ML17212A044
| ML17212A044 | |
| Person / Time | |
|---|---|
| Site: | Millstone  |
| Issue date: | 06/29/2017 |
| From: | Dominion Nuclear Connecticut |
| To: | Office of Nuclear Reactor Regulation |
| Shared Package | |
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| References | |
| 17-208 | |
| Download: ML17212A044 (134) | |
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Millstone Power Station Unit 2 Safety Analysis Report Chapter 3 MPS2 UFSAR 3-i Rev. 35CHAPTER 3-REACTOR Table of ContentsSection Title Page3.1
SUMMARY
DESCRIPTION..............................................................................3.1-13.1.1References...................................................................................................3.1-33.2DESIGN BASES.................................................................................................3.2-13.2.1Mechanical Design Bases...........................................................................
3.2-13.2.1.1Fuel Assembly Design Bases......................................................................3.2-13.2.1.2AREVA Fuel Rod Cladding Design Bases.................................................3.2-2 3.2.1.3Control Element Assembly Design Bases..................................................3.2-23.2.1.4Reactor Internals Design Bases..................................................................
3.2-33.2.1.5CEDM/RVLMS (HJTC) Pressure Housing Design Bases.........................
3.2-53.2.2Nuclear Design Bases.................................................................................3.2-6 3.2.3Thermal and Hydraulic Design Basis.........................................................
3.2-83.2.4References...................................................................................................3.2-8
3.3 MECHANICAL
DESIGN...................................................................................
3.3-13.3.1Core Mechanical Design.............................................................................3.3-13.3.1.1AREVA Fuel Rod.......................................................................................3.3-13.3.1.1.1Fuel Rod Mechanical Criteria.....................................................................3.3-1 3.3.1.1.2Fuel Rod Design Analyses..........................................................................3.3-33.3.1.2(Deleted).....................................................................................................3.3-63.3.1.3AREVA Fuel Assembly..............................................................................3.3-6 3.3.1.3.1Design Summary.........................................................................................
3.3-63.3.1.3.2Fuel Assembly Mechanical Criteria...........................................................
3.3-93.3.1.4Fuel Assembly Holddown Device............................................................3.3-12 3.3.1.5Control Element Assembly.......................................................................3.3-123.3.1.6Neutron Source Design.............................................................................
3.3-133.3.1.7In-Core Instruments..................................................................................3.3-133.3.1.8Heated Junction Thermocouples...............................................................
3.3-143.3.2Reactor Internal Structures.......................................................................
3.3-143.3.2.1Core Support Assembly............................................................................
3.3-153.3.2.2Core Support Barrel..................................................................................
3.3-153.3.2.3Core Support Plate and Support Columns................................................
3.3-163.3.2.4Core Shroud..............................................................................................3.3-163.3.2.5Flow Skirt.................................................................................................3.3-163.3.2.6Upper Guide Structure Assembly.............................................................3.3-16 3.3.3Control Element Drive Mechanism..........................................................3.3-173.3.3.1Design.......................................................................................................
3.3-17 MPS2 UFSAR Table of Contents (Continued)
Section Title Page 3-ii Rev. 353.3.3.2Control Element Drive Mechanism Pressure Housing.............................3.3-183.3.3.2.1Heated Junction Thermocouple Pressure Boundary.................................3.3-193.3.3.3Magnetic Jack Assembly..........................................................................3.3-193.3.3.4Position Indication....................................................................................
3.3-193.3.3.5Control Element Assembly Disconnect....................................................3.3-20 3.3.3.6Test Program.............................................................................................3.3-20 3.3.4References.................................................................................................3.3-20
3.4 NUCLEAR
DESIGN AND EVALUATION......................................................
3.4-13.4.1General Summary.......................................................................................3.4-13.4.2Core Description.........................................................................................3.4-1 3.4.3Nuclear Core Design...................................................................................3.4-13.4.3.1Analytical Methodology.............................................................................
3.4-23.4.3.2Physics Characteristics...............................................................................
3.4-23.4.3.2.1Power Distribution Considerations.............................................................
3.4-23.4.3.2.2Control Rod Reactivity Requirements........................................................3.4-23.4.3.2.3Moderator Temperature Coefficient Considerations..................................
3.4-33.4.4Post-Reload Startup Testing.......................................................................
3.4-33.4.5Reactor Stability.........................................................................................
3.4-43.4.5.1General........................................................................................................3.4-4 3.4.5.2Detection of Oscillations............................................................................
3.4-43.4.5.3Control of Oscillations................................................................................
3.4-53.4.5.4Operating Experience.................................................................................
3.4-63.4.5.5Method of Analysis.....................................................................................
3.4-63.4.5.5.1Radial Xenon Oscillations..........................................................................
3.4-73.4.5.5.2Azimuthal Xenon Oscillations....................................................................
3.4-73.4.5.5.3Axial Xenon Oscillations............................................................................
3.4-73.4.6References...................................................................................................3.4-8 3.5THERMAL-HYDRAULIC DESIGN..................................................................3.5-13.5.1Design Bases...............................................................................................3.5-13.5.1.1Thermal Design...........................................................................................3.5-13.5.1.2Hydraulic Stability......................................................................................3.5-13.5.1.3Coolant Flow Rate, Distribution and Void Fraction...................................
3.5-13.5.2Thermal and Hydraulic Characteristics of the Design................................
3.5-23.5.2.1Fuel Temperatures......................................................................................3.5-2 3.5.2.1.1Fuel Cladding Temperatures.......................................................................
3.5-23.5.2.1.2Fuel Pellet Temperatures............................................................................
3.5-23.5.2.1.3UO2 Thermal Conductivity........................................................................
3.5-3 MPS2 UFSAR Table of Contents (Continued)
Section Title Page 3-iii Rev. 353.5.2.1.4Gap Conductance........................................................................................3.5-33.5.2.2Departure from Nucleate Boiling Ratio......................................................3.5-33.5.2.2.1Departure from Nucleate Boiling...............................................................
3.5-33.5.2.2.2Hot Channel Factors...................................................................................
3.5-33.5.2.2.3Effects of Rod Bow on DNBR...................................................................
3.5-53.5.2.3Void Fraction and Distribution...................................................................
3.5-53.5.2.4Coolant Flow Distribution..........................................................................3.5-53.5.2.4.1Coolant Flow Distribution and Bypass Flow..............................................
3.5-53.5.2.4.2Core Flow Distribution...............................................................................
3.5-63.5.2.5Pressure Losses and Hydraulic Loads........................................................3.5-6 3.5.2.5.1Pressure Losses...........................................................................................
3.5-63.5.2.5.2Hydraulic Loads..........................................................................................3.5-73.5.2.6Correlation and Physical Data....................................................................
3.5-73.5.2.7Plant Parameters for Thermal-Hydraulic Design........................................
3.5-73.5.2.8Summary of Thermal and Hydraulic Parameters.......................................
3.5-83.5.3Thermal And Hydraulic Evaluation............................................................3.5-83.5.3.1Analytical Techniques and Uncertainties...................................................
3.5-83.5.3.1.1XCOBRA-IIIC DNBR Analyses................................................................
3.5-83.5.3.1.2Parameter Uncertainties..............................................................................
3.5-83.5.3.2Hydraulic Instability Analysis....................................................................3.5-83.5.3.3Core Hydraulics........................................................................................3.5-11 3.5.3.3.1Fuel Assembly Pressure Drop Coefficients..............................................
3.5-113.5.3.3.2Guide Tube Bypass Flow and Heating Analysis......................................
3.5-123.5.3.3.3Control Element Assembly Insertion Time Analysis...............................3.5-13 3.5.3.3.4Fuel Assembly Liftoff...............................................................................
3.5-133.5.4Tests And Inspections...............................................................................
3.5-143.5.4.1Reactor Testing.........................................................................................
3.5-143.5.4.2AREVA DNB and Hydraulic Testing......................................................3.5-143.5.4.2.1DNB Testing.............................................................................................
3.5-143.5.4.2.2Fuel Assembly Hydraulic Testing............................................................
3.5-143.5.5References.................................................................................................3.5-15 3.A ANALYSIS OF REACTOR VESSEL INTERNALS........................................
3.A-13.A.1Seismic Analysis........................................................................................3.A-13.A.1.1Introduction................................................................................................3.A-13.A.1.2Method of Analysis....................................................................................
3.A-13.A.1.2.1General.......................................................................................................3.A-13.A.1.2.2Mathematical Models................................................................................
3.A-13.A.1.2.3Natural Frequencies and Normal Modes...................................................
3.A-33.A.1.2.4 Response Calculations..............................................................................
3.A-4 MPS2 UFSAR Table of Contents (Continued)
Section Title Page 3-iv Rev. 353.A.1.3Results........................................................................................................3.A-53.A.1.4Conclusion.................................................................................................3.A-53.A.2Normal Operating Analysis.......................................................................3.A-5 3.A.3Loss of Coolant Accident Analysis...........................................................3.A-73.A.3.1Discussion..................................................................................................3.A-73.A.3.2Analysis Codes........................................................................................
3.A-103.A.4Effects of Thermal Shield Removal.........................................................3.A-113.A.5Leak-Before-Break Analysis...................................................................
3.A-113.A.6References................................................................................................
3.A-12 MPS2 UFSAR 3-v Rev. 35CHAPTER 3-REACTOR List of Tables Number Title3.2-1Stress Limits for Reacto r Vessel Internal Structures3.3-1Mechanical Design Parameters
- 3.3-2Pressurized Water Reac tor Primary Coolant Water Chemistry Recommended Specifications3.4-1Fuel Characteristics for a Representative Reload Core3.4-2Neutronics Characteristics for a Representative Reload Core 3.4-3Representative Shutdown Margin Requirements 3.5-1Nominal Reactor and Fuel Design Parameters 3.5-2Design Operating Hydraulic Loads on Vessel Internals 3.5-3Uncertainty Sources for DNBR Calculations (DELETED)3.A-1Natural Frequencies for Vertical Seismic Analys is Mathematical Model3.A-2Seismic Stresses in Critical Reactor Internals Components for the Design Basis Earthquake MPS2 UFSARNOTE: REFER TO THE CONTROLLED PLANT DRAWING FOR THE LATEST REVISION.
3-vi Rev. 35CHAPTER 3 - REACTOR List of Figures Figure Title3.1-1Reactor Vertical Arrangement3.1-2Reactor Core Cross Section 3.3-1Fuel Rod Assembly 3.3-2AAREVA - Reload Fuel Assembly Batch "S" and Prior 3.3-2BAREVA - Reload Fuel Assembly Batch "T" and Later 3.3-3AAREVA - Reload Fuel Assembly Components Batch "S" and Prior3.3-3BAREVA - Reload Fuel Assembly Components Batch "T" and Later3.3-4ABi-Metallic Fuel Spacer Assembly 3.3-4BHTP Fuel Space Assembly 3.3-5Fuel Assembly Hold Down Device 3.3-6Control Element Assembly 3.3-7Control Element Assembly Materials 3.3-8Control Element Assemblies Group and Number Designation 3.3-9Core Orientation 3.3-10In-Core Instrumentation Assembly 3.3-11Reactor Internals Assembly 3.3-12Pressure Vessel-Core S upport Barrel Snubber Assembly3.3-13Core Shroud Assembly 3.3-14Upper Guide Structure Assembly 3.3-15Control Element Drive Mechanism (Magnetic Jack) 3.3-16(Left Blank Intentionally) 3.3-17Heated Junction Thermocouple Probe Pressure Boundary Installation3.3-18Typical Heated Junction Thermoc ouple Probe Assembly Installation3.3-19Placement of Natural Uranium Replacement Fuel Rods and Fuel Assembly Orientation Relative to the Core Baffle for Cycle 193.4-1Representative Full Core Loading Pattern 3.4-2Representative Quarter Core Loading Pattern MPS2 UFSAR List of Figures (Continued)NOTE: REFER TO THE CONTROLLED PLANT DRAWING FOR THE LATEST REVISION.
Figure Title 3-vii Rev. 353.4-3Representative BOC and EOC Exposure Distribution3.4-4Representative Boron Letdown, HFP, ARO3.4-5Representative Normalized Power Dist ributions, Hot Full Power, Equilibrium Xenon, 150 MWD/MTU3.4-6Representative Normalized Power Di stribution, Hot Full Power, Equilibrium Xenon, 18,020 MWD/MTU3.A-1Representative Node Locations - Horizontal Mathematical Model3.A-2Mathematical Model - Ho rizontal Seismic Analysis3.A-3Mathematical Model - Vertical Seismic Analysis3.A-4Core Support Barrel Upper Flange - Finite Element Model3.A-5Core Support Barrel Lower Flange - Finite Element Model3.A-6Lateral Seismic Model - Mode 1, 3.065 CPS 3.A-7Lateral Seismic Model - Mode 2, 5.118 CPS 3.A-8Lateral Seismic Model - Mode 2, 5.118 CPS 3.A-9Reactor Vessel Flange Verti cal Response Spectrum (1% Damping)3.A-10ASHSD Finite Element Model of the Co re Support Barrel/Thermal Shield System3.A-11Vertical Shock Model 3.A-12Lateral Shock Mode 3.A-13SAMMSOR DYNASOR Finite Elemen t Model of Core Support Barrel MPS2 UFSAR3.1-1Rev. 35CHAPTER 3 - REACTOR 3.1
SUMMARY
DESCRIPTION The reactor is of the pressurized water type using two reactor coolant loops. A vertical cross section of the reactor is shown in Figure 3.1-1. The reactor core is comp osed of 217 fuel assemblies, 73 control element a ssemblies (CEA) and four neutr on source assemblies. The fuel assemblies are arranged to approximate a right circ ular cylinder with an e quivalent diameter of 136 inches and an active length of 136.7 inches. The fuel assemblies are co mprised of a structure and fuel and poison rods. The st ructure, which provides for 176 ro d positions, consists of five guide tubes attached to spacer grids and is encl osed at the top and botto m by end fittings. Each of the guide tubes replaces four fu el rod positions and provides a channel which guides the control element over its entire le ngth of travel. In selected fuel asse mblies the central guide tube houses in-core instrumentation. The reactor is currently fueled by assemblies produced by AREVA.The fuel is low enrichment UO 2 in the form of ceramic pellets an d encapsulated in zircaloy tubes. These tubes are seal welded as hermetic enclosures.Figure 3.1-2 shows a view of the reactor core cross section a nd some dimensional relations between fuel assemblies, fuel rods and CEA guide tubes.
The reactor internals support and orient the fuel assemblies and CEAs, absorb the static and dynamic loads and transmit the loads to the reactor vessel flange, provide a passage way for the reactor coolant, and guide in-core instrumentation. The internals will safely perform their func tion during normal operating, upset and emergency conditions. The internals are designe d to safely withstand the forces due to dead weight, pressure differential, flow impingement, temperature differential, vibrat ions and seismic acceleration. All reactor components are consider ed category 1 for seismic desi gn. The reactor internals design limits deflection where required by function. Where necessary, components have been subjected to fatigue analysis. Where appropriate, the effect of neutron irradi ation on the mate rials concerned is included in the design evaluation. The effects of shock loadings on the in ternals is included in the design analysis.
Reactivity control is provided by two independent systems: The control element drive system (CEDS) and the chemical and volume control system (CVCS). The CEDS controls short term reactivity changes and is used for rapid shutdown.
The CVCS is used to co mpensate for long term reactivity changes and can make the reactor subc ritical without the bene fit of the CEDS. The design of the core and the reactor protective system (RPS) prevents fuel damage limits from being exceeded for any single malfunction in either of the reactivity control systems.
The CEAs consist of five poison rods (control elements) assembled in a square array, with one rod in the center. The rods are connect ed to a spider casting which is coupled to the control element drive mechanism (CEDM) shaft.
There are a total of 73 CEAs.
Some CEAs are mechanically connected in pairs and are known as dual CEAs.
MPS2 UFSAR3.1-2Rev. 35 Both dual and single CEAs are maneuvered by magnetic jack type CEDM's mounted on the reactor vessel head.The maximum reactivity worth of the CEAs and the associated reactivity a ddition rate are limited by core, CEA and CEDS design to prevent sudden large reactivity in creases. The design restraints are such that reactivity increases will not result in violation of th e fuel damage limits, rupture of the reactor coolant pressure boundary (RCPB), or disruption of the core or other internals sufficient to impair the effectiveness of emergency cooling.
The three-batch fuel management scheme is employed, where approximate ly 40 percent of the core is replaced at each refueling. Sufficient margin is provided to ensure that peak burnups of the individual fuel assemblies ar e within acceptable limits.
The nuclear design of the core will ensure that th e combined response of all reactivity coefficients to an increase in reactor thermal power yields a net decrease in reactiv ity and that CEAs are moved in groups to satisfy the requirements of shutdown, power level changes and operational maneuvering. The control systems are designed to produce power dist ributions that are within the acceptable limits on overall nuc lear heat flux factor (F N Q) and departure from nucleate boiling ratio (DNBR). The RPS and administ rative controls ensure that these limits are not exceeded.
The reactor coolant enters the upper section of the reactor vessel through f our inlet nozzles, flows downward between the reactor vessel shell and the core barrel, and passes th rough the flow skirt and into the lower plenum where the flow dist ribution is equalized. The coolant then flows upward through the core removing h eat from the fuel rods, exits from the reactor vessel through two outlet nozzles and passes thr ough the tube side of the verti cal "U" tube steam generators where heat is transferred to the secondary system. The reactor coolant pumps (RCPs) return the coolant to the reactor vessel.
The principal objective of the th ermal-hydraulic design is to a void fuel damage during normal operation and anticipated transients. It is recogni zed that there is a small probability of limited fuel damage in certain situations as discussed in Chapter 14.In order to meet the objective of the thermal-hydraulic design the following design limits are established, but violation of either is not necessarily equivalent to fuel damage:a.There is a high confidence level that departure from nucl ear boiling (DNB) is avoided during normal operation and anticipated transients. This is achieved by confirming the minimum DNBR calculated according to the HTP correlation (Reference 3.1-1) is greater than the 95/95 limit for the correlation;b.The melting point of the UO 2 fuel is not reached during normal operation or anticipated transients.
The RPS and the reactor c ontrol system (RCS) provide for automatic reactor trip or corrective actions before these design limits are exceeded.
MPS2 UFSAR3.1-3Rev. 35The core design bases are presented in Section 3.2; the core mechanical design is discussed in Section 3.3; the nuclear design of the core is discussed in Section 3.4; and the thermal and hydraulic design is discussed in Section 3.5.
3.
1.1 REFERENCES
3.1-1EMF-92-153(P)(A) Rev. 1, "H TP: Departure From Nucleate Boiling Correlation for High Thermal Performance Fuel," Siem ens Power Corporation, January 2005.
MPS-2 FSAR April 1998 Rev. 26.2FIGURE 3.1-1REACTOR VE RTICAL ARRANGEMENT MPS-2 FSAR April 1998 Rev. 26.2FIGURE 3.1-2REACTOR CORE CROSS SECTION MPS2 UFSAR3.2-1Rev. 35
3.2 DESIGN
BASES The full power thermal rating of the core is 2,700 MWt. The physics a nd thermal and hydraulic information presented in this secti on is based on this core power level.
3.2.1 MECHANICAL
DESIGN BASES 3.2.1.1 Fuel Assembly Design Bases The design bases for evaluating the structural integrity of AREVA fuel assemblies are:
A.Fuel Assembly Handling The fuel assembly is evaluated for dynamic ax ial loads of approximately 2.5 times the fuel assembly weight.B.For All Applied Loads for Normal Oper ation and Anticipated Operational EventsFuel assembly component strength is evaluated against either prototype testing or elastic stress analysis. When the stress analysis method is used, the stress limits presented in the ASME Boiler and Pressure Vessel Code,Section III, Division 1, are used as a guide.
The stress design limits for structural components are:
P m 1.0S m P m + P b 1.5S m P + Q 3.0S m where: P m is the primary membrane stress intensity P b is the primary bending stress intensityP is the primary stress intensity Q is the secondary stress intensity The design stress, S m is identified in the ASME Boiler and Pressure Vessel Code for austenitic stainless steel as a function of temperature. In the case of Zircaloy, which is not specifically identified in the ASME Boiler and Pressure Vess el Code, the design stress is identified as the lesser of two-thirds the yield stress, S y , or one-third the ultimate stress, S u.The ASME Boiler and Pressure Vessel Code de fines the stress intensity based on the maximum shear stress theory. The stress intensity is equal to one-half the largest algebraic difference between two principal stresses.
MPS2 UFSAR3.2-2Rev. 35 Primary stresses are deve loped by the imposed loading which is necessary to satisfy the laws of equilibrium between external and internal forces and moments. The basic characteristic of a primary stress is that it is not self-limiting. If a primary stress exceeds the yield strength of the material through the entire wall thickness, the prevention of fail ure is entirely dependent on the strain-hardening properties of the material.
Secondary stresses are developed by the self-constrai nt of a structure. It must satisfy an imposed strain pattern rather than being in equilibrium with an external load. The basic characteristic of a secondary stress is that it is self-limiting. Local yi elding and minor distor tions can satisfy the discontinuity conditions due to thermal expansions which cause the stress to occur.C.Loads during Postulated Accidents Deflection or failure of components shall not interfere with reactor shutdown or emergency cooling of the fuel rods.
The fuel assembly structural component stresses under faulted conditions are evaluated using primarily the methods outlined in Appendix F of the ASME Boiler and Pressure Vessel Code,Section III. The current methods utilize the limits provided fo r elastic system analysis.
The design stress intensity value (S m) is defined the same as fo r normal operating conditions.
Spacer grid crush load strength is based on the 95% confidence le vel on the true mean as taken from test measurements on unirradiated production grids at (or corrected to) operating temperature.
3.2.1.2 AREVA Fuel Rod Cladding Design BasesA discussion of the AREVA fuel ro d cladding is given as part of the AREVA fuel rod discussion in Section 3.3.1.1.
3.2.1.3 Control Element Assembly Design Bases The CEA has been designed to ensure that the stress intensitie s in the individual structural components do not exceed the allowable limits for the appropriate material established in Section III of the ASME Boiler and Pressure Vessel Code. The exceptions to this criterion are that (a) the Inconel 625 cladding is permit ted to sustain plastic strain up to 3 percent due to irradiation induced expansion of the filler materials, and (b) because th e ASME Code does not apply to springs, the allowable stresses fo r the CEA springs are based on valu es which have been proven in practice.The CEA stress analyses consider the following load sources:a.Internal pressure build up due to the effect of irradiation on B 4 C (production of helium).
MPS2 UFSAR3.2-3Rev. 35b.External pressure of r eactor coolant (in the computation for determining the maximum stress in the cladding due to inte rnal pressure, no inte rnal pressure is assumed).c.Dynamic stresses produced by seismic loading.d.Dynamic loads produced by stepping motion of the magnetic jack.
e.Mechanical and hydraulic loads produced during SCRAM.f.Cladding loads produced by differential expansion between clad and filler materials.In addition to the comparison of calculated stress levels with allowable stresses, the fatigue damage produced by significant cyclic stresses is also determined.
It is a design requirement that the calculated cumulative damage factor for any location may not be equal to or greater than 1.0.
The fatigue usage factor calculations are based on the fatigue curves (str ess range vs. number of cycles) contained in Section III of the ASME Boiler and Pressure Vessel Code.
3.2.1.4 Reactor Internals Design Bases The reactor vessel internals are designed to meet the loading c onditions and the design limits specified below. The materials used in fabrication of the reactor internal structures are primarily Type 304 stainless steel. Th e flow skirt is fabricated from Inconel. Welded connections are used where feasible; however, in locations where mechanical connections are required, structural fasteners are used which are designed to remain captured in the event of a single failure.
Structural fastener material is typically a high strength austenitic stainless steel; however, in less critical applications, Type 316 stai nless steel is employed. Hardfacing, of Stellite material, is used at wear points. The effe ct of irradiation on the properties of the materials is considered in the design of the reactor internal structures.A.Categorization and Combination of Loadings1.Normal Operating and Upset Conditions The reactor vessel internals are designed to perform their functi ons safely without shutdown. The combination of design lo adings for these conditions are the following:Normal operating temperature differencesNormal operating pressure differences
Low impingement loads Weights, reactions and superimposed loads MPS2 UFSAR3.2-4Rev. 35Vibration loads Shock loads (including OBE)Transient loadings of frequent occurrences not requiring shutdown Handling loads2.Emergency Conditions The internals are designed to permit an acceptable amount of local yielding while experiencing the loadings listed above with the SSE load replacing the OBE load.3.Faulted Conditions Permanent deformation of the reactor internal structur es is permitted. The loadings for these conditions include all the loadings listed for emerge ncy conditions plus the loadings resulting from the postulated LOCA.B.Design Limits Reactor internal compone nts are designed to ensure that th e stress levels and deflections are within an accept able range. The stress values for core support structures are not greater than those given in the May 1972 draft of Section III of the ASME Boiler and Pressure Vessel Code, Subsection NG, including Appendix F, "Rules for Evaluation of Faulted Conditions." St ress limits for the reactor ve ssel core support structures are presented in Table 3.2-1. In addition, to properly pe rform their functions, the r eactor internal structures will satisfy the deformation limits listed below
.1.Under design loadings plus operating ba sis earthquake forces or normal operating loadings plus SSE forces, de flections will be limited so that the CEAs can function and adequate core cooling is preserved.2.Under normal operating loadings plus SS E forces plus pipe rupture loadings resulting from a break of the largest line connect to the primary system piping, deflections will be limited so that the core will be held in place, adequate core cooling is preserved, and all CEAs can be inserted. Those deflections which would influence CEA movement will be limited to less than 80 percen t of the deflections required to prevent CEA insertion.3.Under normal operating loadings plus SSE forces plus the maximum pipe rupture loadings resulting from the full spectrum of pipe breaks, deflections will be limited so that the core will be held in place and adequate core cooling is preserved.
Although CEA insertion is not required for a safe and orderly shutdown for break sizes greater than the largest line connected to the primary system piping, calculations show that the CEAs will be insertable for larger breaks except for a MPS2 UFSAR3.2-5Rev. 35 few CEAs located near the vessel outlet nozzle which is feeding the postulated rupture.3.2.1.5 CEDM/RVLMS (HJTC) Pressu re Housing Design Bases The control element drive mechanism and Reactor Vessel Level Monitoring System (RVLMS) pressure housings form part of the reactor coolant boundary and are, therefore, designed to meet the stress requirements consistent with those of the reactor ve ssel closure head. The limiting stresses in the CEDM's and RV LMS pressure boundary components due to the design, Level A, Level B, Level C, Level D and Test conditions satisfy ASME Boiler Pressure Vessel Code,Section III, Subsection NB plus Appendix 1 and Section II, Pa rt D, 1998 Edition through 2000 Addenda, including Code Case N-4-12 for the CEDM motor housing material.The CEDMs and the RVLMs are designed to function normally during and after exposure to normal operating conditions plus the design basis earthquake (DBE). Under normal operating conditions, plus DBE, plus pipe rupture loadings, de flections of the CEDM will be limited so that the CEAs can be inserted afte r exposure to these conditions.
Those deflections, which could influence CEA movement, will be limited to less than 80 percent of the deflections required to prevent CEA movement. The RVLMS and the adja cent CEDMs do not cont act each other with maximum lateral displacement of the pressure housings.
Loading Combinations ASME Code SubsectionDesign ConditionP m S m NB-3221 P 1 1.5S m P 1 + P b < 1.5S m Level A and Level BP 1 + P b + Q 3S m NB-3222 and NB3223(Normal and Upset)U 1Level C ConditionP m greater of [1.2S m , S y]NB-3224(Emergency)P 1 + P b greater of [1.8S m , 1.5S y]Level D ConditionP m lesser of [2.4S m , 0.7S u]Paragraph F-1330 or F-1340, Appendix F(Faulted)P 1 + P b lesser of [3.6S m , 1.05S u]Test ConditionsP m 0.9S y NB-3226 P m + P b 1.35S y when P m 0.67S y or P m + P b (2.15 S y - 1.2P m) when 0.67S y < P m 0.9S y Design ConditionP m S m NB-3221 MPS2 UFSAR3.2-6Rev. 35Where P m = General primary membrane stress intensity P 1 = Primary local membrane stress intensity P 1 + P b = Primary membrane plus bending stress intensity P 1 + P b + Q = Primary plus sec ondary stress intensity S m = Design stress intensity S y = Yield strength S u = Tensile strength U = Cumulative fatigue usage factor
3.2.2 NUCLEAR
DESIGN BASES The initial full power thermal rati ng of the core is 2700 MWt. It is upon this power level that the physics and thermal and hydraulic information presented in this section are based. The design basis for the nuclear design of the fuel and reactivity control systems are: a.Excess Reactivity and Fuel BurnupThe excess reactivity provided for each cycle is based on the depletion characteristics of the fuel and burnabl e poison and the desired burnup for each cycle. The desired burnup is based on the ec onomic analysis of both the fuel cost and the projected operating load demand cycle for the plant. The average burnup in
the core is chosen so as to insure that the peak assembly burnup is not greater than 56,000 MWD/MTU for Batch N, 52,500 MWD/MTU for Batch P, and 57,400
MWD/MTU for Batch R and later
.b.Core Design Lifetime and Fuel Replacement ProgramThe core design lifetime and fuel repl acement program are based on a th ree region core with approximately 40 percent of the fuel assemblies replaced at each refueling.c.Negative Reactivity Feedback and Reactivity Coefficients The negative reactivity feedback provided by the design is based on the requirement of General Design Criterion (GDC) 11. In the power operating range, the inherent combined response of the reactivity feedback characteristics (fuel temperature coefficient (FTC), moderator temperature coefficient (MTC), moderator void coefficient (MVC), and moderator pressure coefficient (MPC)) to
an increase in reactor thermal power will be a decrease in reactivity.Shear Stress 0.6S m NB-3227.2Loading Combinations ASME Code Subsection MPS2 UFSAR3.2-7Rev. 35d.Burnable Poison Requirements The burnable poison reactivity worth provided in the design will be suff icient to ensure that moderator coefficients of reactivity have magnitudes and algebraic signs consistent with the requirements for negative reactivity feedback and acceptable consequence in the event of postulated accidents or anticipated operational occurrences, viewed in conjunction with the supplied protective equipment.e.Stability Criteria The design of the reactor and the instrume ntation and control systems is based on meeting the requirements of GDC 12 with respect to spatial oscillations and stability. Sufficient CEA rod worth will be availabl e to suppress xenon-induced power oscillations.f.Maximum Controlled Reactivity Insertion Rates The maximum reactivity addition rates are limited by core, CEA, and reactor regulating system (RRS) design based on pr eventing increases in reactivity which would result in the violation of specified acceptable fuel design limits, damage to the reactor pressure boundary, or disruption of the core or other internals sufficient to impair the effectiveness of emergency core cooling.g.Power Distribution Control Acceptable operation of the reactor in the absence of an accidental transient depends on maintaining a relationship among many parameters, some of which depend on the power distribution. In the ab sence of an accidental transient the power distribution is controlled such th at in conjunction with other controlled parameters, limiting conditi ons of operation (LCO) are not violated. LCO are not less conservative than the initial conditions used in the accident analyses in Chapter 14. LCO and limiting safety system settings (LSSS) are determined such that specified acceptable fuel design limits are not violated as a result of anticipated operational occurrences and such that specified predicted acceptable consequence are not exceeded for other postulated accidents.h.Shutdown Margins and Stuck Rod Criteria The amount of reactivity available from insertion of withdr awn CEAs is required to be sufficient, under all pow er operating conditions, to en sure that the reactor can be brought to at least 3.6 percent subcritical from th e existing condition, including the effects of cooldown to an average coolant temperature of 5 32°F, even when the highest worth CEA fails to in sert. This criteria is exclusive of any safety allowance and is consistent with the most pessimistic analysis in Chapter 14.
MPS2 UFSAR3.2-8Rev. 35i.Chemical Shim Control The chemical and volume control system (CVCS) (Section 9.2) is used to adjust dissolved boron concentration in the moderator. After a reactor shutdown, this system is able to compensate for the reactivity changes as sociated with xenon decay and reactor coolant te mperature decrease to ambi ent temperature. It also provides adequate shutdown mar gin during refueling. This system also has the capability of controlling long term reactivity changes due to fuel burnup, and reactivity changes during xenon transients resulting from changes in reactor load independently of the CEAs. In particular, any xenon transient may be accommodated at any time in the fuel cycle.
3.2.3 THERMAL
AND HYDRAU LIC DESIGN BASISAvoidance of thermally induced fuel damage during normal steady state and anticipated transient operation is the principal thermal and hydraulic design basis. It is recognized that there is a small probability of limited fuel damage in certain unl ikely accident situations discussed in Chapter 14.
The following corollary design ba sis are established, but violati on of them is not necessarily equivalent to fuel damage.a.A limit corresponding to 95% proba bility with 95% confidence (Reference 3.2-1) is set on the departure from nucleat e boiling ratio (DNBR) during normal operation and any anticipated transients as calculated according to the HTP correlation.b.The peak temperature of the fuel will be less than the melti ng point during normal operation and anticipated transients.
The reactor control and protecti on system will provide for automatic reactor trip or other corrective action before thes e design limits are exceeded.
The core hydraulic resistance was considered in establishing the ope rational limits curves provided in Figures 4.5-4 and 4.5-5, and the Low Temperature Overpressure Protection (LTOP) System described in Section 7.4.8. As fuel design changes, effects on the flow resistance will be
evaluated to determine the impact.
3.
2.4 REFERENCES
3.2-1EMF-92-153(P)(A) Rev. 1, "H TP: Departure From Nucleate Boiling Correlation for High Thermal Performance Fuel," Siem ens Power Corporation, January 2005.
MPS2 UFSAR3.2-9Rev. 35TABLE 3.2-1 STRESS LIMITS FOR REACTOR VESSEL INTERNAL STRUCTURESOperating ConditionsStress Categories and Limits of Stress Intensities1.Normal and UpsetFigure NG 3221.1 including notes2.EmergencyFigure NG 3224.1 including notes3.FaultedAppendix F, Rules for Evaluating Faulted Conditions MPS2 UFSAR3.3-1Rev. 35
3.3 MECHANICAL
DESIGN The reactor core and internals are shown in Figure 3.3-1. A cross section of the reactor core and internals is shown in Figure 3.1-2. Mechanical design features of the reactor internals, the control element drive mechanisms (CEDM) and the core are described below. Mechanical design parameters are listed in Table 3.3-1.
3.3.1 CORE MECHANICAL DESIGN The core approximates a right circular cylinder with an equivalent diameter of 136 inches and an active height of 136.7 in ches. It is made up of Zircaloy-4 clad fuel r ods containing slightly enriched uranium in the form of sintered UO 2 pellets and UO 2-Gd 2 O 3 pellets. The fuel rods are grouped into 217 assemblies.
Short term reactivity control is provided by 73 control element assemblies (CEA). The CEAs are guided within the core by the guide tubes which are integral parts of the fuel assemblies.
3.3.1.1 AREVA Fuel Rod The detailed fuel rod design (see Figure 3.3-1) establishes such parameters as pellet diameter and length, density, cladding-pellet di ametral gap, fission gas plenum size, and rod pre-pressurization level. The design also considers effects and physical properties of fuel rod components which vary with burnup.
The integrity of the fuel rods is ensured by designing to prevent excessive fuel temperatures, excessive internal rod gas pressures, and excessi ve cladding stresses and strains. This end is achieved by designing the fuel r ods to satisfy the design crit eria during norma l operation and anticipated operational occurrences over the fuel lifetime. Fo r each design criteria, the performance of the most lim iting fuel rod shall not exceed the specified limits.
Fuel rods are designed to function throughout the design life of the fuel based upon the reactor operating conditions designated below without loss of mechanical integrity, significant dimensional distortion, or releas e of fuel or fission products.
The assemblies were evaluated for a peak assembly burnup of 56,000 MDW/MTU for Batch N, 52,500 MWD/MTU for Batch P, and 57,400 MWD/MTU for Batch R and later.The Millstone Unit 2 Cycle 19 reload core included four fuel assemblies with natural uranium replacement fuel rods with an anti-rotation feature designed to pr event spinning of the rod during operations. The four assemblies containing replacement rods, and the conditions under which they were evaluated for use, are discussed in Section 3.3.1.3.1, "Design Summary".
3.3.1.1.1 Fuel Rod Mechanical Criteria The cladding primary and secondary stresses sh all meet the 1977 ASME Boiler and Pressure Vessel Code Section III (Reference 3
.3-1) requirements summarized below:
MPS2 UFSAR3.3-2Rev. 35 Primary stresses are deve loped by the imposed loading which is necessary to satisfy the laws of equilibrium between external and internal forces and moments. The basic characteristic of a primary stress is that it is not self-limiting. If a primary stress exceeds the yield strength of the material through the entire thickness, the prevention of failure is entirely dependent on the strain-hardening properties of the material.
Secondary stresses are developed by th e self constraint of a structur
- e. It must satisfy an imposed strain pattern rather than being in equilibrium with an external load. The basic characteristic of a secondary stress is that it is self-limiting. Local yi elding and minor distor tions can satisfy the discontinuity conditions due to thermal expansions which cause the stress to occur.Cladding circumferential strain shall not exceed the design limit through end-of-life (EOL).The total uniform strain, elastic and plastic shal l not exceed the design limit during a transient.
The strain analysis was performed with th e RODEX2 (Reference 3.3-2) RAMPEX codes benchmarked to available power ramp test data, i.e., INTERRAMP, OVERRAMP, and SUPERRAMP.
The fuel rod shall be designed such that at a rod average burnup when substantial axial consolidation has occurred, the total clad creep deformation shall not exceed the initial minimum diametral fuel cladding gap. This will prevent pellet hangups allowi ng the plenum spring to close axial gaps until densification is substantially co mplete, thus preventing the formation of pellet column gaps of sufficient size for clad flattening.
The fuel rod pressure at EOL shall not exceed the criteria appr oved by the NRC (Ref. 3.3-3). A review of departure from nucleate boiling ratio (DNBR) limits fo r condition III or IV postulated accidents events is required for fuel rods that exceed nominal system pressure. When fuel rod pressure is predicted to exceed system pressu re, the pellet-cladding gap shall not increase for steady or increasing power conditi ons. Analysis approved by the NRC has shown that the fuel rod gas pressure can safely exceed system pressu re without causing any da mage to the cladding.Total cladding wall thinning due to generalized external and internal corrosion shall not exceed a value which will impair mechanic al performance over the projected fuel rod design lifetime under the most adverse projected power conditions within coolant ch emistry limits recommendations of Table 3.3-2. It will also assure that the metal/oxide interface temperature will re main well below Stress Intensity Limits (Parameter)Yield StrengthUltimate Tensile Strength Primary Membrane (P m)< 2/3 S y< 1/3 S u Primary Membrane Plus Primary Bending (P m + P b)< 1.0 S y< 0.5 S u Primary Plus Secondary (P + Q)
< 2.0 S y< 1.0 S u MPS2 UFSAR3.3-3Rev. 35the level where large increases in corrosion, due to the insulating eff ect of the oxide, would adversely affect the mechanic al behavior of the cladding.The cumulative usage factor for cyclic stresses for all important cyclic loading conditions shall not exceed the design limit.
The clearance between the upper and lower tie plate shall be able to accommodate the maximum differential fuel rod and fuel assembly growth to the designed burnup.
The centerline temperature of th e hottest pellet shall be below the melting temperature. Fuel centerline temperature is calculated at overpower conditions to verify that fuel pellet overheating does not occur during normal operation and anticipated operational occurrences.
3.3.1.1.2 Fuel Rod Design Analyses Each design analysis was performed with AREVA methodology which invol ves a well defined selection of appropria te data and parameters, a nd the latest approved versio ns of computer codes.
This methodology , as required, ha s been submitted to the Nucl ear Regulatory Commission (NRC) and approved. The analysis is performed in accordance with the methods described in AREVA's "Qualification of Exxon Nuclear Fuel Fo r Extended Burnup" (Reference 3.3-3).The cladding steady state stress analysis was performed by cons idering primary and secondary membrane and bending stresses due to hydrostatic pressure, flow-induced vibration, spacer contact, pellet cladding interaction (PCI), thermal and mechanic al bow and thermal gradients. Stresses were calculated for the various combinations of the following conditions:a.beginning of life (BOL) and EOLb.cold and hot conditionsc.at mid-span and at spacer locationsd.at both the inner and outer surfaces of the claddingThe analysis was performed for the various sources of stress, in cluding pressure, thermal, s pacer contact, PCI, and rod bow. The app licable stresses at each orthogonal direction were combined to calculate the maximum stress intensities which are compared to the ASME design criteria. The results of the analysis indicate that all stress values are within acceptable design limits for both BOL and EOL, hot and cold condi tions. The EOL stresses have ample margin for both the hot and cold condition stresses.
The cladding steady state strain is evaluated w ith the RODEX2 code, which has been approved by the NRC (Reference 3.3-2). The code consider s the thermal-hydraulic environment at the cladding surface, the pressure insi de the cladding, and the thermal, mechanical and compositional state of the fuel and cladding. Pellet density, sw elling, densification, and fission gas release or absorption models, and cladding a nd pellet diameters ar e input to RODEX2 to provide the most MPS2 UFSAR3.3-4Rev. 35conservative strain calculation or subsequent ramping or collapse calculations for the reference fuel rod design. The major fuel rod performanc e characteristics modeled by the RODEX2 code are:a.Radial Thermal Conduc tion and Gap Conductanceb.Fuel Swelling, Densification, Cracking, and Crack Healingc.Gas Release and Absorption d.Cladding Creep Deformation and Irradiation-Induced Growthe.Cladding Corrosionf.PCI g.Free Rod Volume and Gas Pressure The calculations are performed on a time incremental basis with conditions updated at each calculated increment so that th e power history and path dependent processes can be modeled. The axial dependence of the power a nd burnup distributions are handled by dividing the fuel rod into a number of axial and radial regions. Power distributions can be ch anged at any desired time, and the coolant and cladding temperatures are readjusted in all the region
- s. All the performance models, e.g., giving the defo rmations of the fuel and cladding a nd gas release, ar e calculated at successive times during each period of assume d constant power generation. The calculated cladding strain is reviewed throughout the life of the fuel and both the maximum circumferential strain and the maximum strain increment are comp ared with the design criteria. The calculated strain did not exceed the strain limit. Both the maximum strain and the positive strain increment are below the design limit strain.
The ramping strain and the fati gue evaluation of the fuel rod were evaluated. The ramps are assumed to occur anytime duri ng the irradiation and may reac h the maximum peaking factor allowed by the limits of operation. The ramps ar e analyzed either from cold shutdown or from a variety of hot powered st arting conditions. The approach to ra ted power at the beginning of each reactor cycle is performed to satisfy the AREVA maneuvering and conditioning recommendations. The clad response during ramp ing power changes is calculated with the RAMPEX code. This code calculates the PCI duri ng a power ramp for one axial node at a time.
The initial conditions are obtai ned from RODEX2 output. The RAMPEX code considers the thermal condition of the rod in its flow channel, and the mechanical interactions that result from fuel and cladding creep at any desired axial section in the rod during the power ramp. As compared to RODEX2, RAMPEX a dditionally models the pellet cl adding axial stress interaction, primary creep with strain hardening, the effects of pellet chips, and localized stresses due to ridging.The RAMPEX code provides the hoop stress and the stress intensity. The stress results of the ramping analysis are used to ev aluate the cladding fatigue damage through life due to the cyclic MPS2 UFSAR3.3-5Rev. 35 power variations. The fatigue an alysis is based on the O'Donnel and Langer (Reference 3.3-4) design curve. The cyclic amplitudes of the maximum local stress intensity, as determined by RAMPEX over the power cycling range, are compared with this curve to determine the allowed cycles for each stress range. This result is combined with the projected number of duty cycles to determine a fatigue usage factor.
All of the reactor cycle (startup) ramp stresses were within the design limit.
Creep collapse calculations are performed with RODEX2 a nd COLAPX codes. The RODEX2 code determines the cladding temperature and in ternal pressure history based on a model which accounts for changes in fuel r od volumes, fuel densification a nd swelling, and fill gas absorption.
The reactor coolant, fuel rod internal temperature, and pressure histories generated by the RODEX2 analysis are input to th e COLAPX code along with a cons ervative statisti cal estimate of initial cladding ovality and the fast flux history. The COLAPX code calculates, by large deflection theory, the ovality of the claddi ng as a function of time while the uniform cladding creepdown is obtained by the RODEX2 an alysis. The cladding ovality increa se and creepdown are summed, at a rod average burnup when substantia l axial consolidation has occurre d, to show that they remain less than the initial minimum pell et clad gap. Measurem ents of highly densif ying irradiated fuel have demonstrated that pellet densification is essentially complete by the time the fuel has attained this burnup so that furthe r creepdown after this phase will not result in significant pellet to pellet gaps. The combined radial creepdown wa s shown to meet the desi gn criteria. This will prevent pellet hangups due to cla dding creep, allowing the plenum sp ring to close axial gaps until densification is substantially complete, and thus assures that clad collapse will not occur. The pitch of the plenum spring is less than the spacing calculated for stiffening rings in a cylindrical shell under external pressure which will pr event clad collapse in the plenum area.
Calculation of the gas pressure within a fuel r od is performed with the RODEX2 code. The initial fill gas is found by calcul ating the initial free volume and using the ideal gas law, along with input values for fill gas pressure a nd reference fill gas temperatur
- e. The free gaseous fission product yield is calculated for each ax ial region and the total yield obta ined by summing the axial region contributions. The power of each history used was multiplied fo r each cycle by a factor required for the highest projected rod power to reach the F r limit plus uncertainties. The calculations show that for all power histories analy zed, the rod internal gas pressure will remain below the criteria approved by the NRC (Reference 3.3-3) for us e in extended burnup gas pressure analysis.The waterside corrosion of fuel rods is evaluated with the MATPRO-11 (Reference 3.3-5) correlation. The MATPRO-11 mode l is a two-stage corrosion ra te model which is cubic in dependence on oxide thickness until a transition to a subsequent linear dependence occurs. To calculate the rate changes as a function of both oxide thickness a nd the operating conditions of the fuel rod, the MATPRO model is incorporated into AREVA's RODEX2 fuel performance code.
The RODEX2 code determines the temperature incr ease of the water along the fuel rod assuming heat balance within a channel for the prescribed mass flow and inlet temperature. The radial temperature drops are evaluated successively between the water, the oxide surface, the metal/
oxide interface, and the inside of the cladding using RODEX2 correlations and methods. To account for the change in corros ion rate due to the changing oxi de layer and thermal conditions, the code includes an update in cl adding temperature at ev ery calculation step. This is an iterative process due to the continuously changing oxide thickness. Conditio ns are also revised at times MPS2 UFSAR3.3-6Rev. 35where new power or flow conditions are prescribed. The MATPRO model incorporated in RODEX2 is benchmarked via an overall enhanc ement factor to oxide thickness data from assemblies in seven separate reactors. Each data point represents the maximum thickness measured along a rod length. The enha ncement factor is based on a be st fit regression analysis of the data. A final multiplier is also applied which envelopes the data. The waterside corrosion in the cladding was evaluated with RODEX2 for the steady state strain analysis. A best-fit corrosion amplification factor was applied to the MATPRO model along with a final mu ltiplier to bound the measured data on AREVA standard cladding.
The maximum calculated oxide thickness was below the design limit.
Fuel rod and fuel assembly growth projected to occur during irradiation was based on conservative design curves establ ished from measured irradiati on growth data. The rod growth minus the assembly growth plus tolerances was compared with the clearance within the assembly for fuel rod growth. Differential thermal expansion between the fuel rods and guide tubes was also considered. There is space between the uppe r and lower tie plates to accommodate the maximum differential growth out to a rod burnup of 62,000 MWd/MTU.The pellet centerline temperatur e calculation was performed with the RODEX2 code. Fuel pellet centerline temperatures were calculated at overpower conditions. The high power cycle of each power history was modified to include a spike in each cycle. This spike increased the maximum power of a pellet in the rod up to F T Q. Pellet melting temperature is a function of burnup.
Considering a conservative peak pellet bur nup to determine the minimum pellet melting temperature at EOL, the maximu m pellet centerline temperatur e is well below both BOL and EOL limits.
3.3.1.2 (Deleted)3.3.1.3 AREVA Fuel Assembly 3.3.1.3.1 Design SummaryThe AREVA fuel assemblies are 14 by 14 arrays containing 176 fuel rods in a cage structure of 5 guide tubes and 9 spacer grids. Both the guide tubes and the fuel r od cladding are made of Zircaloy-4 for low neutron absorp tion and high corrosion resistance.
The fuel assembly upper tie plates are stainless steel cast ings with Inconel holddown spri ngs. The fuel assembly upper tie plate is mechanically locked to the guide tubes and may be easil y removed to allo w inspection of irradiated fuel rods. For Reload T (Cycle
- 15) and beyond, lower tie plates are the FUEL GUARD debris resistant design.
In Reloads M, N, and P (Cycles 10-12), eight of the nine spacers in each fuel assembly are made of a Zircaloy-4 structure with Inconel-718 springs (i.e., bi-metallic spacer). The ninth spacer, located just above the lower tie plate, is made of Inconel-718 and, using features of the AREVA High Thermal Performance (HTP) spacer design, has been adapted to provide fuel assembly debris resistance.
MPS2 UFSAR3.3-7Rev. 35The fuel assembly design for Re loads R and S (Cycles 13 and 14) has all nine spacers of the bimetallic design. Additionally, in this design a longer solid fu el rod lower end cap is used. The longer end cap serves to raise th e fuel rod cladding above the debr is trapping regi on of the ninth (bottom) spacer.
In Reloads T through X (Cycles 15-18), the Hi gh Thermal Performance (HTP) fuel assembly design was implemented in which all nine spacers are of the Zi rcaloy-4 HTP desi gn. This design retained the longer, solid fuel rod lower end cap. The fuel assembly design for Relo ad Y (Cycle 19) and later utili zed eight Zircaloy-4 HTP spacers and replaces the ninth, bottom spacer with an Inconel High Mechanical Performance (HMP) spacer. The HMP spacer is similar to the HTP spacer, except that it is constructed of Inconel-718 and the flow channels are parallel to the fuel. Drawings of the AREVA fuel assemblies are given in Figure 3.3-2A and Figure 3.3-3A. Fuel assembly drawings for Reload T (Cycle 15) and beyond are in cluded in Figur es 3.3-2B and 3.3-3B.The analysis has shown that the AREVA reload fuel assemblies will meet the design criteria:a.The maximum steady state cladding strain is well below the design limit.b.The maximum steady state cladding stress meets the ASME Boiler and Pressure Vessel Code requirements.c.The transient strain is within the circumferential limit.
d.Cladding creep collapse is precluded.e.The fuel rod internal pressure at the EOL remains below the criteria approved by the NRC (Ref. 3.3-3).f.The maximum clad oxidation is below the design limit.
g.The cladding fatigue usage factor is well below the design limit.h.There is space between the upper and lower tie plate to accommodate fuel rod growth.i.Pellet centerline temperatures remain below the design criteria.j.The fuel assembly growth is within the space available between the upper and lower core plates in the reactor core.
k.The assembly holddown springs will prevent bundle lift-off.
MPS2 UFSAR3.3-8Rev. 35The fuel rods consist of short cylindrical UO 2 pellets or UO 2-Gd 2 O 3 pellets contained in Zircaloy-4 tubular cladding. Zircaloy-4 end caps are welded to each end to give a hermetic seal.
The fuel rod upper plenum contains a high stre ngth alloy compression spring to prevent fuel column separation during fabric ation and shipping, and during in core operation.
The rods are pressurized with helium to improve heat transfer and reduce clad creep ovality.
The fuel assembly structure cons ists of an upper tie plate assembly, lower tie plate, guide tubes and spacer grids, which together pr ovide the support for the fuel rods.
The lower tie plate is a machined stainless steel castin g which provides the lo wer end support for the guide tubes. The Zircaloy guide tubes are attached to the lowe r tie plate by means of Inconel cap screws. The FUELGUARD TM lower tie plate, in cluded in Reload T and beyond provides protection to the fuel from de bris in the primary coolant.
The upper tie plate assembly latc hes to and provides the upper e nd support for the guide tubes. The upper tie plate assembly consists of a st ainless steel grid structure and reaction plate containing five Inconel X-750 ho lddown springs. The springs are located around Inconel locking nuts and sleeves which mechanically attach to the guide tubes and pilot into the reactor alignment plate. The springs are partially shrouded on the outside di ameter by stainless st eel cups to prevent flow induced spring vibration.
The guide tubes, in conjunction wi th the spacers and tie plates, form the structural skeleton of the fuel assembly and provide cha nnels for insertion of the cont rol rods. The guide tubes are fabricated from Zircaloy-4 tubing and are fully annealed. The center tube is of uniform diameter whereas the outer four gui de tubes have a reduced diameter section at the bot tom which produces a dashpot action to decelerate the dropped CEAs.
An end plug is welded to the lowe r end of the guide tube and is drilled and threaded to accept the lower cap screws. At the upper end, the guide tube is crimped into an external stainless steel locking sleeve which engages the upper tie plate assembly. The upper tie plate assembly is locked to the guide tube end fittings a nd can be unlocked for re constitution or for fu el examination using special tools.A stainless steel sleeve assembly with a chrome plated inside diameter is inse rted in the top end of the guide tube assembly. This sleeve protects th e guide tube from control rod fretting and wear when the rod is in the withdrawn/ready position. The sleeve is mechanically captured by the upper tie plate.
Fuel rod pitch and position is maintained by nine spacer grids. The spacers are axially positioned so that the assemblies will be compatible with existing fuel assemblies.
The bi-metallic spacers used in Reloads M through S (Cyc les 10-14) are formed by an interlocking rectangular grid of Zircaloy-4 structural strips (see Figure 3.3-4A). Inconel-718 spring strips are mechanically secured within th ese strips. The Zircaloy-4 structural strips are welded at all intersections and to the side plates. Dimples formed in the structural strips center the MPS2 UFSAR3.3-9Rev. 35 fuel rod within the cell and along with the springs provide a positive but compliant support for each rod, sufficient to prevent fretting vibration.
In Reloads M, N, and P (Cycles 10-12), the debris resistant Incone l HTP spacer grid in the ninth, bottom location is located just above the lower tie plate. It is formed by an interlocking rectangular grid of Inconel-718 stri ps. The strips are welded at all intersections and to the side plates. The spacer is positioned on top of the lower tie plate with the strip intersections directly above the tie plate flow holes. This reduces the size of debris that may pass through the flow holes thereby reducing the possibility of fretting against the cladding.
Reloads R and S (Cycles 13 and
- 14) use a similar debris resistant concept with the Inconel HTP spacer replaced by a bimetallic spacer coupled with a longer lo wer end cap on the fuel rods. The HTP spacers for Reloads T through X (Cycles 15-18) are all Zircaloy-4 (Figure 3.3-4B). The strips are welded at the inters ections and side plates. The stru cture of the Zircaloy-4 strips provides the rod support.
In Reload Y (Cycle 19) and later, all Zircaloy-4 HTP spacers ar e used in eight locations. The Inconel-718 HMP spacer is used in the ninth, bottom location. The Inconel-718 HMP bottom spacer is similar in design to the HTP spacers except for the flow channels, which are not canted.
The Millstone Unit 2 Cycle 19 reload core included four fuel assemblies with natural uranium replacement fuel rods with an anti-rotation feature designed to pr event spinning of the rod during operations. The four assemblies containing replacement rods were installed in symmetric, peripheral core locations against the baffle as shown in Figure 3.3-19 (Reference 3.3-9). The core locations into which the assemblies were placed where P-1, A-8, H-21, and Y-14 (see Figure 3.4-1). The replacement rods installed unde r these conditions were evaluated against established mechanical, nuclear, and thermal/hydraulic design cr iteria for Millstone Unit 2 fuel, and were determined to be compliant with their design and licensing bases (Reference 3.3-10).
3.3.1.3.2 Fuel Assembly Mechanical Criteria The structural integrity of the fu el assemblies is assured by set ting design limits on stresses and deformations due to various handl ing operational and accident loads
. These limits are applied to the design and evaluation of upper and lower tie plates, grid spacers, guide tubes, holddown springs, and locking hardware.
The design bases for evaluating the structural integrity of the fuel assemblies are:
a.Fuel Assembly Handling - Dynamic axial loads appr oximately 2.5 times assembly weight.b.For All Applied Loads for Normal Operat ion and Anticipated Operational Events -The fuel assembly component structural design criteria are established for the two primary material categories, austenitic stainless steels (tie plates), and Zircaloy (guide tubes, grids, spacer sleeves). Th e stress categories and strength theory for MPS2 UFSAR3.3-10Rev. 35austenitic stainless steel presented in the ASME Boiler and Pressure Ve ssel Code,Section III (Reference 3.3-1) are used as a general guide.Steady state stress limi ts are given in FSAR Section 3.3.1.1.1. Stress nomenclature is per the ASME Boiler and Pressure Vessel Code,Section III.c.Loads During Postulated Accidents - Deflection or failure of components shall not interfere with reactor shutdown or emer gency cooling of the fuel rods during postulated seismic and loss of cool ant accident (LOCA) occurrences.
The assembly structural component stress es under faulted condi tions are evaluated using primarily the methods outlined in Appendix F of the ASME Boiler and Pressure Ve ssel Code,Section III.The design basis for the guide tube wear sleeves is that the fuel assembly shall not be damaged by CEA induced fretting-wear. Flow tests at reactor conditions of prototypic fuel and guide tube wear sleeve assemblies have been used in establ ishing the performance of the CEA wear sleeve combination.
The holddown springs, as compressed by the upper core plate during re actor operation, shall provide a net positive downward force during st eady state operation, based on the most adverse combination of compone nt dimensional and ma terial property tolerances. In addition, the holddown springs are designed to accommodate the additional lo ad associated with a pump overspeed transient (re sulting in possible temporary liftoff of the fuel asse mblies), and to continue to ensure fuel assembly holddown following such an occurrence.
The fuel assembly growth pl us BOL length shall not exceed the minimum space between the upper and lower core plates in the reactor cold condition (70
°F). The reactor cold condition is limiting since the expansion coefficient of the stainless steel core barrel is greater than the coefficient of expansion of the Zircaloy guide tubes.
The spacer assembly is designed to withstand the thermal and irradiation induced differential expansion between the fuel r ods and guide tubes and to wi thstand the design handling and accident loads discussed above. Th e debris resistant Inconel-718 HTP spacer used in the ninth, bottom location for reloads M, N a nd P (Cycles 10-12) was positioned such that the internal strip intersections are directly above the lower tie plate flow holes, t hus reducing the size of debris which could pass through the lower tie plate.
In Reloads R and S (Cycles 13 and 14), the In conel-718 HTP spacer grid at the ninth, bottom location was replaced with a bimetallic spacer which is raised off the upper surface of the lower tie plate. The gap between the upper surface of the lower tie plate and the lower surface of the bimetallic spacer is spanned by a long fuel rod end cap of solid Zircaloy-4.
The Zircaloy-4 HTP spacer grid is used in all nine locations in Reloads T through X (Cycles 15-18). This design is typically referred to as the 'HTP Fuel Assembly'. This spacer grid design MPS2 UFSAR3.3-11Rev. 35 provides improved DNB performanc e, structural strength, and fret ting resistance. The long fuel rod end cap is maintained in the HTP Fuel Assembly.In Reload Y (Cycle 19) and later, the Zircaloy-4 HTP spacer grid is used in eight locations and an Inconel HMP spacer grid is used in the ninth, bottom location. This design retains the long fuel rod lower end cap and is typically referred to as the 'HTP
+HMP Fuel Assembly'. The HTP+HMP design has improved structural strength, and fretting resistance compared to the HTP design.
The design analysis is based upon reactor operating conditions. Typically, these conditions are:
Nominal Core Thermal Power = 2700 MW Nominal Coolant Pressure = 2250 psia
Maximum Flow for Fuel Assembly Liftoff = 422,466 gallons per minute (at 480
°F) Maximum Core Coolant Inlet Temp erature at Nominal Power = 549
°F Total Average Linear Power = 6.206 kW/ft The power histories used in the design analysis are designed to achiev e a peak assembly burnup of 56,000 MWD/MTU for Batch N, 52,500 MDW/MTU for Batch P, and 57,400 MDW/MTU for Batch R and later.
Conservative rod local peaking factors are used which result in a peak rod burnup of 62,000 MWd/MTU. Each of the rod design histories foll ows the single hottest r od in the first cycle operation, the hottest rod in second cycle operation, etc.
Fuel assembly components must be able to wi thstand anticipated seis mic and LOCA forces.
These may result from bundle vibrati on and impact due to a seismic or LOCA event. An analysis was performed for the previous reloads to determine the maxi mum bundle displacements and the maximum spacer grid forces expected during postulated accidents for Mill stone 2. The loads and displacements analysis, which was performed by CE (Reference 3.3-6), considered the safe shutdown earthquake (SSE) and limiting Branch Line LOCA events, a nd the dynamic properties of the AREVA reload fuel assemblies. The re sulting fuel assembly displacements and the combined seismic and LOCA grid spacer impact forces were provided to AREVA.The loads and displacements were conservatively adjusted for the Batch R design due to the optimization of the fuel rod. The fuel weight was increased and the assembly stiffness was decreased. The spacer impact loads and the fuel assembly maximu m deflections were conservatively recalculated from the reference analysis values. The spacer strength margin, the guide tube stresses, and the fuel rod stresses were calculated for the adjusted loads.
Calculated stresses at the appropria te deflections were combined wi th the steady state stresses and compared with the ASME Design Criteria for faulted conditions. This limit is 0.7 times the ultimate strength for the primary stress combinati ons as compared to 0.5 times ultimate for steady state loadings. This criteria was met fo r both the fuel rods and the guide tubes.
MPS2 UFSAR3.3-12Rev. 35 The calculated grid spacer load s during each accident and the combined loads were compared with the allowable grid spacer strength at operating temperature. The loads evaluated were the maximum projected one-sided impact load and the maximum through grid load. The maximum allowable crushing load is the 95 percent confidence lower limi t of the true mean of the distribution of crush test measur ements. The allowable through gr id strength is well above the maximum through grid load. It is also above the maximum one-sid ed impact load. For Reload R and beyond, the seismic/LOCA calculations were reviewed and determined to be bounding.
3.3.1.4 Fuel Assembly Holddown Device A fuel assembly holddown device ha s been incorporated to preven t the possibility of lifting the fuel assembly by hydraulic forces under all normal flow conditions with temperature greater than 500°F. The holddown device consists of a spring-load ed plate which is integral to the fuel assembly. The springs are compre ssed as the upper guide structure is lowered into place. The added spring load, together with the weight of the fuel assembly, prevents possible axial motion of the fuel assembly during operating conditions.
The holddown device is incorporat ed into the upper end fitting a nd features a movable holddown plate which acts on the underside of the fuel alignment plate (ref er to Figure 3.3-5). The movable plate is loaded by coil springs which are loca ted around the upper end fitting posts. The springs are positioned at the upper end of the assembly so that the spring load combines with the assembly weight in counteracting the upward hydr aulic forces. The spri ngs are sized and the spring preload selected, such th at a net downward force will be maintained for all normal and anticipated transient flow and te mperature conditions. It should be noted that the movable plate also serves as the lifting surface during handling of the fuel assembly.
Assembly holddown was previously analyzed in Section 3.6.1 of Reference 3.3-8. The analysis has been reperformed for Batch T a nd beyond fuel and is conservative.
3.3.1.5 Control Element Assembly CEAs are provided by Combustion Engineering (CE) and AREVA. The CEA (shown in Figure 3.3-6) is comprised of five Inconel tubes 0.948 inch in diameter. All tube s contain neutron poison materials with the distri bution of the poison materials as depicted in Figure 3.3-7. Each tube is sealed by welded end caps. A gas expa nsion space is provided to limit maximum tube stress due to internal pressure developed by th e release of gas and mo isture from the boron carbide. The overall length of the CEA is provided in Table 3.3-1. Four tubes are assembled in a square array around the centrally lo cated fifth tube. The tubes are we lded to an upper end fitting.
The upper end fittings are attached to a spider hub wh ich couples the CEA to the drive mechanism through the extension shaft.
Mechanical reactivity control is achieved by operational maneuvering of groups of single CEAs.
The dual CEA is made up of two si ngle CEAs connected to separate grippers attached to single extension shaft. The arrangement of the CEAs in the core is shown in Figures 3.3-8 and 3.3-9.
MPS2 UFSAR3.3-13Rev. 35 There are 49 single CEAs and 12 dual CEAs all operated by a tota l of 61 CEDMs. Considering the 12 dual CEAs as 24 single CEAs gives an overall number of 73 CEAs in the core.A buffer (deceleration dashpot) system is used for slowing down the CEAs at the end of a reactor trip. The buffering action is accomplished by guide tubes which have a reduc ed diameter in the lower section. When the tip of a CEA falls into the buffer regi on, the pressure buil dup in the lower guide tube supplies the force to slow down the CE A. The velocity is decreased to a level which will minimize impact. The final impact is furthe r cushioned by a coil spring arrangement mounted around the center CEA finger.
The four outer guide tube s have the reduced diamet er lower section (dashpot
). There is no dashpot in the center guide tube. There ar e four bleed holes above the da shpot region for the four outer guide tubes. For the center guide tube, these four bleed holes ar e at a lower elevation. For all guide tubes, there is a small drain hole at the bottom. The CEA tip is filled with a Silver-Indium-Cadmium alloy. This replaces the B 4C to avoid the change of buffer characteristics that B 4 C radiation-induced swel ling might bring about.
The design parameters have been optimized to establish the best combination of buffer stroke and buffer annulus. A significant analytical effort has shown that the pressure buildup and the impact loads are not damaging to the syst em. In addition, a test program has confirme d the feasibility of the system. It has demonstrated that the buffer will work under the worst expected tolerance condition.
3.3.1.6 Neutron Source Design For Cycle 18 and beyond, the reactor core will not utilize neutron s ources. It has been determined that during startups without neutron sources, there will continue to be a suff icient ne utron count rate at each of the four Wide Range (WR) Excore fission dete ctors due to the high burnup fuel assemblies that will be positioned on the core periphery. For Cycle 17 and earlier, four neut ron sources were installed in the reactor core. They were held in vacant CEA guide tubes by mean s of an externally loaded sp ring reacting between the upper fuel alignment plate and the top of the fuel assembly. The cladding of the neutron source rods is of a free standing design. The internal pressure is always less th an reactor operating pressure.
Internal gaps and clearances are provided to allow for differenti al expansion between the source material and cladding.
3.3.1.7 In-Core InstrumentsThe in-core instruments (refer to Section 7.5.4) are located in the in-core instrumentation assembly (Figure 3.3-10). The in-c ore instrumentated thimble su pport frame and gui de tubes are supported by the upper guide structure (UGS) assembly. The tubes are condui ts which protect the in-core instruments and guide th em during removal and inserti on operations. The thimble support frame supports the 43 in-core thimble assemblies and acts as an elevat or to lift the thimbles from the core into the UGS during the refueling operation.
MPS2 UFSAR3.3-14Rev. 35 3.3.1.8 Heated Junction Thermocouples The heated junction thermocoupl e (HJTC) system is composed of two channels of HJTC instruments. Each HJ TC instrument channel is manufactured into a probe assembly consisting of eight HJTC sensors, a seal plug, and electrical connectors (Figure 7.5-6). The eight HJTC sensors are physically independent and loca ted at eight levels from the reactor vessel head to the fuel alignment plate.
The probe assembly is housed in a stainless steel suppor t tube structure that protects the sensors from flow loads and serves as the guide path for the sensors. Figure 3.3-18 describes the locations of the HJTC probe assemblies.HJTC Probes and Support Tubes in Upper Guide Structure The HJTC probes and support t ubes are installed inside two-part length CEA shrouds which protect the support tubes from norma l operating cross-flow loads as well as blowdown loads. The support tubes are latched to the bottom of th e CEA shroud and permanently tensioned by means of a threaded spanner nut at the top. Operating lo ads are far less than the preload developed by the tensioning operation. Therefore, the support tubes will not be affect ed by thermal or flow loads.
The support tubes are designed to account for all tolerance conditi ons so that proper clearances will be assured. Physically, the support tubes are similar in mass and size to a typical control element assembly drive shaft, wh ich would reside in the same ar ea of the upper guide structure.
The presence or absence of the HJTC probes within the support tubes will in no way affect the integrity of the support tubes, the UGS, the pressure boundary, and will have no significant effect upon the hydraulic conditions within the reactor vessel head.
3.3.2 REACTOR
INTERNAL STRUCTURES The reactor internals are designe d to support and orient the reac tor core fuel assemblies and CEAs, absorb the CEA dynamic loads and transmit these and other loads to the reactor vessel flange, provide flow paths for the reactor coolant, and guide in-core instrumentation.
The internals are designed to sa fely perform their f unction during all steady state conditions and during normal operating transients.
The internals are designed to sa fely withstand the forces due to deadweight, handling, system pressure, flow impingement, temperature diff erential, vibration and seismic acceleration. All reac tor components are considered Class 1 for seismic design. The reactor internals design limits de flection where required by function. In most cases the design of reactor internals components is limited by stress, not deflection.
For the CEA shroud which is the most limiting internal component for deflection, the allowable de sign deflection limit is 0.5 inch.
This limit is two-thirds of the conservatively established loss-of-function deformation limit, 0.75 inch and applies to a break whose equivalent diameter is no larger than the largest line connected to the primary coolan t line. The structural com ponents satisfy stress values given in Section III of the ASME Pressure Vessel Code. Certain component s have been subjected to a fatigue analysis. Where appropriate, the effect of neutron irradiation on the materials concerned is included in the design evaluation.
MPS2 UFSAR3.3-15Rev. 35The components of the reactor internals are divided into four major parts consisting of the core support barrel, the lower core support structure (including the core shroud), the UGS (including the CEA shrouds, the in-core instrumentation guide tubes and the HJTC s upport tubes). The flow skirt, although functioning as an integral part of the coolant flow path is separate from the internals and is affixed to the bottom head of th e pressure vessel. These components are shown in Figure 3.1-1 and 3.3-11. The in-core instrume ntation is describe d in Section 7.5.4.
Dynamic system analysis methods and procedures which have be en used to determine dynamic responses of reactor internals have been pr ovided in CE, Report CENPD-42, "Topical Report of Dynamic Analysis of Reactor Vessel Internals under Loss-of-Coolant Acci dent Conditions with Application of Analysis to CE 800 MWe Class Reactors".
3.3.2.1 Core Support Assembly The major support member of the reactor internals is the core support assembly. This assembled structure consists of the core support barrel, the lowe r support structure, and the core shroud. The major materials for the assembly is T ype 304 stainless steel.
The core support assembly is s upported at its upper end by the uppe r flange of the core support barrel which rests on a ledge in the reactor vessel flange.
The lower flange of the core support barrel supports and positi ons the lower support structure.
The lower support structure pr ovides support for the core by means of a core support plate supported by columns resting on beam assemblies. The core s upport plate provides support and orientation for the fuel assemblies. The core shroud which provides lateral support for the fuel assemblies is also supported by th e core support plate. The lower end attaches the core barrel to the pressure vessel.
3.3.2.2 Core Support Barrel The core support barrel is a right circular cylinder with a nominal inside diameter of 148 inches and a minimum wall thickness of 1.75 inch. It is suspended by a 4 in ch thick flange from a ledge on the pressure vessel. The core support barrel, in turn, supports the lower support structure upon which the fuel assemblies rest. Pr ess fitted into the flange of the core support barrel are four alignment keys located 90 degrees apart. The reactor vessel, closure head and upper guide structure assembly flanges are sl otted in locations corresponding to the alignment key locations to provide proper alignment between these components in the vessel flange region.
Since the core support barrel is over 27 feet long and is supporte d only at its upper end, it is possible that coolant flow could induce vibrations in th e structure. Therefor e, amplitude limiting devices, or snubbers are installe d on the outside of the core s upport barrel near the bottom end.
The snubbers consist of six equally spaced double lugs ar ound the circumference and are the grooves of a "tongue-and groove" asse mbly; the pressure vessel l ugs are the tongues. Minimizing the clearance between the two mating pieces limits the amplitude of a ny vibration. During assembly, as the internals are lowered into the ve ssel, the pressure vessel tongues engage the core support grooves in an axial direction. With this design, the intern als may be viewed as a beam MPS2 UFSAR3.3-16Rev. 35 with supports at the furthest extr emities. Radial and ax ial expansion of the co re support barrel are accommodated, but lateral movement of the core support barrel is restri cted by this design. The pressure vessel tongues have bolted, lock welded Inconel X shims and th e core support barrel grooves are hardfaced with Stellite to minimize wear. The snubber as sembly is shown in Figure 3.3-12.
3.3.2.3 Core Support Plate and Support Columns The core support plate is a 147 inch diameter, 2 inch thick, T ype 304 stainless steel plate into which the necessary flow distributor holes for the fuel assemblies have been machined. Fuel assembly locating pins (four fo r each assembly) are shrunk-fit into this plate. Columns and support beams are located between this plate and the bottom of the core support barrel in order to provide support for this plate and transmit the core load to the bottom flange of the core support barrel.3.3.2.4 Core Shroud The core shroud provides an envel ope for the core and limits the amount of coolant bypass flow.
The shroud (Figure 3.3-13) consists of two Type 304 stainless st eel ring sections, aligned by means of radial shear pins and attached to the core support plate by Type 348 stainless steel tie rods. A gap is maintained between the core shr oud outer perimeter and the core support barrel in order to provide some coolant flow upward between the core sh roud and core support barrel, thereby minimizing thermal stresses in the core shroud and eliminating stagnant pockets.
3.3.2.5 Flow Skirt The Inconel flow skirt is a right circular cylinder, perforated wi th 2-1 1/16 inch diameter holes, and reinforced at the top and bottom with stif fening rings. The flow sk irt is used to reduce inequalities in core inlet flow distributions and to prevent formation of large vortices in the lower plenum. The skirt provides a nearly equalized pressure distri bution across the bottom of the core support barrel. The skirt is suppor ted by nine equally spaced machin ed sections which are welded to the bottom of the pressure vessel.
3.3.2.6 Upper Guide Structure AssemblyThis assembly (Figure 3.3-14) consists of the upper support structure, 69 CEA shrouds, a fuel assembly alignment plate and an expansion compensating ring. The UGS assembly aligns and laterally supports the upper end of the fuel assemb lies, maintains the CEA spacing, prevents fuel assemblies from being lifted out of position during a severe acci dent condition and protects the CEAs from the effect of coolan t crossflow in the upper plenum.
The UGS is handled as one unit during installation and refueling.
The upper end of the assembly is a structure consis ting of a support plate welded to a grid array of 24 inch deep beams and a 24 inch deep cylinder which encl oses and is welded to the ends of the beams. The periphery of the pl ate contains four accurately ma chined and located alignment keyways, equally spaced at 90 degree intervals, which engage the core barrel alignment keys. The MPS2 UFSAR3.3-17Rev. 35 reactor vessel closure head flange is slotted to engage the upper ends of the alignment keys in the core barrel. This system of keys and slots provides an accurate me ans of aligning the core with the closure head. The grid aligns and supports the upper end of CEA shrouds.
The CEA shrouds extend from the fuel assembly al ignment plate to an elev ation about three feet above the UGS support plate. There are 57 single-type shrouds. These cons ist of cylindrical upper sections welded to integral bot tom sections, which are shaped to provide flow passages for the coolant passing through the alignment plate while shrouding the CEAs from cross-flow. There are also 12 dual-type shrouds which in configuration consist of two single-type shrouds connected by a rectangular secti on shaped to accommodate the dual CEAs. The shrouds are bolted to the fuel assembly alignment plate. At the UGS support plat e, the single shrouds are connected to the plate by spanner nuts which permit axia l adjustment. The spanner nuts ar e tightened to proper torque and lockwelded. The dual shrouds are at tached to the upper plate by welding.
The fuel assembly alignment plat e is designed to align the upper ends of the fuel assemblies and to support and align the lower ends of the CEA shrouds.Precision machined and located holes in the fuel assembly alignment plate align the fuel assemblies. The fuel assembly ali gnment plate also has four equally spaced slots on its outer edge which engage with Stellite hard faced pins protruding from the co re shroud to limit lateral motion of the UGS assembly during operation. The fuel alignment plate bears th e upward force of the fuel assembly holddown devices. Th is force is transmitted from the alignment plate through the CEA shrouds to the UGS support plate and hence to the expansion compensating ring.
The expansion compensating ring bear s on the flange at the top of the assembly to accommodate axial differential thermal expans ion between the core barrel flange, UGS flange and pressure vessel flange support edge and head flange recess.
The UGS assembly also supports the in-core inst rumentation thimble suppor t frame, guide tubes, and HJTC support tubes.
All integral connections in the reactor internals are designed within the stress intensity limits listed in Tables N-422 and N-416.1 of Section III of the ASME code for normal and upset conditions. For emergency and faulted conditions, the design limits are as given in Table 3.2-1.
3.3.3 CONTROL
ELEMENT DRIVE MECHANISM 3.3.3.1 Design The CEDM is of the magnetic jack type drive. Each CEDM is cap able of withdrawing, inserting, holding or tripping the CEA from any point within its 137-inch st roke. The design of the CEDM is shown in Figure 3.3-15 and is identical to th at for Maine Y ankee (A EC Docket Number 50-309) and Calvert Cliffs Units 1 and 2 (A EC Docket Numbers. 50-317 and 50-318).
The CEDM drives the CEA within the reactor co re and indicates the position of the CEA with respect to the core. The speed at which the CEA is inserted or withdrawn from the core is MPS2 UFSAR3.3-18Rev. 35 consistent with the reactivity change requireme nts during reactor operati on. For conditions that require a rapid shutdown of the reactor, the CEDM coils of the shutdown and regulating CEAs are deenergized, allowing the CEA and the supporting CEDM components to drop into the core by gravity. The CEA drop time is 2.75 seconds, where drop time is de fined as the interval between the time power is removed from the CEDM coils and the time th e CEA has reached 90 percent of its fully inserted position. The reactivity is reduced during such a drop at a rate sufficient to control the core under any opera ting transient or accident conditi on. The CEA accelerates to about 11 ft/sec and is decelerated at the end of the drop by the buffer se ction of the CEA guide tubes. Drive down capability following a reactor trip is not required for safe ty purposes. The safety analyses of Chapter 14 assume th e CEA of highest reacti vity worth sticks in the fully withdrawn position. A drive down feature would introduce the pos sibility of a failure which would prevent power from being remove d from the CEDMs during a trip, wh ich would lead to a reduction in plant safety.There are 69 CEDM nozzles on top of the reactor vessel closure head. Eight of the 69 nozzles were used for the part length CEAs in Cycle 1, six of which are no longer used, and two of which are used for HJTC/RVLMS instru mentation. There are 61 CEDMs in current use. The six spare nozzles are capped with adapters. Each CEDM is connected to a CEA by a locked coupling. The weight of the CEAs and CEDMs is carried by the vessel head.
The CEDM is designed to handl e dual, single or part lengt h CEAs. The maximum operating speed capability of the CEDMs is 40 inches per minut e for single CEAs and 20 inches per minute for dual CEAs.
3.3.3.2 Control Element Drive Mechanism Pressure HousingThe CEDM housing is attached to the reactor vessel head nozzle by means of a threaded joint and seal welded. The CEDM nozzles are made of Inconel Alloy 690 to minimize Primary Water Stress Corrosion Cracking. The CEDM pressure housings including the ma gnetic coil ja ck assemblies were replaced as part of the replacement reactor vessel closure head project.
The CEDM upper housing design and fabrication conform to the requirements of the ASME Boiler and Pressure Vessel Code,Section III, 1998 Edition th rough 2000 Addenda. The housing is designed for steady state conditions as well as all anticipated pr essure and thermal transients. Once the CEDM housing is seal welded to th e head nozzle, it need no t be removed since all servicing of the CEDM is perf ormed from the top of the CEDM housing. This opening is closed by means of an upper housing and an omega seal weld. The CEDM pr essure housing is capable of being vented after major coolant refills of the reactor coolant system (RCS), such as after a refueling and after react or coolant pump (RCP) maintenance. However, venting of the CEDM pressure housing is no longer necessary after majo r refills of the Reactor Coolant System (RCS), since a vacuum refill me thod is used. The vacuum refill process involves a partial vacuum in the RCS while at mid-loop level and then slowly refilling the RCS.
MPS2 UFSAR3.3-19Rev. 35 3.3.3.2.1 Heated Junction Thermoc ouple Pressure BoundaryThe HJTC probe assemblies are located at the two original locations (CEDMs 11 and 13) on the replacement reactor vess el closure head. The HJTC pressure boundary also known as the Reactor Vessel Level Monitoring System (RVLMS) pressure housing assembly consists of upper pressure housing tube, upper flange type Grayloc connection and lowe r housing. The lower housing is joined to the reactor vessel head nozzle by means of a threaded joint and an omega seal weld. The pressure boundary at the top of the RVLMS pressu re housing is maintained by a quick disconnect Grayloc type flange (See Figure 3.3-17). The components are designed to ASME Section III, B&PV Code 1998 Edition through 2000 Addenda.
The pressure and thermal loads associated with normal operation and tr ansient conditions have been included in stress analyses performed in accordance with ASME BPVC criteri
- a. All stresses are within allowable limits.
3.3.3.3 Magnetic Jack Assembly The magnetic jack motor assembly is an integr al unit which fi ts into the CEDM housing through an opening in the top of the hous ing. This unit carries the motor tube, lift and hold pawls and magnets. The drive power is supplied by electri cal coils positioned around the CEDM housing.
The CEDMs are cooled by ai r supplied at 900 CFM at 95
°F (maximum) to each CEDM. The design of the control element drive mechanism is such that loss of cooling air will not prevent the CEDM from releasing the CEA. Th e ability of the CEDM to release the rods is not dependent on the cooling flow provided by th e CEDM Cooling System. Cooling function is only to ensure reliability of the CEDM coil stack. Following insertion of the CEDM motor assembly, the upper pressure housing is threaded into the CEDM motor housing and seal welded. This upper pressure housing encloses the CEDM extension shaft and supports the shroud assembly. The reed switch assembly is supported by the shroud assembly.
The lifting operation consis ts of magnetically operated step movements. Two sets of mechanical latches (one holding, one lifting) are utilized engaging a notched drive shaft. To prevent excessive latch wear, a means has been provided to unload th e lifting latches during the engaging and disengaging operations.
The magnetic force is obtained from large DC magnet coils m ounted on the outside of the motor tube.
Power for the electromagnets is obtained from one of two separate supplies. A control programmer actuates the stepping cycle and obtains the CEA lo cation by a forward or reverse stepping sequence. CEDM hold for shutdown and regulating CEAs is obtained by energizing a hold coil at a reduced current while all other coils are deenergized. The full length CEAs are tripped upon interruption of elec trical power to all coils.
3.3.3.4 Position IndicationThree separate means are provided for tr ansmitting CEA position indication.
MPS2 UFSAR3.3-20Rev. 35 The first method utilizes the el ectrical pulses from the magnetic coil power programmer. The second method utilizes reed swit ches and a voltage divider ne twork mounted on the CEDM to provide an output voltage proportional to CEA position. The thir d method utilizes three pairs of reed switches spaced at discre te locations within a position transmitter assembly. A permanent magnet built into the drive shaft actuates the reed switches one at a time as it passes by them. CEA position instrumentation is discu ssed in detail in Section 7.5.3.
3.3.3.5 Control Element Assembly Disconnect The CEA is connected to the drive shaft extensio n with an internal colle t-type coupling at its lower end. (Coupling is performed be fore the vessel head is installe d). In order to disengage the CEA from the drive shaft extension, a tool is att ached to the top end of the drive shaft when the reactor vessel head has been removed.
By pulling up on the spring-loaded operating rod in th e center of the drive shaft, a tapered plunger is withdrawn from the center of the collet-type gripper causing it to collapse due to axial pressure from the CEA, thus permitting removal of the c oupler from the CEA. Releasing the operating rod plunger after the coupler has been withdrawn from the CEA expands the coupler to a diameter that prevents recoupling to the CEA.
3.3.3.6 Test Program A test program has been conducte d to verify the adequacy of the magnetic jack CEDM. The program is described in Section 1.5.4.
3.
3.4 REFERENCES
3.3-1ASME Boiler and Pressure Vessel Code,Section III, 1977 Edition, ASME New York, NY.3.3-2K. R. Merckx, "RODEX2 - Fuel Rod Th ermal-Mechanical Response Evaluation Model," XN-NF-81-58 (NP)(A), Revi sion 2, March 1985 and Supplements. 3.3-3"Qualification of Exxon Nuclear Fuel for Extended Burnup (PWR)," XN-NF-82-06 (NP)(A), Revision 1, Suppl ements 2, 4, 5, October 1985.3.3-4W. J. O'Donnel and B. F. Langer, "Fat igue Design Bases for Zircaloy Components,"
Nuclear Science and Engineering, Volume 20, January 1964.
3.3-5MATPRO Version, "A Handbook of Material Properties for Use in the Analysis of Light Water Reactor Fuel Rod Behavior," TREE-NUREG 1008, December 1976.3.3-6J. C. Winslow (CE) to T. J. Honan (NU), CE Letter, "Seismic and Branch Line LOCA Analysis of SPC Reload Fuel for Millstone 2," NU-88-043 (March 31, 1988).
MPS2 UFSAR3.3-21Rev. 353.3-7"PWR Primary Water Chemistry Guidelines
," Revision 2, Electric Power Research Institute (EPRI) Final Report, EPRI NP7077, dated November 1990.3.3-8ANF-88-88(P), Rev. 1, "Design Report for Millstone Point Unit 2 Reload ANF-1,"
August 29, 1988.3.3-9AREVA Contract Requirements Document Number 89-9070921-001-AREVA Contract No. J37MIL219B, January 28, 2008.
3.3-10AREVA Document 51-9074000-000, "Complianc e Document - Replacement Fuel Rod -
Millstone 2 Fuel Failure Mitigation," March 5, 2008.
MPS2 UFSAR3.3-22Rev. 35TABLE 3.3-1 MECHANICAL DESIGN PARAMETERS
- Fuel Assembly Geometry 14 by 14 Assembly Pitch, inches 8.180 Assembly Envelope, inches 8.160 Rod Pitch, inches 0.580Number of Grids per Assembly 9 Approximate Assembly Weight, lb.
1280/1313
- Fuel Rod to Fuel Rod Outside Dimension, inches 7.980 Fuel Rod and Pellet Clad OD, inches 0.440Clad thickness, inches 0.031/0.028
- Pellet Diameter, inches 0.3700/0.3770
- Pellet Length, inches 0.425/0.435 *Pellet Density (% Theoretical) 94.0/95.0/95.35
- Active Stack Length, Cold, inches 136.7Control Rod Guide Tube Number per assembly 4Tube ID, above dashpot, inches 1.035 Wall Thickness, inches 0.040Instrumentation Tube Number per Assembly 1 Tube ID, inches 1.035 Wall Thickness, inches 0.040 Spacer Grid Material Zircaloy-4 / Inconel-718 MPS2 UFSAR3.3-23Rev. 35Number per Assembly 8/1 for Batches N, P 9/0 for Batch R, S9/0 for Batch T - X
8/1 Batch Y and beyond Sleeves (Wear)
Material SS/Chrome Plate Burnable Poison Rod Active Length, inches 124.7 + UO 2 blankets Material Gd 2 O 3 / U0 2 Pellet Diameter, inches 0.3700/0.3770 Clad Material Zircaloy-4 Clad ID, inches 0.378/0.384 Clad OD, inches 0.440 Clad Thickness, (nominal) inches 0.031/0.028Diametral Gap, (cold, nominal), inches 0.008/0.007
Pellet Length, inches 0.545Control Element Assembly Number 73 Number of Absorber Elements per Assembly 5 Type Cylindrical Rods Clad Material Inconel 625 Clad Thickness, inches 0.036 Clad OD, inches 0.948 Poison Material B 4C & Ag-In-CD Corner Element Pitch, inches 4.64Total CEA Length, inches 161.31- CE / 161.25 - AREVA Poison Length, inches 132 -CE / 133.5 - AREVACEA Dry Weight, lb. 95 - CE / 85 - AREVA MPS2 UFSAR3.3-24Rev. 35Total Operating Assembly Dry Weight, lb. Single 210 - CE / 200 - AREVADual 334 - CE / 314 AREVACore Arrangement Number of Fuel Assemblies in Core Total 217Number of Single CEAs 49 Number of Dual CEAs 12 CEA Pitch, minimum, inches 11.57 Spacing Between Fuel Assemblies, Fuel Rod Surface to Surface, inches0.200
Spacing, Outer Fuel Rod Surface to Core Shroud, inches0.18 Hydraulic Diameter, Nominal Channel, feet 0.04445Total Flow Area (Excluding Guide Tubes), square feet 53.5Total Core Area, square feet 101.1 Core Equivalent Diameter, inches 136 Core Circumscribed Diameter, inches 143.1 Core Volume, liters 32,526 Total Fuel Loading, MTU (Typical) 83.65 Total Heat Transfer Area, square feet 50,117*Applicable to Batches N, P/applicable to Batch R and subsequent Batches.**Applicable to Batches N, P/applicable to Batc hes R, S/applicable to Batch T and subsequent Batches.
MPS2 UFSAR3.3-25Rev. 35TABLE 3.3-2 PRESSURIZED WATER REACTOR PRIMARY COOLANT WATER CHEMISTRY RECOMMENDED SPECIFICATIONS Conductivity (µS/cm at 25
°C) Relative to Lithium and Boron concentration.
pH at 25°C Determined by the concentrati on of boric acid and lithium present. Consistent with th e Primary Chemistry Control Program.(4) Dissolved Oxygen, at power < 0.1 ppm (1) (2) (3)
Chloride < 0.15 ppm Fluoride < 0.10 ppm Hydrogen 25-50 cc (STP)/KgH 2 O Suspended Solids 0.35 ppm prior to reactor startup Li Consistent with the Prim ary Chemistry Control Program.
(4) Boron, as boric acid 0-2620 ppm (5) NOTES:(1)The temperature at which th e Oxygen limit applies is > 250
°F.(2)The at power operation residual Oxyg en concentration control value is 0.005 ppm
.(3)During plant startup, Hydrazine may be used to control disso lved Oxygen concentration at 0.1 ppm.(4)During power operation lithium is c oordinated with boron to maintain a pH (t) of 7.0, but 7.4, consistent with the Primary Chemistry Control Program.
Lithium is added to the RCS during plant startup, but prior to reactor criticality, and is in specification per the Primary Chemistry Control Program within 72 hours8.333333e-4 days <br />0.02 hours <br />1.190476e-4 weeks <br />2.7396e-5 months <br /> after criticality. Lithium may be removed from the reactor coolant immediately before, or dur ing, shutdown periods to aid in the cleanup of corrosion products. By eval uation, a maximum lithium concentration of 4.5 ppm is permissible with a target lithium concentration of 4.3 ppm for 100% power operations.(5)RCS boron concentration is maintained as necessary to ensu re core reactivity or shutdown margin requirements are met. Although the RC S and related auxiliary systems containing reactor coolant are designed for a maximum concentration of 2620 ppm boron, it should be noted the design basis for the TSP baskets in the containment sump assumes the RCS, SITs, and RWST are at a maximu m boron concentration of 2400 ppm.
MPS-2 FSAR April 1998 Rev. 24.8 FIGURE 3.3-1 FUEL ROD ASSEMBLY UPPER END CAP PLENUM SPRING DISHED PELLETS FUEL CLADDING
.440 CLADDING OD
.3770 PELLET DIAMETER
.028 CLADDING WALL 136.70 ACTIVE FUEL LENGTH 136.70 ACTIVE FUEL LENGTH 146.25 MPS-2 FSAR Rev. 24.8FIGURE 3.3-2AAREVA -
RELOAD FUEL ASSEMBLY BATCH "S" AND PRIOR 9X TYPICAL SPACER UPPER TIE PLATE ASSEMBLY LOWER TIE PLATE MPS-2 FSAR Rev. 24.8FIGURE 3.3-2BAREVA - RELOAD FUEL ASSEMBLY BATCH "T" AND LATER UPPER TIE PLATE LOWER TIE PLATE MPS-2 FSAR Rev. 24.8FIGURE 3.3-3AAREVA - RELOAD FUEL ASSEM BLY COMPONENTS BATCH "S" AND PRIOR 9X TYPICAL SPACER UPPER TIE PLATE ASSEMBLY LOWER TIE PLATE MPS-2 FSAR Rev. 24.8FIGURE 3.3-3BAREVA - RELOAD FUEL ASSEM BLY COMPONENTS BATCH "T" AND LATER UPPER TIE PLATE SPACER 136.70 ACTIVE FUEL LENGTH FUEL ROD LOWER TIE PLATE MPS-2 FSAR Rev. 21FIGURE 3.3-4ABI-METALLIC FUEL SPACER ASSEMBLY GUIDE TUBE L OC ATI O N FUEL ROD SPACER SIDEPLATE SPRING STRIP MPS-2 FSAR Rev. 21FIGURE 3.3-4BHTP FUEL SPACER ASSEMBLY GUIDE TUBE FUEL ROD MPS-2 FSAR Rev. 26.2FIGURE 3.3-5FUEL ASSEMBLY HOLD DOWN DEVICE LOCKING NUT UPPER REACTION PLATE FUEL ALIGNMENT
PLATE SPRING UPPER TIE
PLATE MPS-2 FSAR Rev. 30.2FIGURE 3.3-6CONTROL ELEMENT ASSEMBLY MPS-2 FSAR Rev. 30.2FIGURE 3.3-7CONTROL ELEMEN T ASSEMBLY MATERIALS MPS-2 FSAR April 1998 Rev. 26.2FIGURE 3.3-8CONTROL ELEMENT ASSEMBLIES GROUP AND NUMBER DESIGNATION MPS-2 FSARApril 1998Rev. 26.2FIGURE 3.3-9CORE ORIENTATION Outlet Nozzle Alignment Key 4 Equally Spaced Inlet Nozzle See Figure 3.3-8 for Identification of Core
Arrangement and CEA
Groups Fuel Assembly CEDM CEA Building North Reactor Vessel Core Support Barrel Elevation View MPS-2 FSARApril 1998Rev. 26.2FIGURE 3.3-10IN-CORE INSTRUMENTATION ASSEMBLY 90 180 270 0 MPS-2 FSARApril 1990Rev. 26.2FIGURE 3.3-11REACTOR INTERNALS ASSEMBLY Upper Guide Structure Support Plate CEA Shroud In-Core Instrumentation
Guide Tube Core Support
Barrel Core Support
Assembly Snubber Core Support Plate Core Shroud Fuel Aignment
Plate Aignment Pins Outlet Nozzle Alignment Key Expansion
Compensating Ring MPS-2 FSARApril 1990Rev. 24.8FIGURE 3.3-12PRESSURE VESSEL-CORE SUPPORT BARREL SNUBBER ASSEMBLY CENTER SUPPORT BARREL HARD-FACED SURFACE BOLT (12 REQ'D PER ASSEMBLY)CORE STABILIZING
LUG PRESSURE VESSEL SNUBBER SPACER BLOCK SHIM (2 REQ'D PER ASSEMBLY)
PIN (4 REQ'D PER ASSEMBLY)
BOLT (4 REQ'D PER ASSEMBLY)
MPS-2 FSARApril 1990Rev. 24.8FIGURE 3.3-13CORE SHROUD ASSEMBLY MPS-2 FSARApril 1990 Rev. 26.2FIGURE 3.3-14UPPER GUIDE STRUCTURE ASSEMBLY MPS-2 FSARApril 1990 Rev. 26.2 FIGURE 3.3-15 CONTROL ELEMENT DRIVE MECHANISM (MAGNETIC JACK)
MPS-2 FSARApril 1998 Rev. 26.2 FIGURE 3.3-16 (LEFT BLANK INTENTIONALLY)
MPS-2 FSAR Rev. 23.3FIGURE 3.3-17HEATED JUNCTION THERMOCOUPLE PROBE PRESSURE BO UNDARY INSTALLATION MPS-2 FSARApril 1990 Rev. 26.2 FIGURE 3.3-18 T YPICAL HEATED JUNCTION THERMOCOUPLE PR OBE ASSEMBLY INSTALLATION MPS-2 FSAR Rev. 27.4 FIGURE 3.3-19 PLACEMENT OF NATURAL URANIUM REPLACEMENT FUEL RODS AND FUEL ASSEMBLY ORIENTATION RELATIVE TO THE COR E BAFFLE FOR CYCLE 19 MPS2 UFSAR3.4-1Rev. 35
3.4 NUCLEAR
DESIGN AND EVALUATION
3.4.1 GENERAL
SUMMARY
This section summarizes the nuclear characteristics of the core a nd discusses the design parameters which are of significan ce to the performance of the core in normal transient and steady state operational conditions. A discussion of the nuclear design methods employed and comparisons with experiment s which support the use of these methods is included.
The numerical values presented ar e based on a representative core design. Sufficient analyses are completed each cycle to ensure that actual reload batches k eep operating parameters within design limits, accommodate essential reactivity require ments with the cont rol system provided, and meet other requirements for safe operation.
3.4.2 CORE DESCRIPTIONThe Millstone Unit 2 reactor consists of 217 assemblies, each having a 14 by 14 fuel rod array. The assemblies are composed of up to 176 fuel rods, four control rod guide tubes, and one center control rod guide tube/instrum ent tube. The fuel rods consist of slightly enriched UO 2 or UO 2-Gd 2 O 3 pellets inserted into Zircaloy tubes. The control rod guide tubes and instrument tubes are also made of Zircaloy. Each AREVA assembly contains nine spacers. A description of the AREVA supplied fuel design and design methods is contained in References 3.4-1, 3.4-2 and 3.4-3.
A representative loading pattern is shown in Figure 3.4-1 and is expressed in terms of previous cycle core locations and fuel assembly identifi ers. A summary of fuel characteristics for a representative core design is presented in Table 3.4-1. Figure 3.4-2 pres ents representative quarter core assembly movements. Representati ve beginning of cycle (BOC) and end of cycle (EOC) assembly exposures are shown in a quarter core representation in Figure 3.4-3.A representative low radial leakag e fuel management plan results in scatter loading of the fresh fuel throughout the core. Some fresh assemblies loaded in the co re interior contain gadolinia-bearing fuel in order to control power peaki ng and reduce the initial boron concentration to maintain the moderator temperature coefficient (MTC) within its Technical Specification limit.
The exposed fuel is also scatter loaded in th e center in a manner to control the power peaking.
3.4.3 NUCLEAR
CORE DESIGN The nuclear design bases for core design are as follows:a.The design shall permit operation within the Te chnical Specificat ions for Millstone Unit 2 Nuclear Plant.b.The design Cycle length (EFP D) shall be determined on the basis of an estimated Cycle energy and previous Cycle energy window.
MPS2 UFSAR3.4-2Rev. 35c.The loading pattern shall be designed to achieve power distributions and control rod reactivity worths according to the following constraints:1.The peak linear heat rate (LHR) and the peaking factor Fr shall not exceed Technical Specifications limits in a ny single fuel rod throughout the cycle under nominal full power operating conditions.2.The SCRAM worth of all rods minus the most reactive rod s hall exceed the shutdown requirement.
The neutronic design methods used to ensure the above requireme nts are consistent with those described in Reference 3.4-4.
3.4.3.1 Analytical Methodology The neutronics methods us ed in the core analysis are descri bed in Reference 3.4-4. The neutronic design analysis for each reload core is performed using the PRISM reactor simulator code. Full-core depletion calculations perf ormed with PRISM are used to determine the core wide power distribution in three dimensions and to r econstruct the individual rod power and burnup distributions. Thermal-hydraulic f eedback and axial exposure distribution effects are explicitly accounted for in the PRISM cal culations. The CASMO/MICBURN a ssembly depletion model is used to generate the microscopic cr oss section input to the PRISM code.
3.4.3.2 Physics Characteristics The neutronics characteristics of a representative reload core are presented in Table 3.4-2. The safety analysis for each cycl e is applicable for a specified previous cycle energy window. A representative HFP letdown curve is shown in Figure 3.4-4.
3.4.3.2.1 Power Distribution Considerations Representative calculated power maps are shown in Figures 3.4-5 and 3.4-6 for BOC (equilibrium xenon), and EOC conditions, respectively. The power distributions were obtained from a three-dimensional neutronics model with moderator density and Doppler feedback effects incorporated. The Technical Specification limi ts on Fr and LHR are 1.69 and 15.1 kW/ft, respectively.
3.4.3.2.2 Control Rod Reactivity Requirements A representative shutdown margin evaluation is given in T able 3.4-3. The Millstone Unit 2 Technical Specifications require a minimum shutdown margin of 3,600 pcm.
MPS2 UFSAR3.4-3Rev. 35 3.4.3.2.3 Moderator Temperature Coefficient ConsiderationsThe Technical Specifications require that the MTC be less than +7 pcm/
°F at or below 70 percent of rated thermal power, less than +4 pcm/
°F above 70 percent power a nd greater than -32 pcm/
°F at 100 percent of rated thermal power. Representa tive MTC calculation resu lts are presented in Table 3.4-2.
3.4.4 POST-RELOAD STARTUP TESTINGStartup tests will be performed at the beginning of each reload cy cle to obtain the as-built core characteristics and to verify Technical Specification and core physics design parameters. The reload startup physics test progr am is based on ANSI-19.6-1 (Ref erence 3.4-9). The Startup Test Activity Reduction (STAR) Program (Reference 3.4-10) provides an alternative to the ANSI-19.6-1 test program provided that sp ecific criteria for the reload core design and construction are satisfied. The STAR Program criteria are established in station procedures and include additional applicability require ments for core design, fuel and contro l element assembly (CEA) fabrication, CEA lifetime monitoring, refu eling and startup testing.
The reload startup physics test progr am shall consist of the following:a.Critical Boron Concentration - HZP, Control Rods Withdrawn.b.Critical Boron Concentration - HZP, Control Rod Group(s) of at least 1% reactivity are fully inserted in the core.
1c.Control Rod Group Worths - HZP, two or more control rod groups shall be measured which are well dist ributed radially and repres ent a predicted total worth of at least 3% reactivity.
1 d.Isothermal Temperature Coefficient - HZP.e.Flux Symmetry - between 0 and 30% of full power.f.Power Distribution - between 40 and 75% of full power.
g.Isothermal Temperature Coefficient - greater than 70% of full power.
h.Power Distribution - greater than 90% of full power.
i.Critical Boron Concentration - greater than 90% of full power.
j.HZP to full power reactivity difference.
1.This test may be eliminated if performing the STAR Program per Reference 3.4-10.
MPS2 UFSAR3.4-4Rev. 35
3.4.5 REACTOR
STABILITY 3.4.5.1 General Xenon induced spatial oscillations on the Millstone Unit 2 core fall into three classes or modes. These are referred to as axial os cillations, azimuthal oscillations, and radial oscillations. An axial oscillation is one in which the axial power distri bution periodically shifts to the top and bottom of the core. An azimuthal oscillation is one in which the X-Y power distribution periodically shifts from one side of the core to the other. A ra dial oscillation is one in which the X-Y power distribution periodically shifts inward and outward from the center of the core to the periphery.
Xenon stability analyses indicate that a numb er of general statem ents can be made:a.The time scale on which the oscillations occur is long, and a ny induced oscillations typically exhibit a pe riod of 25 to 30 hours3.472222e-4 days <br />0.00833 hours <br />4.960317e-5 weeks <br />1.1415e-5 months <br />.b.As long as the initial power peaking asso ciated with the pert urbation initiating the oscillation is within the limiting conditions for operation, specified acceptable fuel design limits will not be approached for a period of hours allowing an operator time to decide upon and take appropriate remedial action prior to the time when allowable peaking factors would be exceeded.c.The core will be stable to radial mode oscillations at all tim es in the burnup cycle.d.The core will be stable to azimuthal m ode oscillations at all times in the burnup cycle.e.All possible modes of undamped oscillations can be detected by both exactor and in-core instrumentation as discussed below.
3.4.5.2 Detection of Oscillations Primary reliance for the det ection of any xenon oscillations is placed on the exactor flux monitoring instrumentation. The power range exco re neutron detectors (one axial pair per quadrant) are us ed to monitor the symmetry of power distributions and are located at distinct azimuthal and axial positions. These detectors are sensitive primarily to the power density variations produced by peripheral fu el assemblies in the vicinity of the detectors. All possible xenon induced spatial oscillations will affect the power densitie s of the peripheral fuel assemblies in the core.
In addition, the in-core instrumentation provides information which will be used in the early stages of cycle operation to confirm predicted correlations between indications from the excore detectors and the space-depende nt flux distribution within th e core. Later on, during normal operation, the in-core detector syst em provides information which ma y be used to supplement that available from the excore detectors.
MPS2 UFSAR3.4-5Rev. 35 3.4.5.3 Control of Oscillations Since the reactor will not be ope rated under conditions that imply instability with respect to azimuthal xenon oscillation, no sp ecial protective system features are needed to accommodate asimuthal mode oscillations. Regardless, a maximu m azimuthal power tilt is prescribed in the Te chnical Specifications along with prescribed operating re strictions in the event that the azimuthal power tilt limit is exceeded.As described earlier, the power range excore neutron detectors ar e used to monitor the azimuthal symmetry of the power distributions since they are located at distinct locations in the X-Y plane. Should the excore detectors indicate different readings in the azimuthal direct ion, a tilt in the core power distribution would be indicated. When the tilt exceeds a preset magnitude an alarm will occur. In the event of an alarm, the orientation of the tilt will be determined and, on the basis of orientation, the proper CEA's will be manually adjusted to redu ce the magnitude of the tilt.
The features provided for azimuthal xenon osci llation control are:a.instrumentation for monitoring azimuthal power tilt.
b.administrative limits on azimuthal power tilt.The excore detectors are used to monitor the ax ial power distribution and to detect deviations from the equilibrium distribut ion such as those which woul d occur during an axial xenon oscillation. This is done by monito ring variations in the external axial shape index, a parameter derived from the excore detector readings which is related to the axial power distribution. Control of axial xenon oscillation is accomp lished utilizing Regulating Bank 7. When it is determined that the axial shape index may exceed the boundaries of a sp ecified control band about the equilibrium value, this bank is slowly in serted and eventually withdrawn over a period of several hours. The core is then stabilized until a new oscillation develops.
The features provided for axial xenon control and protection are:a.equipment for monitoring axial shape index.
b.administrative limits on axial power distribution, external axial shape index.c.an axial shape index reactor trip (local power density - high).d.use of Regulating Bank 7 for cont rol of axial power distribution.
MPS2 UFSAR3.4-6Rev. 35 3.4.5.4 Operating Experience Recent core designs for Millstone Unit 2 (Cycles 10 and beyond) ha ve been developed to include longer fuel cycles along with low radial leakage fuel management. These current designs scatter load fresh fuel assemblies throughout the interi or of the core with the highest burnup fuel assemblies being loaded along the core periphery. Core designs prior to Cycle 10 operation were not of a low radial leakage desi gn due to the loading of fresh fu el assemblies along the core periphery.With respect to xenon oscillations in the radial and azimuthal direct ions, studies indicate that core designs of a low radial leak age design (i.e., highest burnup a ssemblies loaded on the core periphery with fresh fuel assemblies scatter loaded about the core interior) are more stable than those designs which load fres h fuel assemblies along the core periphery. Therefore, the conclusions regarding xenon oscillations in th e radial and azimuthal directions, which are presented in Section 3.4.5.5, remain app licable to current plant operations.With regard to axial xenon oscillations, the core near end-of-cycle may be naturally unstable in the absence of any control rod action even if low leakage core designs are utilized. But axial xenon oscillations are sufficiently slow (the period of osc illation being 25 to 30 hours3.472222e-4 days <br />0.00833 hours <br />4.960317e-5 weeks <br />1.1415e-5 months <br />) so that there would be sufficient time to control the oscillations. In addition, automatic protection is provided if operator action is not taken to remedy the situati on. Regulating Bank 7 CEA's are utilized for controlling axial xenon oscillations.
3.4.5.5 Method of Analysis The classic method for assessing spatial xenon oscillations is that developed by Randall and St. John (Reference 3.4-5) which c onsists of expanding small pert urbations of the flux and xenon concentrations about equilibrium values in eigenfunctions of th e system with equilibrium xenon present. However, it is necessary to extend this simple linear analysis to treat cores which are nonuniform because of fuel zoning, depletion, and CEA patterns, fo r example. Such extensions have been worked out and are reported in Re ferences 3.4-6 and 3.4-8. In this extension, the eigenvalue separations between th e excited state of interest a nd the fundamental are computed numerically for symmetrical fl ux shapes. For nonsymmetrical flux shapes, the eigenvalue separation can usually be obtained indirectly from the dominance ratio 1/0 , computed during the iteration cycle of the spatial calculation.
Numerical space time calcu lations are performed in the required number of spatial dimensions for the various modes as checkpoints for the predictions for the extended Randall-St.
John treatment described above.
MPS2 UFSAR3.4-7Rev. 35 3.4.5.5.1 Radial Xenon OscillationsTo confirm that the radial oscillation mode is extremely stable, a space-time calculation was run for a reflected, zoned core 1 1 feet in diameter without including the damping effects of the negative power coefficient. The initial perturbati on was a poison worth of 0.4 percent in reactivity placed in the central 20 percent in the core for 1 hour1.157407e-5 days <br />2.777778e-4 hours <br />1.653439e-6 weeks <br />3.805e-7 months <br />. Following removal of the perturbation, the resulting oscillation was fo llowed in 4-hour time steps for a pe riod of 80 hours9.259259e-4 days <br />0.0222 hours <br />1.322751e-4 weeks <br />3.044e-5 months <br />. Results show that the oscillation died out very rapidly with a da mping factor of about minus 0.06 per hour. When this damping coefficient is correct ed for a finite time mesh by the formula in Reference 3.4-7, it is more strongly convergent. On this ba sis, it is concluded that radial oscillation instability will not occur.This conclusion is of particular significance because it means that there is no type of oscillation where the inner portions of the core act independently of the periphe ral portions of the core whose behavior is most closely followe d by the excore flux detectors. Ra dial mode oscillations, even though highly damped, would be mani fested as periodic variation in the excore flux power signal while the delta-T power signals re mained constant. Primary reliance is placed on the excore flux detectors for the detection of any xenon oscillations.
3.4.5.5.2 Azimuthal Xenon Oscillations Analyses indicate that the eige nvalue separation between the first asimut hal harmonic and the fundamental is a bout 0.86 percent in . The calculated damping coefficient for the first azimuthal mode is minus 0.016 per hour, and the higher modes will be even more strongly damped. Furthermore, the Doppler coefficien t applicable to the Mi llstone Unit 2 reactor is calculated to be approximately minus 1.36 x 10
-3 /(kW/ft) which is sufficiently negative to ensure stability of all the azimuthal modes.
3.4.5.5.3 Axial Xenon Oscillations As checkpoints for the predictions for the modified Randall-St. John appr oach, numerical spatial time calculations have been pe rformed for the axial case at bot h beginning and end-of-cycle. The fuel and poison burnup distributions were obtained by depletion with soluble boron control so that the power distribution was strongly flattened. Spatial Doppler fee dback was included in these calculations. The initial perturbati on used to excite the oscillations was a 50 percent insertion into the top of the core of a 1.5 percent reactivity CEA bank for 1 hour1.157407e-5 days <br />2.777778e-4 hours <br />1.653439e-6 weeks <br />3.805e-7 months <br />. The damping factor for this case was calculated to be about +0.02 per hour; however, when co rrected for fin ite time mesh intervals by the methods of Reference 3.4-7, the da mping factor is increased to approximately
+0.04. When this damping factor is plotted at the appropriate eige nvalue separation for this mode at end-of-cycle, it is apparent that good agreement is obtained with the modified Randall-St. John prediction.
MPS2 UFSAR3.4-8Rev. 35Calculations performed with both Doppler and moderator reactivity feedback have resulted in damping factors which are essentially the same as those obtained with Doppler feedback alone.
This result suggests that the constant power condition which a pplies to the axial oscillations results in a very weak moderator feedback since the moderator density dist ribution is fixed at the top and bottom of the core and only the de nsity distribution in between can change.
For the calculated Doppler coefficient of minus 1.36 x 10
-3 /(kW/ft), the damping factor toward the end of the burnup cycle is positive. Thus, within the uncertainties in predicting power coefficients and uncertainties in the analyses, there is a pr ediction of unstable axial xenon oscillations in the absence of any control action. These oscillations are sufficiently slow (the period of oscillation being 25 to 30 hours3.472222e-4 days <br />0.00833 hours <br />4.960317e-5 weeks <br />1.1415e-5 months <br />) so that there would be sufficient time to control the oscillations. In addition, automati c protection is provided if operato r action is not taken to remedy the situation. Regulating Bank 7 CEA's are utili zed for controlling axial xenon oscillations.
3.
4.6 REFERENCES
3.4-1"Generic Mechanical Design Report E xxon Nuclear 14 x 14 Fuel Assemblies for Combustion Engineering Reactors," XN-NF-82-09(A), Exxon Nuclear Company, Richland, WA 99352, November 1982.3.4-2"Design Report for Millstone Point Unit 2 Reload ANF-1," ANF-88-088(P), Rev. 1, Advanced Nuclear Fuels Corpor ation, Richland, WA 99352, August 1988.3.4-3"Millstone Unit 2 Mechanical Design Report for Increased Peaking" EMF-91-245(P), Siemens Nuclear Power Corporation, January 1992.3.4-4EMF-96-029(P)(A) Volumes 1 and 2, "Reactor Analysis System for PWRs Volume 1 -
Mehodology Description, V olume 2 - Benc hmarking Results", Siemens Power Corporation, January 1997.3.4-5Randall, D., "Xenon Spatial Oscillations," Nucleonics, 16, 3, pages 82-86 (1958).3.4-6Stacey, Jr., W. M., "Linear Analysis of Xenon Spatial Oscillations
," Nuclear Sci. Eng., 30, pages 453-455 (1967).3.4-7Poncelet, C. G., "The Effect of a Finite Time Step Length on Calculated Spatial Xenon Stability Characteristics in Large PWR's" Trans. ANS, 10, 2, page 571 (1967).3.4-8CEND-TP-26., Diatch, P.B.3.4-9ANSI/ANS-19.6-1 "Reload Startup Physics Tests for Pressurized Water Reactors,"
2005.3.4-10WCAP-16011-P-A, Revision 0, "Startup Te st Activity Reduction Program," February 2005.
MPS2 UFSARMPS2 UFSAR3.4-9Rev. 35TABLE 3.4-1 FUEL CHARACTERISTICS FOR A REPRESENTATIVE RELOAD COREFuel TypesN1N2N3N4P1P2P3P4P5R1R2R3R4R5R6Central Zone Assem-bly Average Enrich-ment (w/o)3.943.903.873.823.873.863.843.813.764.494.494.474.394.334.42Number Gadolinia
Bearing Rods0612160481216048121612Nominal Density (%
TD)949494949494949494959595959595Pellet OD (inches)0.3700.3700.3700.3700.3700.3700.3700.3700.3700.3770.3770.3770.3770.3770.377Clad OD (inches)0.4400.4400.4400.4400.4400.4400.4400.4400.4400.4400.4400.4400.4400.4400.440 Diametral Gap (inches)0.00800.00800.00800.00800.00800.00800.00800.00800.00800.0070.0070.0070.0070.0070.007Clad Thickness (inches)0.0310.0310.0310.0310.0310.0310.0310.0310.0310.0280.0280.0280.0280.0280.028Rod Pitch (inches) 0.5800.5800.5800.5800.5800.5800.5800.5800.5800.5800.5800.5800.5800.5800.580Spacer MaterialBime-tallic Bime-tallic Bime-tallic Bime-tallic Bime-tallic Bime-tallic Bime-tallic Bime-tallic Bime-tallic Bime-tallic Bime-tallic Bime-tallic Bime-tallic Bime-tallic Bime-tallicFuel SupplierAREVAAREVAAREVAAREVAAREVAAREVAAREVAAREVAAREVAAREVAAREVAAREVAAREVAAREVAAREVAFuel Stack Height Nominal (inches)136.7136.7136.7136.7136.7136.7136.7136.7136.7136.7136.7136.7136.7136.7136.7Number of Assem-blies82082588128368888484Regionwise Loading (MTU)3.047.603.039.433.043.044.553.0313.583.193.193.193.1718.981.59 MPS2 UFSAR3.4-10Rev. 35 (a) Including uncertainties.TABLE 3.4-2 NEUTRONICS CHARACTERISTICS FOR A REPRESENTATIVE RELOAD CORE
<characteristic
> B OC E O C Critical Boron (ppm): HZP, ARO, No Xenon1453
---Critical Boron (ppm): HFP, ARO, Equilibrium Xenon 1024 0Moderator Temperature Coefficient (pcm/
°F): HZP+2.0-10.4Moderator Temperature Coefficient (pcm/
°F): HFP-6.0-23.3Doppler Coefficient (pcm/
°F)-1.17-1.33Boron Worth (pcm/ppm): HZP-8.8-10.8Boron Worth (pcm/ppm): HFP-8.4-10.4LHR (kW/ft) HFP (a)12.811.6Delayed Neutron Fraction0.00640.0054HFP, PDIL Worth (pcm)157241 N-1 Rod Worth, HZP (pcm)62717696Excess Shutdown Margin (pcm): HFP124323Excess Shutdown Margin (pcm): HZP140751 MPS2 UFSAR3.4-11Rev. 35TABLE 3.4-3 REPRESENTATIVE SHUTDOWN MARGIN REQUIREMENTSControl Rod Worth (pcm)
<parameter>BOC: HZP BOC: HFPEOC: HZP EOC: HFPARI931593151045010450N-16271627176967696PDIL21161572862241[(N-1) - PDIL]
- 0.93740550343516710Reactivity Insertion (pcm)
<parameter>BOC: HZP BOC: HFPEOC: HZP EOC: HFPPower Defect0150702515 Void050050Flux Redistribution02220222Total Requirements0177902787Shutdown Margin (pcm)
<parameter>BOC: HZP BOC: HFPEOC: HZP EOC: HFP[(N-1)
- PDIL]
- 0.9 - Total3740372443513923Required Shutdown 3600360036003600 Excess Shutdown Margin140124751323 MPS-2 FSAR June 2000 FIGURE 3.4-1 REPR ESENTATIVE FULL COR E LOADING PATTERN MPS-2 FSAR June 2000FIGURE 3.4-2REPRESENTATIVE QUARTER CORE LOADING PATTERN MPS-2 FSAR June 2000FIGURE 3.4-3REPRESENTATIVE BOC AND EOC EXPOSURE DISTRIBUTION MPS-2 FSAR June 2000FIGURE 3.4-4REPRESENTATIVE BORON LETDOWN, HFP, ARO MPS-2 FSAR June 2000FIGURE 3.4-5 REPRESENTATIVE NORMALIZ ED POWER DISTRIBUTIONS, HOT FULL POWER, EQUILIBRIUM XENON, 150 MWD/MTU MPS-2 FSAR Rev. 32FIGURE 3.4-6REPRESENTATIVE NORMALI ZED POWER DISTRIBUTION, HOT FULL POWER, EQUILIBRIUM XENON, 18,020 MWD/MTU
MPS2 UFSAR3.5-1Rev. 35 3.5 THERMAL-HYDRAULIC DESIGN This section presents thermal and hydraulic analysis of the reactor core, analytical methods utilized, and experiment al work supporting the analytical t echniques. The prime objective of the thermal and hydraulic design of the reactor is the assurance that th e core can meet normal steady state and anticipated transient performance requirements without exceeding the design bases. A summary of the significant reactor and fuel parameters used in the thermal and hydraulic design and analysis is presented in Table 3.5-1.
3.5.1 DESIGN
BASES 3.5.1.1 Thermal DesignAvoidance of thermally induced fuel damage during any norma l steady state and anticipated transient operation is the principal thermal and hydraulic design basis. Th e following limits are established, but violati on of them will not necessarily result in fuel damage. The Reactor Protection System will provide for automatic reactor trip or other corrective action before these design limits are exceeded.a.Avoidance of departure from nucleate boiling (DNB) for the limiting rod in the core with 95 percent probability at a 95 percent confidence level.b.Limitation of the peak temperature of the fuel to less than the melting point during normal operation and anticipated transients.
Since the departure from nucleate boiling ratio (DNBR) criter ion ensures that the cladding temperature remains close to the coolant te mperature, no additiona l criteria for cladding temperature are required for normal operation and anticipated transients. For design basis accident conditions (loss of coolant accidents (LOCA)), under wh ich the DNBR criterion does not apply, cladding temperatures are calculated to ensure that they remain below 2200
°F, which is the peak clad temperature criterion of 10 CFR 50 Appendix K. For other postu lated accidents, fuel failure is assumed to occur if the calculate d DNBR is below the DNB correlation 95/95 limit.
3.5.1.2 Hydraulic StabilityOperating conditions shall not lead to flow inst ability during normal stea dy state and anticipated transient operation.
3.5.1.3 Coolant Flow Rate, Distribution and Void Fraction A lower limit on the total primary coolant flow rate, called "design" flow, is set to assure that the core is adequately cooled when uncertainties in system resistance, pum p head, and core bypass flow are taken in the adverse direction. By design of the reactor inte rnal flow passages, this flow is distributed to the core such that the core is adequately cooled with al l permissible core power distributions. The hydraulic loads fo r the design of the internals ar e based on the upper limit of the MPS2 UFSAR3.5-2Rev. 35flow. The upper limit is obtained in a similar manner as the design flow but with the uncertainties taken in the opposite direction.To ensure that sufficient coolant flow reaches the fuel, the amount of c oolant flow which bypasses the core through the guide tubes mu st not excessively reduce the active core flow. The guide tube coolant flow must, however, be sufficient to ensu re that coolant in the guide tubes will not boil and ensure adequate cooling of the CEA fingers. The CEA drop time in the guide tubes must also meet the criterion of 90 percent insertion within 2.75 seconds to ensure that scram performance is in accordance with plant Technical Specifications.
Although the coolant velocity, its distribution, and the coolant voids affect the thermal margin, design limits need not be applied to these parameters because they are not themselves limiting with respect to thermal margin. These parameters are included in the thermal margin analyses and thus affect the thermal margin to the design limits.
3.5.2 THERMAL
AND HYDRAULIC CHAR ACTERISTICS OF THE DESIGN 3.5.2.1 Fuel TemperaturesThe RODEX2 code (Reference 3.5-1) incorpor ates models to desc ribe the thermal and mechanical behavior of the fuel rod in a fl ow channel including the gas release, swelling, densification, and cracking in the pellet; the gap conductance; the radial thermal conduction; the free volume and gas pressu re internal to the fuel rod; the fuel and claddi ng deformations; and the cladding corrosion as a function of burnup. The calculations are performed on a time-incremental basis with conditions being updated at each calculated increment.
3.5.2.1.1 Fuel Cladding TemperaturesThe RODEX2 thermal-hydraulic model (Reference 3
.5-1) calculates the lowest cladding surface temperature based on one of two h eat transfer regimes; i.e., forc ed convection and fully developed nucleate boiling. The forced conv ection and fully developed nucleate boiling heat transfer correlations in RODEX2 were developed by Kays and Thom et al., respectively.
3.5.2.1.2 Fuel Pellet Temperatures The RODEX2 radial temperature distribution model begins with the standard diff er ential equation of heat conduction (Poisson Equation) for an isotropic solid with internal heat generation. The equation is written in cylindrical coordinates assuming that the thermal conduc tivity of the fuel is a function of fuel temperature, but is independent of position. With additional assu mptions of axial symmetry, negligible heat conduction in the axial directi on, and steady state conditions, a one-dimensional (i.e., radial) steady state form of the equa tion is derived and employed.
The minimum power level required to produce center line melt in Zircaloy clad uranium fuel rods is defined as the Fuel Centerline Melt Linear Heat Rate (FCMLHR) limit and is expressed in kW/
ft. This FCMLHR is determined using the methodology of Reference 3.5-22. A conservative cycle specific FCMLHR limit is used for Millstone Unit 2. The maximum LHR for normal MPS2 UFSAR3.5-3Rev. 35 operation and anticipated transients is typically well below the c onservative FCMLHR limit. It should be noted that a gadolinia-bearing fuel rod will, for a given LHGR , operate with a higher fuel temperature than an all-uranium-bearing fuel rod. Gadolinia rods are specifically analyzed to centerline melt criteria.
3.5.2.1.3 UO 2 Thermal ConductivityLyon's expression for thermal conduc tivity of the fuel is used in RODEX2. Tw o corrections are applied: one for density and one to account for the gadol inia content in the fuel.
3.5.2.1.4 Gap Conductance The RODEX2 gap conductance mode l is based on that proposed by Kjaerheim and Rolstad. The total gap conductance has three components: (1) gas conductance, (2) radiation, and (3) fuel/
cladding solid-to-solid contact.
3.5.2.2 Departure from Nucleate Boiling Ratio DNBRs are calculated using approved correlat ions. An approved core thermal-hydraulic computer code is used to determine the flow and enthalpy distribution in the core and the local conditions in the hot channel for use in the DNB correlation.
3.5.2.2.1 Departure from Nucleate BoilingThe XCOBRA-IIIC (Reference 3.5-2) computer c ode is employed to evaluate the thermal-hydraulic conditions in the various assemblies and in the subchannels of the limiting assembly.
Heat, mass, and momentum fluxes between the inter-rod flow cha nnels are explic itly calculated.
Fuel and reactor design conditions employed in these calculations are given in Table 3.5-1.
The calculations include a stat istically determined engineer ing factor to account for manufacturing tolerances, thermal expansion and densification effect
- s. The engineering factor is applied to the local heat fl ux in the calculation of DNBR.In-reactor densification results in a shortening of the fuel column. At power levels typical of DNBR-limiting rods, thermal expansion tends to offset the densification effect. The XCOBRA-IIIC model does not specifically model changes in stack length due to thermal expansion and densification.
The HTP DNB correlation, demonstr ated to be applicable to the AREVA 14 by 14 reload fuel assemblies for CE reactors, is described in Reference 3.5-3. A minimum allowable limit corresponding to 95% probability with 95%
confidence is set on the DNBR during normal operation and any anticipated transients.
3.5.2.2.2 Hot Channel Factors Hot channel factors for heat flux and enthalpy rise, F q and F r:
MPS2 UFSAR3.5-4Rev. 35The total hot channel factors for heat flux and enthalpy rise are defined as the maximum-to-core average ratios of these quantities. The heat fl ux hot channel factor (F q) considers the local maximum linear heat generation ra te at a point (the ho t spot), and the enthalpy rise hot channel factor (F r) involves the maximum integrat ed linear heat generation ra te along a channel (the hot channel).Engineering hot channel factor, F E: The engineering hot channel factor is used to evaluate the maximum linear heat generation rate in the core. This subfactor is determined by statistically combining the fabrication uncertainties for fuel pellet diameter, density, and enrichment, as well as the ef fect of densification. A conservative value of 1.03 is used. The effect of variations in fabrication tolerances is considered in the analysis. To account for manufacturing uncertainties and densification, the peak rod heat flux is increased by 3% in the calculation of DNBR.
3.5.2.2.2.1 Nuclear Peaking Factors Assembly and rod peaking factors and axial power distributions are input into the XCOBRA-IIIC code. Departure from nucleate boili ng is dependent on the local rod heat flux and the local fluid conditions within the channel.The effect of asymmetries in core power distri bution (specifically azimuth al power tilt) is not directly taken into account in the XCOBRA-IIIC thermal-hydraulic calculations. The effects of azimuthal power tilt are accounted for in the generation (verifi cation) of the TM/LP trip and LPD trip monitoring setpoints through the m easurement of radial peaking factors.
3.5.2.2.2.2 Rod Bowing Factor As the fuel assembly burnup increases, the gaps between fuel rods change. Decreased rod-to-rod gaps can occur, which can reduce the DNB ratio.
Penalties are calculated as a function of burnup and applied to the DNBR or peak linear power as appropriate.
3.5.2.2.2.3 Inlet Flow Distribution FactorInlet flow maldistribution is treated in the XC OBRA-IIIC model by applyi ng a generic inlet flow penalty to the limiting assemb ly and its crossflow neighbors.
3.5.2.2.2.4 Flow Mixing FactorThe effects of both pressure-driven and turbul ent flow mixing between channels on the hot channel enthalpy rise are calculated by the XCOBRA-IIIC comput er code. The turbulent flow mixing is modeled empiri cally and is based on the reduction of the data from hot mixing tests using XCOBRA-IIIC.
MPS2 UFSAR3.5-5Rev. 35 The geometry of the channels su rrounding the hot channel and the radial power distribution affect the lateral enthalpy transport for both the pressure-driven and tu rbulent flow mixing.
3.5.2.2.3 Effects of Rod Bow on DNBRIn accordance with AREVA rod bow methodology (Reference 3.5-4), the magnitude of rod bow for assemblies of the type used in Millstone Un it 2 has been estimated.
Significant impact on the DNBR due to rod bow does not occur until the gap closures exceed 50 percent. The maximum design exposure for AREVA reload fuel in Millstone Unit 2 is signifi cantly less than that at which 50 percent closure occurs; theref ore, rod bow does not significa ntly impact the minimum DNBR (MDNBR). A further consequence of the small amount of rod bow for AREVA fuel is that total power peaking is not si gnificantly impacted.
3.5.2.3 Void Fraction and Distribution The XCOBRA-IIIC model calculates the local thermal and hydraulic conditions for input to the DNB correlation. While local conditions of enthalpy, quality, flow rate and pressure are associated with a code-calcula ted local void fraction, the void fr action is not input to the DNB correlation. The DNB correlation is approved over a local quality range, but it is not a direct function of void fraction. Th erefore, there is no expl icit limit set on averag e or local void fraction beyond that implied in the test conditions used to develop the DNB correlation.
3.5.2.4 Coolant Flow Distribution 3.5.2.4.1 Coolant Flow Distribution and Bypass Flow The minimum primary coolant flow rate at fu ll power conditions is given in Ta ble 3.5-1.Tracing the coolant flow path in Figure 3.1-1, the c oolant enters the four inlet nozzles and flows into the annular plenum between the reactor vessel and core support barrel. It then flows down the annulus between the reactor vessel and core barrel and up through the flow skirt to the plenum below the core lower support structure. The skirt and lower support structure help to even out the inlet flow distribution to the core. The coolant passes through the openings in the lower core plate and flows axially through the fuel assemblies.
A portion of the coolant passes through the lower core plate and into the guide tubes in the fuel assemblies. The fuel assembly alignm ent plate is not drilled through in guide tube locations without CE As; therefore, core bypass flow is limited in these guide tubes. After passing th rough the core, the coolant flow s into the region outside the control element assembly shrouds.
From this region, the coolant fl ows across the control element assembly shrouds and passes out th rough the outlet sleeves on the core barrel to the outlet nozzles.
The coolant which does not contact any fuel rods is termed core bypass cool ant. The following are the principal core bypass routes:a.Direct inlet to outlet coolant flow at the joint between the core support barrel sleeve and reactor vess el nozzle.
MPS2 UFSAR3.5-6Rev. 35b.Coolant flow into the guide tubes in the fuel assemblies.c.Coolant flow in the region between the core support barrel and core shroud.d.Coolant flow from the inlet nozzle re gion through the alignment keyways to the vessel head region.Table 3.5-1 gives the "best estimat e" value for the core bypass flow rate as a fraction of the total primary flow rate. Ta king into a ccount the core bypass flow rate, th e core flow rate, which is the effective flow rate for heat transfer, can be calculated from the total primary coolant flow rate.
3.5.2.4.2 Core Flow Distribution The core flow distribution (CFD) analysis is performed to assess cross flow between assemblies in the core for use in subsequent MDNBR subcha nnel analyses. A full core model provides cross-flow boundary conditions to a full assembly m odel at the assembly boundaries. MDNBRs are computed from a full assembly simulation.
In the analysis, each fuel assemb ly in the Millstone Unit 2 core is modeled as a hydraulic channel.
The calculations are performed wi th the XCOBRA-IIIC computer code (Reference 3.5-2). Cross flow between adjacent assemblies in the open lattice core is di rectly modeled. The single-phase loss coefficients are used in the CFD analyses to hydraulically characterize the assemblies in the core.This computational procedure is designed to evaluate thermal-hydraulic conditions during boiling and non-boiling conditions. One-dimens ional, two phase se parated, slip flow is assumed in the XCOBRA-IIIC calculati on. These assumptions are valid onl y if the cross flow between connecting channels is small compared to the axial velocities in the individual channels. Because small cross flow does exist, mathematical models have to be postulated for both turbulent and diversion cross-flow mixing. Mode ls of the two-phase state are also defined in terms of void fraction, which is a function of enthalpy, flow rate, heat flux, pr essure, and axial position. This computational procedure is not applicable when large blockages exist in the fuel bundles since this leads to considerable cross flow wh ich cannot be adequately represented by the one-dimensional analysis.Table 3.5-1 summarizes the reactor and fuel desi gn parameters used in these CFD calculations and subsequent MDNBR analyses.
3.5.2.5 Pressure Losses and Hydraulic Loads 3.5.2.5.1 Pressure LossesThe fuel assembly irrecoverable pressure los ses have been calculated using standard loss coefficient methods and results from model tests. The pressure loss across the AREVA fuel assembly was determined based on the re sults of Reference 3.5-5 and analyses.
MPS2 UFSAR3.5-7Rev. 35 3.5.2.5.2 Hydraulic Loads 3.5.2.5.2.1 Hydraulic Loads on Vessel Internal Components The design hydraulic loads for the internal com ponents for steady state operating conditions are listed in Table 3.5-2. These loads were derived fr om analysis and from reactor flow model and component test results. All hydraulic loads in Table 3.5-2 are ba sed on the maximum expected system flow rate and a coolant temperature of 500
°F. When these hydraulic loads are used in the structural analysis, they are adjusted for coolant temperatur
- e. The worst condition (i.e., coolant temperature) is not necessarily the same for each internal component; therefore, the loads are adjusted to reflect the difference in coolant temperature. This is done to ensure the design hydraulic stresses are acceptable duri ng start-up and during power operation.The types of loads considered in the analysis ar e: (1) steady-state drag and impingement loads, and (2) fluctuating loads induced by pump pressure pu lsations, turbulence, and vortex shedding.
All of these loads are not exerte d on each internal component, but each component sees at least one of the loads. Table 3.5-2 lists the components and type of loads that are exerted on them.
3.5.2.5.2.2 Core Hydraulic Loads/Fuel Assembly Liftoff The holddown spring force and the a ssembly weight force prevent th e fuel assembly from lifting off the core support plate duri ng reactor steady-state operation, based on the most adverse combination of compone nt dimensional and ma terial property tolerances. In addition, the holddown springs are designed to accommodate the additional lo ad associated with a pump overspeed transient (re sulting in possible temporary liftoff of the fuel asse mblies), and to continue to ensure fuel assembly holddow n following such occurrences.
The limiting reactor steady-state conditions are the 4 th pump startup conditions. These corres pond to the minimum temperature and maximum pressure and coolant fl ow for reactor startup. Thermal expansion of the reactor vessel and fuel assembly is also considered.
3.5.2.6 Correlation and Physical DataReference 3.5-1 describes the correlations and physical data employed in heat transfer calculations performed by RODEX2. Reference 3.5-7 describes the co rrelations and physical data employed in the hydraulic calc ulations performed by XCOBRA-I IIC. Reference 3.5-3 describes the correlations and physical data employed in the DNB correlation.
3.5.2.7 Plant Parameters for Thermal-Hydraulic Design The plant parameters considered include total primary coolant flow rate, vessel inlet temperature, primary pressure, and core thermal power. Two se ts of thermal-hydraulic conditions are defined:
nominal conditions and design co nditions. Nominal plant conditions represent the best estimate for the primary coolant flow ra te, pressure, and vessel inlet temperature and do not include allowances for instrument errors. Design plant conditions represent the lower limit on primary flow rate when uncertainties in system resistan ce and pump head are included, and represent the upper limit on vessel inlet temperature when design margins on st eam generator performance are MPS2 UFSAR3.5-8Rev. 35 included. Furthermore, the varia tions which occur during steady state operation in the power, pressure, and inlet temperature due to controlle r deadband and instrument error are considered with the design plant parameters. During steady state operation, the possible variations in these parameters define an operatin g envelope. One combination of these parameters gives the MDNBR, and this combination is utilized in Chapter 14 as the initial condi tions in transient and accident analysis. Table 3.5-1 lists the nominal plant parameters.
3.5.2.8 Summary of Thermal a nd Hydraulic ParametersThe thermal and hydraulic parameters for the reactor are listed in Ta ble 3.5-1.
3.5.3 THERMAL
AND HYDRAULIC EVALUATION 3.5.3.1 Analytical Techniques and Uncertainties 3.5.3.1.1 XCOBRA-IIIC DNBR Analyses The thermal-hydraulic simulations employed to evaluate the MDNBR were performed in accordance with AREVA's Nuclear Regulatory Commission (NRC) approved thermal-hydraulic methodology for mixed co res (Reference 3.5-8).
The MDNBR performanc e of the core during anticipated transi ents will be demonstrated to meet the thermal-hydraulic design crite rion on DNBR through th e performance of tran sient analysis of the limiting events. The results of this analysis are included in Chapter 14.
3.5.3.1.2 Parameter UncertaintiesTables 14.0.7-2 through 14.0.7-5 identif y parameter uncertainties included in the AREVA thermal and hydraulic and DNB methodology. Plant instrument calibrat ion procedures and related specification requirements are designed so that these uncertainties do not increase.
3.5.3.2 Hydraulic Instability Analysis Boiling flows may be suscepti ble to thermohydrodynamic instabili ties. These instabilities are undesirable in reactors since th ey may cause a change in ther mohydraulic conditions that may lead to a reduction in the DNB h eat flux or to undesired forced vi brations of core components. However, unlike in Boiling Water Reactors (BWRs), hydraulic stability is not a concern in PWR cores. This statement, which is discussed below, is supported by the literature and the state of the art on instabilities occurring in two-phase flow systems.Instabilities in vertical up-flow of a two-phase mixture in a heated channel can be broadly classified into several categories. Of these, th e following relevant instabilities are discussed.1.Flow Excursion MPS2 UFSAR3.5-9Rev. 35Also called Ledinegg Instability, this is well described in Ref. 3.5-10. This instability occurs when the slope of th e boiling channel pressure drop-flow rate curve (internal characteristic) becomes sm aller than the slope of the loop supply pressure drop-flow rate curve (external characteristic), i.e., where P is the pressure drop and G is the mass flow rate.In this manner, a negative flow perturba tion will be amplified as the internal pressure drop becomes larger than the exte rnal at the perturbed flow and the flow decelerates further until a stable point is reached.
If the core is considered as a single averag e channel, the external pressure and flow characteristics as s een by the core exhibit due to the pump characteristics. Th is negative slope is stabilizing.
On the other hand, considering flow in a single limiting bundle, the other parallel flow paths impose a flat pressure drop versus flow relation where d(P)/dG = 0.
While this situation is less stable than th e average core assumption, it is mitigated by the cross flow and mixing between th is limiting bundle and the neighboring bundles. Ref. 3.5-11 shows experimentally a definite stabilizing influence of cross flow mixing.
The internal pressure drop ve rsus flow characteristics were shown to satisfy the Ledinegg stability criterion for a wide range of conditions in the LOFT reactor (Ref. 3.5-12) which closely approximates a PWR core during nominal and worst case operating conditions.
Therefore, in conclusion, Ledinegg Inst ability is not a concern in PWR cores.2.Density Wave InstabilityDynamic instabilities may occur even when the static stab ility criterion is satisfied (pressure drop increases wh en flow increases). For a density wave dynamic dP ()dG---------------
internal dP ()dG--------------------external<dP ()dG--------------------external0
<dP ()dG--------------------internal0
>
MPS2 UFSAR3.5-10Rev. 35instability, consider an inlet flow increase perturbing the initial value. The rate of enthalpy rise and density eff ects will travel up the channel, and the pressure drop increase is delayed. In the case of a sinusoi dal inlet flow pertur bation of particular frequency, the lagging pressure drop respons e is such that its instantaneous value supports the growth of the initial perturbation (Ref. 3.5-13). Such unstable behavior requires the delayed portion of the total pressure drop (in the two-phase region) to be large compared with the si ngle-phase pressure dr op. The onset of this instability depends on the operating conditi ons and the distribution of pressure drop along the channel, as well as the external loop char acteristics. A vast body of literature and several computer programs for the analysis of density waves exists mainly for BWR concerns (see for exampl e the collection of papers in Ref.
3.5-14). Inferences from BWR experience ar e drawn to dismiss the possibility of density wave instabilities in a PWR core:*Unlike a BWR, there is no riser sect ion contributing significantly to the 2-phase pressure drop.*For a single limiting channel with a constant pressure drop boundary condition, the cross flow in a PWR core has a stabilizing effect.
- Density wave oscillations are known to be stabilized with increasing pressure (decreasing enthalpy and density diff erence between the two phases). No unstable density wave os cillations could be obtained for pressures higher than 1200 psia (Ref. 3.5-15).*BWR oscillations occur when the saturated boiling boundary is low (elevation <<4 feet). For a PWR, such boiling boundary can be achieved at nominal flow rates by more than doubling the power, which leaves a
considerable stability margin even for the worst case transient.*Considering the nuclear coupling, the void-reactivity coefficient in a PWR is reduced when the coolant is borated. Such reduction in the void-reactivity coefficient is stabiliz ing to this mode of oscillation.*For a density wave coupled with an out-of-phase neut ron flux oscillation mode, the larg e subcritical reactivity of the first flux harmonic stabilizes this mode of hydraulic-neu tronic oscillation. This is due to the PWR core being small compared with typical BWR cores.
The LOFT reactor stability study also addr essed the density wave oscillations and concluded that these ar e not likely (Ref. 3.5-12).In conclusion, Density Wave Instability is not a concern in PWR cores.3.Flow Pattern Transition Instability MPS2 UFSAR3.5-11Rev. 35 The term "Flow Pattern Instability" is us ed in the literature in two connotations.
The first refers to the slug flow pattern where a particular elevation in a heated channel experiences a succes sion of high void and low void flows as a vapor slug passes through (Ref. 3.5-12). As a vapor slug clears the channel exit, the average void content in the channel is temporaril y reduced and vice versa resulting in pressure drop and flow rate oscillations.
In a worst case condition in a PWR, slug flow may occur in a small number of ch annels near the exit. No significant oscillatory response is expect ed, particularly since the slug formation is limited to a short segment near the ex it of the hot channels.The more common meaning of the "Flow Pattern Transition Instab ility" refers to unstable transitions between bubbly and a nnular flow (Ref. 3.5-10). A flow rate perturbation decreasing the flow rate and increasing the voi d fraction will result in flow transition from bubbly-slug to annul ar pattern. The annular flow is characterized with lower pressure drop, which results in accelerating the flow. The increase in flow rate brings the void fraction back below the value required to support annular flow. Thus the transiti on back to bubbly-slug regime takes place.
Extensive work has been done on flow pa ttern transition (see for example Ref.
3.5-16). Most work was limited to pressu res of 1000 psia and below where these transitions are more distin ct. At higher pressures, Ho sler (Ref. 3.5-17) notes for 1400 and 2000 psia, that the flow appear s more homogeneous with no reliable observation of pattern transition.Weisman et. al. (Ref. 3.5-18) observed no premature DNB due to bubbly-to-slug flow transition which they expected as the range of tested void fractions covers the transition range. Hosler (Ref. 3.5-17), on the other hand, noted that CHF occurred via a film dryout mechanism in established annular flow, which is far from the transition boundary to bubbly-slug pattern.
In conclusion, Flow Pattern Transition Instability is not a concern in PWR cores.
3.5.3.3 Core Hydraulics 3.5.3.3.1 Fuel Assembly Pressure Drop CoefficientsPressure drop coefficients for the AREVA reload fuel presented are deri ved from pressure drop tests performed in AREVA's portable hydraulic test facility (Reference 3.5-5). The pressure drop coefficients are for the liquid phase and ar e referenced to the bare rod flow area.For reload Batches M (Cycle 10), N (Cycle 11), a nd P (Cycle 12) the pressure drop coefficient for the lower tie plate/spacer combination includes the effects of a debris resistant spacer. The reload Batch R (Cycle 13) and S (Cycle 14) fuel assemblies implemented an alternative debris resistant design which has a slightly lowe r pressure drop across the lower tie plate/spacer combination compared to the Batch P and prio r fuel assemblies. As a result, for Cycles 13 and 14, this results MPS2 UFSAR3.5-12Rev. 35 in higher inlet flows to the Ba tch R and S assemblies and a re duction in flow to the surrounding Batch M, N, and P reinsert assemblies.Due to crossflow effects, the decr eased flow will equilibrate with adjacent assemblies within the next one or two spacers.
Limiting MDNBRs occur toward the top of the core. Therefore, the slight redistribution in the inlet flows, due to the new lower tie plate and adjacent spacer, will not affect calculated MDNBRs.
The Batch T and later HTP fuel assemblies have a lower total pr essure drop than the previous bimetallic fuel assemblies (i.e
., Batch S and prior). A thermal hy draulic compatibility analysis was performed in Reference 3.5-23 for HTP fuel assemblies co-resident with bimetallic fuel assemblies in the Millstone Unit 2 core. This analysis demonstrates that the two fuel assembly types are compatible. Of note is that the core pressure drop would decrease by approximately 1.5% from the all bimetallic core (Cycle 14) to an all HTP core. The core pressure drop decrease from Cycle 14 to Cycle 15 will be approximately 0.69% since the Cycle 15 core has 80 HTP (Batch T) fuel assemblies and 137 bimetallic (Batch N, P, and R) fuel assemblies. Use of the HMP spacer in the lowermost position (R eload Y and later) has a negligible effect on core differential pressure.
3.5.3.3.2 Guide Tube Bypass Flow and Heating Analysis The guide tube thermal-hydrauli c design calculations are perfor med to demonstrate adequate cooling of the CEA fingers an d to ensure that bypass flow through the guide tubes does not unduly reduce core flow.
Flow enters the guide tube thro ugh the weep holes and cap screw and exits through the top of the guide tube. In the Millstone Unit 2 core, there are 81 assemblies under CEA positions. Of these, 73 assemblies are under ac tive CEA positions. The CEA fingers extend a short distance into the guide tube in these 73 assemblies at the all-rods-out (ARO) posit ion which provides a substantial reduction in the guide tube bypass flow. The remaining eight assemb lies were originally under the part length CEAs which have be en removed. In these eight assemblies, the flow is unimpeded, since the last flow plugging devices were removed in Cycle 12. The assembly guide tubes of 91 assemblies project a short distan ce into close fitting sockets in the upper alignment plate. The resulting flow annulus represen ts a significant resistance to gu ide tube bypass flow in these assemblies. The remaining 45 core locations are instrument tube locations. In these locations, the peripheral guide tubes also proj ect a short distance into clos e fitting sockets in the upper alignment plate. The center guide tube contains instrumentation which produces a flow annulus which in turn reduces the flow in the center guide tubes.
The guide tube model employed in the flow and heating calculations uses loss coefficients to determine the guide tube flow path hydraulic losse
- s. The core pressure drop at rated power and flow is employed as the driv ing force for flow through the guide tube. The model permits calculation of the guide tube configurations described above.
The guide tube thermal model includes the effects of coolant heating by gamma deposition and neutron deceleration. The effects of heating due to neutron abso rption and gamma deposition in the inserted control rod are MPS2 UFSAR3.5-13Rev. 35 evaluated. Heat transfer through the guide tube wa ll to the coolant in the surrounding assembly is accounted for in the model.
Calculations were performed to assess th e maximum expected guide tube bypass flow (Reference 3.5-6). At hot full pow er (HFP), ARO configuration was se lected as that resulting in the greatest bypass flow. The total core bypass flow, including flow through the guide tubes in this instance, was determined to be le ss than 4.0 percent of vessel flow. The result conf irms that guide tube bypass flow does not unduly reduce core flow.To assess the adequacy of guide tube cooli ng, a simulation was also performed for a single assembly with the CEA fully inserted at HF P conditions. The fully inserted CEA fingers substantially increase the hydraulic resistance in the guide t ube, and also represent a significant heat source. The exit coolant temp erature is well below saturation.
Heat transfer through the guide tube wall provides a significant part of the cooling.
Based on the results desc ribed above, it is concluded that ample guide tube cooling is afforded by the current design, and that bypass flow remains within acceptable limits.
3.5.3.3.3 Control Element Assembly Insertion Time AnalysisA large data base of CEA insertion time measurements has been obtained at a CE plant similar to Millstone Unit 2, with fuel identical in pertinent guide tube design characteristics to the Millstone Unit 2 AREVA reload fuel. The measurements sp an a time period during which reload quantities of AREVA fuel resided in the core. Statistical analysis (Reference 3.5-
- 6) of this data indicates that the CEA 90 percent insertion time is equal to or less than 2.5 seconds, which is well below the maximum acceptable 90 percent insertion time of 2.75 seconds specified in the Technical Specifications.
Over 500 CEA insertion time measurements from nine different tests were analyzed. The measurements reflect the time required to reach 90 percent inserti on from interrupti on of power to the CEA drive mechanism. Approximately six stan dard deviations separate the mean of the measured CEA insertion time data from the 2.75 second maxi mum allowable for Millstone Unit 2.
With over 500 data points, higher order statistics may also be applied to the data to conclude that the rod drop time will be equal to or less than the greatest time measured in the tests with a probability of 99 percent at a 99 percent confidence level. The gr eatest rod drop time in the tests, as noted above, was 2.50 seconds. The AREVA assembli es are, therefore, expected to conform to the maximum CEA 90 percent insertion time of 2.75 seconds with a substantial margin.
3.5.3.3.4 Fuel Assembly Liftoff The hydraulic lift force on the fuel assembly wa s calculated (Reference 3.5-6) using the drag coefficient for a 14 by 14 fuel assembly with bimeta llic spacer grids. This value differed slightly for Reload Batches M, N, and P (Cycles 10, 11, and 12). The replacement of a bimetallic spacer with a debris resistant Inconel HTP spacer increased the drag while the thermal rounding of the MPS2 UFSAR3.5-14Rev. 35 leading edges of the remaining bimetallic spacers decreased the drag. The overall effect was a slight increase in drag force. The total of th e buoyancy and hydraulic lift fo rces was calculated to be 1194 pounds. The assembly we ight and spring force totals 1801 lbs, thus providing a 607 pound holddown margin. This margin, wh ich is more than half of th e worst case steady state lift force, will envelope any minor variation due to the spacer modifications. It will also provide holddown during and after a 20% pump overspeed re sulting in a 44% lift force increase. For Reload Batch R (Cycle 13) and Batch S, the fuel assembly wei ght increased by approximately 40 pounds and a bimetallic sp acer replaced the Inconel HTP spacer, increasing the margin to liftoff.
A similar analysis was performed for the Relo ad T design. The use of HMP spacers beginning with Reload Y has a negligible effect on lift.
The maximum shear stress of 84,062 psi in the holddown springs occurs in the cold reactor condition. This is below the de sign criterion of 100,000 psi. The stress at reactor operating conditions is 74,188 psi, which is below the cr iterion of 90,000 psi at operating temperature.
Irradiation may cause some stress relaxation of the Inc onel X-750 holddown springs while causing irradiation induced growth of the fuel assemblies. The as sembly growth results in higher spring deflection which offsets any radiation induced relaxation of the springs. The springs are partially shrouded in spring cups, which minimi ze flow-induced vibrati on of the springs and prevent potential fretting wear.
3.5.4 TESTS
AND INSPECTIONS 3.5.4.1 Reactor Testing Thermal-hydraulic design cr iteria are verified during plant star tup testing. This is accomplished by measuring the primary intrinsi c parameters (e.g., levels, pres sures, temperatures, flows, neutron fluence and diff erential pressures) and calculating th e non-measurable and extrinsic parameters (e.g., power level, core peaking factors). During the operating cy cle, various thermal-hydraulic parameters are pe riodically monitored to ensure compliance with the Technical Specifications.
3.5.4.2 AREVA DNB and Hydraulic Testing 3.5.4.2.1 DNB Testing Details of the testing supporting the HTP DNB correlation are contained in Reference 3.5-3.
3.5.4.2.2 Fuel Assembly Hydraulic Testing Single-phase hydraulic characteristics of the AREVA Millstone Unit 2 fuel assembly were experimentally determined by hydraulic tests (R eference 3.5-5) performed in AREVA's Portable Hydraulic Test Facility (PHTF).
The pressure drop testing characte rized the component loss/flow coef ficients of the lower tie plate (including the inlet hardware), spacers, and th e upper tie plate (includi ng the exit hardware).
MPS2 UFSAR3.5-15Rev. 35Differential pressure meas urements were taken over a range of Reynolds Numbers (N Re). These data were used to drive empirica l relationships, which describe th e single-phase pre ssure drops of the Millstone Unit 2 fuel assembly and its components.These test data from Reference 3.5-5 were used to calculate the Batch M, N, and P lower tie plate, spacer, and upper tie plate pressure drop coefficients, and the bare rod friction factor. Additional test data and analyses were used to determine the Batch R lower tie plate pressure drop coefficient correlations. The loss/flow coefficients derived from these tests and calculations are all referenced to the bare rod Reynolds Number.
3.
5.5 REFERENCES
3.5-1XN-NF-81-58(P)(A), Revision 2, and Supplements 1 and 2, "RODEX2 Fuel Rod Thermal-Mechanical Response Evaluation Model," March 1984.3.5-2XN-NF-75-21(P)(A), Revision 2, "XCOBRA-IIIC: A Comput er Code to Determine the Distribution of Coolant During Steady-State and Transient Core Operation," January 1986.3.5-3EMF-92-153(P)(A) Rev. 1, "H TP: Departure From Nucleate Boiling Correlation for High Thermal Performance Fuel," Siem ens Power Corporation, January 2005.3.5-4XN-75-32(P)(A), Supplements 1, 2, 3, and 4, "Computational Procedure for Evaluating Fuel Rod Bowing," October 1983.3.5-5ANF-89-018(P), "Single-Phase Hydraulic Flow Te st of ANF Millstone-2 Fuel Assembly," January 1989.3.5-6ANF-88-088(P), Revision 1, "Design Report for Millst one Point Unit 2, Reload ANF-1," August 1988.3.5-7BNWL-1695, "COBRA-IIIC: A Digital Computer Program for Steady-State and Transient Thermal-Hydraulic Analysis of Rod Bundle Nuclear Fuel Elements," March 1973.3.5-8XN-NF-82-21(P)(A), Revision 1, "Appl ication of Exxon Nuclear Company PWR Thermal Margin Methodology to Mixed Co re Configurations," September 1983.
3.5-9 EMF-2135, Revision 0, "Millstone Unit 2 Cycle 13 Extended Shutdown Safety Analysis Report," January 1999.3.5-10J. A. Boure, A. E. Bergles, and L. S. Tong, "Review of Two-Phase Flow Instability," ASME Paper 71-HT-42, August 1971.
MPS2 UFSAR3.5-16Rev. 353.5-11S. Kakac et. al., "Sustained and Transient Boiling Flow Instabilities in a Cross-Connected Four-Parallel-Channel Upfl ow System," Fifth International Heat Transfer Conf., pp. 235-239, Tokyo, Japan (September 1974).3.5-12S. A. Eide, "Instability Study for LOFT for L2-1, L2-2 and L2-3 Pretest Steady State Operating Conditions," RE-A-78-096, Idaho National Engineering Laboratory, November 1978.3.5-13J. March-Leuba, "Density-Wave Instabilities in Boiling Wa ter Reactors," Oak Ridge National Laboratory Report ORNL/TM-12130 (September 1992).3.5-14Proceedings of the International Workshop on Boiling Water Reactor Stability, Committee on the Safety of Nuclear Reactors Installations, OECD Nuclear Energy Agency, Holtsville , NY (October 1990).3.5-15H. S. Kao, C. D. Morgan, and W. B. Parker , "Prediction of Flow Oscillation in Reactor Core Channel," Trans. ANS Vol. 16, pp. 212-213 (1973).3.5-16A. E. Bergles and M. Suo, "Investigation of Boiling Wa ter Fl ow Regimes at High Pressure," Dynatech Corp. NYO-3304-8 (February 1966).3.5-17E. R. Hosler, "Flow Patterns in High Pressure Two-Phase (Steam-Water) Flow with Heat Addition," 9th National Heat Transfer Conferrence, Chemical Engineering Progress Symposium Series, Number 82, Vol. 64, pp. 54-66 (August 1967).3.5-18Weisman et. al., "Experiment al Determination of the Depa rture from Nucleate Boiling in Large Rod Bundles at High Pressure," 9th National Heat Transfer Conference, Chemical Engineering Progress Symposium Series, Number 82, Vol. 64, pp. 114-125 (August 1967).3.5-19Reference Deleted3.5-20Letter, R. I. Wescott (SPC) to C. H. Wu (NU), "T ransmitt al of Bases for New Uncertainties in the Setpoint Analysis for Millstone Unit 2," RIW:97:049, February 27, 1998.3.5-21Reference Deleted by FSARCR 06-MP2-016.3.5-22"Qualification of Exxon Nuclear Fuel for Extended Burnup,"
XN-NF-82-06(P)(A) Revision 1 and Supplements 2, 4 and 5, Exxon Nuclear Company, October 1986.
3.5-23EMF-2664, Rev. 0, "Millstone Unit 2 Therma l Hydraulic Compatibility Analysis,"
January 2002.
MPS2 UFSAR3.5-17Rev. 35TABLE 3.5-1 NOMINAL REACTOR AND FUEL DESIGN PARAMETERS Design and Operating ParametersValue Core Rated Power 2700 MWt Fraction of Heat Ge nerated in Fuel 0.975 Primary System Pressure 2250 psiaCore Inlet Temperature 549°F Reactor Coolant Flow (Minimum) 360,000 gpm aa.Flow reductions to 349,200 gpm are comp ensated for by reductions in the F r T and linear heat rate limits.Assembly Pitch8.18 inches Bypass Flow Fraction (Best Estimate)0.0303Average Linear Heat Rate6.206 kW/ftTotal Number of Assemblies217 Fuel Parameter sDesign and Operating ParametersValueFuel Rod OD0.440 inches Guide Tube OD (above dashpot)1.115 inches Rod Array14 by 14Rod Pitch0.580 inchesNumber of Fuel Rods/Assembly176 Number of Guide Tubes/Assembly5Active Fuel Length136.7 inchesTotal Fuel Rod Assembly Length146.25 inches Number of Spacers9 MPS2 UFSARMPS2 UFSAR3.5-18Rev. 35TABLE 3.5-2 DESIGN OPERATING HYDRAULIC LOADS ON VESSEL INTERNALS ComponentLoad DescriptionLoad Value Core Support BarreRadial pressure differ ential directed inward opposite inlet duct 40 psi Core Support Barrel and Upper Guide Structure Uplift load 480,000 pounds Flow SkirtRadial pressure differential directed inward 6.0 psi average, 10.2 psi maximum, over 40° sector Bottom PlatePressure differential load directed upward 43,400 pounds Core Support PlatePressure differential load directed upward 43,100 pounds Fuel Assembly Uplift load1194 lbs at 120% flow Core Shroud Radial load directed out ward 20.8 psi at bottom, 0.0 psi at topUpper Guide StructurePressure differential load directed upward 148,000 pounds Fuel Alignment PlatePressure diff erential load directed upward 89,600 pounds Upper Guide PlatePressure differen tial load directed downward 132,000 pounds CEA Shrouds Lateral drag load 4,200 pounds (dual CEA) 1,100 pounds (single CEA)
MPS2 UFSAR3.5-19Rev. 35TABLE 3.5-3 UNCERTAINTY SOURCES FOR DNBR CALCULATIONS (DELETED)
MPS2 UFSAR3.A-1Rev. 35 3.A ANALYSIS OF REACTOR VESSEL INTERNALS 3.A.1 SEISMIC ANALYSIS 3.A.1.1 IntroductionDynamic analyses of the reacto r vessel internals for both hori zontal and vertical seismic excitation were conducted to prov ide further bases for assessing the adequacy of their seismic design. These analyses were perf ormed in conjunction with the dyna mic seismic analyses of the reactor coolant system (RCS) which is discus sed in Appendix 4.A. The following paragraphs provide a discussion of the anal ytical procedures used for th e reactor internals, including a description of the mathematical models. Significant results are listed and compared to the results obtained from applicati on of the design loads.
3.A.1.2 Method of Analysis 3.A.1.2.1 General The procedure used in conducting th e seismic analysis of the reacto r internals consisted basically of three steps. The first step involved the formulation of a ma thematical model. The natural frequencies and mode sh ape of the model were determined during th e second step. The response of the model to the seismic excitation was determined in the third step. In this analysis, the horizontal and vertical components of the seismic excitation were considered separately and the maximum responses added to ob tain conservative results.
3.A.1.2.2 Mathematical Models For the dynamic analysis of the re actor internals, equivalent mu lti-mass mathematical models were developed to represent the system. Since the seismic input ex citation of the reactor internals was obtaine d in the form of acceleration time history of the reactor vessel flange, only the internals are included in the model. The coupling eff ect of the internals' response on the vessel flange acceleration was accounted for by including a simplified representation of the reactor internals with the model of the RCS. This is discussed in Appendix 4.A.
Since the horizontal and vertical responses were treated as uncoupled, separate horizont al and vertical models were developed to more efficiently account for the structural differences in these directions. A sketch of the internals showing the relati ve node locations for the horiz ontal model is presented in Figure 3.A-1. Figures 3.A-2 and 3.A-3 show the ideal ized horizontal and vertical models. Since the structural details provide fo r no vertical load transfer betw een the upper guide structure (UGS) and core or core shroud, the vert ical response of the UGS is in dependent of the rest of the internals. Consequently, the vert ical model was divided into two submodels. Model I consists of the core support barrel/thermal shield (CSB/TS), lower support structure, core shroud and core mass; Model II consists of the UGS.
The mathematical models of the internals are constructed in terms of lumped masses and elastic beam elements. At appropriate lo cations within the internals, poi nts (nodes) are chosen to lump the weights of the structure.
Between these nodes, properties ar e calculated for moments of MPS2 UFSAR3.A-2Rev. 35inertia, cross-section areas, effective shear areas, and le ngths. The salient detail s of the models are discussed below.
3.A.1.2.2.1 Hydrodynamic Effects The dynamic analysis of r eactor internals presents some spec ial problems due to their immersion in a confined fluid. It has been shown both analytically and experimentally (Reference 3.A-1) that immersion of a body in a dense flui d medium lowers its natural freq uency and significantly alters its vibratory response as compar ed to that in air. The effect is more pronounced where the confining boundaries of the fluid are in close proximity to the vibrating body as is the case for the reactor internals. The method of accounting for the effects of a surroundi ng fluid on a vibrating system has been to ascribe to the system additional or "hydrodynamic mass."
This "hydrodynamic mass" decreases the frequencies of the system, but is not directly involved in the inertia force effects. Th e hydrodynamic mass of an immersed system is a f unction of the dimensions of the real mass and the space between the real mass and confining boundary.Hydrodynamic mass effects for mo ving cylinders in a water annulus are discussed in References 3.A-1 and 3.A-2. The results of th ese references are applied to the internals structures to obtain the total (structural plus hydrodynamic) mass matrix which was then used in the evaluation of the natural frequencies and m ode shapes for the model.
3.A.1.2.2.2 Fuel Assemblies For the horizontal model, the fu el assemblies are treated as vibrating in unison. The member properties for the beam elements representing the fuel assemblies we re derived from the results of experimental tests of the fuel assembly load deflection characteristics and natural frequency.
3.A.1.2.2.3 Core Support Barrel FlangesTo obtain accurate lateral and vertical stiffnesses of the upper a nd lower flanges, finite element analyses of these two regions were performed. As s hown in Figures 3.A-4 and 3.A-5, the flanges were modeled with quadrilateral and triangular ring elements. Asym metric loads, equivalent to lateral shear loads and bending moments, and symmetric axial loads were applied and the resulting displacements ca lculated. These results were then used to derive the e quivalent member properties for the flanges.
3.A.1.2.2.4 Control Element Assembly Shrouds For the horizontal model, the control element asse mbly (CEA) shrouds are tr eated as vibrating in unison and are modeled as guided cantilever beams in parallel. To acc ount for the decreased lateral stif fness of the UGS due to local bending of the fuel alignment plate, a short member with properties approximating the local bending stiffness of the fuel ali gnment plate is included at the bottom of the CEA shrouds. Since the stiffness of the UGS support plate is large compared to that of the shrouds, the CEA shrouds are assumed to be rigidly connected to the UGS support plate.
MPS2 UFSAR3.A-3Rev. 35 3.A.1.2.2.5 Thermal Shield Supports For the horizontal model, the th ermal shield supports are modele d as horizontal members. The member properties of the beam elements repr esenting the positioning pins were based on the radial stif fness of the circumfe rential set of pins. Likewise, th e properties of the beam member representing the support lugs were based on the tangential stiffness of the circumferential set of lugs. For the vertical model, th e equivalent cross-section area of the bar element representing the support lugs was based on the axial bending stiffness of the circumferential set of lugs. For both the horizontal and vertical models, the stiffness of the thermal shield supports includes the effect of local deformation of the core support barrel.
3.A.1.2.2.6 Upper Guide Structure Support Plate and Lower Support Structure Grid BeamsThese grid beam structures were modeled as plane grids. Displace ments due to vert ical (out of plane) loads applied at the be am junctions were calculated through the use of the STRUDL computer code (Reference 3.A-3). Average stiffness values based on these results yielded equivalent member cross-secti on areas for the vertical model.
3.A.1.2.3 Natural Frequencies and Normal ModesThe mass and beam element properties of the models were utilized in STAR, a computer program from the MRI/STARDYNE Analysis System programs (Reference 3.A-4) to obtain the natural frequencies and mode shapes. This system utilizes the "stiffness matrix" method of structural analysis. The natural frequencies and mode shapes are extracted from th e system of equations.
[K-W n 2 M]n = 0 where: K = Model stiffness matrix M = Model mass matrix W n = Natural circular frequency for the n th moden = Normal mode shape matrix for n th mode The mass matrix, M , includes the hydrodynamic and structural masses.
The natural frequencies and mode shapes calculated for the first 3 modes for the horizontal model are presented in Figures 3.A-6 through 3.A-8. The natural frequencie s calculated for the vertical model are presented in Table 3.A
-1. The modal data shown is typical and is presented for illustrative purposes. The effect of additional higher modes was included in the response analyses.
MPS2 UFSAR3.A-4Rev. 35 3.A.1.2.4 Response Calculations 3.A.1.2.4.1 Horizontal Direction The time history analysis technique was utilized to obtain th e response of the internals for the horizontal seismic excitation. The horizontal excitation was specifi ed as the acceleration time history of the reactor vessel fl ange, resulting from the operationa l basis earthquake (OBE) (OBE = 0.09g gr ound acceleration). The fla nge excitation resulting from the design basis earthquake (DBE) (DBE = 0.17g ground acceleration) was conservatively specified as 0.17/0.09 times that for the OBE.
The time history response analysis was perfo rmed utilizing the MRI STARDYNE System/
DYNRE 1 Computer Program. This program utilizes the "Normal Mode Method" to obtain time history response of linear elas tic structure. Detail s of the program a nd the "Normal Mode Method" are presented in Refe rences 3.A-4, 3.A-5 and 3.A-6.
Input to DYNRE 1 consisted of the modal data as determined in Section 3A.2.3, the modal damping factors, and the forcing function time history. This analysis used the modal data for all modes with frequencies below 100 cps. This in cluded the first 14 modes. Contributions from higher modes are negligible.
The modal damping factors were obtained by the method of "Mass Mode Weighting" which gives: where:n = Modal damping factor M i = Structural mass of mass node i lil = Absolute value of the mode shape as mass mode ii = Damping associated with pass point i The damping factor assigned to the nodes representing the fuel assemblies was 5 percent. This is a conservative value derived from proprietary experimental results.
A value of 1 percent was used for the other nodes.
The output from the DYNRE 1 code consists of the noda l displacement, velocity, and acceleration time history relative to the base. The member bendi ng moments and shears were obtained from the STAR code (Reference 3.A-5) and were derived from the DYNRE 1 nodal displacement vectors at the times of peak response.nM iiniM iin--------------
-=
MPS2 UFSAR3.A-5Rev. 35 3.A.1.2.4.2 Vertical DirectionThe response of the reactor internals to the vertical excitation was obtained by the response spectrum technique. Because of th e high natural frequencies and resu lting low levels of responses for the vertical direction, the more conservative spectrum response analysis results were used instead of time history results. The response spectrum utilized was derived from the vertical acceleration time history at th e reactor vessel flange. The sp ectrum curv e is presented in Figure 3.A-9.
An acceleration level corresponding to the natural frequency of each mode was selected from the spectrum curve. The response spect rum technique uses these acceler ation values to determine the inertia forces, accelerations, and displacements of each mode. The results for each mode were conservatively combined on the basis of absolute values. For the vertical models, the first seven modes were included in the results.
3.A.1.3 Results Combined results for the horizontal and vertical dynamic seismic analyses are presented in Table 3.A-2 in terms of stresses at critical locations in the reactor internals for the DBE. Table 3.A-2 also lists the seis mic stresses which result from application of the design loads specified for the DBE. A comparison shows the resu lts of the dynamic analysis to be less severe.
3.A.1.4 ConclusionIt is concluded that the seismic loads specified for the design of the inte rnals are adequate. All seismic loads calculated by the dynamic seismic analysis are less than th e design loads specified by the DBE.
3.A.2 NORMAL OPERATING ANALYSIS Design analyses were performe d on the reactor internals for normal operating conditions to demonstrate that the mechan ical design bases were satisfied. These design calculations included appropriate vibration anal yses of the component assemblies.
The flow induced vibration of the CSB/TS, during normal operation, was characterized as a forced response to deterministic and random pressure fluctuations in the coolant. Me thods were developed fo r predicting the response of components to the hydraulic forcing functions.
Emphasis was placed on analysis and design of those components which were particularly critical and susceptible to vibratory ex citation, such as the thermal sh ield. Using a top supported, as opposed to a bottom supported, thermal shield design improves stability as it eliminates a free edge in the flow path. Increas ing the number of upper supports and lower jackscrews, in the specific manner chosen, provides a much stiffer structure and the use of an all-welded shield eliminates local flexibilities and relative motio n at bolted joints. Analytical studies show the thermal shield to be stable on its support system when exposed to the axial annular flow encountered during normal operation. The snubber design is based upon limiting the motion of the core support barrel under condi tions of hydraulically induced vi brations. The snubbers are at MPS2 UFSAR3.A-6Rev. 35 the position of maximum amplitude for the funda mental lateral bending mode of the barrel, thereby restricting motion of the barrel at the most efficien t position. The ci rcumferential distribution of snubbers assures restraint regardless of the di rection of response.
The random hydraulic forcing f unction was developed by analyti cal and experimental methods.
An analytical expression was de veloped to define the turbulent pressure fluctuation for fully developed flow. This expression was modified, based upon the result of scale model testing, to account for the fact that flow in the downcomer was not fully de veloped. Based upon test results, an expression was developed to define the spatial dependenc y of the turbulent pressure fluctuations. In addition, experi mentally adjusted analytical expressions were developed to define; the peak value of the pressure spectral density associated with the turbulence and; the maximum area of coherence, in terms of the boundary layer displacement, across which the random pressure fluctu ations are in phase.
The natural frequencies and mode shapes of the CSB/TS sy stem were obtained using the axisymmetric shell finite elem ent computer program, ASHSD (Reference 3.A-7). This computer program is capable of obtaining natural frequencies and mode sh apes of complex axisymmetric shells; e.g., arbitrary meri dional shape, varying thickness, bran ches, multi-materials, orthotropic material properties, etc. To em ploy the ASHSD code, the CSB/TS were modeled as a series of conical shell frustrums joined at their nodal point circles. The length of each element, throughout the ASHSD model, was a fraction of the shell de cay length. Since rapid changes in the stress pattern occur in regions of structural discontinuity, the nodal poi nt circles were more closely spaced in such regions. The finite element model of the CSB/TS system included representation of the core support barrel upper and lower flanges, sections of different wall thickness, and thermal shield support lugs and j ackscrews. Elements with orthot ropic material properties were utilized to provide equiva lent axisymmetric models of the structural stif fness and constraints to relative motion between the core support barrel and thermal shield provided by the thermal shield support lugs and jackscrews. Those modes which re flect the mass of the lower support structure, core shroud and fuel were simulated by the addition of concentrated masses at specific nodes in the core support barrel fla nge finite element model.Applying Hamilton's Variational Principle to the conical shell el ements an equation of motion was formulated for each degree of freedom of the system. An inverse iterati on technique was utilized in the program to obtain solutions to the charact eristic equation, wh ich was based on a diagonalized form of a consistent mass matrix and stiffness matr ix developed using the finite element method. Four degrees of freedom - radial di splacement, circumfere ntial displacement, vertical displacement, and meridi onal rotation - were taken into account in the analysis, giving rise to coupled mode shapes and corresponding frequencies. Eval uation of the reduction of these frequencies for the system immersed in coolant was made by means of the "virtual mass" method outlined in Reference 3.A-2.
The random response analysis cons iders the response of the CSB/
TS system to the turbulent downcomer flow during steady-st ate operation. The random forcing function is assumed to be a wide-band stationary random pro cess with a pressure spectral density equal to the peak value associated with the turbulence.
The rms vibration level of the CS B/TS system was obtained based upon a damped, single degree of freedom analysis assuming the rms random pressure fluctuation MPS2 UFSAR3.A-7Rev. 35 to be spatially invariant. The analysis demonstrates that the anticipated rms response of the CSBÚTS system is low. Snubber load s were derived using an anal ytical technique originally developed by a Combustion Engin eering (CE) consultant usi ng the random loads discussed above. Modeling the reactor vessel snubbers and core support barrel system as a single degree of freedom spring-mass system, the number and magnitude of snubber, core support ba rrel impacts was calculated based upon the res ponse of the system to random excitation. The sn ubbers were designed, based upon this loading re quirement, to meet the cyclic strength requirements specified in Section III of the ASME Boiler and Pressure Vessel Code.
The forced response of the reactor internals to deterministic loading was evaluated by classical analytical methods, using lumped mass and continuous elastic struct ural models. These calculated responses were used to verify the structural integrity of the reactor ve ssel internals to normal operating vibratory excitation. Components were de sign analyzed to assure that there were no adverse effects from dominant ex citation frequencies, such as pump rotational and blade passing frequencies.
3.A.3 LOSS OF COOLANT ACCIDENT ANALYSIS 3.A.3.1 DiscussionA dynamic analysis (Reference 3.A-8) has been perf ormed to determine the structural response of the reactor vessel internals to th e transient loss of c oolant accident (LOCA) loading. The analysis determined the shell, beam and rigid body mo tions of the internal s using established computerized structural response analyses. The finite-element computer code, ASHSD (Reference 3.A-7) was used to ca lculate the time-dependent beam and shell response of the CSB/TS system to the transient LOCA loading. The finite-element computer code SAMMSOR-DYNASOR (Reference 3.A-9) was used to eval uate the core support barrel's potential for buckling when loaded by a net external radial pressure resulting from an outlet line break. The structural response of the reactor internals to vertical and transver se loads resulting from inlet and outlet breaks, was determined using the spring-mass computer code, SHOCK (Reference 3.A-10).
The time and space depende nt pressure loads used in the above analysis were the result of a detailed hydraulic blowdown analysis. The pressure fluctuations we re determined for each node in the hydraulic model for inlet a nd outlet line breaks. The pressure time histories at these nodal locations were then decomposed into the Four ier harmonics which define the circumferential pressure distribution at the nodal elevations. Where the hydrau lic model nodes did not correspond to those of the structural mode l, the hydraulic model pressure co mponents were interpolated to provide the required loading information.
The finite element computer c ode, ASHSD, was used to calcul ate the dynamic response of the CSB/TS to transient LOCA loading resulting from an inlet break. To employ the ASHSD code, the CSB/TS were modeled as a series of conical shell frustrums (elements) joined at their nodal point circles. Applying Hamilton's Variational Principle to the conical shell elements a damped equation of motion was formulated for each degree of freedom of the system. F our degrees of freedom - radial displacement , circumferential displacement, vertical displacement and meridional rotation - were taken into account in the analysis, giving rise to coupled modes. The MPS2 UFSAR3.A-8Rev. 35differential equations of motions were solved numerically using a step integration procedure. To ensure computational stability of the numerical solution, the inte gration time step was chosen such that it is small compared to the shortest period of the finite el ement system. The model developed for the CSB/TS system is shown in Figure 3.A-10. Th e length of each element, throughout the analytical model, was a fraction of the shell decay length.
Since rapid changes in the stress pattern occur in regions of structural discontinuity, the nodal po int circles were more closely spaced in such regions. The finite el ement model of the CSB/TS system included representation of the core support barrel upper and lower flanges, sections of different wall thickness, and thermal shield support lugs and j ackscrews. Elements with orthotropic material properties were utilized to provide equivalent axis ymmetric models of the structural stiffness and constraints to relative motion be tween the core support barrel a nd thermal shield provided by the thermal shield support lugs and jackscrews. Those modes which re flect the mass of the lower support structure, core shroud a nd fuel were stimulated by the addition of conc entrated masses at specific nodes in the core support barrel flange finite element model.
In performing the dynamic analysis of the CSB/TS system, the transient load harmonics were applied in two successive phas es to account for time-depe ndent boundary conditions at the snubbers. The first phase used th ose harmonics which excite the beam modes, whereas the second phase used those harmonics which excite the shell modes. During th e first phase, the lower end of the core support barrel was unrestrained. Within a very few milliseconds, the clearances between the core support barrel and reac tor vessel snubbers were closed and for the remainder of the LOCA transient, the core support barrel was restrained radially at the snubber level. Transient responses were computed throughout each loading phase.
The ASHSD code computed the nodal point displ acement, resultant shell forces, shell stresses and maximum principle stresses as functions of time. The maximum principle stresses at the internal and external surfaces of the CSB/TS were determined from the bending and membrane components during each phase of transient loading. Stress intens ity levels calcul ated from the principle stresses were comb ined with normal operating and seismic induced stresses for comparison with design criteria.
Accurate representation and analysis of the CSB/
TS shell structures wa s obtained through use of the finite element code ASHSD. Accurate representation of the remainder of the internals (i.e., fuel, core shroud, CEAs, UGS, lower support st ructure, etc.) was obtained using the SHOCK code.The SHOCK code determines the response of structures which ar e represented as lumped-mass systems and subjected to arbitrary loading functions. The code solves the differential equations of motion for each mass by a numerical step-integration procedure.
The lumped mass model can represent a vertically or laterally responding system subject to arbitrary loading functions and initial conditions. Options are available for de scribing steady state loads, preloads, input accelerations, linear a nd nonlinear springs (incl uding tension and compressi on only springs) gaps, and structural and viscous damping.
The reactor internals were developed in terms of a spring-mass system for both vertical and lateral directions; see Figures 3.A-1 1 and 3.A-12. For both models, the spring rates were generally MPS2 UFSAR3.A-9Rev. 35 evaluated using strength of material techniques. However, in complex areas such as at the core support barrel flanges and UGS support flange, the stiffness was derived from finite element model analyses. The lumped mass weights were generally based upon the mass distribution of the uniform support structures, but included at appr opriate nodes, local ma sses such as snubber blocks, fuel end fittings , thermal shield lugs, etc. The net re sult was a lumped-mass system having the same distribution of mass as the actual structure. To simulate the effect associated with the internals oscillating laterally in the water filled vessel, a distributed virtual mass was calculated based upon the procedure outlined in Reference 3.A-8 (which includes the annulus effect) and was added to the structural lumped-mass system, to provide an analytical model with a dynamic response quantitatively similar to the actual internals. In the case of the vertical model, the hydraulic effect is notably one of reducing the effe ctive weight of the reactor internals and this effect was included in the structural lumped-mass system.
The SHOCK code provided excellent facility for modeling cleara nces, preloads and component interfaces. In the lateral model, the core supp ort barrel, reactor vess el snubber clearance was simulated by a nonlinear spring wh ich accounted for the increased resistance to core support barrel motion when snubbing occurred. In the vertical model, nonlinear springs in the form of compression only springs, were used extensively to simulate preload and interface conditions, such as exist between the UGS support plate and co re support barrel upper flange; at the fuel hold-down spring; at the fuel, core support plate interface and at the co re shroud, core support plate interface. Tension only springs were used to simulate the effect of the core shroud tie rods.
In both the vertical and latera l SHOCK models, damping was vari ed throughout the system to simulate structural and hydraulic frictional effects within the reactor internals. The effect of hydraulic drag in the vertical mo del was simulated by a force time-history applied to the fuel lower end-fitting. Vertical loads were used directly from the de tailed hydraulic analysis, whereas lateral loads were obtained by inte grating those harmonics which excite the beam modes to obtain the net lateral load on the CSB/TS system.The SHOCK code calculated the vertical and lateral response of the system in terms of displacements, velocities and accelerations and internal force, moments and shear s as related to each model. These quantities were sufficient to permit calculation of membrane and where appropriate bending stresses for comparison with design criteria.The finite-element code SAMMSOR-DYNASOR was used to dete rmine the dynamic response of the core support barrel, with initially imperfect geometry, to a net external radial pressure resulting from an outlet line break. The above an alysis has the capability of determining the nonlinear dynamic response of axisymmetric shells with initial imperf ections subjected to arbitrarily varying load configurations.Since SAMMSOR-DYNASOR is a finite-element program, a model was developed, Figure 3.A-13, of the core support barrel usi ng axisymmetric finite-elements similar to those used for the ASHSD analysis. As was for the ASHSD model, the SAMMSOR-DYNASOR finite-element lengths were considerably less than the decay length of the core support barrel. The boundary condition at the core support barrel flange was considered fixed, whereas at the core support barrel lower flange radial displacements were restrained. These boundary conditions represented MPS2 UFSAR3.A-10Rev. 35 the restraint due to the expansi on compensating ring and pressure vessel head at the top and the snubbers and lower support structure at the bottom. For conservatism, the stiffening effects of the fuel alignment plate, core shroud an d core support plate were neglected.
Since the basic phenomenon in buckling is nonlinear instability, the init ial deviation of the structure from a perfect geometry greatly affects its response. The initial imperfection was applied to the core support barrel by means of a pse udo-load so developed to provide the maximum imperfection over each of the de sired number of circumferentia l harmonics. The actual transient loading in terms of its harmonics was applied to the initially "i mperfect" geometry core support barrel and the response obtained fo r each of the imperfection harm onics for the combined loading harmonics.
3.A.3.2 Analysis CodesASHSD (Reference 3.A-7) is a structural finite-element computer code developed at the University of California, Berkeley, and supporte d in part by the Nationa l Science Foundation. It performs dynamic analyses of co mplex axisymmetric structures subjected to arbitrary dynamic loadings or base accelerations. The frequencies of free vibrations as calculated by ASHSD compare well to those calculated by th e equations of Herma nn-Mirshy and Flugge, References 3.A-11 and 3.A-12, respectively. The authors also make comparisons with available experimental results (Reference 3.A
-13) of free vibrat ions of cylindrical shells. The resulting comparison is good. Comparison of the numerical solution (Reference 3.A-14) of the dynamic response of a shell to suddenly applied loads a nd the finite-element (A SHSD) solution of the same problem are in good agreement. The response of a shell to a moving axisymmetric pressure load was evaluated by ASHSD and analytically (Reference 3.A-15) with the results being in good agreement.SAMMSOR-DYNASOR (Reference 3.A-9) is a finite-element computer code developed at Texas A&M University and supported in part by a NA SA grant from the Manned Spacecraft Center, Houston, Texas. This code has the capability of determining th e nonlinear dynami c response of axisymmetric shells subjected to arbitrary dynamic loads. Asymme trical dynamic buckling can be investigated using this program. The program has been extensiv ely tested, using problems the solutions to which have been reported by other researchers, in order to estab lish the validity of the codes. Among these are a shallow shell with axisym metric loading as described in Reference 3.A-
- 16. Identical results are obtained with those of Reference 3.A-17 for the analytical evaluation of blast loadings on a cylindrical shell. Calc ulations made by SAMMSOR-DYNASOR for the symmetric buckling of a shallow spherical cap is in good agreement with the analyses of References 3.A-18 and 3.A-19 a nd the experimental data of References 3.A-20 and 3.A-21.
SABOR DRASTIC, (Reference 3.A-22) is a structural fini te-element computer code developed at the Aeroelastic and Structures Research Laboratory, Department of Aeronautics at the Massachusetts Institute of Technology. The work was administ ered by the Air Force Systems Command with technical monitoring by the Aerospace Corp. SABOR 5 - DRASTIC is the end result of combining a finite-dif ference solution procedure and a fi nite-element program to permit predicting the transient response of complex shells of revolution wh ich are subjected to arbitrary transient loadings. Comparisons w ith reliable independent analytic al predictions (notably finite-MPS2 UFSAR3.A-11Rev. 35difference transient response solutions submitted by AVCO) co nfirm the accuracy and reliability of the SABOR 5 -DRASTIC dynamic response pr edictions. An experiment and accompanying analysis were performed by the Aerospace Corp. (Reference 3.A-23) to verify the ability of the code to account for a complex ge ometry shell of revolution subj ected to transient asymmetric loads. Loads were applied by means of well-defined explosive charges.
Based upon the results of dynamic strain measurements made on the test st ructure, it is evident that the SABOR 5 -
DRASTIC code is capable of solving complex dynamic shell structure problems successfully.
In developing the above finite-element com puter codes, (i.e., ASHSD, SAMMSOR-DYNASOR, SABOR 5 - DRASTIC) the authors have independently verified their codes with respect to the results of other established struct ural programs, classical solutions and as possible to experimental data. The correlations demonstr ate that the above programs ar e capable of solving complex dynamic shell structure problems successfully and that the fini te-element method of modeling provides accurate representation of the st ructural phenomena. The SABOR 5 - DRASTIC code, which has had extensive and successful analytical and experimental correlation (Reference 3.A-6) for transient (explosive) asymmetr ic loading, was used to analyze a core support barrel structure with short-term loading. The resu lts of this well-verified program are identical to these of the finite-element codes ASHSD and SAMMSOR-DYNASOR (which are used in the LOCA analysis) for the same core suppor t barrel problem, demonstrating th e ability of these programs to adequately represent and evaluate the effect of a transient load on an axisymmetric structure like the core support barrel.
3.A.4 EFFECTS OF THERMAL SHIELD REMOVAL Following the discovery of the thermal shield s upport degradation at the end of Cycle 5 in July, 1983, the thermal shield was removed. A detailed insp ection of the core barrel revealed damage at two thermal shield support lug locations. Repairs to the core barrel comp rised of drilling crack arrestor holes at the ends of through-wall cracks and rem oval by machining of non through-wall cracks.Analytical evaluations and asse ssments were performed to dem onstrate continued structural adequacy of the reactor intern als without the thermal shield for all design loading conditions.
Special attention was paid to the co re barrel to justify the repairs.
A description of the repairs to the core barrel, analyses, and significant results is given in Reference 3.A-24.
In conclusion, there was no significant change in the loads and the stresses in the internal structures remained within the ASME Code allowables.
3.A.5 LEAK-BEFORE-BREAK ANALYSIS Leak-Before-Break (LBB) analyses for the reactor coolant system (RCS) main coolant loops, for the pressurizer surge line, and unisolable RCS portions of the safety injection and shutdown cooling piping, which demonstrated that the pr obability of fluid syst em piping rupture was extremely low, was reviewed and approved by the commission. (S ee References 3.A-25 through 3.A-29.) Subsequent to the com mission review and approval, we ld overlays were applied to dissimilar metal we lds (DMWs) at the shutdown cooling, the safety inje ction and the pressurizer MPS2 UFSAR3.A-12Rev. 35surge nozzles. A revised LBB analysis was performed for these nozzles (see Reference 3.A-30). Accordingly, pursuant to revised GDC 4, the dynamic effects associated with pipe ruptures in the above piping segments, including the effects of pipe whipping and discharging fluids have been excluded from the design basis of the following reactor vessel a nd reactor internals components:
Core barrel snubbers, core barrel stabilizer blocks Reactor vessel core support ledge Reactor Cavity Seal Plate, Neutron Shielding 3.A.6 REFERENCES3.A-1Fritz, R. J., and Kiss, E., "The Vibration Response of a Cantilevered Cylinder Surrounded by an Annular Fluid," KAPL-M-6539, February 1966.3.A-2Kiss, E., "Analysis of the Fundamental Vibration Frequency of a Radial Va ne Internal Steam Generator Structure," ANL-7685, Proc eedings of Conference on Flow-Induced Vibrations in Reactor System Components, May 1970, Argonne National Laboratory, Argonne, IL.3.A-3ICES STRUDL-II, The Structural De sign Language Engineering Users' Manual.3.A-4"MRI/STARDYNE - Static and Dynamic St ructural Analysis System: User Information Manual," Control Data Corporation, June 1, 1970.3.A-5MRI/STARDYNE User Manual, Computer Methods Department, Mechanics Research, Inc., Los Angeles, California, January 1, 1970.
3.A-6Hurty, W. C., and Rubinstein, M. F., "Dynamics of Structures," Chapter 8, Prentice Hall, Inc., Englewood Cliffs, New Jersey , 1964.3.A-7Ghosh, S., Wilson, E., "Dynamic Stress Analysis of Axisymmetric St ructures Under Arbitrary Loading," Dept. No. EERC 69-10, University of California, Berkeley, September 1969.3.A-8CENPD-42, "Topical Report on Dynamic Analysis of Reactor Ve ssel Internals Under Loss of Coolant Accident Conditions with Application of Analys is to C-E 800 Mw(e) Class Reactors," August 1972.3.A-9Tillerson, J. R., Haisler, W. E., "SA MMSOR II - A Finite Element Program to Determine Stiffness and Mass Matrices of Shells-of- Revolution," Texas A&M University, TEES-RPT-70-18, October 1970. "DYNASOR II - A Finite Element Program for the Dynamic Nonlinear Analysis of Shells-of-Revolution," Texas A&M University, TEES-RPT-70-19, October 1970.3.A-10Gabrielson, V. K., "SHOCK - A Computer Code for Solving Lumped-Mass Dynamic Systems," SCL-DR-65-34, January 1966.
MPS2 UFSAR3.A-13Rev. 353.A-11Hermann, G., Mirshy, I., "Three Dimensi onal Shell Theory Analysis of Axially Symmetric Motions of Cylinders
," Journal of Applied Mechanics, Trans. ASME, Vol. 78, P. 563-568, 1956.3.A-12Flugge, W., "Stresses in Shells," Third Printing, Springer-Verlag, New York, 1966.3.A-13Koval, L. R., Cranch, E. I., "On the Free Vibrations of Thin Cylindrical Shells Subjected to Initial Torque," Proceedings of the U. S. National Congress of Applied Mechanics, P. 11, 1962.3.A-14Reismann, H., and Padloy, J., "Forced, Axis ymmetric Motions of Cylindrical Shells,"
Journal of the Franklin Institute, Vol. 284, Number 5, November 1967.
3.A-15Tang, Sing-Chih, "Response of a Finite Tube to Moving Pr essure," Journal Engineering Mechanics Division, ASCE, V ol. 93, Number EM3, June 1967.3.A-16Klein, S., and Sylvester, R. J., "The Li near Elastic Dynamic Analysis of Shells of Revolution by the Matrix Displacement Method," Air Force Slight Dynamics Laboratory, TR-66-80, 1966, P. 299-329.3.A-17Johnson, D. E., Grief, R., "Dynamic Response of a Cylindrical Shell: Tw o Numerical Methods," AIAA Journal, Vol. 4, Number 3, March 1966, P. 486-494.3.A-18Huang, N. C., "Axisymmetric Dynamic Snap-through of Elastic Clamped Shallow Spherical Shells," AIAA Journal, Vol. 7, Number 2, February 1969, P. 215-220.3.A-19Stephen, W. B., and Fulton, R. E., "Axisymmetric St at ic and Dynamic Buckling of Spherical Caps due to Centrally Distributed Pressures," Paper 69-89, AIAA Journal, 1969.3.A-20Lock, M. H., Okrebo, S., and Whittier, J.
S., "Experiment of the Snapping of a Shallow Dome Under a St ep Pressure Loading," AIAA Journal, Vol. 6, No. 7, July 1968, P. 1320-1326.3.A-21Stricklin, J. A., and Martinez, J. E., "Dynamic Buckling of Clamped Spherical Caps Under St ep Pressure Loadings," AIAA Journal, Vol. 7, Number 6, June 1969, P. 1212-1213.3.A-22Kotanchik, J. J., et al., "The Transient Linear Elastic Response Analysis of Complex Thin Shells of Revolution Subjected to Ar bitrary External Loadings, by the Finite-Element Program SABOR 5 - DRASTIC," AD-709-189, Massa chusetts Institute of Technology, April 1970.3.A-23Klein, S., "A Static and Dynamic Finite Element Shell Analysis with Experimental Ve rification," International Journal for Numerical Methods in Engineering, Vol. 3, P.
299-315, 1971.
MPS2 UFSAR3.A-14Rev. 353.A-24"Thermal Shield Damage Recovery Program - Final Report," Northeast Nuclear Energy Company, Millstone Nuclear Power Station, Unit Number 2, Docket No. 50-336, License No. DPR-65, December, 1983.3.A-25NRC Letter from D. G. McDonald, Jr. to M. L. Bowling, Jr., "Rev ised Evaluation of the Primary Cold Leg Piping Leak - Before-Break Analysis for the Millstone Nuclear Power Station, Unit Number 2," dated November 9, 1998.3.A-26NRC Letter from D. G. McDonald, Jr. to M. L. Bowling, Jr., "Application of Leak -
Before-Break Status to the Portions of the Safety Injection and Shutdown Cooling System for the Millstone Nuclear Power Sta tion, Unit Number 2," dated November 9, 1998.3.A-27NRC Letter from B. Eaton to R. P. Necci, "Staff Review of the Submittal by Northeast Nuclear Energy Company to Apply Leak-Before-Break Status to the Pressurizer Surge Line, Millstone Nuclear Power Station, Unit 2," dated May 4, 1999.3.A-28NRC Letter from G.S. Vissing to J.F. Opeka, "Application of Reactor Coolant System Leak-Before-Break Analysis," dated September 1, 1992.3.A-29Federal Register/Vol ume 53, No. 66/April 6, 1988, "10 CFR Part 50 Leak Before Break Te chnology Solicitation of Public Co mment on Additiona l Applications."3.A-30Structural Integrity Associates Report: 0901238.401, Revision 0, dated: December 2010, Updated Leak-Before-Break Evaluation of Weld Overlaid Hot Leg Surge, Shutdown Cooling and Safety Injection Nozzles for Millstone Nuclear Power Station, Unit 2.
MPS2 UFSAR3.A-15Rev. 35TABLE 3.A-1 NATURAL FREQUENCIES FOR VERTICAL SEISMIC ANALYSIS MATHEMATICAL MODELMode NumberSub-Model I Frequency, cpsSub-Model II Frequency, cps121.6072.98267.75404.093124.59-MPS2 UFSAR3.A-16Rev. 35TABLE 3.A-2 SEISMIC STRESSES IN CRITICAL REACTOR INTERNALS COMPONENTS FOR THE DESIGN BASIS EARTHQUAKE Structural ComponentLocationStress Mode Design Load StressDynamic Analysis Stress Core Support BarrelUpper Section of BarrelTension & Bending1,129 psi746 psiLower Core SupportBeam FlangeBending5,278 psi929 psi CEA Shrouds:
SingleEnd of ShroudTension & Bending3,548 psi1,295 psiCEA Shrouds: DualEnd of ShroudTension & Bending2,762 psi697 psiUpper Grid BeamsCenter of BeamBending1,652 psi127 psi Upper Guide Structure Flange Junction of Flange &
Barrel CylinderTension & Bending2,823 psi146 psi MPS-2 FSAR April 1990 Rev. 26.2FIGURE 3A-1REPRESENTATIVE NODE LOCATIONS - HORIZONTAL MATHEMATICAL MODEL MPS-2 FSAR April 1990 Rev. 26.2FIGURE 3A-2MATHEMATICAL MODEL - HORIZONTAL SEISMIC ANALYSIS MPS-2 FSAR April 1990 Rev. 26.2FIGURE 3A-3MATHEMATICAL MODEL - VERTICAL SEISMIC ANALYSIS MPS-2 FSAR April 1990 Rev. 26.2FIGURE 3A-4CORE SUPPORT BARREL UPPER FLANGE - FINITE ELEMENT MODEL MPS-2 FSAR April 1990 Rev. 26.2FIGURE 3A-5CORE SUPPORT BARREL LOWER FLANGE -
FINITE ELEMENT MODEL MPS-2 FSAR April 1990 Rev. 26.2FIGURE 3A-6LATERAL SEISMI C MODEL - MODE 1, 3.065 CPS MPS-2 FSAR April 1990 Rev. 26.2FIGURE 3A-7LATERAL SEISMIC MODEL - MODE 2, 5.118 CPS MPS-2 FSAR April 1990 Rev. 26.2FIGURE 3A-8LATERAL SEISMIC MODEL - MODE 2, 5.118 CPS MPS-2 FSAR April 1990 Rev. 26.2FIGURE 3A-9REACTOR VESSEL FLANGE VERTIC AL RESPONSE SPECTRUM (1% DAMPING)
MPS-2 FSAR April 1990 Rev. 26.2FIGURE 3A-10ASHSD FINITE ELEMENT MODEL OF THE CORE SUPPORT BARREL/THERMAL SHIELD SYSTEM MPS-2 FSAR April 1990 Rev. 26.2FIGURE 3A-11 VERTICAL SHOCK MODEL MPS-2 FSAR April 1990 Rev. 26.2FIGURE 3A-12LATERAL SHOCK MODE MPS-2 FSAR April 1990 Rev. 26.2FIGURE 3A-13SAMMSOR DYNASOR FINITE ELEMENT MODEL OF CORE SUPPORT BARREL