ML17212A044

Final Safety Analysis Report, Rev. 35, Chapter 3, Reactor

ML17212A044
Person / Time
Site:	Millstone
Issue date:	06/29/2017
From:	Dominion Nuclear Connecticut
To:	Office of Nuclear Reactor Regulation
Shared Package
ML17212A038	List: ML17212A039 ML17212A041 ML17212A042 ML17212A044 ML17212A045 ML17212A047 ML17212A048 ML17212A049 ML17212A050 ML17212A051 ML17212A052 ML17212A053 ML17212A054 ML17212A056 ML17212A058 ML17212A059 ML17212A060 ML17212A063 ML17212A064 ML17212A065 ML17212A066 ML17212A067 ML17212A068 ML17212A069 ML17212A070 ML17212A071 ML17212A072 ML17212A073 ML17212A074 ML17212A076 ML17212A077 ML17212A079 ML17212A080 ML17212A081 ML17212A082 ML17212A083 ML17212A084 ML17212A086 ML17212A087 ML17212A088 ML17212A089 ML18295A107 ML18295A114 ML18295A123 ML18295A125
References
17-208
Download: ML17212A044 (134)

Text

Millstone Power Station Unit 2 Safety Analysis Report Chapter 3

Table of Contents tion Title Page

SUMMARY

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

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

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

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

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

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

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

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

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

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

tion Title Page

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

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

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

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

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

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

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

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

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

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

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

-1 Natural Frequencies for Vertical Seismic Analysis Mathematical Model

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

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

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

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

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

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

-1 Representative Node Locations - Horizontal Mathematical Model

-2 Mathematical Model - Horizontal Seismic Analysis

-3 Mathematical Model - Vertical Seismic Analysis

-4 Core Support Barrel Upper Flange - Finite Element Model

-5 Core Support Barrel Lower Flange - Finite Element Model

-6 Lateral Seismic Model - Mode 1, 3.065 CPS

-7 Lateral Seismic Model - Mode 2, 5.118 CPS

-8 Lateral Seismic Model - Mode 2, 5.118 CPS

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

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

-11 Vertical Shock Model

-12 Lateral Shock Mode

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

SUMMARY

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

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

se tubes are seal welded as hermetic enclosures.

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

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

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

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

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

3.1-1 Rev. 35

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

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

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

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

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

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

a. There is a high confidence level that departure from nuclear boiling (DNB) is avoided during normal operation and anticipated transients. This is achieved by confirming the minimum DNBR calculated according to the HTP correlation (Reference 3.1-1) is greater than the 95/95 limit for the correlation;

b. The melting point of the UO2 fuel is not reached during normal operation or anticipated transients.

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

3.1-2 Rev. 35

raulic design is discussed in Section 3.5.

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

3.1-3 Rev. 35

MPS-2 FSAR FIGURE 3.1-1 REACTOR VERTICAL ARRANGEMENT

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

ALIGNMENT PIN

~ UPPER 40' 11 7h" GUIDE

~

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

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

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

.t FlCNY SKIRT April 1998 Rev. 26.2

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

~~UIVALENT VESSEl.

13~ETER.

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

~TT TT

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

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

~~~nOD

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

..tHua RODS J

0.200 11 WATER GAP April 1998 Rev. 26.2

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

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

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

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

stress design limits for structural components are:

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

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

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

3.2-1 Rev. 35

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

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

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

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

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

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

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

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

CEA stress analyses consider the following load sources:

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

3.2-2 Rev. 35

assumed).

c. Dynamic stresses produced by seismic loading.

d. Dynamic loads produced by stepping motion of the magnetic jack.

e. Mechanical and hydraulic loads produced during SCRAM.

f. Cladding loads produced by differential expansion between clad and filler materials.

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

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

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

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

Categorization and Combination of Loadings

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

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

Shock loads (including OBE)

Transient loadings of frequent occurrences not requiring shutdown Handling loads

2. Emergency Conditions The internals are designed to permit an acceptable amount of local yielding while experiencing the loadings listed above with the SSE load replacing the OBE load.

3. Faulted Conditions Permanent deformation of the reactor internal structures is permitted. The loadings for these conditions include all the loadings listed for emergency conditions plus the loadings resulting from the postulated LOCA.

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

1. Under design loadings plus operating basis earthquake forces or normal operating loadings plus SSE forces, deflections will be limited so that the CEAs can function and adequate core cooling is preserved.

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

3. Under normal operating loadings plus SSE forces plus the maximum pipe rupture loadings resulting from the full spectrum of pipe breaks, deflections will be limited so that the core will be held in place and adequate core cooling is preserved.

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

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

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

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

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

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

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

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

a. Excess Reactivity and Fuel Burnup The excess reactivity provided for each cycle is based on the depletion characteristics of the fuel and burnable poison and the desired burnup for each cycle. The desired burnup is based on the economic analysis of both the fuel cost and the projected operating load demand cycle for the plant. The average burnup in the core is chosen so as to insure that the peak assembly burnup is not greater than 56,000 MWD/MTU for Batch N, 52,500 MWD/MTU for Batch P, and 57,400 MWD/MTU for Batch R and later.

b. Core Design Lifetime and Fuel Replacement Program The core design lifetime and fuel replacement program are based on a three region core with approximately 40 percent of the fuel assemblies replaced at each refueling.

c. Negative Reactivity Feedback and Reactivity Coefficients The negative reactivity feedback provided by the design is based on the requirement of General Design Criterion (GDC) 11. In the power operating range, the inherent combined response of the reactivity feedback characteristics (fuel temperature coefficient (FTC), moderator temperature coefficient (MTC),

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

3.2-6 Rev. 35

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

e. Stability Criteria The design of the reactor and the instrumentation and control systems is based on meeting the requirements of GDC 12 with respect to spatial oscillations and stability. Sufficient CEA rod worth will be available to suppress xenon-induced power oscillations.

f. Maximum Controlled Reactivity Insertion Rates The maximum reactivity addition rates are limited by core, CEA, and reactor regulating system (RRS) design based on preventing increases in reactivity which would result in the violation of specified acceptable fuel design limits, damage to the reactor pressure boundary, or disruption of the core or other internals sufficient to impair the effectiveness of emergency core cooling.

g. Power Distribution Control Acceptable operation of the reactor in the absence of an accidental transient depends on maintaining a relationship among many parameters, some of which depend on the power distribution. In the absence of an accidental transient the power distribution is controlled such that in conjunction with other controlled parameters, limiting conditions of operation (LCO) are not violated. LCO are not less conservative than the initial conditions used in the accident analyses in Chapter 14. LCO and limiting safety system settings (LSSS) are determined such that specified acceptable fuel design limits are not violated as a result of anticipated operational occurrences and such that specified predicted acceptable consequence are not exceeded for other postulated accidents.

h. Shutdown Margins and Stuck Rod Criteria The amount of reactivity available from insertion of withdrawn CEAs is required to be sufficient, under all power operating conditions, to ensure that the reactor can be brought to at least 3.6 percent subcritical from the existing condition, including the effects of cooldown to an average coolant temperature of 532°F, even when the highest worth CEA fails to insert. This criteria is exclusive of any safety allowance and is consistent with the most pessimistic analysis in Chapter 14.

3.2-7 Rev. 35

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

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

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

a. A limit corresponding to 95% probability with 95% confidence (Reference 3.2-1) is set on the departure from nucleate boiling ratio (DNBR) during normal operation and any anticipated transients as calculated according to the HTP correlation.

b. The peak temperature of the fuel will be less than the melting point during normal operation and anticipated transients.

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

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

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

3.2-8 Rev. 35

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

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

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

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

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

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

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

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

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

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

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

3.3-1 Rev. 35

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

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

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

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

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

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

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

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

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

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

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

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

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

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

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

a. beginning of life (BOL) and EOL

b. cold and hot conditions

c. at mid-span and at spacer locations

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

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

a. Radial Thermal Conduction and Gap Conductance

b. Fuel Swelling, Densification, Cracking, and Crack Healing

c. Gas Release and Absorption

d. Cladding Creep Deformation and Irradiation-Induced Growth

e. Cladding Corrosion

f. PCI

g. Free Rod Volume and Gas Pressure calculations are performed on a time incremental basis with conditions updated at each ulated increment so that the power history and path dependent processes can be modeled. The l dependence of the power and burnup distributions are handled by dividing the fuel rod into a ber of axial and radial regions. Power distributions can be changed at any desired time, and coolant and cladding temperatures are readjusted in all the regions. All the performance dels, e.g., giving the deformations of the fuel and cladding and gas release, are calculated at cessive times during each period of assumed constant power generation. The calculated ding strain is reviewed throughout the life of the fuel and both the maximum circumferential in and the maximum strain increment are compared with the design criteria. The calculated in did not exceed the strain limit. Both the maximum strain and the positive strain increment below the design limit strain.

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

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

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

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

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

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

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

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

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

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

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

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

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

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

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

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

3.3-6 Rev. 35

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

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

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

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

a. The maximum steady state cladding strain is well below the design limit.

b. The maximum steady state cladding stress meets the ASME Boiler and Pressure Vessel Code requirements.

c. The transient strain is within the circumferential limit.

d. Cladding creep collapse is precluded.

e. The fuel rod internal pressure at the EOL remains below the criteria approved by the NRC (Ref. 3.3-3).

f. The maximum clad oxidation is below the design limit.

g. The cladding fatigue usage factor is well below the design limit.

h. There is space between the upper and lower tie plate to accommodate fuel rod growth.

i. Pellet centerline temperatures remain below the design criteria.

j. The fuel assembly growth is within the space available between the upper and lower core plates in the reactor core.

k. The assembly holddown springs will prevent bundle lift-off.

3.3-7 Rev. 35

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

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

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

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

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

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

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

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

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

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

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

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

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

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

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

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

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

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

a. Fuel Assembly Handling - Dynamic axial loads approximately 2.5 times assembly weight.

b. For All Applied Loads for Normal Operation and Anticipated Operational Events -

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

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

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

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

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

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

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

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

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

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

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

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

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

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

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

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

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

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

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

3.3-11 Rev. 35

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

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

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

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

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

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

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

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

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

3.3-12 Rev. 35

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

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

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

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

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

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

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

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

3.3-13 Rev. 35

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

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

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

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

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

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

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

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

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

3.3-14 Rev. 35

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

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

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

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

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

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

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

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

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

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

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

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

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

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

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

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

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

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

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

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

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

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

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

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

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

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

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

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

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

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

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

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

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

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

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

3.3-18 Rev. 35

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

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

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

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

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

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

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

3.3-19 Rev. 35

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

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

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

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

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

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

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

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

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

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

3.3-20 Rev. 35

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

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

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

Millstone 2 Fuel Failure Mitigation, March 5, 2008.

3.3-21 Rev. 35

TABLE 3.3-1 MECHANICAL DESIGN PARAMETERS

• l Assembly Geometry 14 by 14 Assembly Pitch, inches 8.180 Assembly Envelope, inches 8.160 Rod Pitch, inches 0.580 Number of Grids per Assembly 9 Approximate Assembly Weight, lb. 1280/1313

• Fuel Rod to Fuel Rod Outside Dimension, inches 7.980 l Rod and Pellet Clad OD, inches 0.440 Clad thickness, inches 0.031/0.028

• Pellet Diameter, inches 0.3700/0.3770

• Pellet Length, inches 0.425/0.435

• Pellet Density (% Theoretical) 94.0/95.0/95.35 **

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

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

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

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

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

3.3-24 Rev. 35

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

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

TES:

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

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

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

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

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

3.3-25 Rev. 35

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

.440 CLADDING OD

.3770 PELLET DIAMETER

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

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

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

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

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

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

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

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

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

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

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

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

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

, * "-53 ,"

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

"I' ~

6 I~ &-15 ~1a M7

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

8- *fJ

~ 3081 8-8 U6 I---

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

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

-ll -- ~

J.~1 ,..--

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

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

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

,V ... "

~

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

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

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

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

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

180 270 April 1998 Rev. 26.2

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

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

SHIM (2 REQ'D PER ASSEMBLY)

PIN (4 REQ'D PER ASSEMBLY)

BOLT (4 REQ'D PER ASSEMBLY)

PRESSURE VESSEL April 1990 Rev. 24.8

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

.J Lower Segment April 1990 Rev. 24.8

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

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

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

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

ELECTRICAl CO~DUIT LIFT COil DRJVING LATCH '

COIL DRIVING LATCHES PULlDOWN COil MOTOR TUBE LOAD

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

April 1990 Rev. 26.2

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

April 1998 Rev. 26.2

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

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

L =~ CS EX rr' ::~.~ w

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

NOZZl.E (C*~ i't.ANiSi l.C.I. Pt.Ai !

L.OWe~!O PCStT10N

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

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

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

April 1990 Rev. 26.2

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

1 GENERAL

SUMMARY

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

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

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

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

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

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

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

a. The design shall permit operation within the Technical Specifications for Millstone Unit 2 Nuclear Plant.

b. The design Cycle length (EFPD) shall be determined on the basis of an estimated Cycle energy and previous Cycle energy window.

3.4-1 Rev. 35

1. The peak linear heat rate (LHR) and the peaking factor Fr shall not exceed Technical Specifications limits in any single fuel rod throughout the cycle under nominal full power operating conditions.

2. The SCRAM worth of all rods minus the most reactive rod shall exceed the shutdown requirement.

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

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

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

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

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

3.4-2 Rev. 35

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

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

reload startup physics test program shall consist of the following:

a. Critical Boron Concentration - HZP, Control Rods Withdrawn.

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

reactivity are fully inserted in the core. 1

c. Control Rod Group Worths - HZP, two or more control rod groups shall be measured which are well distributed radially and represent a predicted total worth of at least 3% reactivity. 1

d. Isothermal Temperature Coefficient - HZP.

e. Flux Symmetry - between 0 and 30% of full power.

f. Power Distribution - between 40 and 75% of full power.

g. Isothermal Temperature Coefficient - greater than 70% of full power.

h. Power Distribution - greater than 90% of full power.

i. Critical Boron Concentration - greater than 90% of full power.

j. HZP to full power reactivity difference.

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

3.4-3 Rev. 35

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

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

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

a. The time scale on which the oscillations occur is long, and any induced oscillations typically exhibit a period of 25 to 30 hours3.472222e-4 days <br />0.00833 hours <br />4.960317e-5 weeks <br />1.1415e-5 months <br />.

b. As long as the initial power peaking associated with the perturbation initiating the oscillation is within the limiting conditions for operation, specified acceptable fuel design limits will not be approached for a period of hours allowing an operator time to decide upon and take appropriate remedial action prior to the time when allowable peaking factors would be exceeded.

c. The core will be stable to radial mode oscillations at all times in the burnup cycle.

d. The core will be stable to azimuthal mode oscillations at all times in the burnup cycle.

e. All possible modes of undamped oscillations can be detected by both exactor and in-core instrumentation as discussed below.

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

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

3.4-4 Rev. 35

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

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

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

features provided for azimuthal xenon oscillation control are:

a. instrumentation for monitoring azimuthal power tilt.

b. administrative limits on azimuthal power tilt.

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

features provided for axial xenon control and protection are:

a. equipment for monitoring axial shape index.

b. administrative limits on axial power distribution, external axial shape index.

c. an axial shape index reactor trip (local power density - high).

d. use of Regulating Bank 7 for control of axial power distribution.

3.4-5 Rev. 35

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

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

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

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

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

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

3.4-6 Rev. 35

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

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

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

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

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

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

3.4-7 Rev. 35

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

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

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

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

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

Siemens Nuclear Power Corporation, January 1992.

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

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

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

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

30, pages 453-455 (1967).

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

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

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

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

3.4-8 Rev. 35

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

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

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

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

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

Nominal (inches)

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

RELOAD CORE

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

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

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

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

3.4-10 Rev. 35

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

BOC: BOC: EOC: EOC:

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

[(N-1) - PDIL]

• 0.9 3740 5503 4351 6710 ctivity Insertion (pcm)

BOC: BOC: EOC: EOC:

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

BOC: BOC: EOC: EOC:

<parameter> HZP HFP HZP HFP

[(N-1)

• PDIL]

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

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

• 11 13 15 16 17 18 18 20 21 y

I=J N2B1 N191 N32/

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

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

~

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

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

f--- I'"'-

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

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

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

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

N18 N3S/

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

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

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

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

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

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

• 32.683 46.522 0.000 20.109 20.166 3 ....107 0.000 16.108 33.070 40.056 61 112 62 113 31.943 34.813 38.624 39.900 12 14 FIGURE 3.4-3 REPRESENTATIVE BOC AND EOC EXPOSURE DISTRIBUTION

-- ~

June 2000 June 2000

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

1100

- 1(XX) '" I'..

E -,

a..

-Q.

!IX)

"~

(

0 fa)

.J

.J ~,

0 -

700 t

.J

( -,

e 0

c 600 0 SOO 0

"~

(

L 0

0 400 '" "\

r\.

r\

Q)

m ~

1"\

X()

i\.

100

" \.

o r\

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

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

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

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

• 1.317 1.564 1.083 1.151 1.258 1.573 1.130 1.222 1.302 1.587 1.115 1.218 1.120 1.574 0.385 0.846 43 1f4 44 as 45

• N4 4' Jl5 4' P3 48 P2 49 114 D 0.971 1.236 0.924 1.277 1.115 0.823 0.346 1.047 1.536 0.geO 1.595 1.218 1.068 0.684 50 R6 51 P4 52 R5 53 PI 54 R3 55 114 e 1.327 1.107 1.210 1.212 1.120 0.347 Re9ioft 110. I ruel 1.577 1.1'7 1.541 1.313 1.574 0.686 A**-=bly Average Power 56 K2

• 57 R2 58 P5 59 JU 60 .2 A** embly Peak Pin Power

• 0.765 0.945 1.216 1.578 0.782 1.046 1.027 1.57" 0.385 0.845 Pr

• 1.600 (e 11) 61 If2 62 1f3 PEAK LRR (kW/ft)

• 12.8 (1 6 B 16) 0.349 0.264 0.630 0.547 12 ~.

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

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

June 2000 June 2000

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

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

~

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

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

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

• 0.736 0.883 1.063 1.302 0.'40 0.985 0.859 1.262 0.386 0.758 "r

• 1.517 (E 15) 61 112 62 .3 PEAX LHR (kW/ft)

• 11.6 (21 E 15) 0.384 0.293 0.621 0.547 12 14 FIGURE 3.4-6 REPRESENT:ATIV~_~O~~~_I..I:l:~D~9WER DISTRIBUTION HFP, EQUILIBRIUM XENON, EOC Rev. 32 June 2000

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

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

a. Avoidance of departure from nucleate boiling (DNB) for the limiting rod in the core with 95 percent probability at a 95 percent confidence level.

b. Limitation of the peak temperature of the fuel to less than the melting point during normal operation and anticipated transients.

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

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

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

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

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

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

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

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

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

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

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

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

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

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

ding solid-to-solid contact.

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

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

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

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

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

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

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

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

3.5-3 Rev. 35

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

ineering hot channel factor, FE:

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

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

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

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

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

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

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

3.5-4 Rev. 35

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

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

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

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

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

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

3.5-5 Rev. 35

c. Coolant flow in the region between the core support barrel and core shroud.

d. Coolant flow from the inlet nozzle region through the alignment keyways to the vessel head region.

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

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

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

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

-dimensional analysis.

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

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

3.5-6 Rev. 35

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

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

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

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

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

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

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

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

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

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

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

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

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

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

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

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

1. Flow Excursion 3.5-8 Rev. 35

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

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

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

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

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

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

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

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

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

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

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

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

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

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

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

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

• Unlike a BWR, there is no riser section contributing significantly to the 2-phase pressure drop.

• For a single limiting channel with a constant pressure drop boundary condition, the cross flow in a PWR core has a stabilizing effect.

• Density wave oscillations are known to be stabilized with increasing pressure (decreasing enthalpy and density difference between the two phases). No unstable density wave oscillations could be obtained for pressures higher than 1200 psia (Ref. 3.5-15).

• BWR oscillations occur when the saturated boiling boundary is low (elevation <<4 feet). For a PWR, such boiling boundary can be achieved at nominal flow rates by more than doubling the power, which leaves a considerable stability margin even for the worst case transient.

• Considering the nuclear coupling, the void-reactivity coefficient in a PWR is reduced when the coolant is borated. Such reduction in the void-reactivity coefficient is stabilizing to this mode of oscillation.

• For a density wave coupled with an out-of-phase neutron flux oscillation mode, the large subcritical reactivity of the first flux harmonic stabilizes this mode of hydraulic-neutronic oscillation. This is due to the PWR core being small compared with typical BWR cores.

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

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

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

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

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

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

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

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

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

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

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

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

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

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

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

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

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

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

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

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

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

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

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

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

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

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

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

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

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

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

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

3.5-14 Rev. 35

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

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

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

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

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

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

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

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

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

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

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

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

3.5-15 Rev. 35

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

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

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

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

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

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

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

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

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

21 Reference Deleted by FSARCR 06-MP2-016.

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

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

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

3.5-16 Rev. 35

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

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

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

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

.1 SEISMIC ANALYSIS

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

.1.2 Method of Analysis

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

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

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

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

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

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

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

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

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

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

3.A-2 Rev. 35

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

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

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

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

[K-Wn2 M]n = 0 where:

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

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

3.A-3 Rev. 35

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

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

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

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

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

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

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

M i in where:

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

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

3.A-4 Rev. 35

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

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

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

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

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

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

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

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

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

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

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

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

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

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

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

.3 LOSS OF COOLANT ACCIDENT ANALYSIS

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

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

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

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

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

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

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

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

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

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

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

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

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

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

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

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

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

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

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

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

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

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

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

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

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

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

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

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

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

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

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

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

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

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

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

.6 REFERENCES

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

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

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

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

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

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

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

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

Class Reactors, August 1972.

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

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

3.A-12 Rev. 35

78, P. 563-568, 1956.

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

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

11, 1962.

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

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

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

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

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

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

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

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

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

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

299-315, 1971.

3.A-13 Rev. 35

License No. DPR-65, December, 1983.

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

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

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

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

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

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

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

3.A-14 Rev. 35

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

3.A-15 Rev. 35

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

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

• MASS NOPE o MASSLESS NODE April 1990 Rev. 26.2

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

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

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

STRUCTURE April 1990 Rev. 26.2

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

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

~ 45 44 UPPER 43 MODEL I

• GUIDE ~ 42 CORE STRUCT URE III SUPPORT

• 41 FUEL BARREL 37 ~ 39'AlIGNMENT 40 PLATE 35 4 33 30 31 428 26 Z7 24 I

MODEL I 22 23 20 THERMAL 16 17 SHltLD

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

... I 4 3

42 1

April 1990 Rev. 26.2

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

~Z April 1990 Rev. 26.2

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

,\

\ \

\

\, \ \

\ \

.Jz R

\

, \\. \

~-I-

~i- ...

~--

1---...

~-

~--

~--

~--

1'/ i'\/ ~ ---

v ---

~

v " --........

~

\.

Y

\

\

April 1990 Rev. 26.2

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

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

J I

• I t

I I

f I I T 1 I *I I 1 f I I I  ; ,

• I I I

\,

\

J I I 1

*I 1 I

I I I

+ 1 I I T I I iL.

t I 1 I

I I L_ *f April 1990 Rev. 26.2

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

r I

• I I

~

I I

I April 1990 Rev. 26.2

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

• 0.20

~

~

D:: 0.15

~

LIJ U

U 4:

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

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

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

8~.S I

85 l~Det.i1A csa R,In.

Snubber Elevation NoJe LaNer

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

5- 12y 4-3- 125

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

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

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

t Expansion

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

~

~

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

.~

1\

Sprang S ructure Plate FI ange ~CSB

.t Control L Upper Flange

."" Element ~

}Cs0rTlpresslo n Only Shroud

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

\ i .l

~ Thermal I .Shield

~,

£ Core

~ Shroud 1.

t

~

~

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

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

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

/ Core Region Lower Barrel April 1990 Rev. 26.2