical shim. A comparison of the burnup distribution is also presented in
Figure 3.2.1-25.
In both cases two calculated values are given which show the effect of a rod
program interchange during life.(2)
3.2.1-14 Revised 04/17/2013 The ability to evaluate power distributions in complex designs is
demonstrated in the Reference 20 qualification of advanced design methods.
These methods have been qualified against actual measured power
distributions from the moveable incore detector system. The qualification
report of Reference 20 indicates good agreement of average assembly power
distribution between measured and predicted results.
RCC Assembly Worth Analysis In the control rod calculations performed by PDQ, the RCC assemblies are
represented by internal boundary conditions ('s) in the fast and thermal groups. These boundary conditions applied to the unit cell in which the
absorber rod, its clad and the associated water are homogenized. The
values of these 's are determined to make the calculated rod worth of a single fuel assembly equal to that calculated by a more refined model.
The better model represents each absorber rod explicitly and is used to
analyze an extensive set of critical measurements. Approximately 30
different critical measurements were made for uniform and cluster arrays
of absorber rods with different enrichments, rod diameters, water-to-uranium ratios and boron concentrations.
In the analysis of these measurements, the rods were represented by a
theoretically determined thermal boundary condition and by a diffusion
region in the single fast group. The fast absorption cross section was
empirically determined from the measured rod worth to give agreement
between analytical and experimental results. The development of this
calculation scheme for rod worth and a description of the measurements is
given in Reference 8. Figures 3.2.1-26 and 3.2.1-27 are reproduced from
this reference to show the fast absorption cross section as a function of
the radius of the absorber which fits the experimental measurements for
cluster and uniform cases, respectively. The solid lines were obtained by
at least square fitting of the experimental data.
The ability of the current design methods to accurately predict control
rod worth is demonstrated in References 19 and 20. Reference 19 compares
measured and predicted worths for Turkey Point Unit 4 Cycles 12,13,and 14;
while Reference 20 presents similar comparisons for four different plants
with a total of seven cycles. The mean difference between measured and
predicted worth in the operating reactor shows the calculation to be well
within the uncertainty of 7 percent.
3.2.1-15 Revised 04/17/2013 Moderator Coefficient Analysis Inasmuch as the safe operation of any nuclear unit is closely associated with
the ability to predict the behavior of that unit, correlation of analysis with
experiment is presented to show that the moderator temperature coefficient is
quite predictable. Measurements were made during the startup and operation of
the SELNI core to get data for a core controlled by chemical shim. During the
startup, the core was heated from room to operating temperature at a constant
boron concentration of 1600 ppm. Figure 3.2.1-28 shows the results of the
moderator coefficient measurements taken during this core heatup, and also the
comparable calculated values. The calculations were performed with the
one-dimensional AIM-5 code with LEOPARD input constants as described for neutron
multiplication calculations. The agreement between calculation and experiment
is good over the entire temperature range.
In order to measure the moderator coefficient at different boron concentrations, control rods were traded for boron during the hot, no power startup tests. This
procedure permitted moderator coefficient measurements to be made over a range
of boron concentrations from 1300 to 1800ppm. The method of analysis for the
case of trading rods for boron is, of necessity, different from the method
discussed above. The AIM-5 code was again used, but an axial calculation was
performed with an homogenized bank of absorber used to represent the moving
RCCA. The results of analysis and measurement are shown in Figure 3.2.1-29.
The calculations were performed in the same manner as the measurement; i.e., the
control group was inserted as boron was removed.
When the control group was fully inserted, further boron removal was compensated for by insertion of all rods banked. PDQ analyses were also performed for the
all rods in and all rods out end points and the results are given in Figure
3.2.1-29. It can be seen that the one-dimensional calculations in which rods
are represented by a homogenized absorber predict the measured data very well.
The effect of burnup on the moderator coefficient has been measured in the core evaluation program performed on Yankee Core I.
(9) Yankee Core I was controlled by cruciform blade rods, and so it was necessary to separate the effect of control
rods from the effect of burnup on the moderator coefficient. Figure 3.2.1-30
illustrates these components and the agreement between analysis and measurement.
The effect of rods was evaluated by treating the rods as an equivalent
absorption area (approximation 1 in Figure 3.2.1-30) with a correlation for the
effects of resonance absorption (approximation 2 in Figure 3.2.1-30). The
results of the analysis lie within the experimental uncertainty and the burnup
effect on the moderator coefficient results in a more negative coefficient with
increasing burnup.
3.2.1-16 Revised 04/17/2013 The isothermal temperature coefficient (ITC) is defined as the change in reactivity per 0 F change in moderator and fuel temperature. This quantity is significant in HZP measurements because it can be measured directly and it is
used to determine compliance with the moderator temperate coefficient Technical
Specification.
The ITCs are measured by a series of heatup and cooldown sequences over a small change in reactor coolant system temperature. ITCs were predicted by uniformly
varying the core temperature by +/-5 0 F about the HZP temperature in the PHOENIX-P/ANC design methods of Reference 20. The qualification presented in this
reference analyzed a variety of control rod positions in addition to the basic
ARO configuration. Specific comparisons for Turkey Point are presented in
Reference 19.
Doppler and Power Coefficient Analysis
As the fuel pellet temperature increases with power, the resonance absorption in U-238 increases due to Doppler broadening of the resonances. In order to
predict the reduction in reactivity caused by this effect, it is necessary to
know the temperature of the fuel as a function of power level, the position of
burnup of fuel in the core, as well as the radial distribution of temperature
within the individual fuel rods. However, uncertainties arise during operation
at power which make it difficult to predict accurately the temperature of the
fuel pellet. For example, pellets do not remain intact (i.e., uncracked) and in
a concentric relationship with the clad, as has been observed from the Yankee
spent fuel analysis.
(10) In addition, the composition of gases in the gap changes with the burnup because of diffusion of fission product gases to the gap. This
generally results in an uncertainty in the temperature drop across the gap as a
function of power level and burnup.
A semi-empirical model has been developed for calculating the effective fuel
temperature (T eff) based on fitting the measured power coefficients of the Yankee, Saxton, BR-3 and SELNI reactor cores. The measured power coefficient
1/k k/P can be written (1)
The first term in the product on the right side of the Equation (1) is the Doppler coefficient which can be computed without knowing the heat transfer
behavior of the fuel pellet or the relationship of T eff and power. The second term on the right side of Equation (1) can then be related to the measured
values of power coefficients.
3.2.1-17 Revised 04/17/2013 P effT effTk k 1 pk k 1 In this manner an empirical expression for the effective fuel temperature is obtained which makes it possible to relate T eff to power, and thus calculate the power coefficient.
The method of analysis described in the preceding paragraph assumes accuracy of
prediction of the Doppler coefficient as a function of the effective fuel
temperature. This assumption indicates that the behavior of the U-238 resonance
integral with a change in the fuel temperature is well known. Data is presented
here to support this assumption. A correlation has been developed for the U-238
resonance integral which is known as the metal-oxide correlation.
(5) This correlation has been found to agree with Hellstrand's uranium metal (11) and uranium dioxide (12) correlations for isolated rods. The correlation is also consistent with Hellstrand's temperature correlations.
(13) Thus, a single correlation replaces the four Hellstrand correlations. The metal-oxide
correlation is
R.I.28 = 2.16X + 1.48 + (0.0279X - 0.0537) T eff 1/2 where T eff is in degrees Kelvin and
so = scattering cross section of the fuel (10.7 barns for uranium and 3.8 barns for oxygen)
N o 28 = U-238 number density in the fuel region
1 o = mean chord length in the fuel
D = shielding factor (calculated by Sauer's Method)
(14)
P o = 1 - P c (P c is tabulated in Reference 15)
This form of the resonance integral is not strictly rigorous, but its validity
is demonstrated in Figure 3.2.1-31 where it is compared with Hellstrand's
results for different temperatures.
(5)
An extensive evaluation of power coefficient measurements has been made for the
Yankee, Saxton, Br-3 and SELNI cores.
3.2.1-18 Revised 04/17/2013 1/2 28 o N o 1 D oP 28 o N so x
The results of these measurements are given in Figure 3.2.1-32 which shows the change in the effective fuel temperature per k W/ft as a function of core average k W/ft. From this data an empirical equation for T eff has been developed which will predict T eff as a function of power level.
(16) This equation for T eff is given below. _
_ T eff (P/P o) = 0.55 T fuel + (q") q" + 1.571 P/P o T o (clad + film) + T coolant where
P/P o = fraction of full power
T fuel = difference between maximum and surface fuel pellet temperature (function of power)
_
(q") = Empirical parameter dependent upon average heat flux
= ratio of the cold diametrical gap to the inner diameter of the clad
_
q" = average surface heat flux to the pellet
T o (clad + film) = temperature drop across clad and film (function of power)
T coolant = average temperature of the coolant (function of power)
The empirically determined is given in Figure 3.2.1-33 as a function of pellet surface heat flux. The difference in the effective temperature obtained from the experimental data of Figure 3.2.1-32 and from the correlation employing
Figure 3.2.1-33 is shown in Figure 3.2.1-34 as a function of surface heat flux.
It can be seen that even though there is some scatter in the experimental data (Figure 3.2.1-32), all the experimental points fall into a small band when the
T eff correlation is used. The most scattered experimental data points deviate from the predicted value (solid line) by no more than +
80 F. It is concluded that the T eff correlation can predict T eff at any power level to within +
80 F which constitutes less than +
5% of the effective fuel temperature at full power.
Although the experimental data discussed above continues to be the basis for
currents methods, enhanced modeling schemes have been developed for the
calculation of the effective fuel temperatures used in the PHOENIX-P/ANC nuclear
design system. The model also includes 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 effective
temperatures of U-238 and Pu-240 are obtained by appropriate radial weighting of
the temperature distribution. These modeling enhancements have lead to overall
high accuracy of the PHOENIX-P/ANC nuclear design system as demonstrated in
Reference 20.
3.2.1-19 Revised 04/17/2013 C26 Comparison of Predicted and Measured Boron Concentrations For historical purposes, core startup data obtained from operating power
reactors prior to Turkey Point operation are shown in Table 3.2.1-6. Comparison
of the predicted and measured critical boron concentration indicates differences
of about 50 to 150 ppm boron, or approximately a 5% overestimate, for which
allowance has been made in design calculations.
To demonstrate the ability of the current design methods, measured boron
concentration at startup was collected from operating reactors. The analysis
included in Reference 20 used data from four different reactors with a total of
seven cycles. Specific comparisons for Turkey Point Unit 4 are presented in
Reference 19. Differences between measured and predictions for the HZP, ARO
critical boron concentration are within the review criteria of 50 ppm for this
parameter.
3.2.1-20 Revised 04/17/2013 3.
2.1 REFERENCES
- 1. WCAP-7208 (1968), "Power Distribution Control of Westinghouse Pressurized Water Reactors," PROPRIETARY. A NON-PROPRIETARY version of this report is
- 2. McGaugh, J.D., and Chastain, R.H., "Power Density and Burnup Distributions in Yankee Core I," WCAP-6051 (1963), NON-PROPRIETARY.
- 3. Wood, P.M. Bassler, E.A., et al, "Use of Burnable Poison Rods in Westinghouse Pressurized Water Reactors," WCAP-7113 (October 1967),
NON-PROPRIETARY.
- 4. WCAP-9000-L Revision 1 (1969), "Nuclear Design of Westinghouse Pressurized Water Reactors with Burnable Poison Rods", PROPRIETARY. A NON-PROPRIETARY
version of this report is WCAP-7806.
- 5. Strawbridge, L.E., "Calculations of Lattice Parameters and Criticality for Uniform Water Moderated Lattices," WCAP-3269-25 (1963) NON-PROPRIETARY.
- 6. Eich, W.J., and Kovacik, W.P., "Reactivity and Neutron Flux Distribution Studies in Multi-region Loaded Reactor Cores," WCAP-1433 (1961),
NON-PROPRIETARY.
- 7. Barry, R.F., "The Revised Leopard Code - A Spectrum Dependent Non-Spatial Depletion Program," WCAP-2759, March 1965, PROPRIETARY.
- 8. Sha, W.T., "An Analysis of Reactivity Worth of the Rod Cluster Control (RCC) Elements and Local Water-Hole Power Density Peaking," WCAP-3269-47
(1965), NON-PROPRIETARY.
- 9. Poncelet, C.G., "Effects of Fuel Burnup on Reactivity and Reactivity Coefficients in Yankee Core I," WCAP-6076 (1965) NON-PROPRIETARY.
- 10. "Yankee Core Elevation Program Quarterly Progress Report for the Period Ending June 30, 1963," WCAP-6055 (1963), NON-PROPRIETARY.
- 11. Hellstrand, E., and Lundgren, G., "The Resonance Integral for Uranium Metal and Oxide," Nuclear Science and Engineering 12, 435, (1962).
- 12. Hellstrand, E., J. Applied Physics 28, 1493 (1957).
3.2.1-21 Revised 04/17/2013 3.
2.1 REFERENCES
(Continued)
- 13. Hellstrand, E., Blomberg, P., and Horner, S., "The Temperature Coefficient of the Resonance Integral for Uranium Metal and Oxide," Nuclear Science and Engineering 8, 497 (1960).
- 14. Sauer, A., "Approximate Escape Probabilities, "Nuclear Science and Engineering" 16, 329 (1963).
- 15. Case, K.M., deHoffman, F., and Placzek, G., "Introduction to the Theory of Neutron Diffusion," (1953).
- 16. Sha, W.T., "An Experimental Evaluation of the Power Coefficient in Slightly Enriched PWR Cores," WCAP-3269-40 (1965), NON-PROPRIETARY.
- 17. Gordon E. Edison (NRC) to W.F. Conway (FPL), "Turkey Point Units 3 and 4 Issuance of Amendments RE: Refueling Shutdown Margin", dated July 18, 1988.
- 18. R.W. Miller, et al., "Relaxation of Constant Axial Offset Control, F Q Surveillance Technical Specification," WCAP-10216-P-A,Revision 1 dated June 1983.
- 19. Nuclear Physics Methodology for Reload Design of Turkey Point and St.
Lucie Nuclear Plants, FPL Report No. NF-TR-95-01, January 1995 (NRC
Approved).
- 20. Nguyen, T.O., et al., "Qualification of the Phoenix-P/ANC Nuclear Design
System for Pressurized Water Reactor Cores," WCAP-11596-P-A (Proprietary),
June 1988.
- 21. Davidson, S.L. (Ed), et al, "Extended Burnup Evaluation of Westinghouse Fuel," WCAP-10125-P-A (Proprietary), December 1985, and Bahr, K.E., "Extended Burnup Evaluation of Westinghouse Fuel, Revision to Design Criteria," WCAP-10125-P-A, Addendum 1, Revision 1-A, MAY 2005.
3.2.1-22 Revised 04/17/2013 C26C26 TABLE 3.2.1-1 Sheet 1 of 3
NUCLEAR DESIGN DATA (FIRST CYCLE)
STRUCTURAL CHARACTERISTICS
- 1. Fuel Weight (UO 2), lbs. 176,000
- 2. Zircaloy Weight, lbs.
34,900
- 3. Core Diameter, inches 119.7
- 4. Core height, inches 144 Reflector Thickness and Composition
- 5. Top - Water Plus Steel 10 in.
- 6. Bottom - Water Plus Steel 10 in.
- 7. Side - Water Plus Steel 15 in. 8. H 2 O/U Volume Ratio (cold) 4.18
- 9. Number of Fuel Assemblies 157
- 10. UO 2 Rods per Assembly 204
PERFORMANCE CHARACTERISTICS
- 11. Core Heat Output, MWt (initial rating)*
2200
- 12. Heat Output, MWt 2300 (maximum calculated turbine rating)*
- 13. Fuel Burnup, MWD/MTU
First Cycle (Average) 13,000
Enrichments, w/o*
- 14. Region 1 1.85
- 15. Region 2 2.55
- 16. Region 3 3.10
- 17. Equilibrium 3.10
- 18. Nuclear Heat Flux Hot Channel Factor F N* 3.13 q 19. Nuclear Enthalpy Rise Hot Channel Factor, F NH
- 1.75
- Corresponding values for current cycles are given in Appendices 14A and 14B
for Units 3 and 4 respectively.
Revised 02/18/11 C25C25 TABLE 3.2.1-1 Sheet 2 of 3
CONTROL CHARACTERISTICS FOR FIRST CYCLE
Effective Multiplication (Beginning of Life)
With Burnable Poison Rods in (First Cycle)
- 20. Cold, No Power, Clean 1.180
- 21. Hot, No Power, Clean 1.138
- 22. Hot, Full Power, Clean 1.119
- 23. Hot, Full Power, Xe and Sm Equilibrium 1.077 Rod Cluster Control Assemblies
- 24. Material 5% Cd; 15%; 80% Ag
Number of RCC Assemblies
- 25. Full Length 45
- 26. Partial Length*
8
- 27. Number of Absorber Rods per RCC Assembly 20
- 28. Total Rod Worth, BOL, % (See Table 3.2.1-3)
Boron Concentrations (ppm) for 1st Core Cycle
Loading with Burnable Poison Rods
- 29. Refueling Shutdown; Rods Inserted (k = .90) 1950
- 30. Shutdown (k = .99) with Rods Inserted, Clean, Cold 780
- 31. Shutdown (k = .99) with Rods Inserted, Clean, Hot 510
- 32. Shutdown (k = .99) with No Rods Inserted, Clean, Cold 1250
- 33. Shutdown (k = .99) with No Rods Inserted, Clean, Hot 1210
To Control at Hot Full Power, No Rods Inserted
k = 1.0
- 34. Poison Free 1000
- 35. Xenon 720
- 36. Xenon and Samarian 670
- 37. Shutdown, All but one Rod Inserted, Cold (k = .99) 850
- 38. Shutdown, All but one Rod Inserted, Hot (k = .99) 610
- Partial length control rods are not used in the present reloads.
Rev 16 10/99 TABLE 3.2.1-1 Sheet 3 of 3
BURNABLE POISON RODS
- 39. Number and Material*
816 Borated Pyrex Glass**
- 40. Worth Hot k/k,% 6.9
- 41. Worth Cold k/k,% 5.3
RANGE OF KINETIC CHARACTERISTICS FOR FIRST CYCLE*
- 42. Moderator Temperature Coefficient, (k/k)/oF +0.3 x 10
-4 to 3.5 x 10
-4
- 43. Moderator Pressure Coefficient (k/k)/psi -0.3 x 10
-6 to +3.5 x 10
-6
- 44. Moderator Density Coefficient (k/k)/gm/cm 3 -0.1 to 0.3
- 45. Doppler Coefficient, (k/k)/o F -1.0 x 10-5 to -1.6 x 10
-5 46. Delayed Neutron Fraction, % 0.52 to 0.72
- 47. Prompt Neutron Lifetime, sec. 1.4 x 10
-5 to 1.8 x 10
-5 48. Moderator Void Coefficient, (k/k)% void +0.5 x 10
-3 to -2.5 x 10
-3
- Corresponding values for current cycles are given in the Appendices 14A and 14B for Units 3 and 4, respectively.
- ZrB 2 IFBA is the burnable absorber used in present reloads.
Revised 02/18/2011 C25C25 TABLE 3.2.1-2
REACTIVITY REQUIREMENTS FOR CONTROL RODS*
Percent Beginning End Requirements of Life of Life Control
Power Defect (Doppler effect, Temperature change) 1.80 2.30
Operational Maneuvering Band and 0.70 0.30
Control Rod Bite Maximum Void and Redistribution 0.25 0.70 Total Control 2.75 3.30
- Corresponding values for current cycles are given in Appendices 14A and 14B
for Units 3 and 4 respectively.
Rev. 16 10/99 TABLE 3.2.1-3 CALCULATED ROD WORTHS, FOR FIRST CYCLE WITH BURNABLE POISON RODS*
Worth Design Core Rod Less Reactivity Shutdown Condition Configuration Worth 10%** Requirements Margin BOL, HFP 45 rod in 8.14%
44 rods in; 7.03% 6.33% 2.65% 3.68%
Highest Worth Rod Stuck Out
EOL, HFP 45 rods in 8.68%
(1st Cycle)
44 rods in; 7.42% 6.68% 3.58% 3.1%
Highest Worth Rod Stuck Out
BOL = Beginning of Life
EOL = End of Life
HFP = Hot Full Power
- Corresponding values for current cycles are given in Appendices 14A and 14B for Units 3 and 4 respectively.
- Calculated rod worth is reduced by 10% to allow for uncertainties.
Current Cycles (Appendices 14A and 14B) allow 7% uncertainty.
Rev. 16 10/99 TABLE 3.2.1-4 RESULTS OF CALCULATIONS AS A FUNCTION OF
LABORATORY PROVIDING EXPERIMENTAL DATA
Type of No. of Calculated Laboratory Experiment Experiments k +
Westinghouse Atomic Power Critical 16 0.9968 +/- 0.0036
Division (WAPD)
Bettis Atomic Power
Laboratory Critical 14 0.9940 +/- 0.0022
Brookhaven National
Laboratory Exponential 35 0.9964 +/- 0.0051
Hanford Atomic Products
Operation Exponential 20 0.9953 +/- 0.0105
Babcock and Wilcox Critical 26 0.9885 +/- 0.0094
111
TABLE 3.2.1-5 CALCULATED AND MEASURED REACTIVITY EFFECTS OF VOID TUBES
Reactivity Change %k/k
No. of Type of Core Tubes Measured Calculated
Unborated Core 0
4 -0.03 -0.034
16 -0.11 -0.125
196 -1.33 -1.416
Borated Core 0 4 -0.017 -0.020
16 -0.076 -0.085
196 -0.850 -0.942
TABLE 3.2.1-6 CORE STARTUP CRITICAL BORON CONCENTRATION
INDIAN POINT I SELNI SENA SCE (Core B)
Cold Critical Boron (ppm)
Predicted 1949 1910 2040 2380
Measured 1897 1800 1885 2250
Difference 52 110 155 130
Hot Zero Power
Critical Boron (ppm)
Predicted 1967 1910 2110 2570
Measured 1893 1840 1972 2524
Difference 74 70 138 46
Rev. 16 10/99 FLORIDA POWER & LIGHT COMPANY TURKEY POINT PLANT UNITS 3 & 4 PATTERN OF CONTROL ROD CLUSTER BANKS FIGURE 3.2.1-1
THE PART LENGTH RO DS HAVE BEEN REMOVED FROM THE REACTOR, HENCE THIS FIGURE HAS BEEN DELETED FROM THE UPDATED FSAR
FLORIDA POWER & LIGHT COMPANY TURKEY POINT PLANT UNITS 3 & 4 AVERAGE POWER DENSITY (BOL)
PART LENGTH RODS IN FIGURE 3.2.1-2
FLORIDA POWER & LIGHT COMPANY TURKEY POINT PLANT UNITS 3 & 4 AVERAGE POWER DENSITY (BOL),
CONTROL BANK D IN FIGURE 3.2.1-3
THE PART LENGTH RO DS HAVE BEEN REMOVED FROM THE REACTOR, HENCE THIS FIGURE HAS BEEN DELETED FROM THE UPDATED FSAR
FLORIDA POWER & LIGHT COMPANY TURKEY POINT PLANT UNITS 3 & 4 AVERAGE POWER DENSITY (BOL),
PART LENGTH RODS IN PLUS PART LENGTH RODS FIGURE 3.2.1-4
Revised 02/18/2011 FLORIDA POWER & LIGHT COMPANY TURKEY POINT PLANT UNITS 3 & 4 AVERAGE POWER DENSITY (BOL),
NO CONTROL RODS IN FIGURE 3.2.1-5 CYCLE 1 AVERAGE POWER DENSITY BOL, NO Control Rods F xy = 1.41
FLORIDA POWER & LIGHT COMPANY TURKEY POINT PLANT UNITS 3 & 4 Schematic Demonstration of Typical kW/ft Limits FIGURE 3.2.1-6
FLORIDA POWER & LIGHT COMPANY TURKEY POINT PLANT UNITS 3 & 4 BURNABLE POISON CLUSTER LOCATIONS FIGURE 3.2.1-7
FLORIDA POWER & LIGHT COMPANY TURKEY POINT PLANT UNITS 3 & 4 PATTERN OF POISON ROD LOCATIONS FIGURE 3.2.1-8
FLORIDA POWER & LIGHT COMPANY TURKEY POINT PLANT UNITS 3 & 4 MODERATOR TEMPERATURE COEFFICIENT Vs. MODERATOR TEMPERATURE FIGURE 3.2.1-9
FLORIDA POWER & LIGHT COMPANY TURKEY POINT PLANT UNITS 3 & 4 DOPPLER COEFFICIENT Vs. EFFECTIVE FUEL TEMPERATURE (BOL)
FIGURE 3.2.1-10
FLORIDA POWER & LIGHT COMPANY TURKEY POINT PLANT UNITS 3 & 4 POWER COEFFICIENT FIGURE 3.2.1-11
FLORIDA POWER & LIGHT COMPANY TURKEY POINT PLANT UNITS 3 & 4 POWER COEFFICIENT (CLOSED GAP MODEL) FIGURE 3.2.1-12
FLORIDA POWER & LIGHT COMPANY TURKEY POINT PLANT UNITS 3 & 4 PLUTONIUM/URANIUM MASS RATIO AS A FUNCTION of URANIUM-235 DEPLATION FIGURE 3.2.1-13
FLORIDA POWER & LIGHT COMPANY TURKEY POINT PLANT UNITS 3 & 4 FRACTION OF PLUTONIUM-239 IN PLUTONIUM AS A FUNCTION OF URANIUM-235 DEPLATION FIGURE 3.2.1-14
FLORIDA POWER & LIGHT COMPANY TURKEY POINT PLANT UNITS 3 & 4 COMPOSITION OF PLUTONIUM AS A FUNCTION OF URANIUM-235 DEPLATION FIGURE 3.2.1-15
FLORIDA POWER & LIGHT COMPANY TURKEY POINT PLANT UNITS 3 & 4 PLAN OF CRITICAL EXPERIMENT (UNBORATED CASE)
FIGURE 3.2.1-16
FLORIDA POWER & LIGHT COMPANY TURKEY POINT PLANT UNITS 3 & 4 PLAN OF CRITICAL EXPERIMENT (BORATED CASE)
FIGURE 3.2.1-17
FLORIDA POWER & LIGHT COMPANY TURKEY POINT PLANT UNITS 3 & 4 BORATED POWER DISTRIBUTION COMPARISON FIGURE 3.2.1-18
FLORIDA POWER & LIGHT COMPANY TURKEY POINT PLANT UNITS 3 & 4 UNBORATED POWER DISTRIBUTION COMPARISON FIGURE 3.2.1-19
FLORIDA POWER & LIGHT COMPANY TURKEY POINT PLANT UNITS 3 & 4 COMPARISON OF EXPERIMENTAL AND CALCULATED POWER DISTRIBUTION USIN G ONE MESH SPACING PER FUEL ROD FIGURE 3.2.1-20
FLORIDA POWER & LIGHT COMPANY TURKEY POINT PLANT UNITS 3 & 4 COMPARISON OF CALCULATED POWER DISTRIBUTION WITH EXPERIMENTAL POWER SCANS - UNBORATED CORE FIGURE 3.2.1-21
FLORIDA POWER & LIGHT COMPANY TURKEY POINT PLANT UNITS 3 & 4 COMPARISON OF CALCULATED POWER DISTRIBUTION WITH EXPERIMENTAL POWER SCANS - BORATED CORE FIGURE 3.2.1-22
FLORIDA POWER & LIGHT COMPANY TURKEY POINT PLANT UNITS 3 & 4 RADIAL FUEL ROD SCAN FIGURE 3.2.1-23
FLORIDA POWER & LIGHT COMPANY TURKEY POINT PLANT UNITS 3 & 4 YANKEE CORE 1 POWER DISTRIBUTION COMPARISON FIGURE 3.2.1-24
FLORIDA POWER & LIGHT COMPANY TURKEY POINT PLANT UNITS 3 & 4 YANKEE CORE 1 BURNUP DISTRIBUTION COMPARISON FIGURE 3.2.1-25
FLORIDA POWER & LIGHT COMPANY TURKEY POINT PLANT UNITS 3 & 4 FAST ABSORPTION OF CLUSTERED ABSORBERS FIGURE 3.2.1-26
FLORIDA POWER & LIGHT COMPANY TURKEY POINT PLANT UNITS 3 & 4 FAST ABSORPTION OF UNIFORMLY DISTRIBUTED ABSORBERS FIGURE 3.2.1-27
FLORIDA POWER & LIGHT COMPANY TURKEY POINT PLANT UNITS 3 & 4 SELNI TEMPERATURE COEFFICIENT Vs.
MODERATOR TEMPERATURE (1600 PPM BORON)
FIGURE 3.2.1-28
FLORIDA POWER & LIGHT COMPANY TURKEY POINT PLANT UNITS 3 & 4 MODERATOR TEMPERATURE COEFFIENT Vs. BORON CONCENTRATION FIGURE 3.2.1-29
FLORIDA POWER & LIGHT COMPANY TURKEY POINT PLANT UNITS 3 & 4 COMPARSION OF CALCULATED AND MEASURED MODERATOR TEMPERATURE COEFFICENT Vs. BURNUP FIGURE 3.2.1-30
FLORIDA POWER & LIGHT COMPANY TURKEY POINT PLANT UNITS 3 & 4 COMPARISON OF RESONANCE INTEGRAL CORRELATIONS FIGURE 3.2.1-31
FLORIDA POWER & LIGHT COMPANY TURKEY POINT PLANT UNITS 3 & 4 FUEL TEMPERATURE CHANGES Vs.
POWER DENSITY FIGURE 3.2.1-32
FLORIDA POWER & LIGHT COMPANY TURKEY POINT PLANT UNITS 3 & 4 ALPHA Vs. HEAT FLUX FIGURE 3.2.1-33
FLORIDA POWER & LIGHT COMPANY TURKEY POINT PLANT UNITS 3 & 4 COMPARISON OF EFFECTIVE FUEL TEMPERATURE WITH CHANGING HEAT FLUX FIGURE 3.2.1-34
3.2.2 THERMAL AND HYDRAULIC DESIGN AND EVALUATION
Thermal and Hydraulic Characteristics of the Design
Thermal Data Fuel and Cladding Temperature The temperature distribution in the pellet is mainly a function of the uranium dioxide thermal conductivity and the local power density. The absolute value
of the temperature distribution is affected by the cladding temperature and
the thermal conductance of the gap between the pellet and the cladding.
The occurrence of nucleate boiling maintains maximum cladding surface temperature below about 657 o F at nominal system pressure. The contact conductance between the fuel pellet and cladding is a function of the contact
pressure and the composition of the gas in the gap and may be calculated by
the following equation:
where h is conductance in Btu/hr ft 2 F P is contact pressure in psi k is the thermal conductivity of the gas mixture in the rod in Btu/hr ft 2 F f is the correction factor for the accommodation coefficient The thermal-hydraulic design assures that the maximum fuel temperature is below the melting point of UO 2 (melting point of 5080 0 F (Reference 3) unirradiated and decreasing by 58 0 F per 10,000 MWD/MTU). To preclude fuel melting, the peak local power experienced during Condition I and 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.
Evaluations for Condition I and II events are used to show that fuel melting
will not occur. The temperature distribution within the fuel pellet is
predominantly a function of the local power density and the UO 2 thermal conductivity. However, the computation of radial fuel temperature
distributions combines crud, oxide, cladding gap and pellet conductances.
3.2.2-1 Revised 04/17/2013
)6-10 x 14.4 ( f k P 0.6 hC26C26C26 The factors which influence these conductances, such as gap size (or contact pressure), internal gas pressure, gas composition, pellet density, burnup and radial power distribution within the pellet, etc., have been combined into a semi-empirical thermal model which includes a model for time dependent fuel densification as given in Reference 4 and used in Reference 67. Reference 76 uses the same thermal model as Reference 67 and incorporates the effects of thermal conductivity degradation (TCD). This thermal model enables the determination of these factors and their net effects on temperature profiles.
The temperature predictions have been compared to inpile fuel temperature measurements (References 5 through 12, 67 and 68), and melt radius data (References 13, 14) with good results.
As described in Reference 15, fuel rod thermal evaluations (fuel centerline, average and surface temperatures) are determined throughout the fuel rod
lifetime with consideration of time dependent densification.
The principal factors which are employed in the determination of the fuel
temperature are discussed below.
UO 2 Thermal Conductivity
The thermal conductivity of uranium dioxide was evaluated from data reported
by Howard et al. (Reference 16); Lucks et al. (Reference 17); Daniel et al.
(Reference 18); Feith (Reference 19); Vogt et al. (Reference 20 ); Nishijima
et al. (Reference 21); Wheeler et al. (Reference 22); Godfrey et al.
(Reference 23); Stora et al. (Reference 24); Bush (Reference 25); Asamoto et
al. (Reference 26); Kruger (Reference 27); and Gyllander (Reference 28).
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 2800ºC Kdt is 93 watts/cm. This conclusion is based on the integral values reported by Gyllander (Reference 28), Lyons et al. (Reference
29), Coplin et al. (Reference 30), Duncan (Reference 13), Bain (Reference 31),
Stora (Reference 32).
The design curve for the thermal conductivity is shown in Figure 3.2.2-1. The
section of the curve at temperatures between 0ºC and 1300ºC is in excellent
agreement with the recommendation of the IAEA panel (Reference 33). The
section of the curve above 1300ºC, is derived for an integral value of 93
watts/cm (References 13,28,32).
3.2.2-2 Revised 04/17/2013 C26 Thermal conductivity for unirradidated UO 2 at 95% theoretical density can be represented best by the following equation:
K=
1 + 8.775 x 10
-13 T 3 11.8 + 0.0238T
Where:
K= watts/cm-
°C T= °C Operational Experience With Westinghouse Cores Fuel operational experience has verified the adequacy of the fuel performance and design bases, as discussed in Reference 34. Fuel experience and testing results, as they become available, are used to improve fuel rod design and manufacturing processes and to assure that design bases and safety criteria are met.
3.2.2-3 Revised 04/17/2013 C26C26 Heat Flux Ratio and Data Correlation Departure from Nucleate Boiling, (DNB), is predicated upon a combination of hydrodynamic and heat transfer phenomena and is affected by the local and upstream conditions including the flux distribution. In reactor design, the heat flux associated with DNB and the location of DNB are both important.
Both the W-3 DNB correlation (37), previously used for LOPAR fuel analysis, and the WRB-1 DNB correlation (36), used for OFA and Upgrade fuel analysis, incorporate local and system parameters in predicting the local DNB heat flux.
DNB correlations, such as WRB-1, account for the non-uniform flux effect, and
the upstream effect which includes a length upstream at which DNB occurs. The
local DNB heat flux ratio (defined as the ratio of the DNB heat flux to the
local heat flux) is indicative of the contingency available in the local heat
flux without reaching DNB.
Definition of Departure from Nucleate Boiling Ratio
The DNB heat flux ratio (DNBR) as applied to typical cells (flow cells with
all walls heated) and thimble cells (flow cells with heated and unheated
walls) is defined as:
DNBR=
q DNB,N q loc Where qDNB,N = qDNB, prediction F And q DNB,prediction is the uniform critical heat flux as predicted by a DNB correlation such as WRB-1.
F is the flux shape factor to account for nonuniform axial heat flux
distributions (Reference 35) with the "C" term modified as in Reference 37.
3.2.2-4 Revised 04/17/2013 C26C26C26C26C26 F = C 0 L DNB q (Z) e -C(L DNB-Z) dz qlocal, (at L DNB) x (1-e-C(L DNB)) L DNB = distance from the inception of local boiling to the point of DNB
.
Z = Distance form the inception of local boiling, measured in the direction of flow.
The empirical constant, C, as presented in (Reference 35) has been revised through the use of non-uniform DNB data However, the revised expression does not significantly influence (<1% deviation from that of Reference 35) the value of the F-factor and the DNB-ratio. It does provide a better prediction
of the location of DNB. The new expression is
C = 0.15 (1-X DNB)4.31 1/inch (G/10 6)0.478 G = Mass Velocity (lb/hr ft
The DNBR as applied to the W-3 DNB correlation when a cold wall is present is:
DNBR = qDNB,N,CW q loc Where: q DNB,N,CW = q DNB,prediction,Dh x CWF F And q DNB,prediction,Dh is the uniform critical heat flux as predicted by the W-3 cold wall DNB Correlation (Reference 37) when not all flow cell walls are heated (thimble cold wall cell). CWF is the cold wall factor.
W-3 Equivalent Uniform Flux DNB Correlation
In determining the F-factor, the value of qlocal at L DNB was measured as z = L DNB, the
3.2.2-5 Revised 04/17/2013 For a uniform flux, F becomes unity so that qDNB,N reduces to q DNB,prediction as expected. The comparison of predictions by using W-3 correlations and the non-uniform DNB data obtained by B&W (38), Winfrith (39) and Fiat are given in Figure 3.2.2-7 and Figure 3.2.2-8. The criterion for determining the
predicted location of DNB is to evaluate the ratio of the predicted DNB flux
to the local heat flux along the length of the channel. The location of the
minimum DNB ratio is considered to be the location of DNB.
W-3 DNB Correlation
The W-3 correlation, and several modifications of it, have been previously
used in Westinghouse critical heat flux calculations. The W-3 correlation was
originally developed from single tube data, (Reference 37) but was
subsequently modified to apply to the 0.422 inch O.D. rod "L" grid (Reference
- 40) rod bundle data. These modifications to the W-3 correlations have been
demonstrated to be adequate for reactor rod bundle design.
3.2.2-6 Revised 04/17/2013 C26 The comparison of the measured to predicted DNB heat flux of this correlation is given in Figure 3.2.2-2. The local flux DNB ratio versus the probability
of not reaching DNB is plotted in Figure 3.2.2-3. This plot indicates that
with a DNBR of 1.3 the probability of not reaching DNB is 95% at a 95%
confidence level.
Rod bundle data without mixing vanes agree very well with the predicted DNB
flux as shown in Figure 3.2.2-4, and rod bundle data with mixing vanes (Figure 3.2.2-5) show on the average an 8% higher value of DNB heat flux than
predicted by the W-3 DNB correlation.
WRB-1 DNB Correlation
The WRB-1 (Reference 36) correlation was developed based exclusively on the large bank of mixing vane grid rod bundle critical heat flux data (in excess of 1100 points) that Westinghouse has collected. The WRB-1 correlation, based on local fluid conditions, represents the rod bundle data with better accuracy over a wide range of variables than the previous correlation used in design (namely the W-3 correlation). This correlation accounts directly for both typical and thimble cold wall cell effects, uniform and nonuniform heat flux
profiles, and variations in rod heated length and in grid spacing.
The applicable range of variables is:
Pressure : 1440 < P < 2490 psia Local Mass Velocity
- 0.9x10 6 < G loc < 3.7 x10 6 lb/ft 2 hr Local Quality
- -0.2 < x loc < 0.3 Heated Length, Inlet to CHF Location : L h < 14 feet Grid Spacing
- 13 < g sp < 32 inch Equivalent Hydraulic Diameter : 0.37 <
d e < 0.60 inch Equivalent Heated Hydraulic Diameter : 0.46 <
d h < 0.59 inch
Figure 3.2.2-6 shows measured critical heat flux plotted against predicted
critical heat flux using the WRB-1 correlation.
The WRB-1 DNB Correlation 95/95 DNBR limit is 1.17 (Reference 36 and 72).
Validation of the WRB-1 correlation applicability to the 15 x 15 upgrade fuel
design is provided in Reference 73.
3.2.2-7 Revised 04/17/2013 C26C26C26 W-3 Alternative Correlation The W-3 Alternative correlations, consisting of ABB-NV and WLOP, are based exclusively on DNB data from rod bundle tests, has a wider applicable range, and is more accurate than the W-3 correlation for prediction of margin to DNB.
It is used for DNBR calculations as an alternative to the W-3 correlation, in supplement to the primary DNB correlation WRB-1.
The ABB-NV correlation was originally developed for fuel designs of Combustion Engineering designed Pressurized Water Reactors (PWR) based on a linear relationship between CHF and local quality. The correlation includes the following parameters: pressure, local mass velocity, local equilibrium quality, distance from grid to CHF location, heated length from inlet to CHF location, and heated hydraulic diameter of the subchannel. Supplemental rod bundle data evaluation confirms that ABB-NV with the 95/95 correlation limit of 1.13 is applicable to the fuel region below the first mixing vane grid of the fuel designs for Westinghouse designed PWR (Reference 74). Figure 3.2.2-9 shows measured critical heat flux plotted against predicted heat flux using the ABB-NV correlation.
The applicable range of the ABB-NV correlation is:
Pressure (psia) 1750 to 2415 Local Mass Velocity (Mlbmihr-ft
- 2) 0.8 to 3.16 Local Quality (fraction) <0.22 Heated Length, inlet to CHF location (in.) 48 (minimum) to 150 Heated Hydraulic Diameter Ratio 0.679 to 1.08 Grid Distance (in.) 7.3 to 24 The WLOP correlation is a modified ABB-NV correlation specifically developed for low pressure conditions and extended flow range to cover low pressure/low flow conditions. Modifications to ABB-NV were made based on test data from rod bundles containing non-mixing vane grids. The WLOP correlation with a 95/95 DNBR limit of 1.18 has also been validated with test data from rod bundles containing mixing vane grids (Reference 75). Figure 3.2.2-10 shows measured critical heat flux plotted against predicted heat flux using the WLOP correlation.
3.2.2-8 Revised 03/11/2016 C28 The applicable range of the WLOP correlation is:
Pressure (psia) 185 to 1800 Local Mass Velocity (Mlbm/hr-ft
- 2) 0.23 to 3.07 Local Quality (fraction) <0.75 Heated Length, inlet to CHF location (in.) 48 (minimum) to 168 Heated Hydraulic Diameter Ratio 0.679 to 1.00 Grid Spacing Term (Reference 72) 27 to 115
To calculate the DNBR of a reactor channel, the values of q DNB,N along q" loc the channel are evaluated and the minimum value is selected as the minimum DNBR incurred in that channel.
Surface Heat Transfer Coefficients
Forced convection heat transfer coefficients are obtained from the
Dittus-Boelter correlation (Reference 41), with the properties evaluated at
bulk fluid conditions:
K C G D0.023 = K hD p 0.4 e 0.8 e where: h = heat transfer coefficient (Btu/hr.-ft 2-°F) D e = equivalent diameter (ft) K = thermal conductivity (Btu/hr.-ft.-
°F) G = mass velocity (lb m/hr.-ft.2) U = dynamic viscosity, (lb m/ft.-hr) C p = heat capacity, (Btu/lb m-°F) This correlation has been shown to be conservative (Reference 42) for rod
bundle geometries with pitch to diameter ratios in the range used by
pressurized water reactors.
3.2.2-9 Revised 04/17/2013 C26C26 The onset of nucleate boiling occurs when the clad wall temperature reaches
the amount of superheat predicted by Thom's (Reference 43) correlation. After
this occurrence, the outer clad wall temperature is determined by:
T sat = [0.072 exp (-P/1260)] (q")
0.5 Where:
T sat = wall superheat, T w - T sat °F q = wall heat flux (Btu/hr.-ft.
- 2) P = Pressure (psia)
T w = outer clad wall temperature, °F T sat = saturation temperature of coolant at P
°F 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 factors consider the local maximum linear heat generation rate at a point, and the enthalpy rise factors involve the maximum integrated value along a channel (the "hot channel").
Definition of Engineering Hot Channel Factor
Each of the total hot channel factors considers a nuclear hot channel factor describing the neutron flux 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 accounting for the influence of the variations of fuel pellet diameter, density, enrichment
and eccentricity; inlet flow distribution; flow redistribution; and flow
mixing.
Heat Flux Engineering Subfactor, E q F The heat flux engineering hot channel factor is used to evaluate the maximum linear heat generation rate in the core. This subfactor is determined by statistically combining the fabrication variations for the fuel pellet diameter, density, enrichment, and has a value of 1.03 at the 95% probability level with 95% confidence. As shown in Reference 44, no DNB penalty need be taken for the short, relatively low intensity heat flux spikes caused by variations in the above parameters, as well as fuel pellet eccentricity and
fuel rod diameter variation.
3.2.2-10 Revised 04/17/2013 Enthalpy Rise Engineering Subfactor, F EH Design values employed in the subchannel code analysis related to the above fabrication variations are based on applicable limiting tolerances such that these design values are met for 95% of the limiting channels at a 95%
confidence level. Measured manufacturing data on Westinghouse fuel show the tolerances used in this evaluation are conservative.
The items considered contributing to the enthalpy rise engineering hot channel factor are discussed below.
a) Pellet diameter, density and enrichment Variations in pellet diameter, density, and enrichment are considered
statistically in establishing the limit DNBRs for the Revised Thermal
Design Procedure (Reference 45) employed in this application.
Uncertainties in these variables are determined from sampling of
manufacturing data.
b) Inlet Flow Maldistribution
Studies performed on 1/7 scale hydraulic reactor models indicate that a conservative design basis is to consider a 5% reduction in the flow to
the hot fuel assembly under isothermal conditions. This inlet flow
reduction in the subchannel code analysis results in an increase of 1%
in the hot channel enthalpy rise.
c) Flow Redistribution
The flow redistribution accounts for the reduction in flow in the hot channel resulting from the high flow resistance in the channel due to the local or bulk boiling. The effect of the nonuniform power distribution is inherently considered in the subchannel code analysis
for every operating condition which is evaluated.
3.2.2-11 Revised 04/17/2013 C26C26C26 d) Flow Mixing The subchannel mixing model incorporated in the subchannel code and used in reactor design is based on experimental data (Reference 46). The
mixing vane incorporated into the spacer grid design induce additional flow mixing between the various flow channels in a fuel assembly and also between adjacent assemblies. This mixing reduces the enthalpy rise in the hot channel resulting from local power peaking or unfavorable
mechanical tolerances.
Table 3.2.2-2 is a tabulation of the design engineering hot channel factors for the first fuel cycle. For current cycles, the effect of variations in flow conditions and fabrication tolerances on the hot channel enthalpy rise is directly considered in the core thermal subchannel analysis under any reactor operating condition. Therefore, this engineering hot channel factor is no longer applied separately.
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, z)dzy,(x,q N 1 z)dz y x(q MAX =power rod averagepower rod hot
=F H o rods all o o H o NH , , where x , y , are the position coordinates of the hot rod.
The way in which F NH is used in the DNB calculation is important. The location of minimum DNBR depends on the axial profile and the value of DNBR depends on the enthalpy rise to that point. Basically, 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 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 which are typical distributions found in hot assemblies. In this manner, worst case axial profiles can be combined with worst case radial distributions for
reference DNB calculations.
3.2.2-12 Revised 04/17/2013 C26C26 It should be noted again that F NH is an integral and is used as such in 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
subchannel code analysis to radial power shapes is discussed in Reference 47.
For operation at a fraction P of full power, the design F NH used is given by:
F NH = 1.248 (1 + 0.3 (1-P)) (optimized fuel)
(1)
F NH = 1.60 (1 + 0.3 (1-P)) (upgrade fuel)
(1) where P is fraction of full power.
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 44), thus allowing greater flexibility in the nuclear
design.
Pressure Drop and Hydraulic Forces
For historical purposes the total pressure loss across the reactor vessel, including the inlet and outlet nozzles, and the pressure drop across the core
for the first cycle are listed in Table 3.2.2-1. This table also includes the
total pressure drop across the core. These pressure drop values include a 10%
uncertainty factor.
(1) Values for current cycles are given in Appendices 14A and 14B.
3.2.2-13 Revised 04/17/2013 C26C26C26C26C26 DNBR Design Methodology The design method employed to meet the DNB design basis for the optimized and
Upgrade fuel assemblies is the Revised Thermal Design Procedure (RTDP),
Reference 45. With the RTDP methodology, uncertainties in plant operating
parameters, nuclear and thermal parameters, fuel fabrication parameters, computer codes and DNB correlation predictions are considered statistically
to obtain DNB uncertainty factors. Based on the DNB uncertainty factors, RTDP design limit DNBR values are determined such that there is at least a 95
percent probability at a 95 percent confidence level that DNB will not occur
on the most limiting fuel rod during normal operation and operational
transients and during transient conditions arising from faults of moderate
frequency (Condition I and II events). Since the parameter uncertainties are
considered in determining the RTDP design limit DNBR values, the plant safety
analyses are performed using input parameters at their nominal values.
The RTDP design limit DNBR value is 1.22 for both typical and thimble cells, for optimized and Upgrade fuel.
To maintain DNBR margin to offset DNB penalties such as those due to fuel rod
bow and transition core, the safety analyses were performed to DNBR limits
higher than the design limit DNBR values. These limits are called the DNBR
safety analysis limits. The difference between the design and safety
analysis limits results in DNBR margin, M:
M-1DNBR Limit Design = DNBR Limit Analysis Safety The net DNBR margin, after consideration of all penalties, is available for operating and design flexibility.
The Standard Thermal Design Procedure (STDP) is used for those analyses where
RTDP is not applicable. In the STDP method, the parameters used in the
analysis are treated in a conservative way from a DNBR standpoint. The
parameter uncertainties are applied directly to the plant safety analyses
input values to give the lowest minimum DNBR. The DNBR limit for STDP is the
appropriate DNB correlation limit increased by sufficient margin to offset
the applicable DNBR penalties.
3.2.2-14 Revised 04/17/2013 C26C26C26C26 The objective of reactor core thermal design is to determine the maximum heat removal capability in all flow subchannels and show that the core safety limits, are not exceeded, using the most conservative power distribution.
The thermal design takes into account local variations in dimensions, power generation, flow redistribution, and mixing. Prior to the uprate to 2644 MWt, the THINC-IV subchannel code was used in thermal/hydraulic analysis (References 47, 50, 51). Commencing with the uprate, the VIPRE-01 code is used the thermal-hydraulic analysis of the core. VIPRE-01 (VIPRE) is a three-dimensional sub-channel code that has been developed to account for hydraulic and nuclear effects on the enthalpy rise in the core and hot channels (Reference 76). VIPRE modeling of a PWR core is based on one-pass modeling approach (Reference 74). In the one-pass modeling, hot channels and their adjacent channels are modeled in detail, while the rest of the core is modeled simultaneously on a relatively coarse mesh. The behavior of the hot assembly is determined by superimposing the power distribution upon inlet flow distribution while allowing for flow mixing and flow distribution between flow channels. Local variations in fuel rod power, fuel rod and pellet fabrication, and turbulent mixing are also considered in determining conditions in the hot channels. Conservation equations of mass, axial and lateral momentum, and energy are solved for the fluid enthalpy, axial flow rate, lateral flow and pressure drop.
Steady State Analysis The VIPRE-01 computer program and subchannel analysis methodology, as approved by the NRC (Reference 74) is used to determine coolant density, mass velocity, enthalpy, vapor void, static pressure, and DNBR distributions within the reactor core hot subchannel under all expected operating conditions. The VIPRE-01 code is described in detail in Reference 74, including models and correlations used.
Experimental Verification Experimental verification of VIPRE-01 is presented in References 76 and 74 The VIPRE-01 analysis methodology is based on a knowledge and understanding of the heat transfer and hydrodynamic behavior of the coolant flow and the mechanical characteristics of the fuel elements. VIPRE-01 analysis provides a realistic evaluation of the core performance and is used in the thermal analyses as described above.
3.2.2-15 Revised 04/17/2013 C26 Transient Analysis The approved VIPRE-01 methodology (Reference 74) was shown to be conservative for transient thermal-hydraulic analysis.
DNB core safety limits are generated as a function of coolant temperature, pressure, core power, and the axial and radial power distributions.
Operation within these DNB safety limits insures that the DNB design basis is
met for both steady-state operation and for 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, specific protection functions are
provided.
Effects of Rod Bow on DNBR
The phenomenon of fuel rod bowing, as described in Reference 52, is accounted
for in the DNBR safety analysis of Condition I and Condition II events for
each plant application. Applicable credits for margin retained in the
evaluation of DNBR are used to offset the effect of rod bow. For the safety
analysis of the Turkey Point units, sufficient DNBR margin was maintained to
accommodate the maximum full flow and low flow DNBR penalties based on
methodology in Reference 53. The maximum rod bow DNBR penalty accounted for
in the design safety analysis is based on an assembly average burnup of
24,000 MWD/MTU. At burnups greater than 24,000 MWD/MTU, credit is taken for the effect of FH burndown, due to the decrease in fissionable isotopes and the buildup of fission product inventory, and no additional rod bow DNBR penalty is required. Reference 69 addresses the burndown credit at burnups
greater than 24,000 MWD/MTU.
3.2.2-16 Revised 04/17/2013 C26C26 Transition Core DNB Methodology The Westinghouse transition core DNB methodology is given in Reference 71.
Using this methodology, transition cores are analyzed as if the entire core
consisted of one assembly type. The resultant DNBRs are then reduced by the
appropriate transition core penalty.
The OFA fuel assembly has a higher mixing vane grid loss coefficient relative
to the Upgrade mixing vane grid loss coefficient. The Upgrade fuel assembly
has Integral Flow Mixer (IFM) grids located in spans between mixing vane
grids, where no grid exists in the OFA assembly. The higher loss
coefficients and the additional grids introduce localized flow redistribution
between the fuel assemblies at various axial zones in a transition core.
Because the localized flow redistribution results in reduced flows to both
fuel types at various axial locations, transition core penalties are applied
to both fuel types. The transition core DNBR penalties are functions of the
number of each fuel assembly type in the core, Reference 71. Sufficient DNBR
margin is maintained in the safety analysis to offset the transition core
penalties.
Effects of DNB on Neighboring Rods
Westinghouse has never observed DNB to occur in a group of neighboring rods
in a rod bundle as a result of DNB in one rod in the bundle. Westinghouse (54) has conducted DNB tests in a 25-rod bundle where physical burnout occurred
with one rod. After this occurrence, the 25 rod test section was used for
several days to obtain more DNB data from the other rods in the bundle. The
burnout and deformation of the rod did not affect the performance of
neighboring rods in the test section during the burnout or the validity of
the subsequent DNB data points as predicted by the W-3 correlation. No
occurrences of flow instability or other abnormal operation were observed.
DNB With Return to Nucleate Boiling
Additional DNB tests have been conducted by Westinghouse (55) in 19 and 21 rod bundles. In these tests, DNB without physical burnout was experienced more
than once on single rods in the bundles for short periods of time. Each
time, a reduction in power of approximately 10% was sufficient to
re-establish nucleate boiling on the surface of the rod. During these and
subsequent tests, no adverse effects were observed on this rod or any other
rod in the bundle as a consequence of operating in DNB.
3.2.2-17 Revised 04/17/2013 Hydrodynamic and Flow Power Coupled Instability Boiling flows may be susceptible to thermohydrodynamic instabilities (Reference 56). These instabilities are undesirable in reactors since they
may cause a change in thermohydraulic 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. Therefore a
thermohydraulic design criterion was developed with states that modes of
operation under Condition I and II events shall not lead to
thermohydrodynamic instabilities.
Two specific types of flow instabilities are considered for Westinghouse PWR
operation. These are the Ledinegg or flow excursion type of static
instability and the density wave type of dynamic instability.
3.2.2-18 Revised 04/17/2013 A Ledineg instability involves a sudden change in flow rate from one steady
state to another. This instability occurs (Reference 56) when the slope of the reactor coolant system pressure drop-flow rate curve (P/ GINTERNAL) becomes algebraically smaller than the loop supply (pump head) pressure drop-
flow rate curve (P/ GEXTERNAL). The criterion for stability is, thus, (P/ GINTERNAL) > (P/ GEXTERNAL). The Westinghouse pump head curve has a negative slope (P/ GEXTERNAL < 0) whereas the reactor coolant system pressure drop-flow curve has a positive slope (P/ GINTERNAL > 0) over the Condition I and Condition II operational ranges. Thus, the Ledinegg instability will not occur. The mechanism of density wave oscillations in a heated channel has been
described by Lahey and Moody (Reference 57). Briefly, an inlet flow
fluctuation produces an enthalpy perturbation. This perturbs the length and
the pressure drop of the single phase region and causes quality or void
perturbations in the two-phase regions which travel up the channel with the
flow. The quality and length perturbations in the two-phase region create
two-phase pressure drop perturbations. However, since the total pressure
drop across the core is maintained by the characteristics of the fluid system
external to the core, then the two-phase pressure drop perturbation feeds
back to the single phase region. These resulting perturbations can be either
attenuated or self-sustained.
A simple method has been developed by Ishii (Reference 58) for parallel
closed channel systems to evaluate whether 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 59, 52, 53), under Condition I and II operation. The results
indicate that a large margin to density wave instability exists, e.g.,
increases on the order of 150 to 200% of rated reactor power would be
required for the predicted inception of this type of instability.
The application of the method of Ishii (Reference 58) to Westinghouse reactor
designs is conservative due to the parallel open channel feature of
Westinghouse PWR cores. For such cores, there is little resistance to
lateral flow leaving the flow channels of high power density channels. This
coupling with cooler channels has led to the opinion that an open channel
configuration is more stable than the above closed channel analysis under the
same boundary conditions.
3.2.2-19 Revised 04/17/2013 Flow stability tests (Reference 62) have been conducted where the closed
channel systems were shown to be less stable than when the same channels were
cross connected several locations. The cross-connections were such that the
resistance to channel to channel cross flow and enthalpy perturbations would
be greater than that which would exist in a PWR core which has a relatively
low resistance to cross flow.
Flow instabilities which have been observed have occurred almost exclusively
in closed channel systems operating at low pressure relative to the
Westinghouse PWR operating pressures. Kao, Morgan and Parker (Reference 63)
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 the data from the rod bundle DNB tests. Many
Westinghouse rod bundles have been tested over wide ranges of operating
conditions with no evidence of premature DNB or of inconsistent data which
might be indicative of flow instabilities in the rod bundle.
In summary, it is concluded that thermohydrodynamic instabilities will not
occur under Condition I and II modes of operation for Westinghouse PWR
reactor designs. A large power margin exists to predicted inception of such
instabilities. Analysis has been performed which shows that minor plant to
plant differences in Westinghouse reactor designs such as fuel assembly
arrays, core power to flow ratios, fuel assembly length, etc. will not result
in gross deterioration of the above power margins.
Effect of Fuel Densification
Fuel densification is the process where geometric dimensions of fuel pellets
under irradiation shrink from their as-built values. The cause of this
shrinkage is the elimination of fine porosity from within the pellet grain
structure. For modern design fuel, the densification process primarily
occurs at beginning of life, but is continuous throughout life. A competing
process, fuel swelling, is dominant after beginning of life and leads to
increased fuel dimensions with burnup. Allowance for the effect of fuel
densification has been made in the design (References 15, 64 and 73).
3.2.2-20 Revised 04/17/2013 Radial shrinkage of the pellet causes an increase in the pellet-to-cladding gap and results in a decrease in gap conductance and a corresponding increase
in fuel temperature and stored energy for a given heat generation rate. This
effect is accounted for in the fuel temperature calculations.
Axial shrinkage of fuel pellets causes an increase in the fuel pellet linear
heat generation rate. This increase is quantified by the stack height
factor, a multiplier on the undensified linear heat generation rate.
An additional potential consequence of axial shrinkage will occur if the fuel
pellets can become hung up in the cladding, which causes a formation of fuel-
free interpellet gap. Removing neutron absorber material from a portion of a
rod, as has occurred in the gap, causes an increase in the neutron density in
the vicinity of the pellets near the gap, both on the rod with the gap and
the adjacent rods. Interaction of these extra neutrons with the fuel pellets
causes power generation spikes on the adjacent rods and the rod with the gap.
The effect of such power spikes is accounted for with the use of a power
spike factor, S(Z), where Z is the axial position in the core.
DNB power tests have shown that local power spikes have no effect on DNB (40). Radial shrinkage of the pellet does not significantly affect the rod average
heat flux at the cladding outer surface. Reference 64 indicates that
process improvements in fuel pellet manufacturing have lead to the production
of fuel pellets which exhibit controlled microstructure with respect to both
grain size and pore size distribution. As a result, the current Westinghouse
fuel used at Turkey Point is stable with respect to densification, and
significant axial pellet column gaps do not occur. Thus, starting with Unit
4 Cycle 16, the elimination of the densification spike factor or the use of a
factor of 1.0 in fuel design evaluation is appropriate.
Thermal-Hydraulic Effect of Reactor Vessel Upper Internals During Refueling Operation of a single residual heat removal loop is permitted for decay heat
removal when fuel is in the reactor vessel and the refueling cavity is
flooded to greater than or equal to 23 feet above the reactor vessel flange.
A loss of this single residual heat removal loop has been evaluated to ensure
that adequate natural circulation cooling can be maintained for decay heat
removal. The analysis utilized GOTHIC thermal-hydraulic analysis software to
evaluate natural circulation cooling conditions with both the reactor vessel
upper internals assembly installed and with the upper internals assembly
removed.
3.2.2-21 Revised 04/17/2013 In each case, stable natural circulation patterns occur such that adequate
heat transfer capability is maintained to prevent fuel damage. The calculated
increase in heat flux in the core volume is only a small fraction of the CHF
for the fuel, and is considered a minimal (essentially negligible) increase.
For more detailed information regarding the analysis, see Reference 68.
For the upper internals assembly installed case, the natural circulation flow path modeled is up from the core to the vessel upper plenum to the refueling cavity via the holes in the upper support plate, the CRDM guide tubes, the head spray flow nozzles, the upper internals hold-down spring gap, and the hot leg gap, with return flow to the core via the downcomer and barrel/baffle bypass. A direct flow path for natural circulation from the core to the vessel upper plenum to the refueling cavity and back to the core via the downcomer and barrel/baffle bypass exists for the upper internals assembly removed case. The presence of these natural circulation flow paths provide assurance that, in the event of a loss of the single residual heat removal loop, the backup decay heat removal capability afforded by the 23 feet of water above the vessel flange can be credited, regardless of whether the upper internals assembly is installed or removed.
3.2.2-22 Revised 04/17/2013 C23 REFERENCES, SECTION 3.2.2
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1967.
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presented at the European Atomic Energy Society Symposium on
Performance Experience of Water-Cooled Power Reactor Fuel, Stockholm, Sweden, October 21-22, 1969.
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WAPD-228, 1960.
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and WCAP-8219-A, March 1975.
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December 1960 and BMI-1518 (Rev.) for May 1961.
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3.2.2-23 Revised 04/17/2013
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RMB-527, 1964, Quoted by IAEA Technical Report Series No. 59, "Thermal
Conductivity of Uranium Dioxide."
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Seventh Conference of Thermal conductivity, P. 467, National Bureau of
Standards, Washington, 1968.
- 23. Godfrey, T. G., et al., "Thermal Conductivity of Uranium Dioxide and Armco Iron by an Improved Radial Heat Flow Technique," ORNL-3556, June
1964.
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February 1965.
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Division of the American Ceramic Society, Pittsburgh, September 1968.
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- 29. Lyons, M. F., et al., "UO 2 Powder and Pellet Thermal Conductivity During Irradiation," GEAP-5100-1, March 1966.
- 30. Coplin, D. H., et al., "The Thermal Conductivity of UO 2 by Direct In-reactor Measurements," GEAP-5100-6, March 1968.
- 31. Bain, A. S., "The Heat Rating Required to Produce Center Melting in Various UO 2 Fuels," ASTM Special Technical Publication, No. 306, pp.
30-46, Philadelphia, 1962.
- 32. Stora, J. P., "In-Reactor Measurements of the Integrated Thermal Conductivity of UO 2 - Effect of Porosity," Trans. ANS, 13, 137-138, 1970.
- 33. 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.
- 34. Slagle, W. H., "Operational Experience with Westinghouse Cores," (through December 31, 1994) WCAP-8183, Revision 23, January 1996.
- 35. Tong, L. S., "Prediction of Departure from Nucleate Boiling for an Axially Non-Uniform Heat Flux Distribution," Journal of Nuclear Energy, Vol. 21, pp. 241-248, (1967).
3.2.2-24 Revised 04/17/2013 C26
- 36. Motley, F. E., Mill, K. W. , et al., "New Westinghouse Correlation WRB-1 for Predicting Critical Heat Flux in Rod Bundles with Mixing Vane
Grids," WCAP-8762-P-A, July 1984.
- 37. Tong, L. S., "Boiling and Critical Heat Flux," TID-25887, 1972.
- 38. Judd, D. F., et al., "Non-Uniform Heat Generation Experimental Program" BAW-3238-7 (1965).
- 39. Lee, D. H. "An Experimental Investigation of Forced Convection Burnout in High Pressure Water, Part IV, Large Diameter Tubes at About 1600
psi", AEEW-R-479, Winfrith, England (1966).
- 40. Motley, F. E., Cadek, F. F., "Applications of Modified Spacer Factor to L. Grid Typical and Cold Wall Cell DNB, "WCAP-7988 (Westinghouse
Proprietary), and WCAP-8030-A (Non-Proprietary), and WCAP-8030-A (Non-Proprietary), October, 1972.
- 41. Dittus, F.W., and Boelter, L.M.K., "Heat Transfer in Automobile Radiators of the Tubular Type," Calif. Univ. Publication in Eng., 2, No. 13, 443461 (1930).
- 42. Weisman, J., "Heat Transfer to Water Flowing Parallel to Tube Bundles," Nucl. Sci. Eng., 6, 78-79 (1959)
- 43. Thom, J.R.S., Walker, W.M., Fallon, T.A. and Reising, G.F.S., "Boiling in Sub-cooled Water During Flowup Heated Tubes or Annuli," Prc. Instn.
Mech. Engrs., 180, Pt. C, 226-46 (1955-66).
- 44. Hill, K. W., Motley, F. E. and Cadek, F. F., "Effect of Local Heat Flux Spikes on DNB in Non-Uniform Heated Rod Bundles," WCAP-8174, August, 1973 (Proprietary) and WCAP-8202, August, 1973 (Nonproprietary).
- 45. Friedland, A. J., Ray, S., "Revised Thermal Design Procedure," WCAP-11397-P-A, April, 1989.
- 46. Cadek, F. F., "Interchannel Thermal Mixing with Mixing Vane Grids," WCAP-7667-P-A (Proprietary), January 1975 and WCAP-7755-A, January
1975.
- 47. Hochreiter, L. E., "Application of the THINC IV Program to PWR Design," WCAP-8054 (Proprietary), October 1973, and WCAP-8195, October 1973.
- 48. McFarlane, A. F., "Power Peaking Factors," WCAP-7912-P-A (Proprietary), January 1975 and WCAP-7912-A, January 1975.
- 49. Turkey Point Units 3 and 4 Uprating NSSS Engineering Report, Westinghouse, WCAP-14291, December 1995.
- 50. Hochreiter, L. E., Chelemer, H., and Chu, P. T., "THINC-IV An Improved Program for Thermal-Hydraulic Analysis of Rod Bundle Cores," WCAP-7956, June 1973.
- 51. Friedland, A. J., and Ray, S., "Improved THINC-IV Modeling for PWR Core Design," CAP-12330-P, August 1989.
- 52. Skaritka, J., (Ed.), "Fuel Rod Bow Evaluation," WCAP-8691, Rev. 1, July 1979.
- 53. "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 9, 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 16, 1982.
3.2.2-25 Revised 04/17/2013
- 54. Weisman, J., Wenzel, A. H., Tong, L. S., Fitzsimmons, D., Thorne, W., and Batch, J., "Experimental Determination of the Departure from Nucleate Boiling in Large Rod Bundles at High Pressure," AIChE, Preprint 29, 9th National Heat Transfer Conference, 1967, Seattle, Washington.
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Semi-International Symposium, Paper #256, 1967, Tokyo, Japan.
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- 57. Lahey, R. T., and Moody, F. J., "The Thermal Hydraulics of a Boiling Water Reactor," American Nuclear Society, 1977.
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Heat Transfer, No. 1976, pp. 616-622.
- 59. Virgil C. Summer FSAR, Docket #50-395.
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Parallel-Channel Upflow System," Proc. of 5th International Heat
Transfer Conference, Tokyo, Sept. 3-7, 1974.
- 63. Kao, H. S., Morgan, C. D., and Parker, W. B., "Prediction of Flow Oscillation in Reactor Core Channel," Trans. ANS, Vol. 16, 1973, pp.
212-213.
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- 65. Davidson, S. L. (Ed), "Vantage + Fuel Assembly Reference Core Report," WCAP-12610-P-A (proprietary), April 1995.
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Revision 0.
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(Westinghouse), June 18, 1986
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Subject:
"Request for Reduction in Fuel Assembly Burnup Limit for Calculation of Maximum
Rod Bow Penalty," June 18, 1986
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1990.
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Design (Proprietary/Non-Proprietary)," February 6, 2004.
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3.2.2-27 Revised 04/17/2013 C26 TABLE 3.2.2-1 THERMAL AND HYDRAULIC DESIGN PARAMETERS
Original EPU Total Primary Heat Output, MWt 2208 2652 Total Reactor Coolant Pump Heat Output, MWt 8 8
Total Core Heat Output, MWt 2200 2644 Total Core Heat Output, Btu/hr 7508.6 x 10 6 9022 x 10 6 Heat Generated in Fuel, %
97.4 97.4 Maximum Thermal Overpower, %
12 20 Nominal System Pressure, psia 2250 2250 Hot Channel Factors (First cycle)
(1) Heat Flux Nuclear, F N q 3.13 Engineering, F E 1.03 Total, F q 3.23 Enthalpy Rise Nuclear, F NH 1.75 Engineering, F EH 1.01 Total, FH 1.77 Coolant Flow Total Flow Rate 1b/hr 101.5 x 10 6 98.1 x 10 6 Average Velocity Along Fuel Rods, ft/sec 14.3 (14.0)
(2) Average Mass Velocity, 1b/hr-ft 2 2.32 x 10 6 (2.28 x 10 6)(2) Coolant Temperature, o F Nominal Inlet 546.2 549.2 Average Rise in Vessel 55.9 67.6 Average Rise in Core 58.3 (59.3)
(2) 71.6 Average in Core 575.4 (575.9) 587.1 Average in Vessel 574.2 583.0 Nominal Outlet of Hot Channel 642.0 (643.2)
(2)
Heat Transfer (First Cycle)
Active Heat Transfer Surface Area, ft 2 42,460 Average Heat Flux, Btu/hr-ft 2 171,600 Maximum Heat Flux, Btu/hr-ft 2 554,200 Maximum Thermal Output, kw/ft 17.9 Maximum Clad Surface Temperature at Nominal Pressure, o F 657 Maximum Average Clad Temperature at Rated
Power, o F 715 Fuel Centerline Temperatures, oF (First Cycle) Maximum at 100% Power 4150 Maximum at 112% Power 4400
DNB Ratio Minimum DNB Ratio at nominal operating Conditions 1.81 Typical Cell 2.069 Thimble Cell 2.055 Pressure Drop, psi (First Cycle)
Across Core 26 Across Vessel, including nozzles 46 Pressure Drop Across the Core, psi (OFA) 22.6 2.3 (Upgrade) 21.7 +
2.2 NOTES :
- 1. Core design values for current cycles are given in Appendices 14A and 14B.
- 2. Values for complete thimble plug removal.
Revised 04/17/2013 C26C26C26C26C26C26C26C26 TABLE 3.2.2-2
ENGINEERING HOT CHANNEL FACTORS
(FIRST CYCLE)
- To Point of Minimum DNB ratio
Rev. 16 10/99 1.03 Bowing and Pitch Diameter, Rod ty Eccentrici and , EnrichmentDensity Diameter, Pellet E q F *0.90
Mixing Flow 1.03
tion Redistribu Flow 1.01
ution MaldistribFlow Inlet 1.08 Bowing and Pitch Diameter, Rod Enrichment Density, Diameter, Pellet EH F 1.01
EHF Resulting
FLORIDA POWER & LIGHT COMPANY TURKEY POINT PLANT UNITS 3 & 4 THERMAL CONDUCTIVITY OF URANIUM DIOXIDE FIGURE 3.2.2-1
FLORIDA POWER & LIGHT COMPANY TURKEY POINT PLANT UNITS 3 & 4 COMPARISON OF W-3 PREDICTION AND UNIFORM FLUX DATA FIGURE 3.2.2-2
Rev. 4-7/86 FLORIDA POWER & LIGHT COMPANY TURKEY POINT PLANT UNITS 3 & 4 W-3 CORRELATION PROBABILITY DISTRIBTION CURVE FIGURE 3.2.2-3
Rev. 4-7/86 FLORIDA POWER & LIGHT COMPANY TURKEY POINT PLANT UNITS 3 & 4 COMPARISON OF W-3 CORRELATION WITH ROD BUNDLE DNB DATA (SIMPLE GRID WITHOUT MIXING VANE)
FIGURE 3.2.2-4
Rev. 4-7/86 FLORIDA POWER & LIGHT COMPANY TURKEY POINT PLANT UNITS 3 & 4 COMPARISON OF W-3 CORRELATION WITH ROD BUNDLE DNB DATA (SIMPLE GRID WITHOUT MIXING VANE)
FIGURE 3.2.2-5
Rev. 4-7/86 FLORIDA POWER & LIGHT COMPANY TURKEY POINT PLANT UNITS 3 & 4 MEASURE Vs. PREDICTED CRITICAL HEAT FLUX-WRB-1 CORRELATION FIGURE 3.2.2-6
Rev. 4-7/86 FLORIDA POWER & LIGHT COMPANY TURKEY POINT PLANT UNITS 3 & 4 COMPARISON OF W-3 PREDICTION AND NON-UNIFORM FLUX DATA FIGURE 3.2.2-7
Rev. 4-7/86 FLORIDA POWER & LIGHT COMPANY TURKEY POINT PLANT UNITS 3 & 4 COMPARISON OF W-3 PREDICTION WITH MEASURED DNB LOCATION FIGURE 3.2.2-8
Revised 04/17/13 FLORIDA POWER & LIGHT COMPANY TURKEY POINT PLANT UNITS 3 & 4 MEASURED VERSUS PREDICTED CRITICAL HEAT FLUX - ABB-NV CORRELATION FIGURE 3.2.2-9 C26 Revised 04/17/2013 FLORIDA POWER & LIGHT COMPANY TURKEY POINT PLANT UNITS 3 & 4 MEASURED VERSUS PREDICTED CRITICAL HEAT FLUX - WLOP CORRELATION FIGURE 3.2.2-10 C26 3.2.3 MECHANICAL DESIGN AND EVALUATION
The reactor core and reactor vessel internals are shown in cross-section in Figure 3.2.3-1 and in elevation in Figure 3.2.3-2. The core, consisting of the fuel assemblies, control rods, source rods and burnable poison rods provides and controls the heat source for the reactor operation. The internals, consisting of the upper and lower core support structure, are designed to support, align, and guide the core components, direct the coolant flow to and from the core components, and to support and guide the in-core instrumentation. A listing of the core mechanical design parameters is given in Table 3.2.3-1.
The seven grid 15x15 Upgrade (Upgrade) assemblies are arranged in a roughly circular cross-sectional pattern in the core. The 15x15 Upgrade and DRFA assemblies are nearly identical in their geometric configuration (number of fuel rods and thimble tubes, fuel rod dimensions, assembly pitch, etc.) with the following exceptions: 1) the addition of Intermediate Flow Mixing Grids (IFMs) and Protective Grid (P-grid); 2) balanced vaned mid grids; 3) shorter end plugs; 4) Debris Filter Bottom Nozzle (DFBN), and; 5) tube-in-tube dashpot. Note that evaluations documented in Section 3.1.1 support the complete or partial removal of thimble plugs from Turkey Point Units 3 & 4.
The enrichment of each region of fuel will vary slightly depending on the energy requirements for a given cycle of operation.
The fuel is in the form of slightly enriched uranium dioxide ceramic pellets.
The pellets are stacked to an active height of 144 inches within Zircaloy-4, ZIRLO or Optimized ZIRLO tubular cladding, which is plugged and seal welded at the ends to encapsulate the fuel. The enrichments of the fuel for the various regions in the first core are given in Table 3.2.3-1. Enrichments for subsequent cycles are given in the cycle specific Reload Characteristics and Parameters report included in Appendices 14A and 14B for Units 3 and 4, respectively. All fuel rods are internally pressurized with helium during fabrication. Heat generated by the fuel is removed by demineralized borated light water which flows upward through the fuel assemblies and acts as both moderator and coolant.
The stress criteria of Article 4 Section III of the ASME code is employed in the design of the fuel assembly with the exception of the fuel clad which is specifically excluded by the code. The criteria for the design of fuel rods are listed in Section 3.1.3. ZIRLO which is used for fabricating the guide thimbles of the fuel assembly and Inconel-718 which is used for fabricating grids and assembly hold down springs are not yet considered as code materials.
In Upgrade and DRFA fuel, intermediate grids are made of ZIRLO and the top, bottom, and protective grids of Inconel-718. The method for establishing design stress intensity values for the materials is consistent with that outlined in the code.
3.2.3-1 Revised 03/11/2016 C28 For historical purposes, the initial core was divided into regions of three different enrichments. The loading arrangement for the initial cycle is indicated on Figure 3.2.3-3. Refueling generally follows an inward loading schedule. However to increase neutron economy and as part of a vessel flux reduction program to resolve pressurized thermal shock concerns, the Turkey Point Units now employ a low leakage loading pattern aimed at reducing the neutron flux at the periphery of the core. This consists of using once-burned and twice-burned assemblies in peripheral core locations and loading a maximum number of fresh assemblies in-board while meeting energy requirements and ensuring the safe operation of the core. The typical reload core is divided into three different regions: feed, once burned and twice burned fuel assemblies of varied enrichments. The loading patterns for current cycles are given in the Reload Characteristics and Parameters report included in Appendices 14A and 14B for Units 3 and 4, respectively.
The control rods, designated as Rod Cluster Control Assemblies (RCCA's),
consist of groups of individual absorber rods which are held together with a spider at the top end and actuated as a group. The absorber rods fit within hollow guide thimbles in the fuel assemblies. The guide thimbles are an integral part of the fuel assemblies and occupy locations within the regular fuel rod pattern. In the withdrawn position, the absorber rods are guided and supported laterally by guide tubes which form an integral part of the upper core support structure. Figure 3.2.3-4 shows a typical RCCA.
As shown in Figure 3.2.3-2, the fuel assemblies are positioned and supported vertically in the core between the upper and lower core plates. The core plates are provided with pins which index into closely fitting mating holes in the fuel assembly top and bottom nozzles. The pins maintain the fuel assembly alignment which permits free movement of the control rods from the fuel assembly into the guide tubes in the upper support structure without binding or restriction between the rods and their guide surfaces.
Operational or seismic loads imposed on the fuel assemblies are transmitted through the core plates to the upper and low support structures and ultimately to the internals support ledge at the pressure vessel flange in the case of vertical loads or to the lower radial support and internals support ledge in the case of horizontal loads. The internals also provide a form fitting baffle surrounding the fuel assemblies which confines the upward flow of coolant in the core area to the fuel bearing region.
3.2.3-2 Revised 03/11/2016 C28 Reactor Internals
Design Description
The reactor internals are designed to support and orient the reactor core
fuel assemblies and RCCA's, absorb the RCCA dynamic loads and transmit these
and other loads to the reactor vessel flange, provide a passageway for the
reactor coolant, and support incore instrumentation. The reactor internals
are shown in Figure 3.2.3-2.
Stresses in reactor internals are computed following the rule established by
Section III of the ASME Code, Article 4.
Materials used for boltings, 316 SS, cold work, and Inconel X750 do not have
allowable and yield stresses in the Code. These bolts are not used as
pressure containing members where leakage and gasketing considerations are
paramount; they are instead used as mechanical attachments. The membrane
stress is kept within the yield stress of the material, including preload.
The basic philosophy is to obtain maximum clamping force and maintain joint
integrity during service.
Fabrication techniques regarding these materials are not different than the
materials considered by the Code except for the cold work of the 316 SS, and
the heat treatment of Inconel X750. The following summarizes these two
materials.
The materials used for bolting of the internals meet non-destructive testing
requirements comparable to requirements of Section III. In addition, all
bolting materials (of diameters less than two inches) have been
ultrasonically tested. Special mechanical property tests are applied to
these bolting materials where their properties were enhanced by strain
hardening of solution treated bars and by precipitation hardening
respectively. Specific written specifications are used for procurement of
these bolts.
3.2.3-3 Revised 04/17/2013 The Type 316 strain hardened bolting conforms to the following properties at room temperature. (The bolting material is strain hardened following a carbide solution treatment):
Bolting ASTM SA193 Type B8M Internals Bolting Requirement (Strain Hardened to Surface Brinell of 320 Max).
Up to 3/4" 3/4" to 1" 1" to 1-1/4" 1-1/4" to 1-1/2" Tensile Strength, psi min. Exceed the actual yield strength by: 15,000 psi for yield strengths of 75,000 psi and less; 10,000 psi for yield strength greater than 75,000 psi. 125,000 115,000 105,000 100,000 Yield Strength, psi min. 60,000 to 85,000 100,000 80,000 65,000 50,000 Elongation in %, Min 30 12 15 20 28 Reduction in Area, % Min 60 35 35 35 45 The strain hardened bolting materials require special processing which results in higher properties than annealed material similar to the approach used for special processing for SA453 Grades 651A, 651B, and 660. The base material chemistry is equivalent to SA193 B8M. The strain hardening concept is included in SA193, however, more restrictive mechanical property limits as noted above are applied.
The use of cold drawn material to enhance mechanical properties is similarly used for Nickel-Copper SB164.
Successful experience with strain hardened austenitic stainless steel (Type 304SS) of yield strength levels of 75,000 Psi Min. YS has been obtained on fuel element cladding used in SAXTON Nuclear Power Plant.
Nickel-Chromium age-hardenable alloys A461 Grade 688 is used for core internal bolting per Code Case 1390. The heat treatment used was optimized for the diameters used and differs from the established ASTM specification.
3.2.3-4 Revised 04/17/2013 The internals are designed to withstand the forces due to weight, preload of fuel assemblies, control rod dynamic loading, vibration, and LOCA blowdown
coincident with earthquake acceleration. The internals are analyzed in a
manner similar to Connecticut Yankee, San Onofre, Zorita, Saxton and Yankee.
Under the loading conditions, including conservative effects of design
earthquake loading, the structure satisfies stress values prescribed in
Section III, ASME Nuclear Vessel Code.
The reactor internals are equipped with bottom-mounted in-core
instrumentation supports. These supports are designed to sustain the
applicable loads outlined above.
The components of the reactor internals are divided into three parts
consisting of the lower core support structure (including the entire core
barrel and thermal shield), the upper core support structure and the in-core
instrumentation support structure.
Lower Core Support Structure
The major containment and support member of the reactor internals is the
lower core support structure, shown in Figure 3.2.3-5. This support
structure assembly consists of the core barrel, the core baffle, the lower
core plate and support columns, the thermal shield, the intermediate diffuser
plate and the bottom support plate which is welded to the core barrel. All
the major material for this structure is Type 304 Stainless Steel. The core
support structure is supported at its upper flange from a ledge in the
reactor vessel head flange and its lower end is restrained in its transverse
movement by a radial support system attached to the vessel wall. Within the
core barrel are axial baffle and former plates which are attached to the core
barrel wall and form the enclosure periphery of the assembled core. 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 provides the necessary flow distributor holes for each
fuel assembly. Fuel assembly locating pins (two for each assembly) are also
inserted into this plate. Columns are placed between this plate and the
bottom support plate of the core barrel in order to provide stiffness to this
plate and transmit the core load to the bottom support plate. Intermediate
between the support plate and lower core support plate is positioned a
perforated plate to diffuse uniformly the coolant flowing into the core.
3.2.3-5 Revised 04/17/2013 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. This bottom support allows for
differential axial growth of the thermal shield with respect to the core
barrel but restricts radial or horizontal movement of the bottom of the
shield. Irradiation baskets in which materials samples can be inserted and
irradiated during reactor operation are attached to the thermal shield. The
irradiation capsule basket supports are welded to the thermal shield so that
they do not extend above the thermal shield. Thus, the basket is not in the
high flow disturbance zone.
The welded attachment to the shield extends the full length of the support
except for small interruptions about one inch long. This type of attachment
has an extremely high natural frequency. The specimens are held in position
within the baskets by a stop on the bottom and a slotted cylindrical spring
at the top which fits against a relief in the basket. The specimen does not
extend through the top of the basket and thus is protected by the basket from
the flow.
The lower core support structure and principally the core barrel serve to
provide passageways and control for the coolant flow. Inlet coolant flow
from the vessel inlet nozzles proceeds down the annulus between the core
barrel and the vessel wall, flows on both sides of the thermal shield, and
then into a plenum at the bottom of the vessel. It then turns and flows
through the lower support plate, passes through the intermediate diffuser
plate and then through the lower core plate. The flow holes in the diffuser
plate and the lower core plate are arranged to give a very uniform entrance
flow distribution to the core. After passing through the core the coolant
enters the area of the upper support structure then flows generally radially
to the core barrel outlet nozzles and directly through the vessel outlet
nozzles.
A small amount of water also flows between the baffle plates and core barrel
to provide additional cooling of the barrel. Similarly, a small amount of
the entering flow is directed into the vessel head plenum and exits through
the vessel outlet nozzles.
Vertically downward loads from weight, fuel assembly preload, control rod
dynamic loading and earthquake acceleration are carried by the lower core
plate partially into the lower core plate support flange on the core barrel
shell and partially through the lower support columns to the bottom support
plate and thence through the core barrel shell to the core barrel flange
supported by the vessel head flange.
3.2.3-6 Revised 04/17/2013 Transverse loads from earthquake acceleration, coolant crossflow, and
vibration are carried by the core barrel shell to be distributed to the lower
radial support to the vessel wall and to the vessel head flange. Transverse
acceleration of the fuel assemblies is transmitted to the core barrel shell
by direct connection of the lower core support plate to the barrel wall and
by a radial support type connections of the upper core plate to slab sided
pins pressed into the core barrel.
The main radial support system of the core barrel is accomplished by "key" and "keyway" joints to the reactor vessel wall. At equally spaced points
around the circumference, an Inconel block is welded to the vessel I.D.
Another Inconel block is bolted to each of these blocks, and has a "keyway" geometry. Opposite each of these is a "key" which is attached to the
internals. At assembly, as the internals are lowered into the vessel, the
keys engage the keyways in the axial direction. With this design, the
internals are provided with a support at the farthest extremity, and may be
viewed as a beam fixed at the top and simply supported at the bottom.
Radial and axial expansions of the core barrel are accommodated, but
transverse movement of the core barrel is restricted by this design. With
this system, cycle stresses in the internal structures are within the ASME
Section III limits.
In the event of downward vertical displacement of the internals, energy
absorbing devices limit the displacement by contacting the vessel bottom
head. The load is transferred through the energy devices to the vessel. The
energy absorbers, cylindrical in shape, are contoured on their bottom surface
to the reactor vessel bottom head geometry. Their number and design are
determined so as to limit the forces imposed to a safe fraction of yield
strength. Assuming a downward vertical displacement, the potential energy of
the system is absorbed mostly by the strain energy of the energy absorbing
devices.
The free fall in the hot condition is on the order of 1/2 inch, and there is
an additional strain displacement in the energy absorbing devices of
approximately 3/4 inch. Alignment features in the internals prevent cocking
of the internals structure during this postulated drop. The RCCA's are
designed to provide assurance of control rod insertion capabilities under
these assumed drop of internals conditions. The drop distance of about 1-1/4
inch is not enough to cause the tips of the shutdown group of RCCA's in the
full withdrawn position to come out of the guide tubes in the fuel
assemblies.
3.2.3-7 Revised 04/17/2013 Upper Core Support Assembly
The upper core support assembly, shown in Figure 3.2.3-6, consists of the top
support plate, deep beam sections, and upper core plate between which are
contained support columns and guide tube assemblies. The support columns
establish the spacing between the top support plate, deep beam sections, and
the upper core plate and are fastened at top and bottom to these plates and
beams. The support columns transmit the mechanical loadings between the two
plates and serve the supplementary function of supporting thermocouple guide
tubes. The guide tube assemblies, shown on Figure 3.2.3-7, sheath and guide
the control rod drive shafts and control rods and provide no other mechanical
functions. They are fastened to the top support plate and are guided by pins
in the upper core plate for proper orientation and support. Additional
guidance for the control rod drive shafts is provided by the control rod
shroud tube which is attached to the upper support plate and guide tube.
The replacement RVCHs for Unit 3 and Unit 4 were fabricated without the CRDM
nozzles and nozzle adapters for the former part length CRDMs that had been
immobilized in the fully withdrawn position (References 18, 19, 23 and 24).
Six CRDM nozzle positions (B-8, F-6, F-10, H-14, K-6 & K-10) were affected.
The two former part length CRDM nozzles at positions H-1 and P-8 are retained
for mounting of the RVLMS probes. At the six affected nozzle locations, the
former part length CRDM lead screw (drive shaft) remained attached to the
RVCH and inserted into the control rod shroud tube when the reactor was
assembled. When the RVCH was removed the lead screw remained with the RVCH.
Deletion of the six part length CRDM nozzles, and thus removal of the
immobilized former part length CRDMs and lead screws from the replacement
RVCH, necessitated installation of a flow restrictor assembly in the control
rod shroud tube. The flow restrictor assembly is shown in Figure 3.2.3-17.
The flow restrictor assembly is bolted into the upper portion of the shroud
tube during a refueling outage using remote tooling. In order to insure the
assembly does not become a loose part in the reactor core, the assembly nut
is torqued onto the mandrel to sufficient preload to resist the imposed
hydraulic and thermal loads and then the nut is captured to prevent
loosening. One exception to this procedure has been evaluated for Unit 3
Cycle 23 for the flow restrictor at core location K-6. This flow restrictor
was torqued on the mandrel, but the mandrel did not seat on the upper core
support plate. This flow restrictor is approximately 0.17 inch short of
being fully seated. The nut on each flow restrictor is captured by the
plastic deformation of the thin ring at the top of the restrictor fingers
component into the flutes of the nut. Prior to installation and following
assembly of all of the parts, the top threads of the mandrel will be
distressed capturing the nut and preventing unintentional disengaging the nut
during installation.
3.2.3-8 Revised 04/17/2013 A further design feature is the j-lock slots in the outer flange. This
feature prevents the unintentional disengagement of the flow restrictor
assembly from the remote installation tooling during installation.
A thermal/hydraulic evaluation of the installed assembly was performed as
part of the design of the flow restrictor (Reference 20). The analysis
confirms that the flow restrictor assembly provides a flow resistance such
that the flow rate through the modified configuration is essentially
equivalent (within 2%) to the original configuration with the part length
CRDM lead screw inserted into the shroud tube. A structural evaluation (Reference 21) of the installed flow restrictor assembly was performed. The
installation design and process was further verified by testing (Reference
22). The evaluation and test results (References 21 and 22) demonstrated the
assembly meets the Code structural requirements for the specified design
conditions and could be successfully installed with its associated remote
installation tooling. It can be concluded the flow restrictor assembly
maintains the core hydraulic flow distribution with the former part length
CRDM lead screws removed under design basis normal, upset and faulted
conditions.
The upper core support assembly, which is removed as a unit during refueling
operation, is positioned in its proper orientation with respect to the lower
support structure by flat-sided pins pressed into the core barrel which in
turn engage in slots in the upper core plate. At an elevation in the core
barrel where the upper core plate is positioned, the flat-sided pins are
located at equal angular positions. Slots are milled into the core plate at
the same positions. As the upper support structure is lowered into the main
internals, the slots in the plate engage the flat-sided pins in the axial
direction. Lateral displacement of the plate and of the upper support
assembly is restricted by this design. Fuel assembly locating pins protrude
from the bottom of the upper core plate and engage the fuel assemblies as the
upper assembly is lowered into place. Proper alignment of the lower core
support structure, the upper core support assembly, the fuel assemblies and
control rods is thereby assured by this system of locating pins and guidance
arrangement. The upper core support assembly is restrained from any axial
movements by a large circumferential spring which rests between the upper
barrel flange and the upper core support assembly and is compressed by the
reactor vessel head flange.
Vertical loads from weight, earthquake acceleration, hydraulic loads and fuel
assembly preload are transmitted through the upper core plate via the support
columns to the deep beams and top support plate and then to the reactor
vessel head.
3.2.3-9 Revised 04/17/2013 Transverse loads from coolant crossflow, earthquake acceleration, and possible vibrations are distributed by the support columns to the top support plate and upper core plate. The top support plate is particularly stiff to minimize deflection.
In-Core Instrumentation Support Structures The in-core 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.
The upper system utilizes the reactor vessel head penetrations. Instru- mentation 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 thermocouples are routed through these port columns and across the upper support plate to positions above their readout locations. The thermocouple conduits are supported from the columns of the upper core support system. The thermocouple conduits are sealed stainless steel tubes.
During the Unit 4 Cycle 27 refueling outage, the following twenty two thimble tubes were replaced/installed: C-12, E-11, G-14, H-1, M-3, J-3, L-11, N-12, F-2, L-5, J-7, G-9, F-13, F-8, F-6, H-4, N-5, C-8, L-9, J-5, L-4, AND N-7.
Thimble tubes H-1 AND M-3 were capped due to not having their isolation valves, casings, fittings, and supporting frame within each respective tube. During the Unit 4 Cycle 29 refueling outage, thimble tubes H-1 and M-3 previously capped during Cycle 27 refueling outage, were restored to operational status. Therefore, following the Cycle 29 refueling outage, all thimble locations are available in Unit 4 for flux mapping. Capped thimble tubes periodically repositioned to minimize tube wall wear. While inspections may result in capping at additional thimble tube locations, the remaining number of detector thimbles will not decrease below the number of required thimbles available for peaking factor verification.
During Unit 3 Cycle 24 refueling outage, F-13, G-7, H-3, L-4,L-9, N-5, N-8, H-13, M-3 and J-12 thimble tube core locations were replaced. During the Unit 3 Cycle 27 refueling outage, D-12, E-11, N-10, B-7, D-10, J-10 and G-9 thimble tube core locations were replaced. While future ECT inspections may result in capping thimble tube locations, the remaining number of detector thimbles will not decrease below the number of required thimbles available for peaking factor verification. Capped thimble tubes are periodically repositioned to minimize tube wall wear.
3.2.3-10 Revised 06/23/2016 C28 In addition to the upper in-core instrumentation, there are reactor vessel
bottom port columns which 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 line. The minimum bend radii are about 90 inches
and the trailing ends of the thimbles (at the seal line) are extracted
approximately 13 feet during refueling of the reactor in order to avoid
interference within the core. The thimbles are closed at the leading ends
and serve as the pressure barrier between the reactor pressurized water and
the containment atmosphere.
Mechanical seals between the retractable thimbles and the surrounding
conduits are provided at the seal table. During normal operation, the
retractable thimbles are stationary in the core and move only during
refueling or for maintenance, at which time a space of approximately 13 feet
above the seal table is cleared for the retraction operation. Section 7.4
contains more information on the layout of the in-core instrumentation
system. Periodic inspections may result in the capping of some thimbles, the
number of in-service thimbles will not decrease below the number of required
thimbles available for peaking factor verification.
The incore instrumentation support structure is designed for adequate support of instrumentation during reactor operation and is rugged enough to resist
damage or distortion under the conditions imposed by handling during the
refueling sequence.
Evaluation of Core Barrel and Thermal Shield
The internals design is based on analysis, test and operational information.
Experience in previous Westinghouse PWR's has been evaluated and information
derived has been considered in this design. For example, Westinghouse uses a
one-piece thermal shield which is attached rigidly to the core barrel at one
end and flexured at the other.
The Connecticut Yankee reactor and the Zorita reactor core barrels are of the
same construction as the Turkey Point reactor core barrel. Deflection
measuring devices employed in the Connecticut Yankee reactor during the
hot-functional test, and deflection and strain gages employed in the Zorita
reactor during the hot-functional test have provided important information
that has been used in the design of the present day internals, including that
for Turkey Point. When the Connecticut Yankee thermal shield was modified to
the same design as for Southern California Edison, it, too, operated
satisfactorily as was evidenced by the examination after the hot-functional
test.
3.2.3-11 Revised 04/17/2013 C26 After these hot-functional tests on all of these reactors, a careful inspection of the internals was provided. All the main structural welds were examined, nozzle interfaces were examined for any differential movement, upper core plate inside supports were examined, the thermal shield attachments to the core barrel including all lockwelds on the devices used to lock the bolt were checked, no malfunctions were found.
Substantial scale model testing was performed by the Westinghouse Atomic Power Division (APD). This included tests which involved a complete full scale fuel assembly which was operated at reactor flow, temperature and pressure conditions. Tests were run on a 1/7th scale model of the Indian Point reactor. Measurements taken from these tests indicate very little shield movement, on the order of a few mils when scaled up to Turkey Point. Strain gage measurements taken on the core barrel also indicate very low stresses.
Testing to determine thermal shield excitation due to inlet flow disturbances has been included. Information gathered from these tests was used in the design of the thermal shield and core barrel. It can be concluded from the testing program and the analyses with the experience gained that the design as employed for Turkey Point is adequate.
Core Components Design Description
Fuel Assembly An upgrade design which incorporates several debris resistant features was introduced beginning in Cycle 25 of Unit 3 and Cycle 26 of Unit 4. The main features of the seven grid 15x15 Upgrade (Upgrade) includes 1) the addition of Intermediate Flow Mixing Grids(IFMs) and a Protective Grid (p-grid); 2) balanced vaned grids; 3) shorter end plug; 4) Debris Filter Bottom Nozzle (DFBN), and; 5) tube-in-tube dashpot. A detailed description of the DRFA is found in Reference 9. The overall configuration of the fuel assemblies is shown in Figure 3.2.3-9. A comparison between the upgrade and the DRFA designs is shown in Figure 3.2.3-9D. The assemblies are square in cross-section, nominally 8.426 inches on a side, and have an overall height of approximately 161.5 inches. The height of the active fuel column is 144 inches.
The upgrade fuel assembly fuel rod cladding, is fabricated from ZIRLO or Optimized ZIRLO material, and the guide tubes, instrument tubes, and mid-span grids are fabricated from ZIRLO to provide added corrosion resistance and fuel reliability. The use of ZIRLO material was submitted to the NRC for review in Reference 16 and received NRC approval. The use of Optimized ZIRLO material was submitted to the NRC for review in Reference 27 and received NRC approval.
In addition, the fuel rod length was increased slightly.
3.2.3-12 Revised 03/11/2016 C28C28 The fuel rods in a fuel assembly are arranged in a square array with 15 rod
locations per side and a nominal centerline-to-centerline pitch of 0.563 inch
between rods. Of the total possible 225 rod locations per assembly, 21 are
occupied by guide thimble tubes; 20 for the RCCA's or burnable absorbers and
one for in-core instrumentation. The remaining 204 locations contain fuel
rods. In addition to fuel rods, a fuel assembly is composed of a top nozzle, a bottom nozzle, 11 grid assemblies (five ZIRLO mixing vane grids, three
ZIRLO IFM grids, two end Inconel-718 grid and one Inconel-718 P-grid), 20
absorber rod guide thimble tubes, and one instrumentation thimble tube.
The guide thimble tubes in conjunction with the grid assemblies and the top
and bottom nozzles comprise the basic structural fuel assembly skeleton. The
top and bottom ends of the guide thimble tubes are fastened to the top and
bottom nozzles respectively. The grid assemblies, in turn, are fastened to
the guide thimble tubes at each location along the height of the fuel
assembly at which lateral support for the fuel rods is required. Within this
skeletal framework the fuel rods are contained and supported and the
rod-to-rod centerline spacing is maintained along the assembly.
Bottom Nozzle
The bottom nozzle is a square box-like structure which controls the coolant
flow distribution to the fuel assembly and functions as the bottom structural
element of the fuel assembly. The nozzle, which is square in cross-section, is fabricated from 304 stainless steel parts consisting of a perforated
plate, 4 angle legs, and four pads or feet. The legs are welded to the plate
to form a plenum and for the inlet coolant to the fuel assembly. The
perforated plate serves as the bottom end support for the fuel rods. The
bottom support surface for the fuel assembly is formed by the four pads which
are welded to the corner angles. Starting in Turkey Point Unit 4 Cycle 13, the bottom nozzle was stiffened with 1/4-inch x 1-inch bars extending between
the legs. Beginning with the Turkey Point Unit 3 Cycle 14 reload (Region 16),
the fuel assemblies contain a cast composite bottom nozzle. The composite
bottom nozzle is functionally interchangeable with the old design.
Coolant flow to the fuel assembly is directed from the plenum in the bottom
nozzle upward to the interior of the fuel assembly through the holes in the
nozzle plate and to the channel between assemblies through the spaces between
the corner legs. The ligaments between the holes of the nozzle plate are
positioned laterally beneath the fuel rods to prevent passage of the rods
beyond this surface.
3.2.3-13 Revised 04/17/2013 C25 The guide thimble tubes, which carry axial loads imposed on the assembly, are
fastened to the bottom nozzle end plate. These loads as well as the weight
of the assembly are distributed through the nozzle to the lower core support
plate. Indexing and positioning of the fuel assembly in the core is
controlled through two holes in diagonally opposite pads which mate with
locating pins in the lower core plate. Lateral loads imposed on the fuel
assembly are also transferred to the core support structures through the
locating pins.
The Upgrade bottom nozzle assembly is essentially the same design as the DRFA
bottom nozzle, except for the flow hole pattern of the nozzle plate. The
diameter of the flow holes is reduced to enhance debris filtering.
The Upgrade and DRFA bottom nozzle designs have a feature which allows it to
be easily removed. As shown in Figure 3.2.3-9B, a locking cup is used to
lock the thimble screw of a guide thimble tube in place, instead of the
lockwire used for the standard LOPAR nozzle design. The reconstitutable
nozzle design facilitates removal of the bottom nozzle and relocking of
thimble screws as the bottom nozzle is reattached.
Top Nozzle
The top nozzle is a box-like structure, which functions as the fuel assembly
upper structural element and forms a plenum space where the heated fuel
assembly discharge coolant is directed toward the flow holes in the upper
core plate. The nozzle is comprised of an adapter plate, enclosure, top
plate, two clamps, four leaf-spring sets, and assorted hardware. All parts
with the exception of the springs and their hold screws are constructed of
Type 304 stainless steel. The screws are made from Inconel 718 or Inconel
600, depending on the fuel region. The springs are made from Inconel 718.
The adaptor plate is square in cross-section, and is perforated by machined
slots to provide for coolant flow through the plate. At assembly, the top
ends of the control guide thimble tubes are fastened to the adaptor. Thus, the adaptor plate acts as the fuel assembly top end plate, and provides a
means of distributing evenly among the guide thimble tubes any axial loads
imposed on the fuel assemblies.
The Upgrade assembly uses the same top nozzle as the DRFA assembly.
The nozzle enclosure is actually a square thin walled tubular shell which
forms the plenum section of the top nozzle. The bottom end of the enclosure
is pinned and welded to the periphery of the adaptor plate, and the top end
is welded to the periphery of the top plate.
3.2.3-14 Revised 04/17/2013 The top plate is square in cross-section with a central hole. The hole
allows clearance for the RCCA absorber rods to pass through the nozzle into
the guide thimbles in the fuel assembly and for coolant exiting from the fuel
assembly to the upper internals area. Two pads containing axial
through-holes which are located on diametrically opposite corners of the top
plate provide a means of positioning and aligning the top of the fuel
assembly. As with the bottom nozzle, alignment pins in the upper core plate
mate with the holes in the top nozzle plate.
Hold down forces of sufficient magnitude to oppose the hydraulic lifting
forces on the fuel assembly are obtained by means of the leaf springs which
are mounted on the top plate. The Upgrade fuel has a slight decrease in
hydraulic resistance to flow compared to the DRFA assembly, primarily due to
lower mid grid loss coefficients. This results in a decreased Upgrade lift
force. Therefore, the current 3-leaf spring top nozzle used with the DRFA
assembly can be used with the Upgrade fuel. The springs are fastened to the
top plate at the two corners where alignment holes are not used and radiate
out from the corners parallel to the sides of the plate. Fastening of the
springs is accomplished with a clamp which fits over the ends of the springs
and two bolts (one per spring) which pass through the clamp and spring, and
thread into the top plate. At assembly, the spring mounting bolts are
torqued sufficiently to preload against the maximum spring load and then
lockwelded to the clamp which is counter-bored to receive the bolt head.
The spring load is obtained through deflection of the spring set by the upper
core plate. The spring form is such that it projects above the fuel assembly
and is depressed by the core plate when the internals are loaded into the
reactor. The free end of the spring is bent downward into a key slot in the
top plate. The free end of the lower spring is captured by the leg of the
upper spring. This is done to guard against loose parts in the reactor in
the event of spring fracture. In addition, the fit between the upper spring
and key slot and between the spring set and the mating slot in the clamp are
sized to prevent rotation of either end of the spring set into the control
rod path in the event of spring fracture.
In addition to its plenum and structural functions, the nozzle provides a
protective housing for components which mate with the fuel assembly. In
handling a fuel assembly with a control rod inserted, the control rod spider
is contained within the nozzle. During operation in the reactor, the nozzle
protects the absorber rods from coolant cross flows in the unsupported span
between the fuel assembly adaptor plate and the end of the guide tube in the
upper internals package. The spiders which support the source rods and
burnable poison rods are all contained within the fuel top nozzle.
3.2.3-15 Revised 04/17/2013 Beginning with the Unit 3 Cycle 12 reload, a reconstitutable top nozzle (RTN) was incorporated into the Turkey Point fuel assembly design. The RTN can be
removed and reattached repeatedly, if necessary, throughout the assembly's
operation. Guide thimble inserts are placed into the circumferentially
grooved holes in the adapter plate to connect the RTN with the fuel assembly.
A lock tube, inserted into the inner diameter of the insert, holds the insert
in place until removal of the RTN is required. The RTN design has the same
flow area and loss coefficients as the previous design, therefore, none of
the core/fuel inputs to the safety analysis were affected by the introduction
of this design. Beginning with Turkey Point Unit 3 Cycle 16, the
Composite(CAST) Top Nozzle was implemented and is functionally
interchangeable with the old design. The keyless/cuspless nozzle features
implemented during Unit 3 Cycle 14 were retained in the modified CAST design.
Guide Thimbles
The control rod guide thimbles in the fuel assembly provide guided channels for the absorber rods during insertion and withdrawal of the control rods.
Beginning with Unit 3 Cycle 25 and Unit 4 Cycle 26, tube-in-tube guide
thimbles will be used. The tube-in-tube design utilizes a separate dashpot
tube assembly that is inserted into the guide thimble assembly pulled to a
press fit over the thimble end plug and bulged into place. The guide thimble
and dashpot are fabricated from ZIRLO tubing, The larger inside diameter at
the top provides a relatively large annular area for rapid insertion during a
reactor trip and to accommodate a small amount of upward cooling flow during
normal operations. The bottom portion of the guide thimble is of reduced
diameter to produce a dashpot action when the absorber rods near the end of
travel in the guide thimbles during a reactor trip.
Flow holes are provided just above the dashpot to permit the entrance of
cooling water during normal operation, and to accommodate the outflow of
water from the dashpot during reactor trip.
The tube-in-tube dashpot is closed at the bottom by means of a welded end
plug. The end plug is fastened to the bottom nozzle during fuel assembly
fabrication.
Grids The grid assemblies consist of individual slotted straps which are assembled and interlocked in an "egg-crate" type arrangement and then furnace brazed or
welded to permanently join the straps at their points of intersection.
Details such as spring fingers, support dimples, mixing vanes, and tabs are
punched and formed in the individual straps prior to assembly.
3.2.3-16 Revised 04/17/2013 Two types of grid assemblies are used in the fuel assembly. One type of these grids having mixing vanes which project from the edges of the straps
into the coolant stream is used in the high heat region of the fuel
assemblies to promote mixing of the coolant. A grid of this type is shown in
Figure 3.2.3-10. The other type of grids, located at the bottom and top ends
of the assembly, are of the nonmixing type. They are similar to the mixing
type with the exception that they contain no mixing vanes on the internal
straps.
In addition to the above, two other grid types are used in the core. These
are the protective grid (P-Grid) and the Intermediate Flow Mixing (IFM) grid.
The Upgrade design includes a P-grid at the bottom of the assembly to provide
an additional debris barrier thereby improving fuel reliability. The P-grid
also provides additional fretting resistance by supporting the bottom of the
fuel rod. The upgrade fuel rods are positioned closer to the bottom and are
modified, with a shorter end bottom end plug compared to the DRFA fuel rods.
Also included in the Upgrade design are Intermediate Flow Mixing (IFM) Grids.
IFM grids are considered nonstructural upper assembly grids which contain
mixing vanes similar to ZIRLO structural grids. Specifically, the IFM grids
are located in the top three mixing vane grid spans. The 15x15 IFM grids
have less than one third the mass of the 15x15 DRFA mid grids. These IFM
grids will have virtually no effect of neutron economy.
As expected, the addition of IFM grids to the fuel assembly increases the
pressure drop across than assembly. In order to alleviate this impact, the
mid grids on the 15x15 Upgrade design were developed to utilize the Low
Pressure Drop (LPD) structural grids in conjunction with these IFM grids.
The LPD grids are reduced in height with chamfered-up stream edges. While
these modifications do not detract from the structural capabilities of the
grid, they significantly reduce the pressure drop of the grids and therefore
allow adding IFM grids with minimal impact.
There are two materials used to construct support grids for the Upgrade and DRFA assemblies. Inconel-718 is used for the top, bottom and protective grid
in the Upgrade assembly. ZIRLO is used for the five intermediate mixing-vane
grids and three IFM grids in the Upgrade assembly. A more detailed
description of the DRFA assembly can be found in the Reload Transition Safety
Report (RTSR) for Turkey Point Units (Reference 2).
The outside straps on all grids contain mixing vanes which, in addition to
their mixing function, aid in guiding the grids and fuel assemblies past
projecting surfaces during handling or loading and unloading the core.
Additional small tabs on the outside straps and the irregular contour of the
straps are also for this purpose.
3.2.3-17 Revised 04/17/2013 Inconel-718 and ZIRLO are used for the grid material because of their corrosion resistance and high strength properties. The Inconel grids are furnace brazed to permanently join the straps at their intersections. After the combined brazing and solution annealing temperature cycle, the grid material is age hardened to obtain the material strength necessary to develop the required grid spring forces. The ZIRLO grid interlocking strap joints and grid-to-sleeve joints are fabricated by laser welding.
Impact tests have been performed at 600F to obtain the dynamic strength data and verify that the ZIRLO grid strength data at reactor operating conditions is structurally acceptable. Some ZIRLO gird designs experience grid crush during the most severe load conditions of a combined seismic/LOCA event. However, crushed grid locations are limited to the periphery of the core and coolable geometry is maintained. Control rod insertability and fuel cladding integrity are also maintained.
Beginning with the Unit 3 Cycle 13 reload, snag-resistant top, middle and bottom grids were introduced. In this design, the outer grid straps are modified to help prevent fuel assembly hangup due to grid strap interference during fuel assembly removal and insertion. This was accomplished by changing the grid strap corner geometry and adding guide tabs on the outer grid strap.
Fuel Rods
The fuel rods consist of uranium dioxide ceramic pellets contained in a slightly cold worked and stress relieved Zircaloy-4, ZIRLO or Optimized ZIRLO tubing which is plugged and seal welded at the ends to encapsulate the fuel. Sufficient void volume and clearances are provided within the rod to accommodate fission gases released from the fuel, differential thermal expansion between the cladding and the fuel, and fuel swelling due to accumulated fission products without over-stressing of the cladding or seal welds. Shifting of the fuel within the cladding is prevented during handling or shipping prior to core loading by a stainless steel helical compression spring which bears on the top of the fuel.
The fuel rods employed in Upgrade and DRFA assemblies are geometrically identical with only slight variations in some design parameters. The Upgrade assembly, utilizes a shorter end plug and the centerline of the pellet stack is approximately 2.103 inches lower than the DRFA fuel. On a cycle-to-cycle and region-to-region basis, fuel enrichment, plenum void volume and initial helium backfill pressure will vary somewhat to accommodate specific cycle design requirements. This fact was also applicable prior to the introduction of Upgrade assemblies. For Unit 3 beginning with cycle-12, the DRFA incorporates axial blankets which consist of low enriched or natural uranium oxide pellets extending 6 inches at the top and bottom of the fuel stack within the fuel rod.
3.2.3-18 Revised 03/11/2016 C28 Unit 4 has axial blankets starting with Cycle 14 and DRFA starting with Cycle
- 13. Starting with Unit 3 Cycle 25 and Unit 4 Cycle 26, the Upgrade fuel will
use 8-inch axial blankets.
During fuel rod assembly, the pellets are stacked in the cladding to the
required fuel height. The compression spring is then inserted into the top
end of the fuel and the end plugs pressed into the ends of the tube and
welded. All fuel rods are internally pressurized with helium during the
welding process. A hold-down force in excess of the weight of the fuel is
obtained by compression of the spring between the top end plug and the top of
the fuel pellet stack. Starting with Unit 3 Cycle 25 and Unit 4 Cycle 26, the Upgrade fuel will use 8-inch axial blankets.
The fuel pellets are right circular cylinders consisting of slightly enriched
uranium-dioxide powder which is compacted by cold pressing and sintering to
the required density. The ends of each pellet are dished slightly to allow
the greater axial expansion at the center of the pellets to be taken up
within the pellets themselves and not in the overall fuel length. The ends
of each Upgrade and DRFA fuel pellet have a small chamfer at the outer
cylinder surface.
For historical purposes, the pellet densities in the initial core were adjusted as shown in Table 3.2.3-1 to compensate for the effects of the
higher burnup of fuel in regions remaining longest in the core. A different
fuel enrichment as listed in Table 3.2.3-1 is used for each of the three
regions in the first core loading. Reload region, as-built fuel enrichments
and pellet densities are provided in the Reload Characteristics and
Parameters report included in Appendices 14A and 14B for Units 3 and 4, respectively.
To prevent the possibility of mixing enrichments during fuel manufacture and
assembly, meticulous process control is exercised.
Process Control
Powder withdrawal from storage can be made by one authorized group only who
direct the powder to correct pellet production line. All pellet production
lines are physically separated from each other and pellets of only a single
enrichment and density are produced in a given production line.
Finished pellets are placed on trays having the same color code as the powder
containers and transferred to segregated storage racks within the confines of
the pelleting area. Physical barriers prevent mixing of pellets of different
densities and enrichments in this storage area. Unused powder and
substandard pellets are returned to storage in the original color coded
containers.
3.2.3-19 Revised 04/17/2013 Each fuel assembly will be identified by means of a serial number engraved on
the upper nozzle. The fuel pellets will be fabricated by a batch process so
that only one enrichment region is processed at any given time. The serial
numbers of the assemblies and corresponding enrichment will be documented by
the manufacturer and verified prior to shipment.
Each assembly will be assigned a core loading position. A record will then
be made of the core loading position, serial number and enrichment. Prior to
core loading, independent checks will be made to ensure that this assignment
is correct.
During initial core loading and subsequent refueling operations, detailed
handling and checkoff procedures will be utilized throughout the sequence.
The initial core will be loaded in accordance with the core loading diagram
similar to Figure 3.2.3-3 which shows the location for each of the three
enrichment types of fuel assemblies in the region. Reload cycle core loading
pattern diagrams are provided in the cycle specific Reload Characteristics
and Parameters report included in Appendices 14A and 14B for Units 3 and 4
respectively.
Rod Cluster Control Assemblies
The rod cluster control assemblies (RCCA) each consist of a group of
individual absorber rods fastened at the top end to a common hub or spider
assembly. These assemblies one of which is shown in Figure 3.2.3-4 are
provided to control the reactivity of the core under operating conditions.
These assemblies consist of rods containing full length absorber material.
The number of RCCA's is specified in Table 3.2.3-1.
The absorber material used in the control rods is silver-indium-cadmium alloy
which is essentially "black" to thermal neutrons and has sufficient
additional resonance absorption to significantly increase its worth. The
alloy is in
the form of extruded single length rods which are sealed in stainless steel
tubes to prevent the rods from coming in direct contact with the coolant.
The overall control rod length is such that when the assembly has been
withdrawn through its full travel, the tip of the absorber rods remain
engaged in the guide thimbles so that alignment between rods and thimbles is
always maintained. Since the rods are long and slender, they are relatively
free to conform to any small misalignments with the guide thimble. Prototype
tests have shown that the RCCA's are very easily inserted and not subject to
binding even under conditions of severe misalignment.
3.2.3-20 Revised 04/17/2013 The spider assembly is in the form of a center hub with radial vanes
containing cylindrical fingers from which the absorber rods are suspended.
Handling detents, and detents for connection to the drive shaft, are machined
into the upper end of the hub. A spring pack is assembled into a skirt
integral to the bottom of the hub to stop the RCCA's and absorb the impact
energy at the end of a trip insertion. A centerpost which holds the spring
pack and its retainer is threaded into the hub within the skirt and welded to
prevent loosening in the service. All components of the spider assembly are
made from Type 304 stainless steel except for the springs which are Inconel
X-750 alloy and the retainer which is of 17-4 Ph material.
The absorber rods are secured to the spider so as to assure trouble free
service. The rods are first threaded into the spider fingers and then pinned
to prevent rotation, after which the pins are welded in place. The end plug
below the pin position is designed with a reduced section to permit flexing
of the rods to correct for small assembly misalignments.
In construction, the silver-indium-cadmium rods are inserted into cold-worked
stainless steel tubing which is then sealed at the bottom and the top by
welded end plugs. Sufficient diametrical and end clearance is provided to
accommodate relative thermal expansions and to limit the internal pressure to
acceptable levels.
The bottom plugs are made bullet-nosed to reduce the hydraulic drag during a
reactor trip and to guide smoothly into the dashpot section of the fuel
assembly guide thimbles. The upper plug is threaded for assembly to the
spider and has a reduced end section to make the joint more flexible.
Stainless steel clad silver-indium-cadmium alloy absorber rods are resistant
to radiation and thermal damage ensuring their effectiveness under all
operating conditions.
Neutron Source Assemblies
Four neutron source assemblies were utilized in the initial core. They
consisted of two secondary source assemblies each, and two primary source
assemblies each. The rods in the source assembly are fastened to a spider at
the top end similar to the RCCA spiders.
For historical purposes, in the core, the neutron source assemblies were
inserted into the RCCA guide thimbles in fuel assemblies at unrodded
locations. The location of these assemblies in the core is shown in Figure
3.2.3-3. Based on the evaluations documented in References 6 and 7 of Section
3.1.1, it has been determined that it is acceptable to remove these sources
from Turkey Point Units 3 & 4.
3.2.3-21 Revised 04/17/2013 Consequently, no startup sources have been used beginning with Cycle 13 in
both units.
The primary and secondary source rods both utilized the same type of cladding
material as the absorber rods (cold-worked type 304 stainless steel tubing, 0.432 in O.D. with 0.019 inch thick walls). The secondary source rods
contain Sb-Be pellets stacked to a height of 121.75 inches. Design criteria
similar to those for the fuel rods are used for the design of the source
rods; ie, the cladding is free standing, internal pressures are always less
than reactor operating pressure, and internal gaps and clearances are
provided to allow for differential expansions between the source material and
cladding.
Thimble Plug Assemblies
Evaluations have been performed to support the complete or partial removal of
thimble plugs from Turkey Point Units 3 & 4. These evaluations have
addressed the effect of thimble plug removal on Core Design, Core Thermal
Hydraulics, Reactor Pressure Vessel System thermal hydraulics and the non-
LOCA and LOCA safety analyses. Based on these evaluations, it has been
determined that it is acceptable to remove all or any combination of thimble
plugs from Turkey Point Units 3 & 4.
The thimble plug assemblies as shown in Figure 3.2.3-10A consist of a flat
base plate with short rods suspended from the bottom surface and a spring
pack assembly. The twenty short rods, called thimble plugs, project into the
upper ends of the guide thimbles to reduce the bypass flow area. Similar
short rods are also used on the source assemblies and burnable poison
assemblies to plug the ends of all vacant fuel assembly guide thimbles. At
installation in the core, the thimble plug assemblies interface with both the
upper core plate and with the fuel assembly top nozzles by resting on the
adaptor plate. The spring pack is compressed by the upper core plate when
the upper internals assembly is lowered into place. Each thimble plug is
permanently attached to the base plate by a nut which is locked to the
threaded end of the plug by a small lock-pin welded to the nut.
All components in the thimble plug assembly, except for the springs, are
constructed from type 304 stainless steel. The springs are wound from an age
hardened nickel base alloy for corrosion resistance and high strength.
Burnable Poison Rod
The burnable poison rods are statically suspended and positioned in vacant
RCC thimble tubes within the fuel assemblies at nonrodded core locations.
The poison rods in each fuel assembly are grouped and attached together at
the top end of the rods by a flat base plate which fits within the fuel
assembly top nozzle and rests on the top adaptor plate.
3.2.3-22 Revised 04/17/2013 The base plate and the poison rods are 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 package
is lowered into the reactor. This ensures that the poison rods cannot be
lifted out of the core by flow forces.
Several types of burnable absorbers have been utilized in Turkey Point Units
3 and 4. Typically, both full-length and part-length borosilicate burnable
poison rods which consist of borosilicate glass tubes contained within type
304 stainless steel cladding were used in LOPAR assemblies. These are plugged
and sealed at both ends to encapsulate the glass. The glass is also supported
along the length of its inside diameter by a thin wall type 304 stainless
steel, tubular inner liner (Figure 3.2.3-11).
The second major type is the Wet Annular Burnable Absorber (3) (WABA) (Figure 3.2.3-11A) which will, as necessary, be used with OFA assemblies. The WABA
consists of an annular aluminum oxide-boron carbide (A1 2 O 3-B 4 C) absorber pellets in two concentric Zircaloy tubes. Coolant flows through the center
holes as well as through the outer annulus between the WABA and the guide
thimble tube. The WABA design provides significantly enhanced nuclear
characteristics compared with the borosilicate absorber rod design. Fuel
cycle benefits result from the reduced parasitic nature of Zircaloy compared
to stainless steel tubes, increased water fraction in the burnable absorber
cell, and a reduced reactivity penalty at the end of each cycle.
The third major type of neutron absorber that can be used is the Hafnium
Vessel Flux Depression (HVFD) absorber (Figure 3.2.3-14). The HVFD consists
of a reduced length annular hafnium absorber axially positioned within
Zircaloy cladding. The primary function of the HVFD is to provide for
reactor vessel flux reduction to satisfy pressurized thermal shock
considerations (1).
The fourth major type of burnable absorber currently utilized is the Integral
Fuel Burnable Absorber (IFBA). The IFBA rods have a thin boride coating on
the cylindrical surface of the fuel pellets along the central portion of the
fuel stack length. In order to offset the effects of the Helium gas release
from the IFBA coating during irradiation, a lower initial Helium backfill
pressure is used in the IFBA rods compared to the non-IFBA fuel rods.
The initial implementation of IFBA underwent a detailed qualification at the
manufacturer and demonstration program to demonstrate integrity during
extended duty. This program included demonstration assemblies which were
loaded in the Unit 4 Cycle 10 core (Reference 13). The mechanical design of
these assemblies was identical to that of the assemblies in that reload
region, except that the demonstration assemblies were of the removable rod
type, discussed below. A detailed description of the IFBA fuel rods is found
in References 14 and 15.
3.2.3-23 Revised 04/17/2013 For historical purposes, the following discussion of the borosilicate glass
BP rod design is retained for completeness. This burnable poison design has
not been used at Turkey Point since Unit 3 Cycle 9 and Unit 4 Cycle 10. The
rods are designed in accordance with the standard fuel rod design criteria;
i.e., the cladding is free standing at reactor operating pressures and
temperatures and sufficient cold void volume is provided within the rods to
limit internal pressures to less than the reactor operating pressure assuming
total release of all helium generated in the glass as a result of the B 10 (n,a) reaction. The large void volume required for the helium is obtained
through the use of glass in tubular form which provides a central void along
the length of the rods. A more detailed discussion of the borosilicate glass
BP rod design is found in WCAP 9000 (4).
Based on available data on properties of Pyrex glass and on nuclear and
thermal calculations for the rods, gross swelling or cracking of the glass
tubing is not expected during operation. Some minor creep of the glass at
the hot spot on the inner surface of the tube is expected to occur but
continues only until the glass comes into contact with the inner liner. The
inner liner is provided to maintain the central void along the length of the
glass and to prevent the glass from slumping or creeping into the void as a
result of softening at the hot spot. The wall thickness of the inner liner
is sized to provide adequate support in the event of slumping but to collapse
locally before rupture of the exterior cladding if large volume changes due
to swelling or cracking should possibly occur. The top end of the inner
liner is open to receive the helium which diffuses out of the glass.
To ensure the integrity of the burnable poison rods, the tubular cladding and
end plugs are procured to the same specifications and standard of quality as
is used for stainless steel fuel rod cladding and end plugs in other
Westinghouse plants. In addition, the end plug seal welds are checked for
integrity by visual inspection and x-ray. The finished rods are helium leak
checked.
Removable Rod Assemblies
Four demonstration assemblies were loaded in Turkey Point Unit 4, Cycle 10.
Each of the assemblies contains twenty-eight demonstration Integral Fuel
Burnable Absorber (IFBA) fuel rods and one hundred and seventy-six unpoisoned
fuel rods.
The mechanical design of the demonstration assemblies is identical to that of
the other fuel assemblies in the reload region, except that the demonstration
assemblies are the "removable rod" type, which will allow removal of some
rods for post-irradiation inspections.
3.2.3-24 Revised 04/17/2013 Similar "removable rod" type Optimized Fuel Assemblies have been used in
previous demonstration assembly programs at Farley, Salem, Beaver Valley and
Point Beach reactors. The mechanical design of the assemblies has been
evaluated and meets the same acceptance criteria as the standard fuel
assembly design for steady state, transient, seismic and LOCA conditions.
The design of the fuel rods contained in the demonstration assemblies is
identical to the fuel rod design of the other fuel rods in the reload region
except that:
- 1. in each of the demonstration assemblies, there are fifty-two removable fuel rods (sixteen removable IFBA rods, 36 removable non-IFBA rods); these removable fuel rods have longer, more
slender top end plugs to facilitate rod removal and a larger
chamfer on their bottom end plugs to ease fuel rod reinsertion;
- 2. for the IFBA fuel rods only (sixteen removable IFBA fuel rods and twelve non-removable IFBA fuel rods per assembly), each fuel stack
contains absorber material coated on the outside diameter of the
U0 2 fuel pellets and distributed uniformly over the entire fuel stack; because the burnable absorber material releases additional
helium into the fuel rod during depletion, the IFBA fuel rods are
prepressurized to 200 psig during manufacture, whereas the
non-IFBA fuel rods in the demonstration assemblies, and the
standard fuel rods throughout the reload region, are
prepressurized 350 psig.
- 3. the core locations of the IFBA demonstration assemblies were chosen such that the IFBA fuel rods are never the lead power rods.
Based on review of the appropriate phase diagram and on destructive examination after one reactor cycle of test rods incorporating coated pellets
essentially identical in material and manufacture method, no adverse chemical
interaction of the absorber material with either cladding or fuel pellet is
predicted for the times and temperatures of operation.
The approved fuel rod model (PAD)
(5) was used to assess in detail the fuel rod design criteria influenced by addition of the absorber material. Based upon a consideration of clad stress, fuel temperatures, and rod internal
pressure, an allowable burnup for the demonstration rods in excess of the
planned burnup of fuel assemblies was calculated. No adverse effects on fuel
rod performance were predicted.
3.2.3-25 Revised 04/17/2013 Evaluation of Core Components Fuel Rod Evaluation
The fission gas release and the associated buildup of internal gas pressure in the fuel rods is calculated by the PAD code (References 10 and 17) based on experimentally determined rates. The increase of internal pressure in the fuel rod due to this phenomenon is included in the determination of the maximum cladding stresses at the end of core life when the fission product gap inventory is a maximum. Per Reference 25, the fission gas release calculated by the PAD model documented in Reference 17 is acceptable to use with thermal conductivity degradation (TCD) since the effects of TCD are implicitly included in the PAD model calibration.
The maximum allowable strain in the cladding, considering the combined effects of internal fission gas pressure, external coolant pressure, fuel pellet swelling and clad creep is limited to less than 1 percent throughout core life. The combined maximum stress intensity meets the criteria based on the ASME code in Addendum 1-A of Reference 26 or the associated stresses are below the yield strength of the material under all normal operating and overpower conditions.
To assure that manufactured fuel rods meet a high standard of excellence from the standpoint of functional requirements, many inspections and tests are performed both on the raw material and the finished product. These tests and inspections include chemical analysis, tensile and ultrasonic testing of fuel tubes, dimensional inspection, ultrasonic test or x-ray of both end plug welds, gamma scanning and helium leak tests.
In the event of cladding defects, the high resistance of uranium dioxide fuel pellets to attack by hot water protects against fuel deterioration or decrease in fuel integrity. Thermal stress in the pellets, while causing some fracture of the bulk material during temperature cycling, does not result in pulverization or gross void formation in the fuel matrix. As shown by operating experience and extensive experimental work in the industry, the thermal design parameters conservatively account for any changes in the thermal performance of the fuel element due to pellet fracture.
The consequences of a breach of cladding are greatly reduced by the ability of uranium dioxide to retain fission products including those which are gaseous or highly volatile. This retentiveness decreases with increasing temperature or fuel burnup, but remains a significant factor even at full power operating temperature in the maximum burnup element.
A survey of fuel elements behavior in high burnup uranium dioxide (6) indicates that for an initial uranium dioxide void volume, which is a function of the fuel density, it is possible to conservatively define the fuel swelling as a function of burnup.
3.2.3-26 Revised 03/11/2016 C28 The evaluation of fuel densification and the treatment of fuel swelling are described by an empirical model developed with data from numerous operating Westinghouse reactors as described in Reference 10 and used in Reference 17.
The integrity of the fuel rod cladding is directly related to cladding stresses and strains under normal and overpower conditions. The combined maximum stress intensity meets the criteria based on the ASME code in Addendum 1-A of Reference 26 or the cladding stress is limited to the yield strength of the material. The steady-state tensile strain is limited to 1.0% and during power increases the cladding tensile strain is limited to 1.0% during the transient.
The cladding stresses at constant local fuel rod power are low. Compressive stresses are created by the pressure differential between the coolant pressure and the rod internal pressure. Tensile stresses could be created once the cladding has come into contact with the pellet which results from fuel swelling and cladding creepdown (thermal and irradiation induced creep as determined by the models of References 10 and 17 and acceptable for use with TCD, per Reference 25). These stresses would be induced by the fuel pellet swelling during irradiation. Fuel swelling can result in small cladding strains (<1%) for expected discharge burnups, but the associated cladding stresses are low because of cladding creep. Furthermore, the 1% strain criterion is extremely conservative for fuel swelling driven cladding strain because the strain rate associated with solid fission product swelling is very slow.
Pellet thermal expansion caused by power increases is considered the only mechanism by which significant stresses and strains can be imposed on the cladding. There are two methods that can be used to evaluate cladding stresses and strains: 1. The maximum cladding stress intensities excluding Pellet Clad Interaction (PCI) induced stress will be evaluated using ASME pressure vessel guidelines. Cladding corrosion is accounted for as a loss of load carrying material. Stresses are combined to calculate a maximum stress intensity which is then compared to criteria on the ASME code (Addendum 1-A of Reference 26). Slow transient power increases can result in large cladding strains without exceeding the yield strength because of cladding creep and stress relaxation. Therefore, the additional limitation of 1% cladding tensile strain during a transient is considered. Together, the ASME based stress criterion, 1% transient strain criterion, 1% steady state criterion, and the fuel melt limit are sufficient to protect the cladding from PCI.
3.2.3-27 Revised 03/11/2016 C28C28
- 2. Radial, tangential, and axial stress components due to pressure differential and fuel cladding contact pressure which can occur during power increases the combined stress components into an effective stress using the maximum-distortion energy theory. The Von Mises criterion is used to evaluate if the yield strength has been exceeded. The effective stress is increased by an allowance for local non-uniformity effects before it is compared to the yield strength. The yield strength of the cladding is a function of the cladding temperature. The yield strength is that appropriate for irradiated cladding since the irradiated properties are attained at low exposure whereas the fuel-cladding contact which can lead to minimum margin to the yield strength occurs at much higher exposure. Slow transient power increases can result in large cladding strains without exceeding the yield strength because of cladding creep and stress relaxation. Therefore, the additional limitation od 1% cladding tensile strain during a transient is considered.
The internal gas pressure of the rods in the reactor is limited to a value below that which would cause the pellet-cladding diametrical gap to increase due to outward cladding creep during steady-state operation, and which would cause extensive DNB propagation to occur. The safety evaluation of the fuel rod internal pressure design basis is presented in Reference 11.
3.2.3-27a Revised 03/11/2016 C28 The fission gas release and the associated buildup of internal gas pressure
inside the fuel rod is determined by the models of References 10 and 17. The
increase of internal pressure in the fuel rod is included in the
determination of the maximum cladding stresses and strains.
Cladding collapse is precluded during fuel rod design lifetime by appropriate
control of pellet characteristics which control in-reactor fuel densification
behavior, as described in References 7 and 12.
The use of chamfered fuel pellets in Optimized Fuel Assemblies results in a hot spot average fuel temperature increase of less than 20 F compared to unchamfered pellets. Evaluation results show that all core design criteria and safety limits (including LOCA and non-LOCA transients) are satisfied when
using chamfered pellets.
Evaluation of Burnable Poison Rods
The burnable poison rods are positioned in the core inside fuel assembly
guide thimbles and held down in place by attachment to a plate assembly
compressed beneath the upper core plate and hence cannot be the source of any
reactivity transient. Due to the low heat generation rate, and the
conservative design of the poison rods, there is no possibility for release
of the poison as a result of helium pressure or clad heating during accident
transients including loss of coolant.
Effects of Vibration and Thermal Cycling on Fuel Assemblies
Analyses of the effect of cyclic deflection of the fuel rods, grid spring
fingers, RCCA's, and burnable poison rods due to hydraulically induced
vibrations and thermal cycling show that the design of the components is good
for an infinite number of cycles.
In the case of the fuel grid spring support, the amplitude of a hydraulically induced motion of the fuel rod is extremely small ( .001) and the stress associated with the motion is significantly small (< 100 psi). Likewise, the reactions at the grid spring due to the motion is much less than the preload
spring force and contact is maintained between the fuel clad and the grid
spring and dimples. Fatigue of the clad and fretting between the clad and
the grid support are not anticipated.
The effect of thermal cycling on the grid-clad support is merely a slight
relative movement between the grid contact surfaces and the clad, which is
gradual in nature during heat-up and cool-down. Since the number of cycles of the occurrence is small over the life of a fuel assembly ( 6 years), negligible wear of the mating parts is expected.
3.2.3-28 Revised 04/17/2013 In-core operation of assemblies in the Yankee Rowe and Saxton reactors using similar clad support have verified the calculated conclusions. Additional
test results under simulated reactor environment in the Westinghouse Reactor
Evaluation Channel also support these conclusions.
The dynamic deflection of the full length control rods and the burnable
poison rods is limited by their fit with the inside diameter of either the
upper portion of the guide thimble or the dashpot. With this limitation, the
occurrence of truly cyclic motion is questionable. However, an assumed
cyclic deflection through the available clearance gap results in an
insignificantly low stress in either clad tubing or in the flexure joint at
the spider or retainer plate. The above consideration assumes the rods are
supported as cantilevers from the spider, or the retainer plate in the case
of the burnable poison rods.
A calculation, assuming the rods are supported by the surface of the dashpots and at the upper end by the spider or retainer, results in a similar
conclusion.
Control Rod Drive Mechanism
Full Length Rods
Design Description
The Control Rod Drive Mechanisms (CRDM) are used for withdrawal and insertion
of the RCCA's into the reactor core and to provide sufficient holding power
for stationary support.
Fast total insertion (reactor trip) is obtained by simply removing the electrical power allowing the rods to fall by gravity.
The complete drive mechanism, shown in Figure 3.2.3-12, consists of the
internal (latch) assembly, the pressure vessel, the operating coil stack, the
drive shaft assembly, and the Rod Position Indicator (RPI) coil stack.
Each assembly is an independent unit which can be dismantled or assembled
separately. Each drive is threaded into an adaptor on top of the reactor
pressure vessel and is connected to the control rod (directly below) by means
of a grooved drive shaft. The upper section of the drive shaft is suspended
from the working components of the drive mechanism. The drive shaft and
control rod remain connected during reactor operation, including tripping of
the rods.
3.2.3-29 Revised 04/17/2013 The replacement RVCH for Unit 3 and Unit 4 were manufactured without spare
CRDM nozzles at locations D-4, D-12, G-7, G-9, J-7, J-9, M-4, and M-12 (Reference 19 and Reference 24). The replacement RVCHs were manufactured
with spare CRDM nozzles at locations B-8 and H-14. The spare CRDM nozzles at
these two locations are closed with a CRDM plug (See Figure 3.2.3-18) that is
threaded onto the CRDM nozzle adapter and seal welded with a canopy seal
weld. The CRDM plug is an ASME Section III, Class 1 component. The former
RVCH spare CRDM nozzles had a Canopy Seal Clamp Assembly (CSCA) installed
around the canopy seal weld. The clamp was not part of the RCS pressure
boundary material. It was a leak deterrent device installed to prevent
leakage in the event the canopy seal weld failed and to provide a compressive
force on the canopy seal weld to help reinforce the canopy seal weld. The
CSCA devices were not reinstalled on the spare CRDM nozzle canopy seal welds
on the replacement RVCHs.
Dummy Cans are hung from the CRDMs surrounding all of the CRDM positions that
formerly were occupied by spare CRDM nozzle penetrations (D-4, D-12, G-7, G-
9, J-7, J-9, M-4, and M-12) or part length CRDM assemblies at positions B-8, F-6, F-10 H-14, K-6 & K-10. The purpose of the Dummy Can is to insure proper
airflow around the CRDMs for CRDM cooling.
Reactor coolant fills the pressure containing parts of the drive mechanism.
All working components and the shaft are immersed in the reactor coolant.
Three magnetic coils, which form a removable electrical unit and surround the
rod drive pressure housing induce magnetic flux through the housing wall to
operate the working components. They move two sets of latches which lift or
lower the grooved drive shaft.
The three magnets are turned on and off in a fixed sequence by solid-state switches for the full length rod assemblies.
The sequencing of the magnets produces step motion over the 144 inches of normal control rod travel.
The mechanism develops a lifting force approximately two times the static lifting load. Therefore, extra lift capacity is available for overcoming
mechanical friction between the moving and the stationary parts. Gravity
provides the drive force for rod insertion and the weight of the whole rod
assembly is available to overcome any resistance.
The mechanisms are designed to operate in water at 650 F and 2485 psig. The temperature at the mechanism head adaptor will be much less than 650 F because it is located in a region where there is limited flow of water from the reactor core, while the pressure is the same as in the reactor pressure
vessel.
3.2.3-30 Revised 04/17/2013 A multi-conductor cable connects the mechanism operating coils to the 125 volt DC power supply. The power supply is described in Section 7.3.2.
Latch Assembly
The latch assembly contains the working components which withdraw and insert
the drive shaft and attached control rod. It is located within the pressure
housing and consists of the pole pieces for three electromagnets. They
actuate two sets of latches which engage the grooved section of the drive
shaft.
The upper set of latches moves up or down to raise or lower the drive rod by
5/8 inch. The lower set of latches has 1/16 inch axial movement to shift the
weight of the control rod from the upper to the lower latches. The housings
are designed in accordance with the requirements for Class A vessels of the
ASME Nuclear Vessel Code.
Rod Drive Mechanism Housing
The pressure vessel consists of the pressure housing and rod travel housing.
The pressure housing is the lower portion of the vessel and contains the
latch assembly. The rod travel housing is the upper portion of the vessel.
It provides space for the drive shaft during its upward movement as the
control rod is withdrawn from the core.
Operating Coil Stack
The operating coil stack is an independent unit which is installed on the
drive mechanism by sliding it over the outside of the pressure housing. It
rests on a pressure housing flange without any mechanical attachment and can
be removed or installed while the reactor is pressurized.
The operator coils (A, B and C) are made of wound copper wire which is
insulated with a double layer of filament type glass yarn.
The design operating temperature of the coils is 232 C (450 F). Coil temperature can be determined by resistance measurement. Forced air cooling
along the outside of the coil stack maintains a coil temperature below 200 C (392 F).
3.2.3-31 Revised 04/17/2013 Drive Shaft Assembly
The main function of the drive shaft is to connect the control rod to the
mechanism latches. Grooves for engagement and lifting by the latches are
located throughout the 144 in. of control rod travel. The grooves are spaced 5/8 inch apart to coincide with the mechanism step length and have 45 angle sides.
The drive shaft is attached to the control rod by the coupling. The coupling
has two flexible arms which engage the grooves in the spider assembly.
A 1/4 inch diameter disconnect rod runs down the inside of the drive shaft.
It utilizes a locking button at its lower end to lock the coupling and
control rod. At its upper end, there is a disconnect assembly for remote
disconnection of the drive shaft assembly from the control rod. During
operation, the drive shaft assembly remains connected to the control rod at
all times.
Position Indicator Coil Stack
The position indicator coil stack slides over the rod travel housing section
of the pressure vessel. It detects drive rod position by means of
cylindrically wound differential transformers which span the normal length of
the rod travel (144 inches).
Drive Mechanism Materials
All parts exposed to reactor coolant, such as the pressure vessel, latch
assembly and drive rod, are made of metals which resist the corrosive action
of the water.
Three types of metals are used exclusively: stainless steels, Inconel
UNS-N07750, and cobalt based alloys. Wherever magnetic flux is carried by
parts exposed to the reactor coolant, stainless steel is used. Cobalt based
alloys are used for the pins and latch tips.
Inconel UNS-N07750 is used for the springs of both latch assemblies and 304
stainless steel is used for all pressure containment. Hard chrome plating
provides wear surfaces on the sliding parts and prevents galling between
mating parts (such as threads) during assembly.
3.2.3-32 Revised 04/17/2013 Outside of the pressure vessel, where the metals are exposed only to the reactor containment environment and cannot contaminate the reactor coolant, carbon and stainless steels are used. Carbon steel, because of its high
permeability, is used for flux return paths around the operating coils. It
is zinc-plated (0.001 to 0.005 inch thick) to prevent corrosion.
Principles of Operation
The drive mechanisms shown schematically in Figure 3.2.3-13 withdraw and
insert their respective control rods as electrical pulses are received by the
operator coils.
ON and OFF sequence, repeated by Silicon Controlled Rectifiers (SCR) in the
power programmer causes either withdrawal or insertion of the control rod.
Position of the control rod is indicated by the differential transformer
action of the position indicator coil stack surrounding the rod travel
housing. The differential transformer output changes as the top of the
ferromagnetic drive shaft assembly moves up the rod travel housing.
Generally during operation, the stationary gripper coil of the drive
mechanisms hold the control rods withdrawn from the core in a static position
until the movable gripper coil is energized.
Control Rod Withdrawal
The control rod is withdrawn by repeating the following sequence:
(1) Movable Gripper Coil - ON
The movable gripper armature raises and swings the movable gripper latches
into the drive shaft groove.
(2) Stationary Gripper Coil - OFF
Gravity causes the stationary gripper latches and armature to move downward
until the load of the drive shaft is transferred to the movable gripper
latches. Simultaneously, the stationary gripper latches swing out of the
shaft groove.
(3) Lift Coil - ON
The 5/8 inch gap between the lift armature and the lift magnet pole closes
and the drive rod rises one step length.
3.2.3-33 Revised 04/17/2013 (4) Stationary Gripper Coil - ON
The stationary gripper armature rises and closes the gap below the stationary
gripper magnetic pole, swings the stationary gripper latches into a drive
shaft groove. The latches contact the shaft and lift it 1/16 inch. The load
is so transferred from the movable to the stationary gripper latches.
(5) Movable Gripper Coil - OFF
The movable gripper armature separates from the lift armature under the force
of three springs and gravity. Three links, pinned to the movable gripper
armature, swing the three movable gripper latches out of the groove.
(6) Lift Coil - OFF
The gap between the lift armature and the lift magnet pole opens. The
movable gripper latches drop 5/8 inch to a position adjacent to the next
groove.
Control Rod Insertion
The sequence for control rod insertion is similar to that for control rod
withdrawal:
(1) Lift Coil - ON
The movable gripper latches are raised to a position adjacent to a shaft
groove.
(2) Movable Gripper Coil - ON
The movable gripper armature rises and swings the movable gripper latches
into a groove.
(3) Stationary Gripper Coil - OFF
The stationary gripper armature moves downward and swings the stationary
gripper latches out of the groove.
(4) Lift Coil - OFF
Gravity separates the lift armature from the lift magnet pole and the control
rod drops down 5/8 inch.
3.2.3-34 Revised 04/17/2013 (5) Stationary Gripper Coil - ON
(6) Movable Gripper Coil - OFF
The sequences described above are termed as one step or one cycle and the
control rod moves 5/8 inch for each cycle. Each sequence can be repeated at
a design rate of up to 72 steps per minute and the control rods can therefore
be withdrawn or inserted at a design rate of up to 45 inches per minute.
Control Rod Tripping
If power to the movable gripper coil is cut off, as for tripping, the
combined weight of the drive shaft and the rod cluster control assembly is
sufficient to move the latches out of the shaft groove. The control rod
falls by gravity into the core. The tripping occurs as the magnetic field, holding the movable gripper armature against the lift magnet, collapses and
the movable gripper armature is forced down by the weight acting upon the
latches.
Reactor Vessel Level Measuring System Probes
The replacement RVCHs, installed by Reference 19 and Reference 24, incorporated a RVLMS nozzle adapter that replaced the two modified part
length CRDMs that served, on the original RVCHs, to accommodate the
installation of the RVLMS thermocouple probes. The replacement RVLMS nozzle
adapters eliminated the modified internal CRDM parts that were required to
insure the thermocouple probe would align with the thermocouple shrouds. The
thermocouple shrouds are a part of the reactor vessel internals that were
installed in place of the control rod guide tubes at the time when the two
part length CRDMs were modified to facilitate installation of the RVLMS. The
shrouds are the receptacle for the probe assembly. The shrouds are designed
in accordance with the guide tube design criteria.
Fuel Assembly and RCCA Mechanical Evaluation
To confirm the mechanical adequacy of the fuel assembly and RCCA's,
functional test programs have been conducted on full scale San Onofre mock-up
version of the fuel assembly and control rods. Additional tests were run on
two full scale prototype assemblies for a twelve-foot active core. One of
the twelve-foot assemblies incorporated stainless steel guide tubes and other
3.2.3-35 Revised 04/17/2013 Reactor Evaluation Center (WREC) Tests
The prototype assemblies were tested under simulated reactor operating conditions (1875 psig, 575 F, and 17.8 fps flow velocity) in the Westinghouse Reactor Evaluation Center (WREC) for a total of more than 6400 hours0.0741 days <br />1.778 hours <br />0.0106 weeks <br />0.00244 months <br />.
Each prototype assembly was subjected to trip cycling equivalent to one or
more plant lifetimes. The test history for each prototype is summarized
below:
PROTOTYPE TEST NUMBER TOTAL TOTAL TOTAL
TIME, OF OF DRIVEN TRIP
HOURS TRIPS TRAVEL, TRAVEL, TRAVEL, FT. FT. FT.
San Onofre, 10 ft.
assy. stainless
steel guide thimbles 4132 1461 38,927 27,217 11,710
12-ft. assembly
stainless steel guide thimbles 1000 600 45,000 38,500 6,500
12-ft. assembly
Zircaloy-4 guide thimbles 1277 600 124,200 117,700 6,500
_____________________________________________________________________________
Each of the three prototype fuel assemblies remained in excellent mechanical condition. No measurable signs of wear on the fuel tubes or control rod
guide tubes were found.
The control rod was also found to be in excellent condition, having maximum wear measured on absorber cladding of approximately 0.001 in.
Loading and Handling Tests Tests simulating the loading of the prototype fuel assembly into a core
location were also successfully conducted to determine that proper provisions
had been made for guidance of the fuel assembly during refueling operation.
Axial and Lateral Bending Tests
In addition, axial and lateral bending tests were performed in order to
simulate mechanical loading of the assembly during refueling operation.
Although the maximum column load expected to be experienced in service is
approximately 1000 lb. the fuel assembly can successfully be loaded to 2200
lb. axially with no damage resulting. This information is also used in the
design of fuel handling equipment to establish the limits for inadvertent
axial loads during refueling.
3.2.3-36 Revised 04/17/2013 CRDM Housing Mechanical Failure Evaluation
An evaluation of the possibility of damage to adjacent control rod drive
mechanism housings in the event of a circumferential or longitudinal failure
of a rod housing located on the vessel head is presented.
A control rod drive mechanism schematic is shown in Figure 3.2.3-12 and
3.2.3-13. The operating coil stack assembly of this mechanism has a 10.8 inch
by 10.8 inch cross section and a 39.55 inch length. The position indicator
coil stack assembly (not shown in this figure) is located above the operating
coil stack assembly. It surrounds the rod travel housing over nearly its
entire length. The rod travel housing outside diameter is 3.8 inches and
the position indicator coil stack assembly consists of a 1/8" thick
stainless steel tube surrounded by a continuous stack of copper wire coils.
This assembly is held together by two end plates (the top end plate is
square), an outer sleeve, and four axial tie rods.
Effect of Rod Travel Housing Longitudinal Failures
Should a longitudinal failure of the rod travel housing occur, the region of
the stainless steel tube opposite the break would be stressed by the reactor
coolant pressure of 2250 psia. The most probable leakage path would be
provided by the radial deformation of the position indicator coil assembly, resulting in the growth of axial flow passages between the rod travel housing
and the stainless steel tube. A radial free water jet is not expected to
occur because of the small clearance between the stainless steel tube and the
rod travel housing, and the considerable resistance of the combination of the
stainless steel tube and the position indicator coils to internal pressure.
Calculations based on the mechanical properties of stainless steel and copper
at reactor operating temperature show that an internal pressure of at least
4000 psia would be necessary for the combination of the stainless steel tube
and the coils to rupture.
Therefore, the combination of stainless steel tube and copper coils stack is
more than adequate to prevent formation of a radial jet following a control
rod housing split which assures the integrity of the adjacent rod housings.
Effect of Rod Travel Housing Circumferential Failures
If circumferential failure of a rod travel housing should occur, the
broken-off section of the housing would be ejected vertically because the
driving force is vertical and the position indicator coil stack assembly and
the drive shaft would tend to guide the broken-off piece upwards during its
travel.
3.2.3-37 Revised 04/17/2013 Travel is limited to less than four feet by the missile shield, thereby limiting the projectile acceleration. When the projectile reaches the
missile shield, it would partially deform the shield and dissipate its
kinetic energy. The water jet from the break would push the broken-off piece
against the missile shield.
If the broken-off piece were short enough to clear the break when fully
ejected, it could rebound after impact with the missile shield. The top end
plates of the position indicator coil stack assemblies would prevent the
broken piece from directly hitting the rod travel housing of a second drive
mechanism. Even if a direct hit by the rebounding piece were to occur, the
low kinetic energy of the rebounding projectile would not be expected to
cause significant damage.
Summary The considerations given above lead to the conclusion that failure of a
control rod housing due to either longitudinal or circumferential cracking
would not cause damage to adjacent housings that would increase the severity
of the initial accident.
3.2.3-38 Revised 04/17/2013 REFERENCES
- 1. Letter from Uhrig, R. E., FP&L to Varga, S. A., NRC,
Subject:
Pressurized Thermal Shock, Letter No L-83-180, March 25, 1983.
- 2. Petrarca, D., et al, "Reload Transition Safety Report For Turkey Point Units 3 & 4," June, 1983.
- 3. Letter from Thomas, C. 0., NRC, to Rahe, E. P., Westinghouse,
Subject:
Acceptance for Referencing of Licensing Topical Report WCAP-10021 (P),
Revision 1, and WCAP-10377 (NP), "Westinghouse Wet Annular Burnable Absorber Evaluation Report," August 9, 1983.
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- 5. WCAP-8720, Addendum-2, "Westinghouse Revised Pad Code Thermal Safety Model," October 27, 1982 (Proprietary).
- 6. Daniel, R. C., et al. "Effects of High Burnup on Zircaloy-Clad Bulk DO 2 Plate Fuel Element Samples," WAPD-263, (September, 1965).
- 7. Oelrich, R. L. and Kersting, P. J., "Assessment of Clad Flattening and Densification Power Spike Factor Elimination in Westinghouse Nuclear Fuel," WCAP-13589, March 1995.
- 8. XN-NF-85-12(P) Ford, K. L., et al "Mechanical Design Report For Turkey Point Units 3 & 4 Hafnium Vessel Flux Depression (HVFD) Cluster Assemblies," Proprietary.
- 9. WCAP-12346, "Turkey Point Units 3 and 4 - 15x15 Debris Resistant Fuel Assembly Design Report," July 1989.
- 10. Weiner, R. A. et. al., "Improved Fuel Performance Models for Westinghouse Fuel Rod Design and Safety Evaluations," WCAP-10851-P-A (Proprietary) and WCAP-11873-A (Non-Proprietary), August 1988.
- 11. Risher, D. H. (Editor), "Safety Analysis for the Revised Fuel Rod Internal Pressure Design Basis," WCAP-8963-P-A (Proprietary) and WCAP-8964-A (Non-Proprietary), August 1978.
- 12. Letter, G. M. Holohan (NRC) to N.J. Liparulo (W),
Subject:
Acceptance for Referencing of Topical Report WCAP-13589, "Assessment of Clad Flattening and Densification Power Spike Factor Elimination in Westinghouse Nuclear Fuel," (TAC NO. M85707), January 30, 1995.
3.2.3-39 Revised 03/11/2016 C28 REFERENCES (Cont'd)
- 13. WCAP-11027, "The Nuclear Design and Core Management of the Turkey Point Unit 4 Power Plant Cycle 11," March 1986.
- 14. WCAP-10445-NP-A. "Reference Core Report Vantage 5 Fuel Assembly,"
S.L.Davidson and W.R. Kramer, September 1995.
- 15. Letter from E. Rahe, Jr. (Westinghouse) to H. Berkow (NRC), NS-NRC 3090,
Subject:
Request for Addendum 1 to WCAP-10445-NP-A, December 1985.
- 16. Davidson, S.L., and Nufer, D.L., eds.,"Vantage+ Fuel Assembly Reference Core Report," WCAP-12610-P-A, April 1995.
- 17. Foster, J. P., Sidener, S., "Westinghouse Improved Performance Analysis and Design Model (PAD 4.0)," WCAP-15063-P-A, Revision 1, with Errata, July 2000.
- 18. PC/M 78-051, "Removal of Part Length Rods and Installation of New Anti-Rotation Devices", 9/15/1978.
19 PC/M 03-057, Rev.1, "Reactor Vessel Closure Head Replacement".
- 20. Framatome Doc.32-5017014, Rev.3 dated 3/9/2004, "Turkey Point 3 and 4 Part-Length CRDM Nozzle Repair Hydraulic Evaluation".
- 21. Framatome Doc. 32-5017272, Rev.2, dated 9/29/2003, "Turkey Point Flow Restrictor Structural Evaluation".
- 22. Framatome Test Report 51-5017418, Rev.2, dated 4/23/2004, "Qualification Plan for the flow Restrictor and Installation Tool".
- 23. PC/M 78-052, "Removal of Part Length Rods & Installation of New Anti-Rotation Devices-Unit 4".
- 24. PC/M 03-058, "Reactor Vessel Closure Head Replacement- Unit 4".
- 25. NRC Safety Evaluation Report, "Turkey Point Units 3 and 4 - Issuance of Amendments Regarding Extended Power Uprate (TAC nos. ME 4907 and ME 4908)," June 15, 2012, (ADAMS Accession Number ML 11293A365).
- 26. Davidson, S.L. (Ed), et al, "Extended Burnup Evaluation of Westinghouse Fuel," WCAP-10125-P-A, December 1985, and Bahr, K.E., "Extended Burnup Evaluation of Westinghouse Fuel, Revision to Design Criteria," WCAP-10125-P-A, Addendum 1-A, Revision 1-A, May 2005.
- 27. Shah, H. H., "Optimized ZIRLOŽ, "WCAP-12610-P-A & CENPD-404-P-A, Addendum 1-A, July 2006.
3.2.3-40 Revised 03/11/2016 C28C28C28 TABLE 3.2.3-1 Sheet 1 of 2
CORE MECHANICAL DESIGN PARAMETERS (1)
Active Portion of the Core Equivalent Diameter, in. 119.7 Active Fuel Height, in - Unit 3 144.00, 144.00, 143.474 Unit 4 144.00, 143.40, 142.80 Length-to-Diameter Ration 1.2 Total Cross Section Area, Ft.
2 78.1 Fuel Assemblies Number 157 Rod Array 15 x 15 Rods per Assembly 204 (2)
Rods Pitch, in. 0.563 Overall Dimensions, in. 8.426 x 8.426 Fuel Weight (as UO 2), pounds 176,000 Total Weight, pounds 225,000 Number of Grids per Assembly 7 Structural Grids 3 Intermediate Flow Mixer(IFM) Grids 1 Protector Grid (P-Grid) Guide Thimble I.D. (Above Dashpot), in. 0.499 (at Dashpot), in. 0.455 Fuel Rods Number 32,028 Outside Diameter, in. 0.422 Diametric Gap, mils 7.5, 7.5, 8.5 Clad Thickness, in. 0.0243 Clad Material Zircaloy-4, ZIRLO or Optimized ZIRLO Overall Length, in. Unit 3 152.235 to <152.880 Unit 4 152.235 to <152.880 Fuel Pellets Material UO 2 sintered Density (% of Theoretical) - First Cycle (3) Region 1 94 (10.3 g/cc)
Region 2 93 (10.19 g/cc) Region 3 92 (10.08 g/cc)(Unit 4-93) Fuel Enrichments w/o - First Cycle (3) Region 1 1.85 Region 2 2.55 Region 3 3.10
Diameter, in. - Unit 3 (Regions 1, 2, 3) 0.3659, 0.3659, 0.3649 Unit 4 (All Regions) 0.3659 Length, in. 0.439
NOTES:
(1) All Dimensions are for cold conditions. (2) Twenty-one rods are omitted: twenty to provide passage for control rods and one to contain in-core instrumentation.
(3) Values for current cycles are given in Appendixes 14A and 14B.
Revised 03/11/2016 C28 TABLE 3.2.3-1 Sheet 2 of 2
Rod Cluster Control Assemblies
Neutron Absorber 5% Cd, 15% In, 80% Ag Cladding Material Type 316L or 304 SS - l Cold Worked
Clad Thickness, in. 0.019 Number of Clusters 45 Full Length 45
Number of Control Rods per Cluster 20
Weight in 60 o F Water Full Length, pounds 147
Length of Rod Control, in. 158.454 (overall) 150.574 (insertion length)
Length of Absorber Section, in. 142.00
Core Structure
Core Barrel, in.
I.D. 133.875 O.D. 137.875
Thermal Shield, in.
I.D. 142.625 O.D. 148.0
Burnable Poison Rods (4) Number 816 Material Borosilicate Glass Outside Diameter, in. 0.4395 Inner Tube, O.D. in. 0.2365 Clad Material S.S.
Inner Tube Material S.S.
Boron Loading (natural) gm/cm 0.0429
of glass rod
Neutron Source Assemblies (5) Primary Source (typical) Pu-Be Secondary Source (typical) Sb-Be
NOTES :
(4) Values for current cycles are given in Appendices 14A and 14B.
(5) Neutron sources are not installed in current cycles.
Rev. 17 FLORIDA POWER & LIGHT COMPANY TURKEY POINT PLANT UNITS 3 & 4 REACTOR CORE CROSS SECTION FIGURE 3.2.3-1
FLORIDA POWER & LIGHT COMPANY TURKEY POINT PLANT UNITS 3 & 4 REACTOR VESSEL INTERNALS FIGURE 3.2.3-2
REV. 15 (4/98)
FLORIDA POWER & LIGHT COMPANY TURKEY POINT PLANT UNITS 3 & 4 THREE REGION CORE LOADING FIGURE 3.2.3-3
FLORIDA POWER & LIGHT COMPANY TURKEY POINT PLANT UNITS 3 & 4 TYPICAL ROD CLUSTER CONTROL ASSEMBLY FIGURE 3.2.3-4
FLORIDA POWER & LIGHT COMPANY TURKEY POINT PLANT UNITS 3 & 4 LOWER CORE SUPPORT ASSEMBLY FIGURE 3.2.3-5
FLORIDA POWER & LIGHT COMPANY TURKEY POINT PLANT UNITS 3 & 4 UPPER CORE SUPPORT ASSEMBLY FIGURE 3.2.3-6
FLORIDA POWER & LIGHT COMPANY TURKEY POINT PLANT UNITS 3 & 4 GUIDE TUBE ASSEMBLY FIGURE 3.2.3-7
FLORIDA POWER & LIGHT COMPANY TURKEY POINT PLANT UNITS 3 & 4 FUEL ASSEMBLY AND CONTROL ASSEMBLY CLUSTER CROSS SECTION FIGURE 3.2.3-8
FLORIDA POWER & LIGHT COMPANY TURKEY POINT PLANT UNIT 3 & 4 FUEL ASSEMBLY OUTLINE 1 of 2 FIGURE 3.2.3-9
FLORIDA POWER & LIGHT COMPANY TURKEY POINT PLANT UNIT 3 & 4 FUEL ASSEMBLY OUTLINE 1 of 2 FIGURE 3.2.3-9
Rev. 2-7/84 FLORIDA POWER & LIGHT COMPANY TURKEY POINT PLANT UNITS 3 & 4 OFA - LOPAR FUEL ASSEMBLY OUTLINES FIGURE 3.2.3-9A
Rev. 4-7/86 FLORIDA POWER & LIGHT COMPANY TURKEY POINT PLANT UNITS 3 & 4 BOTTOM NOZZLE TO THIMBLE TUBE CONNECTION FIGURE 3.2.3-9B
Rev. 16 (10/99)
FLORIDA POWER & LIGHT COMPANY TURKEY POINT PLANT UNITS 3 & 4 15 X 15 OFA/DRFA FUEL ASSEMBLY COMPARISON FIGURE 3.2.3-9C
FLORIDA POWER & LIGHT COMPANY TURKEY POINT PLANT UNITS 3 & 4 SPRING CLIP GRID ASSEMBLY FIGURE 3.2.3-10
Rev 8 7/90 FLORIDA POWER & LIGHT COMPANY TURKEY POINT PLANT UNITS 3 & 4 COMPARISON OF OFA AND LOPAR PLUGGING DEVICE FIGURE 3.2.3-10A
FLORIDA POWER & LIGHT COMPANY TURKEY POINT PLANT UNITS 3 & 4 DETAIL OF BURNABLE POISON ROD FIGURE 3.2.3-11
Rev 16 (10/99)
FLORIDA POWER & LIGHT COMPANY TURKEY POINT PLANT UNITS 3 & 4 WET ANNULAR BURNABLE ABSORBER ROD FIGURE 3.2.3-11A
Revised 04/29/2005 FLORIDA POWER & LIGHT COMPANY TURKEY POINT PLANT UNITS 3 & 4 CONTROL ROD DRIVE MECHANISM ASSEMBLY FIGURE 3.2.3-12
FLORIDA POWER & LIGHT COMPANY TURKEY POINT PLANT UNITS 3 & 4 CONTROL ROD DRIVE MECHANISM SCHEMATIC FIGURE 3.2.3-13
Rev. 4- 7/86 FLORIDA POWER & LIGHT COMPANY TURKEY POINT PLANT UNITS 3 & 4 REDUCED LENGTH HVFD ABSORBER ROD FIGURE 3.2.3-14
REV. 6 (7/88)
FLORIDA POWER & LIGHT COMPANY TURKEY POINT PLANT UNITS 3 & 4 HEAD ADAPTER PLUG DESIGN FIGURE 3.2.3-15
NOTE: This Figure is for historical purposes only. Reference 19 replaced the Unit 3 RVCH.
This design applied to only the Unit 3 RVCH that has been replaced.
Revised 04/29/2005 FLORIDA POWER & LIGHT COMPANY TURKEY POINT PLANT UNITS 3 & 4 CONTROL ROD DRIVE MECHANISM ASSEMBLY FIGURE 3.2.3-16
FINAL SAFETY ANALYSIS REPORT FIGURE 3.2.3-17
REFER TO ENGINEERING DRAWINGS 5613-M-460-1, SHEET 1 5614-M-460-1, SHEET 1 FLORIDA POWER & LIGHT COMPANY TURKEY POINT PLANT UNITS 3 & 4 FLOW RESTRICTOR ASSEMBLY FIGURE 3.2.3-17 05/07/2007C22
FINAL SAFETY ANALYSIS REPORT FIGURE 3.2.3-18
RFER TO ENGIN EERING DRAWINGS 5613-M-460-2, SHEET 6 5614-M-460-2, SHEET 6 FLORIDA POWER & LIGHT COMPANY TURKEY POINT PLANT UNITS 3 & 4 CRDM PLUG FIGURE 3.2.3-18 05/07/2007C22