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Chapter 3 to Davis-Besse PSAR, Reactor. Includes Revisions 1-8
	ML19319C289
Person / Time
Site:	Davis Besse
Issue date:	08/01/1969
From:	; TOLEDO EDISON CO.
To:	;
References
	NUDOCS 8002110800
	Download: ML19319C289 (150)
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D-B t TABLE OF CONTENTS

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Section Page 3 REACTOR 3-1 3.1 DESIGN BASES 3-1 3.1.1 PERFORMANCE OBJECTIVES 3-1 3.1.2 LIMITS 3-1 3.1.2.1 Nuclear Limits 3-1 3.1.2.2 Reactivity Control Limits 3-2 3.1.2.3 Thermal and Hydraulic Limits 3-2 3.1. 2 . k' Mechanical Limits 3-3 3.2 REACTOR DESIGN 3-6 3.2.1 GENERAL

SUMMARY

3-6 3.2.2 NUCLEAR DESIGN AND EVALUATION 3-8 3.2.2/1 Nuclear Characteristics of the Design 3-9 3.2.2.2 Nuclear Evaluation 3-18 3.2.3 THERMAL AND HYDRAULIC DESIGN AND EVALUATION 3-26 3.2.3.1 Thermal and Hydraulic Characteristics 3-26 3.2.3.2 Thermal and Hydraulic Evaluation 3-36 3.2.h MECHANICAL DESIGN LAYOUT 3-57 3.2.h.1 Reactor Internals 3-57 3.2.4.2 Core Components 3-61 3.2.4.3 Control Rod Drive System 3-T3 3.3 . TESTS AND INSPECTIONS 3-88 3.3.1 NUCLEAR TESTS AND INSPECTION 3-88 3.3.1.1 Critical Experiments 3-88

, 3.3.1.2 Zero Power, Approach to Power, and Power Testing 3-88 1

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CONTENTS (Cont'd)

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Section Page 3.3.2 THERMAL AND HYDRAULIC TESTS AND INSPECTION 3-88 3.3.2.1 Reactor Vessel Flow Distribution and Pressure Drop Test 3-88 3.3.3 FUEL ASSEMBLY, CONTROL ROD ASSEMBLY, AND CONTROL ROD DRIVE MECHANICAL TESTS AND INSPECTION 3-89 3.3.3.1 Prototype Testing 3-89 3.3.3.2 Model Testing 3-89 3.3.3.3 Component and/or Material Testing 3-90 3.3.3.k Control Rod Drive Tests and Inspection 3-91 3.3.k INTERNALS TESTS AND INSPECTION 3-92 3.k REFERENCES 3-93 m

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D-B LIST OF TABLES I

Table No. Title Page 3-1 Core Design, Thermal, and Hydraulic Data 3-7 3-2 Nuclear Design Data 3-9 3-3 Excess Reactivity Conditions 3-10 3-4 First Cycle Reactivity Control Distribution 3-11 j 3-5 Shutdown Reactivity Analysis 3-lb 3-6 Soluble Boron Levels and Worth 3-15 l 3-7 pH Chr racteristics i 3-17 3-8 Calculated and Experimental Rod and Rod Assembly Comparison 3-20 3-9 Coefficients of variation 3-30 3-10 DNB Results - Maximum Design Condition 3-32 3-11 DNB Results - Most Probable Condition 3-33 3-12 Hot Channel Performance Vs Pumps in Service 3-35 3-13 Hot Channel Coolant Conditions 3-kl 3-lk DNB Ratios in the Fuel Assembly Channels (W-3) -

Nominal Case 3-55 3-15 DNB Ratios in the Fuel Assembly Channels (W-3) -

Postulated Worst Case (Design) 3-55 3-16 Fuel Assembly Components, Materials, and Dimensions 3-62 3-17 Clad Circumferential Stresses l 3-67 3-18 Control Rod Drive Design Data 3-76

, 3-19 Control Rod Assembly Design Data l 3-83 l

! 3-20 Axial Power Shaping Rod Assembly Design Data 3-85 ]

3-21 Burnable Poison Rod Assembly Design Data 3-86 3-22 Orifice Rod Assembly Data 3-87

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1 D-B LIST OF FIGURES i l

(At Rear of Section) - I Figure No. Title 3-1 Boron Concentration Versus Core Life 3-2 Axial Peak to Average Power Versus Xenon Override Rod Insertion 3-3 Axial Power Profile, Xenon Override Rods 55 Per Cent Inserted 3-h Location of Fuel Assemblies Containing Burnable Poison Rods 3-5 Per Cent Neutron Power Versus Time Following Trip 3-6 Effect of Fuel Temperature (Doppler) on Xenon Oscillations -

Beginning of Life 3-7 Effect of Fuel Temperature (Doppler) on Xenon Oscillations -

Near End of Life 3-8 Control of Axial Oscillation With Partial Rods 3-9 Stability Index Versus Flatness Beginning and F'd of '5 Life (Asimuthal) /

3-10 Population Protected, P, and 1-P Versus DNB Ratio (W-3) 3-11 Power Shape Reflecting Increased Axial Power Peak for ikk-Inch Core 3-12 Distribution of Fuel Rod Peaking 3-13 Possible Fuel Rod DNB's for Maximum Design Conditions -

36,816 - Rod Core 3-lh Possible Fuel Rod DNB's for Most Probable Conditions -

36,816 - Rod Core 3-15 Distribution of Population Protected, P, and 1-P Versus Number of Rods for Most Probable Conditions 3-16 DNB Ratios (W-3) in Hot Unit Cell Versus Reactor Power 3-17 Maximum Hot Channel Exit Quality Versus Reactor Power 3-18 Hot Channel DNB Ratio (W-3) Versus Power for Partial Pump Operation 154 _;

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D-B l FIGURES (Cont'd)

Figure N g Title 3-19 Hot Channel Quality at Point of Minimum DNER Versus Power for Partial Pu=p Operation 3-20 Thermal Conductivity of UO 2

3-21 Fuel Center Temperature at the Hot Spot Versus Linear Power 3-22 Number of Data Points Versus $E/&C 3-23 Hot Channel Factors Versus Per Cent Population Protected 3-2h Burnout Factor (W-3) Versus Population for Various Confidence Levels 3-25 Design Hot Channel and Nominal Channel Exit Qualities Versus Reactor Power (Without Engineering Hot Channel Factors) 3-26 Flow Regime Map for the Hot Unit Cell 3-27 Flow Regime Map for the Hot Control Rod Cell

( 3-28 Flow Regime Map for the Hot Wall Cell 3-29 Flow Regime Map for the Hot Corner Cell 3-30 Hot Channel DNB Ratio (W-3) Versus Power for Various Axial Flux Shapes 3-31 Reactor Coolant System Flow Versus Power 3-32 Hot Channel DNB Ratio (W-3) Versus Power With Reactor System Flov and Energy Mixing as Parameters 3-33 Fuel Center Temperature for Beginning-of-Life Conditic9r 3-34 Fuel Center Temperature for End-of-Life Conditions 3-35 Fuel Te=perature Versus Total Fuel Volume Fraction for Equilibrium Cycle at End of Life 3-36 Typical Reactor Fuel Asse=bly Power Distribution at End of Life Equilibrium Cycle Cond?tions for 1/8 Core 3-37 Per Cent Fission Gas Released ss a Function of the Average Temperature of the UO Fuel 2

3-38 Axial Local to Average Burnup and Instantaneous Power

( Comp,arisons

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D-B FIGURES (Cont'd) g;7 Figure No. Title 3-39 Fission Gas Release for 1 5 and 1.7 max / avg Axial Power Shapes 3-k0 Gas Pressure Inside the Fuel Clad for Various Axial Burnup and Power Shapes 3-kl Nominal Fuel Rod Power Peaks and Cell Exit Enthalpy Rise Ratios 3-42 Maximum Fuel Rod Power Peaks and Cell Exit Enthalpy Rise Ratios 3-h3 Calculated and Design Limit Local Heat Flux Versus Enthalpy in the Hot Unit Cell at the Nominal Condition 3-44 Calculated and Design Limit Local Heat Flux Versus Enthalpy in the Hot Unit Cell at the Design Condition 3-45 Reactor Vessel and Internals - General Arrangement 3-46 Reactor Vessel and Internels - Cross Section 3-h7 Core Flooding Arrangement 3-48 Fuel Assembly 7

3-h9 Orifice Rod Assembly s 3-50 Burnable Poison Rod Assembly 3-51 Control Rod Drive - General Arrangement 3-52 Control Rod Drive - Vertical Section 3-53 control Rod Drive System and Trip Block Diagram 3-54 Control Rod Assembly 3-55 Axial Power Shaping Rod Assembly 4

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D-B I_

3 REACTOR 3.1 DESIGN BASES The Davis-Besse Nuclear Power Station is designed to meet the performance specified in 3.1.1 without exceeding the limits of design and operation speci-fled in 3.1.2.

3.1.1 PERFORMANCE OBJECTIVES The reactor is des'.gned to operate at 2,633 MWtI*} with sufficient. design margins to accommodate transient operation and instrument error without exceeding the pressure at the 'telief valve settings in the reactor coolant system. The ulti-mate operating rower level of the reactor core is expected to be 2,772 MWt.

This section of the report describes only reactor operation at the initial power level.

The fuel rod cladding is designed to maintain its integrity for the antici-pated core life. The effects of gas release, fuel dimensional changes, and corrosion- or irradiation-induced changes in the mechanical properties of cladding are considered in the design of fuel assemblies.

Reactivity is controlled by control rod assemblies (CRA), burnable poison rod assemblies (BPRA), and soluble boron in the coolant. Sufficient CRA worth is available to shut the reactor down (keff g 0.99) in the hot condition at any I time during the life cycle with the most reactive CRA stuck in the fully with-dravn' position. Equipment is provided to add soluble boron to the reactor coolant to insure a similar shutdown capability when the reactor coolant is cooled to ambient temperatures.

The reactivity worth of CRA, and the rate at which reactivity can be added, is limited to insure that credible reactivity accidents cannot cause a tran-sient capable of damaging the reactor coolant system or causing significant fuel failure.

3.1.2 LIMITS 3.1.2.1 Nuclear Limits The core has been designed to the following nuclear limits:

a. Fuel has been designed for a maximum burnup of 55,000 mwd /MTU.

b. The power Doppler coefficient is negative, and the control system is capable of compensating for reactivity changes resulting from nuclear coefficients, either positive or negative.

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(*) Full (rated) core thennal power.

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c. A control system consisting of part length axial power shaping rods is provided to allow the shaping of power axially in the core, there- ~

by thwarting any tendency towards axial instability resulting from a redistribution of xenon.

d. The core vill have sufficient excess reactivity to produce the design power level and lifetime ithout exceeding the control capacity or shutdown margin.

e. Controlled reactivity insertion rates have been limited to 1.09 x 10-N (ak/k)/s for a single regulating CRA group withdrawal, and k.h x 10-6 (Ak/k)/s for soluble boron removal.

f. Reactor control and maneuvering procedures will not produce peak-to-average power distributions greater than those listed in Table 3-1.

The low worth of CRA groups inserted during power operation limits power peaks to acceptable values.

3.1.2.2 Reactivity Control Limits The control system and the operational procedures vill provide adequate control of the core reactivity and power distribution. The following control limits vill be met:

a. Sufficient control vill be available to produce an adequate shutdown margin.

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b. The shutdown margin vill be maintained with the CRA of highest worth stuck out of the core.

c. CRA withdrawal limits the reactivity insertion to 1.09 x 10-4 ( Ak/k)/s on a single regulating group. Bo reactivityinsertionof4.4x10gondilutionisalsolimitedtoa (Ak/k)/s.

3.1.2.3 Thermal and Hydraulic Limits The reactor core is designed to meet the following limiting thermal and hydraulic conditions:

a. No central melting in the fuel at the design overpower (112 per cent).

b. A 99 per cent confidence that at least 99.5 per cent of the fuel rods in the core are in no jeopardy of experiencing a departure from nu-cleate boiling (DNB) during continuous operation at the design over-power. ,

c. Essentially 100 per cent confidence that at least 99.96 per cent of ,

the fuel rods in the core are in no jeopardy of experiencing a DNB '

during continuous operation at rated power.

d. The minimum' allowable DNER during normal operation and expected tran-sients is'l.30 with the W-3 correlation.

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('~ e. The generation of net steam in the hottest core channels is permis-sible, but steam voids vill be lov enough to prevent flow instabili-ties.

The design overpower is the highest credible reactor operating power permitted by the safety system. Normal overpower to trip is significantly less than the design overpower. Core rated power is 2,633 MWt.

3.1.2.h Mechanical Limits 3.1.2.h.1 Reactor Internals The reactor internal components are designed to withstand the stresses result-ing from startup; steady-state operation with two, three, or four reactor cool-ant pumps running; and shutdown conditions. No damage to the reactor internals will occur as a result of loss of pumping power.

Reactor internals vill be fabricated from SA-2h0 (Type 304) material and will be designed within the allowable stress levels permitted by the ASME Code,Section III, for normal reactor operation and transients. Structural integrity of all core support assembly circumferential velds will be assured by compliance with ASME Code, Sections III and IX, radiographic inspection acceptance standards, and welding qualifications.

The core support structure will be designed as a Class I structure, as defined in Appendix 5A of this report, to resist the effects of seismic disturbances.

( The basic design guide for the seismic analysis will be AEC publication TID-702h, " Nuclear Reactors and Earthquakes."

Lateral deflection and torsional rotatien of the lower end of the core support assembly will be limited to prevent excessive movements resulting from seismic disturbance and thus prevent interference with control rod assemblies (CRA).

Core drop in the event of failure of the normal supports vill be limited so that the CRA do not disengage from the fuel assembly guide tubes.

The structural internals will be designed to maintain their functional integrity in the event of a major loss-of-coolant accident as described in 3.2.h.l. The dynamic loading resulting from the pressure oscillations because of a loss-of-coolant accident will not prevent CRA insertion.

Methods of Load Analysis to be Employed for Reactor Internals and Core Static .or dynanic analysis is used as appropriate. In general, dynamic analysis is used for earthquakes and the subcooled portion of the loss-of-coolant acci-dent (LOCA). For the relatively steady-state portion of the LOCA, a static analysis is used. ,

Where it is indicated that substantial coupling, i.e., interrelationship, exists between major components of the Nuclear Steam System (NSS), such as the steam generator, the piping, and the vessel, the dynamic analysis includes the re-sponse of the entire coupled system. However, where coupling is found to be

/' small, the component or groups of components are treated independently of the (s,c overall system. i

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( D-B The dynamic analyris for LOCA uses predicted pressure-time histories as input e to a lumped-mass model. For earthquakes, actual earthquake records nor=alized 'l to appropriate ground motion, are used as input to the model. The output from the analysis is in the form of internal motions (displacementa, velocities, and accelerations), mctions of individual fuel assemblies, impact loads between adj acent fuel assemblies, and impact loads between peripheral fuel assemblies and the core shroud. Motions of the reactor vessel, internals and core have been confirmed using a time history excited lumped mass solution.

In addition, seismic analysis is also performed using a modal superposition and response spectra approach.

For the simultaneous occurrence of LOCA and the maximum earthquake, both time-history excitations are input to the system simultaneously such that maximum structural motions, indicating maximum stresses, are obtained. Outputs are those mentioned above.

The output from the lumped-mass model and additional information such as pres-sure-time histories on separate internals and core components (including con-trol rods) are used to calculate stresses and deflections. These stresses and deflections are compared to the allowable limits for the various loading com-binations as established in Appendix 5A to insure that they are less than these allowables.

3.1.2.h.2 Fuel Assemblies The fuel, assemblies are designed to operate satisfactorily to design burnup ]

,e and to retain adequate integrity at the end of life to permit safe removal from the core.

The assemblies are designed to operate safety during steady-state and transient conditions under the combined effects of flow-induced vibration, cladding strain caused by reactor pressure, fission gas pressure, fuel growth, and differential thermal expansion. The cold-worked Zircaloy h cladding is designed to be free-standing. Fuel rods are held in place by mechanical spacer grids that are de-signed to maintain dimensional control of the fuel rod spacing throughout the design life without impairing cladding integrity. Contact loads are estab-lished to minimize fretting.

The spacer grids are also designed to permit differential thermal expansion of the fuel rods without restraint that would cause distortion of the rods. The fuel assembly upper end fitting and the control rod guide tube in the internals structure are both indexed to the grid plate above the fuel assemblies, thus insuring continuous align =ent of the guide channels for the CRA. The control rod travel is designed so that the rods are always engaged in the fuel assembly guide tubes, thus insuring that CRA can always be inserted. The assembly struc-ture is also designed to withstand handling loads, shipping lot !s, and earth-quake loads.

Stress and strain for all anticipated normal and abnormal operating conditions b will be limited as follows:

a. Stresses not relieved by small material deformation are limited so as _

not to exceed either the yield strength of the material or 75 per

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D-B cent of the stress rupture life of the material. An example of such

( a stress is the circumferential membrane stress in the clad due to internal or external pressure.

b. Stresses relieved by small material defor=ation are permitted to ex-ceed the yield strength. Strain limits for this stress condition are established based on low-cycle fatigue techniques, not to exceed 90%

of material fatigue life. Evaluation of cyclic loading is based on conservative estimates of the number of cycles to be expected. An example of this type of stress is the themal stress resulting from ther=al gradients across the clad thickness.

c. Combinations of these two types of stresses, in addition to the in-dividual treatment outlined above, are evaluated on the low-cycle fatigue basis of item b. Clad plastic strain due to diameter in-creases resulting from fuel swelling, themal ratcheting and creep, and the effects of internal gas pressure, is limited to about 1 per cent.

d. Minimum clad collapse pressure margins will be required as follows:

1. 10 per cent margin over system design pressure, on short-time collapse, at end void.

2. End void must not collapse (must be either freestanding or have adequate support) on a long-time basis.

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3. 10 per cent margin over system operating pressure, on short-time collapse, at hot spot average temperature through the clad vall.

h. Clad must be freestanding at design pressure on a short-time basis at %725 F hot spot average temperature through the clad vall.

3.1.2.4.3 Control Rod Assemblies Absorber material cladding used on the Control Rod Assembly (CRA), Axial Power Shaping Rod Assembly (APSRA), and the Burnable Poison Rod Assembly (EPRA), is designed to the same criteria as the fuel rod clad, as applicable to absorber material characteristics. Clearance is provided between the rods of each of these assemblies and fuel assembly guide tubes to permit coolant flow to limit operating temperature of the absorber materi,als. In addition, this clearance is designed to permit rod motion as required during reactor operation under any conditiop including seismic disturbances. Excessive stress in the CRA com-ponents 415r'ing trip of the rod drive mechanism is prevented by use of conserva-tive design' stress limits and by hydraulic snubbing to minimize shock in the ,

drive mechanism.

l Orifice Rod Assembly (ORA)_

The orifice rod assembly is designed to have adequate clearance when incerted l into the fuel assembly guide tubes to permit coolant flow without unacceptable l

[w mechanical interference between the rod assembly vid guide tubes under any op-erating condition. '

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D-B 3.1.2.4.h Control Rod Drive Mechanisms

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Shim Safety Drive The shim safety control rod drives provide control rod assembly (CRA) insercion and withdrawal rates consistent with the required reactivity changes for reactor operational load changes. This rate is based on the worths of the various rol groups, which have been established to limit power-peaking flux patterns to de-sign values. The maximum reactivity addition rate is specified to limit the magnitude of a possible nuclear excursion resulting frcm a control system or operator malfunction. The nor=al insertion and withdrawal velocity has been established as 30 in./ min.

The drive provides a trip of the CRA which results in a rapid shutdown of the reactor for conditions that cannot be handled otherwise by the reactor control system. The trip set point is based on the results of various reactor emer-gency analyses, including instrument and control delay times and the amount of reactivity that must be inserted before deceleration of the CRA occurs. The maximum travel time for a 2/3 insertion on a trip com=and of a CRA has been established as 1.40 s.

The control rod drives can be coupled and uncoupled to their respective CRA's without any withdrawal movement of the CRA.

All pressure-containing components are designed to meet the requirements of the ASME Code,Section III, Nuclear Vessels, for Class A vessels (summer 1967 rddendum). s

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Materials selected for the control rod drive are capable of operating within the specified reactor environment for the life of the mechanism without any deleterious effects. Adequate clearances are provided between the stationary and moving parts of the control rod drives so that the CRA trip time to full insertion vill not be adversely affected by mechanical interference under all operating conditions and seismic disturbances.

Axial Power Shaping Drive The axial power shaping drives operate similarly to the shim safety drives, ex-cept that the trip function has been eliminated. They have the same insertion and withdrawal velocity of 30 in./ min and can be coupled and uncoupled to their rem ective APSRA without any withdrawal movement of the assembly.

The pressure-containing components are designed to meet the requirements of the ASME Code,Section III, Nuclear Vessels, for Class A vessels (summer 1967 adden-dum). The material and structural design is the same as for the shim safety drive.

3.2 REACTOR DESIGN 3.2.1 GENERAL

SUMMARY

The important core design, thermal, and hydraulic characteristics are tabulated .s in Table 3-1.

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D-B Table 3-1 i

Core Design, Thermal, and Hydraulic Data Reactor Type Pressurized Water Rated Heat Output, MWt 2,633 Vessel Coolant Inlet Temperature, F 557 Vessel Coolant Outlet Temperature, F 608 Core Outlet Tempi >rature, F 609.9 Operating Pressure, psig 2 J85 Core and Fuel Asyemblies Total Number of Fuel Assemblies in Core 177 Number of Fuel Rods per Fuel Assembly 208 Number of Control Rods per Control Rod Assembly 16 Number of In-Core Instrumentation Positions per Fuel Assembly 1 Fuel Rod Outside Diameter, in. 0.h30 Clad Thickness, in.

0.0265 Fuel Rod Pitch, in.

0.568 Fuel Assembly Pitch Spacing, in. 8.587 Unit Cell Metal / Water Ratio (Volume Basis) 0.82 Clad Material Zircaloy-h (Cold-Worked)

Fuel Material UO 2

Form Dished-End Cylindrical Pellets Diameter, in.

0.370 Active Length, in. '

1hh Density, ". of Theoretical 93.5 Heat Transfer and Fluid Flov at Rated Power Total Heat Transfer Surface in Core, ft2 gg,73g AverageHeagFlux, Btu /h-ft

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175,811 Maximum Heat Flux, Btu /h-ft2 538,730 Average Power Density in Core, kW/l 85.5 Average r=al Output, kW/ft of Fuel Rod 5.80 3-7

D-B Table 3-1 (Cont'd) A

)

Maximum Thermal Output, kW/ft of Fuel Rod 17.8 Maximum Clad Surface Temperature, F 650 Average Core Fuel Temperature, F 1,560 Maximum Fuel Central Temperature at Hot Spot, F h,h6h 6

Total Reactor Coolant Flow, lb/h 131.32 x 10 Core Flow Area (Effective for Heat Transfer), ft2 h9,19 Core Coolant Average Velocity, fps 15.7 Coolant Outlet Temperature at Hot Channel, F 648.8 Power Distribution Maximum / Average Power Ratio, Radial x Local (Fah Nuclear) 1.T5 Maximum / Average Power Ratio, Axial (F Nuclear) 1.70 Overall Power Ratio (F Nuclear) 2.98 Power Generated in Fuel and Cladding, % 97.3 Hot Channel Factors Power Peaking Factor (Fq) N 1.011 j Flow Area Reduction Factor (FA Interior Bundle Cells 0.98 Peripheral Bundle Cells 0 97 Local Heat Flux Factor (F ,,) 1.01h Hot Spot Maximum / Average Heat Flux Ratio (F nue and mech) 3.06 DNB Data Design Overpower Ratio 1.12 DNB Ratio at Design Overpower (W-3) 1.50 DNB Ratio at Rated Power (W-3) 1.92 3.2.2 4 NUCLEAR DESIGN AND EVALUATION -

The basic design of the core satisfies the following requirements:

a. Sufficient excess reactivity is provided to achieve the design power 1,y,Vgl,,overthespecifiedfuelcycle. .

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D-B J~ b. Sufficient reactivity control is provided to permit safe reactor ep-eration and shutdown at all times during ccre lifetime.

3.2.2.1 Nuclear Characteristics of the Design 3.2.2.1.1 Excess Reactivity The nuclear design characteristics are given in Table 3-2. The excess reac-tivities associated with various core conditions are tabulated in Table 3-3.

The core will operate for h33 full-power days for the first cycle and will have a 292 full-power day equilibrium cycle. Design limits will be held with re-spect to reactivity control and power distribution. In-core instrumentation will be used to indicate power peaking levels. Single fuel assembly reactivity information is also included in Table 3-3.

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Table 3-2 Nuclear Design Data Fuel Assembly Volume Fractions Fuel 0.303 Moderator 0 580 Zircaloy 0.102 Stcinless Gteel 0.003

!' Void ,

0.012 1.000 Total UO2 (BOL, First Core), Metric Tons 9h.5 Core Dimensions, in.

Equivalent Diameter Active Height 128.9 ihh.0 Unit Cell H O2 to U Atomic Ratio (Fuel Assembly)

Cold Hot 2.85 2.0h Full-Power Lifetime, Days I

First Cycle 433 j Each Succeeding Cycle 292 j Fuel Irradiation, mwd /MTU l

First Cycle Average 13,686 Succeeding Cycle Average

'l65 9.230 s_

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I D-B Table 3-2 (Cont'd)

Feed Enrich =ents, vt % U-235 First Cycle 2.32/2.32/2.68 (by Zene)

Control Data Control Rod Material Ag-In-Cd Number of Full-Length Control Rod Assemblies h9 Number of Xenon (Part-Length) Control Rod Assemblies 8 Total Full-Length Control Rod Worth (ak/k), ". 8.0 Control Rod ladding Material Type 30h SS Table 3-3 Excess Reactivity Conditions Effective Multiplication, kerr Cold, Zero Power, No Burnable Poison 1.276 Hot, Zero Power, No Burnable Poison 1.222 Hot, Rated Power, No Burnable Poison 1.20h Hot, Rated Power, With Burnable Poison 1.119 Hot, Equilibrium Xenon, Rated Power, With Burnable Poison -

1.088 Single Fuel Assembly Hot 0.77 Cold 0.87 (a)First cycle at beginning of life (BOL).

Based on highest probable enrichment of 3.5 veight per cent (c)A center-to-center assembly pitch of 21 inches is required for this k in cold, nonborated water with no xenon or samarium.

77 The minimum critical mass, with and without xenon and samarium poisoning, may be specified as a single assembly or as multiple assemblies in various geo-metric arrays.

purposes. The unit fuel assembly has been investigated for co=parative A single cold, clean assembly containing a maximum probable. enrich-ment of 3.5 weight per cent is suberitical. Two assemblies side-by-side are supercritical except when both equilibrium xenon and sama;-lum are present.

Three assemblies side-by-side are supercritical with both equilibrium xenon and samarium present. -

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D-B 3.2.2.1.2 Reactivity control Distribution Control of excess reactivity is shown in Table 3 h.

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Table 3-h First Cycle Reactivity Control Distribution

% ak/k

1. Controlled by Soluble Boron

a. Moderator Temperature Deficit (70 to 520 F) 3.h

b. Equilibrium Xenon and Samarium 3.5

c. Fuel Burnup and Fission Product Buildup 75

2. Controlled by Burnable Poison Rod Assemblies (BPRA)

Fuel Burnup and Fission Product Buildup 6.0

3. Controlled by Inserted Control Rod Assembl(eg Transient Xenon (Normally Inserted) 0.8

h. Controlled by Movable Control Rod Assemblies (CRA)

a. Doppler Deficit (0 to 100% Rated Power) 1.2

b. Moderator Temperature Deficit (0 to 15% Power at End of Life, 520 to 582 F) 0.8

c. Dilution Control 0.2

d. Shutdown Margin 1.0 Total Movable Control Worth Required 3.2

5. Available Control Rod Assembly Worths

a. Total CRA Worth 8.0

b. Stuck Rod Worth (Rod of Highest Reactivity Value) 2.5

c. Minimum Available CRA Worth 5.5 I

d. Minimum Movable CRA Worth Available h.5 J

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D-B Explanation of Items in Table 3 h

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1. Control by Soluble Boren Baron in solution is used to control'the following relatively slow-moving reactivity changes:

a. The moderator deficit in going from ambient to operating temperatures.

The value shown is for the maximum change which would occur toward the end of the cycle.

b. Equilibrium xenon and samarium.

c. The excess reactivity required for fuel burnup and fission product buildup throughout cycle life.

Figure 3-1 shows the typical variation in boron concentration with life for Cycle 1 and the equilibrium cycle.

2. Centrol by Burnable Poison The 16 control rod holes in 72 of the fuel assemblies not equipped with control rod assemblies will be utilized as locations for burnable poison rods. The 72 element locations are shown in Figure 3 k.

3. Control by Inserted CBA Sufficient rod worth remains inserted in the core during normal opera-tion to overcome the peak xenon transient following a power reduction of 50 per cent of rated power for 90 per cent of the fuel cycle. This override capability facilitates the return to normal operating condi-tions without extended delays. The presence of these rods in the core during operation does not produce power peaks above the design value, and the shutdown margin of the core is not adversely affected. Axial power peak variation, resulting from partial or full insertion of xenon override rods, is described fully in Figures 3-2 and 3-3 The loes of movable reactivity control due to the insertion of this group produces no shutdown difficulties and is reflected in Table 3-5 2

e 168 Amendment No. 2 3-12

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h. Control by Movable CRA 2

(*

\ a. Power level changes (Doppler) and regulation.

b. Between 0 and 15 per cent or rated power, reactivity compensation by CRA may be required as a result of the linear increase of reactor cool-ant temperature from 520 F to the normal operating value,

c. Additional reactivity is held by a group of partially inser ed CRA (25 per cent insertion maximum) to allow periodic rather tha's centin-uous soluble boron dilution. The CRA are inserted to the 25 per cent limit as the boron is diluted. Automatic withdrawal of these CRA dur-ing operation is allowed to a 5 per cent insertion limit where the di-lution procedure is again initiated and this group of CRA is reinserted.

d. A shutdown margin of 1 per cent ak/k below the hot critical condition is also considered as part of the reactivity controlled by CRA.

5 Available Rod Worth A to'.a1 of 3.2 per cent ak/k(*) is required in movable control. Analysis of the 49 full-length CRA under the reference fuel arrangement predicts a total CRA vorth of at least 8.0 per cent ak/k. The stuck-out CRA vorth was also evaluated at a value no larger than 2.5 per cent ak/k.(**) Thia evaluation included selection of the highest worth CRA under the first CRA-out condition. The minimum available CRA vorth of h.5 per cent ak/k(*) is sufficient to meet movable control requirements. The eight axial power shaping (APSRA) rods are vorth from 0.2 to 0.h per cent ak/k. This value is not included in the 8.0 per cent ok/k rod worth reported above.

3.2.2.1.3 Reactivity Shutdown Analysis The ability to shut down the core under hot conditions is illustrated in Table 3-5 In this tabulation both the first and equilibrium cycles are evaluated at the beginning of life (BOL) and the end of life (EOL) for shutdown capability.

Examination of Table 3-5 for Minimum Hot Shutdown Margin (Item h) shows that, with the highest worth CRA stuck out, the core can be maintained in a suberiti-cal condition. Normal conditions indicate a minimum hot shutdown margin of 5.5 per cent ak/k at the end of life.

Under conditions where a cooldown to containment vessel ambient temperature is l6 o required, concentrated soluble boron vill be added to the reactor coolant to produce a shutdown margin of at least 1 per cent ak/k. Beginning-of-life boron levels for several core conditions are listed ip Table 3-6 along with boron worth values. The conditions shewn with ne ' '. A 11ustrate the highest require-ments. _

169 Does not include transient control. See Table 3 I+.

First cycle. See Table 3-h.

3-13 Ame:ndment No. 6 3 (, . ti 2

s Table 3-5

~-

Shutdown Reactivity Analysis

,y First Cycle Equilibrium .

Reactivity Effects, % ak/k BOL EOL BOL EOL

1. Maximum Shutdown CRA Requirement Doppler (100 to 0% Power) 1.2 15 1.2 1.5 Moderator Deficit (15 to 0% Power) 0.0 0.8 0.0 0.8 Total 1.2 2.3 1.2 2.3

2. Maxi =um Available CRA Worth ("} 8.0 8.0 8.0 8.0 Transient Xe Insertion Worth 0.8 0.0 0.8 0.0 Possible Dilution Insertion 0.2 0.2 0.2 0.2

3. Minimum Available CRA Worth All CRA In 7.0 7.8 7.0 7.8 One CRA Stuck-Out(b) h.5 5.3 h.5 5.3

h. Minimum Hot Shutdown Margin All CRA In 5.8 5.5 5.8 55 One CRA Stuck-Out 3.3 3.0 3.3 3.0

" Total worth of 49 CRA.

(b)CRA of highest reactivity value. _s e n

D-B

( Table 3-6 Soluble Boron Levels and Worth (First Cycle)

BOL Boron Core Conditions Levels, orm

1. Cold, k,ff = 0.99 No CRA In 1,316 All CRA In 836 One Stuck CRA 986

2. Hot, Zero Power, k,ff = 0.99 No CRA In 1,28h All CRA In 48h

One Stuck CRA T34

3. Hot, Rated Power, k eff

= 1.00 No CRA In 1,061

4. ~

Hot, Equilibrium Xe and Sm, Rated Power, k,ff = 1.00 No CRA In Til s 5. Boron Worth (f. ak/k)/ ppm Hot Cold 1/100 1/75 3.2.2.1.4 Reactivity Coefficients Reactivity coefficients form the basis for digital studies involving normal and abnormal reactor operating conditions. These coefficients have been in-vestigtsted as part of the analysis of this core and are described below as to function and overall range of values.

a. Doppler Coefficient The Doppler coefficient reflects the change in reactivity as a func-tion of fuel temperature. A rise in fuel temperature results in an increase in the effective absorption cross section of the fuel (the Doppler broadening of the resonance peaks) and a corresponding re- 1 duction in neutron production. The range for the Doppler coefficient i under operating conditions is expected to be -1.1 x 10-5 to -1.7 x 4,',,;1p?I(ak/k)/F.

b. Moderator Void Coefficient i

D ' ' .

l Thel mode' r ~htor void coefficient relates the change in neutron multi-l l

l[} ph cag m .torthe presence,of voids in the. moderator. The expected l

.msnm 3-15 ,

l

D-B range for ',ne void coefficient is -h.0 x 10-k to -3.0 x 10-3 (ak/k)/

per cent void. '

, c. Moderator Pressure Coefficient The coderator pressure coefficient relates the change in moderator density, resulting from a reactor coolant pressure change, to the

orresponding effect on neutron production. This coeffici.ent is

pposite in sign and considerably smaller when compared to the mod-erator temperature coefficient. A typical range of pressure coeffi-cients over a life cycle vould be +h x 10-7 to +3 x 10-6 (ak/k)/ psi,

d. Moderator Temperature Coefficient The moderator temperature coefficient relates a change in neutron multiplication to the change in reactor coolant temperature. Reac-2l tors using soluble boron as a reactivity control have less negative moderator temperature coefficients than do cores controlled solely by movable or fixed CRA. The major temperature effect on the cool-ant is a change in density. An increasing coolant temperature pro-duces a decrease in water density and an equal percentage reduction in boron concentration. The concentration change results in a posi-tive reactivity component by reducing the absorption in the coolant.

The magnitude of this component is proportional to the total reactiv-ity held by soluble boron.

w The moderator temperature coefficient has been calculated for three )

conditions of the hot, clean (no xenon) core with 1,061 ppm boron in '

the moderator.

Zero Power a =0 15 Per Cent Power a = .3 x 10 (ak/k)/F 100 Per Cent Power og = .h x 100 (ak/k)/F Since equilibrium xenon is covered by soluble poison, it follows that at the full-power condition with xenon, the moderator tempera- l l

ture coefficient will be even more negative than shown above.

Th 10-h (ak/k)/F at the end of the equilibrium fuel cycle.derator temper I

e. pH Coefficient Currently, there is no definite correlation which will permit predic-tion of pH reactivity effects. Some of the parameters needing corre-lation are the effects relating pH reactivity change for various op-erating reactors, pH effects versus reactor operating time at power, and changes in effects with varying clad, temperature, and water chemistry. Yankee, Saxton, and Indian Point Station.No. 1 have j

_/

5. .

Amendment No. 2 g 3-16

D-B l'

s experienced reactivity changes at the time of pH changes, but there is no clearcut evidence that pH is the direct reactivity influencing variable without considering other items such as clad materials, fuel assembly crud deposition, system average temperature, and prior sys-tem vater chemistry.

The pH characteristic of this design is shown below in Table 3-7 where the cold values are measured and the hot values are calculated.

Table 3-7 cH Characteristics Boron Concentration, 7 Li, com Tmid' F tra cH Units 0,5 70 1,800 5.0 2.0 70 1,800 5.6 0.5 580 1,200 7.0 2.0 580 1,200 7.5 0.5 580 17 7.2 2.0 580 17 7.8 0.5 70 17 7.9 2.0 70 17 8.5 Saxton experiments have indicated a pH reactivity effect of 0.0016 Ap/ApH unit change with and without locci boiling in the core. Con-sidering the system makeup rate of 35,000 lb/h and the core in the hot condition with 1,200 ppm boron in the coolant, the corresponding changes in pH are 0.02 pH units per hour for boron dilution and 0.05 pH units per hour forILi dilution (starting with 0.5 ppm TLi). Ap-plying the pH vorth value quoted above from Saxton, the total reac-tivity insertion rate for the hot condition is 3.1 x.10-8 Ap/s. This f2 insertion rate of reactivity can be easily compensated by the opera-tor or the automatic control system.

3 . 2. 2.1. 5 Reactivity Insertion Rates Figure 7-6 displays the integrated rod wor $h' 6f three 2.erlapping rod banks as a function of distance withdrawn. The indicated groups are those used in the core during power operation. Using approximately 1.2 per cent Ak/k CRA groups and a 30 in./ min drive speed in conjunction with the reactivity response given in Figure T-6 yields a maximum reactivity insertion rate of 1.1 x 10-0 ( Ak/k)/s. ~

The maximum reactivity insertion rate for soluble boron removal is b.h x 10-6 (Ak/k)/s.

3.2.2.1.6 Power Decay Curves

(' Figure 3-5 displays the beginning-of-life power decay curves for the CRA vorths corresponding to the 1 per cent hot shutdown margin with and without a stuck

][] O ,

3-17 Amendment No. 2 4 4'-

D-B rod. The power decay is initiated by the trip release of the CRA with a 300 /~')

ms delay from initiation to start of CRA motion. The time required for 2/3 v/

rod insertion is 1.h s.

3.2.2.2 Nuclear Evaluation Analytical models and the application of these models are discussed in this section. Core instabilities associated with xenon oscillation are also de-

~

scribed, with data evaluated under reference conditions.

3.2.2.2.1 ,Analytical Models Reactor design calculations are made with a large nu=ber of computer codes.

The choice of vi.ich code set or sets to use depends on which phase of the de-sign is being analyzed. A list of codes used in core analysis with a brief discussion follows in 3.2.2.2.2.

a. Reactivity Calculations Calculation of the reactivity of a pressurized water reactor core is performed in one, two, or three dimensions. The geometric choice depends on the type of calculations to be made. In a clean type of calculation where there are no strong, localized absorbers of a type differing from the rest of the lattice, 1-dimensional analycis is satis factory. This type of problem is handled quite well by the B&W 1-dimensional depletion package code LIFE. LIFE is a composite of MUFT (Ref 3), KATE (Ref h), RIP, WANDA (Ret 5), and a depletion '$

~,j routine. Normally, the MUFT portion is used with 3h-energy groups, an exact treatment of hydrogen, the Greuling-Goertzel approximation for elements of mass less than 10, and Fermi age for all heavier elements. The KATE portion normally uses a Wigner-Wilkins spectrum.

In WANDA, k-energy groups are utilized. Disadvantage factors for input to the thermal group are calculated with the THERMOS (Ref 6) code. This code set has been shown to give reliable results for a reactivity calculation of this type. Recent check calculations on critical experiments have a standard deviation of less than 0 5 per cent Ak/k.

A 1-dimensional analysis of a geometric arrangement, where there are localized strong absorbers such as CRA, requires a preliminary 2-dimensional analysis. The required properties of the 1-dimensional system are then matched to the 2-dimensional analysis. In this man-ner, it is possible to analyze the simpler 1-dimensional system in a depletion survey probl'm e with only a small loss in accuracy.

The 1-dimensional calculations are used as preliminary guides for _

the more detailed 2-dimensional analysis that follows. Values of reactivity coefficients, fuel cycle enrich =ents , lifetimes , and soluble poison concentrations can be found to improve the initial conditions specified for 2-dimensional analysis.

Two-dimensional reactivity calculations are done with either the PDQ ;

(Ref 7) or TURBO (Ref 8) diffusion and/or depletion codes. These . ./

.,- j-18 ,

D-B

( codes have mesh limitations on the size of a configuration which can be shown explicitly and are often studied with quarter core symmetry.

Symmetry is desirable in the design, and no loss in generality occurs.

^

The geometric description includes each fuel assembly and as much detail as is possible, i.e. , usually each unit in the fuel assembly.

Analysis of this type permits detailed power distribution studies as well as reactivity analysis. The power distribution in a large PWR core .which has zone loading cannot be predicted reliably with 1-di-mensional calculations. This is particularly true when local power peaking as a function of power history is of interest. It is nec-essary to study this type of problem with at least a 2-dimensional code, and in some cases, 3-dimensional calculations are necessary.

Use of the 2-dimensional programs requires the generation of group constants as a function of material composition, power history, and geometry. For regions where diffusion theory is valid, MUFT and KATE with THERMOS disadvantage factors are used to generate epither-mal and thermal coefficients. This would apply at a distance of a few mean paths from boundaries or discontinuities in the fuel rod lattice. Discontinuities refer to water channels, instrumentation ports, and CRA guide tubes. The interfaces between regions of dif-ferent enrichment are considered to be boundaries as well as'the outer limit of the core.

To generate coefficients for regions where diffusion theory is inap-propriate several methods are utilized. The arrangement of structural material, water channels and adjacent fuel rod rows can be repre-sented well in slab geometry. The coefficients so generated are utilized in the epithermal energy range. Coefficients for the ther-mal energy range are generated by a slab THERMOS calculation. The regions adjacent to an interface of material of different enrichment are also well represented with the P3MG code.

The arrangement of instrumentation ports and control rod guide tubes lends itself to cylindrical geometry. DTF-IV (Ret 9) is quite ef-fective in the analysis of this arrangement. Input to UTF-IV is from GAM (Ref 10) and THERMOS and KATE. Iteration is required be-tween the codes. The flux shape is calculated by DTF-IV and cross sections by the others. The outer boundary of the core where there is a transition from fuel to reflector and baffle is also represented by the DIF-IV code. The 3-dimensional analysis is accomplished by extending the techniques of 2-dimensional representation.

b. Control Rod Analysis B&W has developed a procedure for analyzing the reactivity worth of ~

small Ag-In-Cd rods in fuel lattices. Verification of this procedure vas made by the epmparative analysis of lh critical experiments with varying rod and rod assembly configurations.(ll,12) Critical lattice geometries were similar to those of the reference core design. Boron concentration' ranged from 1,000 to 1,500 ppm. The Ag-In-Cd rods were arranged in various geometrical configurations which bracket the ref-erence' design. Water holes, simulating withdrawn rods, were included

.. B _eam .... I75

D-B as part of the lattice study. The resulting comparison of the ana-lytical and experimental worths are shown in Table 3-8. Details of

<]

)

the critical configurations are given in References 11 and 12.

Table 3-8 Calculated and Experimental Rod and Rod Assembly Comparison Ag-In-Cd Rod Assembly - Rod Assembly -

Core Assemblies Rods Per H2O Holes Calculated Experimental No. Per Core Assembly Per Core Worth, % Ak/k V L eth, % Ak/k 5-B h 4 252 2.00 1 98 h-F h 9 0 3.3F 3.3h 5-C 2 12 276 2 38 2.35 h-D 1

16 0 1.h3 1.h2 5-D 2 16 284 2.80 2.82 h-E 1 20 0 1.5h 1.52 5-E 2 20 292 3.05 3.01 The mean error in calculating these configurations is shown to be less than 1 per cent. Comparison of the power shape associated with the 16-rod reference assemblies showed good similarity. Point-to-average power had a maximum variation of less than 2 per cent with experimental data.

The analytical method used for this analysis is based on straight dif-fusion theory. Thermal coefficients for a control rod are obtained from THERMOS by flux-weighting. Epithermal coefficients for the upper energy groups are generated by the B&W LIFE program. The resulting coefficients are used in the 2-dimensional code PDQ to obtain the re-g*1 red eigenvalues.

GAKER and LIBPM are used to prepare data for THERMOS. GAKER generates scattering cross sections for hydrogen by the Nelkin technique. LIBPM uses the Brown and St. John free gas model for generating the remain-ing scattering cross sections.

THERMOS is used in two steps. First, the critical fuel cell is ana-lyzed to obtain a velocity-veighted disadvantage factor. This is used in the homogenization of fuel cells and gives a first order correc-tion for spatial and spectral variation. The ratio of flux in the moderator to flux in the fuel was analyzed to within 2 per cent of experimental values using the velocity-weighting technique. The see- -

ond step is to use THERMOS in a calculation where the Ag-In-Cd rod is surrounded by fuel. This is used to generate the flux-weighted con-trol rod cell coefficients as a function of boron concentration. As a check on the validity of the THERMOS approach, extrapolation dis-tances were compared to those given by the Spinks method.(13) The ' -

agreement was within 2.2 per cent for a set of cases wherein the

.gegg. .

176

- if _

3-20

D-B i

( number densities of Ag-In-Cd were varied in a range up to 250 per cent. All other coefficients are generated by LIFE in much the same manner as with THERMOS. The data are used in a 2-dimensional PDQ 1ayout where each fuel rod cell is shown separately.

c. Determination of Reactivity Coefficients

)

This . type of calculation is different from the reactivity analysis only in application, i.e., a series of reactivity calculations being '

required. Coefficients are determined for moderator temperature, voiding, and pressure, and for fuel temperature. These are calcu-lated from small perturbations in the required parameter over the range of possible values of the parameter.

The moderator temperature coefficient is determined as a function of soluble poison concentration and moderator temperature, and fuel tem-perature or Doppler coefficient as a function of fuel temperature.

The coefficient for voiding is calculated by varying the moderator concentration or per cent void.

3.2.2.2.2 Codes for Reactor Calculaticas This section contains a brief description of codes mentioned in the preceding sections.

THERMOS (Ref 6) - This code solves the integral form of the Boltzmann

( Transport Equation for the neutron spectrum as a function of position.

A diagonalized connection to the isotropic transfer matrix has been incorporated allowing a degree of anisotropic scattering.

MUFT (Ref 3) - This program solves the P1 or B1 multigroup equation for the first two Legendre coefficients of the directional neutron flux, and for the isotropic and anisotropic components of the slowing down densities due'to a cosine-shaped neutron source. Coefficients are generated with MUFT for the epithermal energy range.

KATE (Ref h) - The code solves the Wigner-Wilkins differential equation for a homogeneous medium moderated by chemically unbound hydrogen atoms in thermal equilibrium. Coefficibnts for the thermal energy range are generated by KATE.

RIP - This program averages cross sections over an arbitrary group struc-ture, calculates resonance integrals for a set of resolved petAs, and computed L-factors for input to MUFT, P1MG, and P3MG.

WANDA (Ref 5) - This code provides numerical solutions of the 1-dimensional ~

fev-group neutron diffusion equations.

LIFE - This is a 1-dimensional depletion package code which is a combina-tion of MUPf, KATE, RIP, and.WANDA. The combination mechanizes the procedures for using the codes separately.

(

T7;7 m :.

3-21

D-B GAM (Ref 10) - This code is a multigroup coefficient generation program 7;7g that solves the P1 equations and includes anisotropic scattering.

j Inelastic scattering and resonance parameters are also treated by GAM.

P3MG (Ref 2) - The code solves the multienergy transport equation in various geometries. The code is primarily used for epithermal coef-ficient generations.

DTF (Ref 9) - This code solves the multigroup, 1-dimensional Boltzmann transport equation by the method of discrete ordinates. DTF allows multigroup anisotropic scattering as well as up and down scattering.

PDQ (Ref 7) - This program solves the 2-dimensional neutron diffusion-depletion problem with up to five groups. It has a flexible repre-sentation of time-dependent cross sections by means of fit options.

TURBO (Ref 8) - This code is similar in application to the PDQ depletion program. It, however, lacks the great flexibility of the PDQ fit

. options.

CANDLE (Ref 8) - This code is similar to TURBO, but solves the diffusion equations in one dimension.

TNT (Ref 8) - This code is similar in application to TURBO, but is a 3-dimensional code extended from DRACO.

3.2.2.2.3 A Xenon Stability Analysis _/

Model and digital analysis of the core indicated that the tendency towards xenon instability in the axial mode would exist for a given combination of ever.ts.

Therefore, eight part-length Axial Power Shaping Rod Assemblies (APSRA's) have been included in the design. They will be positioned during operation to maintain an acceptable distribution of power for any particular operating condition in the core, thereby reducing the tendency for axial oscil-lations.

The azimuthal stability of the core is dependent on core loading, power density, and the moderator temperature coefficient. In any event, the core vill not be susceptible to diverging azimuthal oscillations. If the loading and power density are lov enough the core vill be inherently stable. If not, then burnable poison is added in the amount necessary to provide the moderator temperature coefficient that vill result in azimuthal stability.

A detailed description of the modal analysis performed on the core is found in B&W Topical Report EAW-10010 " Stability Margin for Xenon Oscillations - Modal Analysis." Digital calculations are continuing and further analysis will be -

reported in Topical Report form. A brief description of the digital method is presented below.

's 178 _

.m ,

g

D-B Digital Analysis Xenon stability studies are continuing with codes in various geometries which have thermal-nuclear iteration capability for both fuel and moderator tempera-ture feedback.

The one-dimensional fsedback code LIFE handles the iteration in the following manner:

Z out ATi = (Tout -T in) = c f PD(Z)dZ (A) 3 in where ATg = temperature change in region "i" PD(Z) = power density in the Z direction Zg,Zout " #*6 * "" * **

and AT c

H (B)

[PD(Z)df o

k. where H is the actise fuel height.

Equation (A) is solved to Tout of region "1." Since Tin is known from core inlet conditions, the average fluid temperature is defined as follows:

T +T

= ut in Tf ]uid 2 i

(C)

The newly computed, region-averaged fluid temperatures are used to compute new fluid densities. These fluid densities are then used to ad,just the number densities for water and acluble poison. Local or bulk boiling is not pemitted in the model, but incorporation would increase stsbility. The average fuel temperature for each axial core region is then computed from the average fluid temperature and power densities as follows:

fuel g * '( i

+T fluid g i

where 5 = average power density of region "i" f(Ep) *;=4' tabular function relating fuel temperature increase and power density obtained from

( auxiliary calculations .

. 179 g

3-23

)

D-B After the new fluid temperatures, moderator densities, and fuel temperatures ]

are obtained, these quantities are used as new LIFE input to obtain a new pown _,,e distribution until a convergence criterion is met.

This analysis used an exact solution in that the spectrum was recalculated for each zone (11 axial zones described the reactor) for each iteration at evcry time step. Thi:: included the effects of the moderator temperature coefficient.

This LIFE packag'e was used to determine the effects of the uncertainty in the power Doppler on the stability of the core. The uncertainty in the Doppler was more than compensated with a reduction in fuel temperature of 500 degrees.

The reference core was analyzed with core average fuel temperatures of 1,h00 F and 900 F. Figure 3-6 compares the cyclic response of these two cases follow-ing the 3-ft insertion and removal (after two hours) of a 1.2 per cent ak/k ro(.

bank near the beginning of life. These studies were made at beginning-of-life boron levels of approximately 1,900 ppm. This level is approximately 800 ppm above the predicted beginning-of-life level and, consequently, reflects a posi-tive moderator te=perature coefficient which is not expected.

Case 1 on Figure 3-6 depicts the behavior of the core if the meat transfer equa-tions were not included in the calculation. Figure 3-7 shows the bffect of fuel temperature toward the end of life. It is easily verified that the 900 F fuel temperature case approached the threshold condition for axial oscilla-tion in this core.

The one-dimensional model was used to determine a method of controlling the core without taking into account the stabilizing effect of the power Doppler. ]

Normally, this would produce a divergent oscillation as shown in Figure 3-8. ' ' ,

A study was completed wherein a 1 per cent ak/k rod bank with a 3-ft-long see-tion of regular control rod material was succes' fully maneuvered to control the core after a perturbation of the power shrpe at a point about 3/h of the 2 vay through Cycle 1. The controlled results .re also shown in Figure 3-8.

The minimum rod motion was 1 foot, and the time step employed wr'.s 4.8 hours0.333 days <br />0.0476 weeks <br />0.011 months <br />.

More precise rod movement over shorter time periods would produce a much smoother power ratio curve. This control mechanism appears to be quite adequate.

Stability in "R-Z" or "X-Y" gecmetry is studied with the HARMONY code, which in either ca.=e can be used with fuel and moderator temperature feedback. This code is used with fitted coefficients to obtain a more complete solution to the perturbed behavior of the reference design.

The digital results are processed by first fitting. power distribution results to an equation of the following form by a least square technique:

2 P' = Ae* sin where P' e m ss power

= stability index T = oscillation period ^bV

) ')

v Amendment No. 2 Y$...$' 3-2h

D-B rh The calculated stability index is then extrapolated to zero-length time steps by the procedures of Reference lh as follows:

T -

T

~

Z=Z ~

1+* ~2 T

r ,

T = <A> T r

<A> = A x +o4 xo where Z = extrapolated stability index T = time step length

$ = average ther=al flux Reasonable agreement has been found between the initial modal analysis and one-dimensional results.

Control of an Axial Osullation The one-dimensional model was used to determine a method of controlling the axial power distribution without considering the stabilizing effects of ther-mal-nnelear feedback. Normally, this calculation vould show a divergent oscil-lation. A study was completed in which a 3-ft section of control rod poison material was successfully maneuvered to control the core after introduction of a perturbation. The minimum rod motion was 12 in. with a h.8-hour minimum period for rod movement. More precise movement over a shorter minimum period would improve control. The procedure vill be extended to two dimensions.

Conclusions The following conclusions have been made as a result' of the studies to date:

a. Instability in the radial direction vill not occur.

b. The core design under examination vill not be susceptible to diverg-ing azimuthal oscillations (Figure 3-9). *

c. Potential axial oscillations vill be thwarted or controlled by the part length control rods.

181 4 '[.. .

@yd g 3-25

D-B 3.2.3 THERMAL AND HYDRAULIC DESIGN AND EVALUATION 3.2.3.1 Thermal and Hydraulic Characteristics 3.2.3.1.1 Fuel Assembly Heat Transfer Design

a. Design Criteria The criterion for the heat transfer design is to be safely below Departu:e From Nucleate Boiling (DNB) at the design overpower (112 per cent of rated power). The analysis is described in detail in 3 2.3.2.2, Statistical Core Design Technique.

The input information for the statistical core design technique and for the evaluation of individual hot channels is as follows:

1. Heat transfer critical heat flux equations and data correlations.

2. Nuclear power factors.

3. Engineering hot channel factors. .

h. Core flow distribution hot channel factors.

5 Maximum reactor overpower.

These inputs have been derived from test data, physical measurements, m and calculations as outlined below.

b. Heat Transfer Equation and Data Correlation

]

The heat transfer relationship used to predict limiting heat transfer conditions is presented in References 15 and 16. The equations are as follows:

1. W-3 uniform flux DNB correlation for single channel with all valls heated:

NB,eu

= {(2.022 - 0.000h302P) 10 6

^ (0.1722 - 0.000098hP) exp (18.177 - 0.00hl329P)x }

C x{0.1h8h-1596x+0.1729xlxl + 1.037 f 2 10 x {1.157 - 0.869x } -

x { 0.266h + 0.8357 exp (-3.151D,)f x{0.8258+0.00078h(H,, .H )f J

Amendment No. 2 3-26

D-B

, . . where Q" = flux, Btu /h-ft 2 P = pressure, psia p

G = mass velocity, lb/h-ft -

X = quality, expressed as fraction D = equivalent diameter, in.

H = enthalpy, Btu /lb

2. W-3 nonunifom flux DNB correlation for single channel with all valls heated:(15)

SbHB,N *SbNB,eu!

where B'N

= DNB heat flux for the nonuniformly heated channel QbNB,eu"'4" * *" "" " '#

C 7

"Q{ocal" DNB [1-eXP(-CA DNB DNB x< [ Q"(Z) exp -C(E DNB~

o C = 0.hh in.~

G j l.72 (10)6 where t

B

= distance from the inception of local boiling to the point of DNB.

Z = distance from the inception of local boiling, measured in the direction of flow.

l83 s-m M

3-27

D-B

3. W-3 un valls:g flux DNB correlation for single channel with 2nhented .m lih kbNB.withunheatedvall b[NB,usingD to replace D

= (1.36 + 0.12 e9X)

-1.92D x (1.2 - 1.6 e h) x (1.33 - 0.237 e s.66x) .

D* = equivalent diameter based on all the vetted perimeter, in.

D = equivalent diameter based on only the heated perimeter, in.

Individual channels are analyzed to determine a DT) ratio, i.e., the ratio of the heat flux at which a DNB is predicted to occur to the heat flux in the channel being investigated. This DNB ratio is re-lated to the data correlation as shown in Figure 3-10. A confidence and population value is associated with every DNB ratio as described in the Statistical Core Design Technique (3.2.3.2.2). A plot of DNB versus P is for a confidence of 99 per cent. The criterion for evaluating the thermal design margin for individual channels or the .m total core is the confidence-population relationship. The DNB ra-ties required to meet the basic criteria or limits are a function of 7g the experimental data and heat transfer correlation used, and vary with the quantity and quality of data. The recommended minimum de-sign DNB ratio for the W-3 correlation is 1.30.

The DNB and population relationship for a design limit of 1.30 in the hot channel corresponds to a 99 per cent confidence that at least 94.3 per cent of the population of all such hot channels is in no jeopardy of experiencing a DNB. The DNB ratios and the fraction of the core in no jeopardy of experiencing a DNB at design conditions are considerably higher than those given in the design limits out-lined in 3.1.2.3 The relationship of sustained DNB to core burnout conditions is dis-cussed in B&W Company Topical Report BAW-10001h, " Analysis of Sustained Departure From Nucleate Boiling Operation." The analysis shows that DNB conditions are very local and there is little likelihood of sudden fuel rod failures or a propagation of DNB conditions.

c. Nuclear Power Factors ~

The heated surfaces in every flow channel in the core are examined for heat flux limits. The heat input to the fuel rods in a coolant channel is determined from a nuclear analysis of the core and fuel assemblies. The results of this analysis are as follows:

\SA

y 3-28

t D-B

1. The nominal nuclear peaking factors for the worst time in core

[' life are Fah =.1.Th Fz = 1.70 Fq = 2.96

2. The design nuclear peaking factors for the worst time in core life are Fah = 1.75 Fz = 1 70 Fq = 2.98 Fah = max / avg total power ratio (radial x local nuclear)

Fz = max / avg axial pcwer ratio (nuclear)

Fq = Fah x Fz (nuclear total) 2 The nominal values are the maximum values calculated with nominal spacing of fuel assemblies. The design values are obtained by ex-amining maximum, nominal, and minimum fuel assembly spacing and de-termining the worst values for the comained effect of flow and rod 7 peaking.

The axial nuclear factor, Fz, is illus,trated in Figure 3-11. The distribution of power expressed as P/P is shown for two conditions of reactor operation. The first condition is an inlet peak with a max / avg value of 1.70 resulting from partial insertion of a CRA group for transient control following a power level change. This condition results in the maximum local heat flux and maximum linear heat rate. The second power shape is a symmetrical cosine which is indicative of the power distribution with xenon override rods with-drawn. The flux peak max / avg value is 1.50 in the center of the active core. Both of these flux shapes have been evaluated for thermal DNB limitations. The limiting condition is the 1.50 cosine power distribution. The inlet peak shape has a larger maximum value.

However, the position of the 1.50 cosine peak farther up the channel results in a less favorable flux to enthalpy relationship. This effect has been demonstrated in DNB tests of nonuniform flux shapes. [y7)

The 1.50 cosine axial shape has been used to determine individual channel DNB limits and to make the associated statistical analysis.

The nuclear factor for total radial x local rod power, Fah, is calcu- -

lated for each rod in the core. A distribution curve of the fraction of the core fuel rods operating above various peaking factors is shown in Figure 3-12 for a typical fuel cycle condition with the maximum fuep god peaking factor of 1.75 i ' 185

} .;. '~

>' 3 il 3-29 Amendment No. 2

D-B

d. Engineeri g Hot Channel Factors

( }

Power peaking factors obtained from the nuclear analysis are based on mechanically perfect fuel assemblies. Engineering hot channel factors are used to describe variations in fuel loading, fuel and clad dimen-sions, and flow channel geometry from perfect physical quantities and dimensions.

The application of hot channel factors is described in detail in 3.2.3.2.2, Statistical Core Design Technique. The factors are de-termined statistically from fuel assembly as-built or specified data where Fq is a heat input factor, Fq" is a local heat flux factor at a hot spot, and FA is a flow area reduction factor describing the varia-tion in coolant channel flow area. Several subfactors are combined statistically to obtain the final values for Fq, Fq", and FA. These subfactors are shown in Table 3-9 The factor, the coefficient of variation, the standard deviation, and the mean value are tabulated.

l Table 3-9 '

Coefficients of Variation CV No. Description o x CV 1 Flow Area q Interior Bundle Cells 0.00190 0.177h'0 0.01072 y Peripheral mund 26 Cells 0.003h6 0.215h6 0.01608 2 Local Rod Diameter 0.0006h7 0.430 0.00151 3 Average Rod Diameter 0.0006h7 0.h30 0.00151 (Die-Drawn, Local and Average

.Same) h Local Fuel Leading 0.00698 Subdensity 0.006h7 0.935 0.00692 l Subfuel Area (Diameter 0.00009h 0.1075 0.00088 l Effect) 5 Average Fuel Loading 0.00557 Subdensity 0.00h85 0.935 0.00519 Sublength 0.2529h 1h4 0.00183 Subfuel Area (Diameter 0.00009h 0.1075 0.00088 Effect) 6 Local Enrichment 0.00h21 2.30 0.00183

~

7 Average Enrichment 0.00h21 C.30 0.00183 CV Coefficient of Variatiop o/5 (Enrichment values are for vorst case , .

o Standard Deviation of Variable normal assay batch; maximum variation x Mean Value of Variable urs for minimum enrichment.)~ } 8[J O '~

3 ? 3-30

D-B

e. Core Flev Distribution Hot Channel Factors The physical arrangement of the reactor vessel internals and nozzles results in a nonuniform distribution cf coolant flow to the various fuel assemblies. Reactor internal structures above and below the active core are designed to minimize unfavorable flow distribution.

A 1/6 scale model test of the reactor and internals was performed to demonstrate the adequacy of the internal arrangements. The results of the test have confirmed the adequacy of the design values used.

Model test results are given in B&W Topical Report BAW-10012.

A flow distribution factor is determined for each fuel assembly loca-tion in the core. The factor is expressed as the ratio of fuel as-sembly flow to average fuel assembly flow. The finite values of the ratio may be greater or less than 1.0 depending on the position of the assembly being evaluated. The flow in the central fuel assem-blies is in general larger than the flov in the outermoht assemblies due to the inherent flow characteristics of the reactor vessel.

The flow distribution factor is related to a particular fuel assembly location and the quantity of heat being produced in the assembly. A flow-to-power comparison is made for all of the fuel assemblies. The worst condition in the hottest fuel assembly is determined by apply-ing model test isothermal flow distribution data and heat input ef-fects at power as outlined in 3.2.3.2.3. Two assumptions for flow distribution have been made in the thermal analysis of the core as follows:

1. For the maximum design condition and for the analysis of the hottest channel, all fuel assemblies receive minimum flow for the worst condition, regardless of assembly power or location.

2. For the most probable design conditions predicted, average flows have been assigned for each fuel assembly consistent with loca-tion and power. The flow factor assumed for the maximum design condition is conservative. Application of vessel flow test data and individual assembly flow factors in the detailed core design vill result in improved statistical statements for the maximum design condition.

f. Maximum Reactor Design Overpower Core perfordhnce is assessed at the maximum design overpower. The selection of the design overpower is based on an analysis of the re-actor protection system as described in Section T. The reactor trip point is 105.5 per cent rated power, and the maximum overpower, which is 112 per cent, vill not be exceeded under any conditions. "

g. Maxi d$' Design Conditions Analysis Su--nry The Statistical Core Design Technique described in 3.2.3 . .as used to analyze the reactor at the maximum design cond:*cns v scribed previously. The total number of fuel rods in the cure thac

.have a possibility of reaching DNB is shown in Figure 3-13 for 00

', y. i o, 3-31 lg7

D-B to 120 per cent overpower for the maximwn design conditions. Point

~

B on Line 1 is the maximum design point for 112 per cent power with the design Fah nuclear of 1 75 and minimum flow to every channel in II')

't '

the core. This Point B forms the basis for this statistical state-ment:

There is a 99 per cent confidence that at least 99 94 per cent of the fuel rods in the core are in no jeopardy of experiencing a departure from nucleate boiling (DNB) during continuous opera-tion at the design overpower of 112 per cent.

At 100 per cent power (2,633 MWt) as shown by Point A, the statisti-cal number of fuel rods in jeopardy is 1.6, resulting in a population protected of 99 996 per cent. The limit imposed by a W-3 DNB ratio of 1.3 is 70 fuel rods in jeopardy or a population protected of 99.810 per cent.

An additional analysis of fuel rods in jeopardy for all the maximum design conditions except fuel assembly flow distribution is shown as Line 2 in Figure 3-13. Each assembly was assumed to receive average flow for the assembly power conditions. The final consideration of flow distribution will result in values within the bounds of the two lines. Statistical results for the maximum design condition calcu-lation, as shown by Figure 3-13, may be summarized as follows in Table 3-10.

m Table 3-10 DNB Results - Maximum Design Ccndition (99% Centidence Level)

Population Hot Channel Power, % Possible Protected, DNB Ratio Point of 2,633 MWt Fah DNB % (W-3)

A 100 1.75 1.6 99.996 1.92 B 112 1.75 20.0 99.9h6 1.50 C 119 1 75 70.0 99.810 1.30 l

h. Most Probable Design Condition Analysis Summary The previous maximum design calculation, as shown by Line 1, Figure 3-13, indicates the total number of rods that may be in jeopardy when it is conservatively assumed that every rod in the core has the mechanical and heat transfer characteristics of a hot channel as de-scribed in 3.2.3.2.2. For example, all channels are analyzed with F A (f1 v are'a factor) less than 1.0, Fq (heat input factor) greater than 1.0, and with minimum fuel assembly flow. It is physically im- T possible for all channels to.6avy. hot channel characteristics. A )

D-B m more realistic indication of the number of fuel rods in jeopardy may be obtained by the application of the statistical heat transfer data to average rod power and mechanical conditions.

An analysis for the most probable conditions has been made based on the average conditions described in 3.2.3.2.2. The results of this analysis are shown in Figure 3-lh. The analysis may be su m rized as follows in Table 3-11.

Table 3-11 DNB Aeeults - Most Probable Condition Population Hot Channel Power, % Possible Protected, DNB Ratio Point of 2,633 MWt Fah DNB % (W-3)

D 100 1.75 <1 >99.997 2.20 E 112 1.75 2.7 99 992 1.78 F 120 1.75 9.5 99.974 1.52 The analysis was made from Point D at 100 per cent power to Point F at 120 per cent power to show the sensitivity of the analysis with power. The worst condition expected is indicated by Point E at 112 per cent power where it is shown statistically that there is a small possibility that 2.7 fuel rods may be subject to a departure from nucleate boiling (DNB). This result forms the basis for the follow-ing statistical statement for the most probable design conditions:

There is at least a 99 per cent confidence that at least 99.992 per cent of the rods in the core are in no jeopardy of experiencing a DNB, even with continuous operation at the design overpower of 112 per cent.

i. Distribution of the Fraction of Fuel Rods Protected The distribution of the fracon (P) of fuel rods that have been shown statistically to be in no jeopardy of a DNB has been calcu-lated for the maximum design and most probable design conditions.

The computer programs used proV4de an output of (N) number of rods and (P) fraction of rods thati will not experience a DNB grouped for ranges of (P). The re$d1ts for the most probable design con-dition are shown in Figure 3.-15 -

The population protected, (P), and the population in jeopardy, (1-P),

are both plotked. The integral of (1-P) and the number of fuel rods gives the numb *er of rods that are in jeopardy for given conditions as shown in Figures 3-13"ind 3-14. The number of rods is obtained from the product of the percentage times the total number of rods being considered (36,816). A typical distribution shown in Figure 3-15 l2 e

~

l 9 3-33 6

t Amendment No. 2

D-B 2 is for the most probable condition analysis of Point E on Figure 3-lh '3 The line shows P and (1-P) at the llE per cent power condition repre- ./

sented by Point E. The integral of N and (1-P) of the curve forms the basis for the statistical statement at the most probable design condition described in (h) above.'

j. Hot Channel Performance Summary for Four Pump Operation The hottest unit cell with all surfaces heated has been examined for hot channel factors, DNB ratios, and quality for a range of reactor powers. The cell has been examined for the maximum value of Fah nuclear of 1.75 The hot channel was assumed to be located in a fuel assembly with 95 per cent of the average fuel assembly flow.

The heat generated in the fuel is 97.3 per cent of the total nuclear heat. The remaining 2.7 per cent is assumed to be generated in the coolant as it proceeds up the channel within the core and is reflected as an increase in AT of the coolant.

Error bands of 65 psi operating pressure and 12 F are reflected in the total core and hot channel thermal margin calculations in the direction producing the lowest DNB ratios or highest qualities. The DNB ratio versus power is shown in Figure 3-16. The DNB ratio in the hot channel at the maximum overpower of 112 per cent is 1.50 which corresponds to a 99 per cent confidence that at least 99.00 per cent of the fuel channels of this type are in no jeopardy of experiencing a DNB. The engineering hot channel factors used in the design analy- %

sis are described in 3.2.3.2.2 and listed below: y Fq = 1.011 F ,, = 1.01h q

F = 0.98 (Interior cells)

FA = 0 97 (Wall cells)

The hot channel exit quality for various powers is shown in Figure 3-17 The combined results may be su=marized for 2,120 psig as follows:

Reactor Power, % DNB Ratio (V-3) Exit Quality, %

~

100 1.92 1.9 105.5 (Trip Setting) 1.72 3.9 112 (Maximum Power) 1 50 6.'6

~'

119 1.30 9.6 128 1.00 13.7 x-

' g, ' %

,;Q Amendment No.-2 3-3h

D-B

(~

\

k. Hot Channel Performance Summary for Partial Pump Flow The power limitations imposed on the reactor due to the loss of one

' or more pumps has been examined by studying DNB ratios and quality in the hot channel for a range of reactor powers. The system param-eters used in the analysis are the same as those discussed in 3.2.3.1.1 J.

A constant reactor vessel temperature at 582.5 F was used to deter-mine inlet temperature. The DNB ratio versus power for the flow con-ditions caused by loss of pu=ps is illustrated in Figure 3-18. The hot channel quality at the minimum DNB ratio point (XDNB) at the same conditions is shown in Figure 3-19.

l2 Two limits have been placed on the analysis. One is the recommended minimum DNB ratio of 1.3 and the other is the quality limits of plus or minus 15% at the point of minimum DNB ratio. Table 3-12 below summarizes the power limitations on reactor operation at 2,120 psig as defined by the hot channel conditions. The overpower margins and system limitations are discussed in h.3.7 Table 3-12 Hot Channel Performance Vs Pumos in Service Reactor Coolant 2 Pumps

/ Pu=us Ocerating 3 Pumns (2 Loops)

Hot Channel DNBR @ Maximum 1.30 1.36 Design Overpower Hot Channel Quality at Minimum 6.95 15.0 DNBR Point, %

Reactor Coolant Flov, % of Rated Th.7 h9.0 Maximum Design Overpower, % of 98.9 75.0 Rated Power (2,633 MW)

.3. 2 . 3 .1. 2 FuelandCladdgsgThermalConditions

a. Fuel A digital computer code is used to calculate the fuel temperature.

The program uses nonuniform volumetric heat generation across the fuel diameter, and external coolant conditions and heat transfer coefficients determined for thermal-hydraulic channel solutions.

The fuel thermal conductivity is varied in a radial direction as S a function of the temperature variation. Values for fuel conduc-I tivity were used as shown in Figure 3-20, a plot of fuel conductivity versus temperature. The heat transfer from the fuel to the clad is 3-35 A Amendment No. 2

D-B calculated with a fuel and clad expansion model proportional t. tem- ,

peratures. The temperature drop is calculated using gas conductivity fg 3 at the beginning-of-life conditions when the gas conductivity is 0.09 >

i Btu /ft-h-F-ft 2. The gas conduction model is used in the calculation until the fuel thermal expansion relative to the clad closes the gap to a dimension equivalent to a contact coefficient. The contact coef-

~

ficient is dependent upon pressure and gas conductivity.

A plot of fuel center temperature versus linear heat rate in kW/ft is shown in Figure 3-21 for beginning-of-life conditions. The linear heat rate at the maximum overpower of 112 per cent is 19.9 kW/ft.

The corresponding center fuel temperature shown in Table 1-2 is 4,720 F. The center and average temperatures at 100 per cent power are h,h6h and 1,560 F as shown in Table 3-1.

The peaking factors used in the calculation are Fah = 1.75 Fz = 1.70 Fn q = 1.03 Fq (nue and mech) = 3.06 A conservative value of 1.03 was assu=ed for the heat flux peaking factor, Fqn. The assigned value corresponds to a 99 per cent populc- "'

tion-protected relationship as described in the statistical technique. ,f

b. Clad The assumptions in the preceding paragraph were applied in the cal-culation of the clad surface temperature at the maximum overpower.

Boiling cond tigns prevail at the hot spot, and the Jens and Lottes relationship 201 for the coolant-to-clad AT for boiling was used to determine the clad temperature. The resulting maximum calculated clad surface temperature is 65h F at a system operating pressure of 2,185 psis.

3.2.3.2 Thermal and Hydraulic Evaluation 3.2.3.2.1 Introduction Su= mary results for the characteristics of the reactor design are presented in 3.2.3.1. The Statistical Core Design Technique employed in the design repre-sents a refinement in the methods for evaluating pressurized water reactors.

Corresponding single hot channel DNB data were presented to relate the nev -

=ethod with previous criteria. A co=prehensive description of the new tech-nique is included in this section to per=it a rapid evaluation of the methods used.

A detailed evaluation and sensitivity analysis of the design has been made by s

! examining the hottest channel in the reactor for DNB ratio, quality, and fuel temperatures. The W-3 correlation has been used in this analysis. -/

3-36

,_,, }92

D-B

,_ 3.2.3.2.2 Statistical Core Design Tachnique k

The core thermal design is based on a Statistical Core Design Technique de-veloped by B&W. The technique offers many substantial improvements over older methods, particularly in design approach, reliability of the result, and math-ematical treatment of the calculation. The method reflects the performance of the entire core in the reaultant power rating and provides insight into the reliability of the calculation. This section discusses the technique in order to provide an understanding of its engineering merit.

The statistical core design technique considers all parameters that affect the safe and reliable operation of the reactor core. By considering each fuel rod, the method rates the reactor on the basis of the performance of the entire core.

The result then vill provide a good measure of the core safety and reliability since the method provides a statistical statement for the total core. This statement also reflects the conservatism or design margin in the calculation.

reactor safe operating power has always been de+ ermined by the ability of the coolant to remove heat from the fuel material. The criterion that best measures this ability is the DNB, which involves the individual parameters of heat flux, coolant temperature rise, and flow area, and their intereffects. The DNB cri-terion is commonly applied through the use of the departure from nucleate boil-ing ratio (DNBR). This is the minimum ratio of the DNB heat flux (as computed by the DNB correlation) to the surface heat flux. The ratio is a measure of the margin between the operating power and the power at which a DNB might be expected to occur in that channel. The DNBR varies over the channel length,

(

and it is the minimum value of the ratio in the channel of interest that is used.

The calculation of DNB heat flux involves the coolant enthalpy rise and cool-ant flow rate. The coolant enthalpy rise is a function of both the heat in-put and the flow rate. It is possible to separate these two effects; the statistical hot channel factors required are a heat input factor, F flow area factor, F A. In addition, a statistical heat flux factor,q, and a Fqn is re-quired; the heat flux factor statistically describes the variation in surface heat flux. The DNBR is most limiting when the burnout heat flux is based on minimum flow area (small FA ) and maximum heat input (large Fq), and e en the surface heat flux is large (large FQ "). The DNB correlation is provided in a best-fit form, i.e. , a form that best fits all of the data on which the cor-relation is based. To afford protection against DNB, the DNB heat flux com-

,uted by the best-fit correlation is divided by a DNB factor (BF) greater than 1.0 to yield the design DNB surface heat flux. The basic relationship DNBR=kNB 1 BF Q

Q"surface xFnQ 1

l l

involves as parameters statistical hot. channel and DNB factors. The Dhu rac-tor (BF) above is usually assigned a value of unity when reporting DNB ratios so that the margin at a given condition is.shown directly by a DNBR greater

[ than 1.0, i.e. ,1.30 in the ho{ channel. l

, 6. y 3-37 dd .

~

D-B Selected heat transfer data are analyzed to obtain a correlation. Since ther-mal and hydraulic data generally are well represented with a Gaussian (normal) 7 distribution (Figure 3-22), mathematical parameters that quantitatively rate ' -=

the correlation can be easily obtained for the histogram. These sa=e mathe-matical parameters are the basis for the statistical burnout factor (BF).

In analyzing a reactor core, the statistical infer =ation required to describe the hot channel subfactors may be obtained from data on the as-built core, from data on similar cores that have been constructed, or from the specified tolerances for the proposed core. Regardless of the source of data, the sub-ractors can be shown graphically (Figures 3-23 and 3-24).

All the plots have the same characteristic shape whether they are for subfac-tors , hot channel factors, or burnout factor. The factor increases with either increasing population or-confidence. The value used for the statistical hot channel and burnout factor is a function of the percentage of confidence de-sired in the result, and the portion of all possibilities desired, as well as the amount of data used in determining the statistical factor. A frequently used assumption in statistical analyses is that the data available represent an infinite sample of that data. The implications of this assumption should be noted. For instance, if limited data are available, such an assumption leads to a somewhat optimistic result. The assumption also implies that more infor-mation exists for a given sample than is indicated by the data; it implies 100 per cent confidence in the end result. The B&W calculational procedure does not make this assumption, but rather uses the specified sample size to yield a result that is much more meaningful and statistically rigorous. The influ-ence of the amount ci data for instance can be illustrated easily as follows: ,

Consider the heat flux factor which has the form ,

s Fn q = 1 -+ Kc Fn q

where Fn q = statistical hot channel factor for heat flux K = statistical multiplying factor F " = standard deviation of the heat flux factor, Q including the effects of all the subfactors If cF Q" = 0.05 for 300 data points, then a K factor of 2.608 is required to protect 99 per cent of the population. The value of the hot channel factor then 16 Fnq = 1 + (2.608 x 0.050) = 1.130h and vill provide 99 per cent confidence for the calculation. If, instead of using the 300 data points, it is assumed that the data represent an infinite ~s sample, then the K factor for 99 per cent of the population is 2.326. The value of the hot channel factor in this case is kh 3-38 .

M

D-B D F ,, = 1 + (2.326 x 0.050) = 1.1163 i S which implies 100 per cent confidence.in th9 c41culation. The values of the K factor used above are taken from SCR-607.l19) The same basic techniques can te used to handle any situation involving variable confidence, population, and number of points.

Having established statistical hot channel factors and statistical DNB factors, we can proceed with the calculation in the classical manner. The statistical f actors are used to determine the minimum fraction of rods protected, or that are in no jeopardy of experiencing a DNB at each nuclear power peaking factor.

Since this fraction is known, the maximum fraction in jeopardy is also known.

It should be recognized that every rod in the core has an associative DNB ratio that is substantially greater than 1.0, even at the design overpower, and that theoretically no rod can have a statistical population factor of 100 per cent, no matter how large its DNB ratio.

Since both the fraction of rods in jeopardy at any particular nuclear power peaking factor and the number of rods operating at that peaking factor are known, the total number of rods in jeopardy in the whole core can be obtained by simple su=mation. The calculation is made as a function of power, and the plot of rods in jeopardy versus reactor overpower is obtained (Figure 3-13).

The su=mation of the fraction of rods in jeopardy at each peaking factor su=med over all peaking factors can be made in a statistically rigorous manner only if the confidence for all populations is identical. If an infinite sample is not assumed, the confidence varies with population. To form this su=mation then, a cons'ervative assumption is required. B&W total core model assumes that the confidence for all rods is equal to that for the least-protected rod, i.e.,

the minimum possible confidence factor is associated with the entire calcula-tion.

The result of the foregoing technique, based on the maximum desien conditions (112 per cent power), is this statistical statement:

There is at least a 99 per cent confidence +. hat at least 99 9h per cent of the rods in the core are in no jeopardy of experiencing a DNB, even with continuous operation at the design overpower.

The maximum design condikions are represented by these assumptions:

a. The maximum design values of Fah (nuclear max / avg total fuel rod heat input) are obtained by examining the maximum, nominal, and mini-mum fuel assembly spacing and determining the worst value for rod peaking.

b. The maximum value of2F (nuclear max / avg axial fuel rod heat input) is determined for the limiting transient or steady-state condition.

c. Every coolant channel in the core is assumed to hrie less than the nominal flow area which is represented by engineering hot channel l2 factors, F , less than 1.0.

A 195 m

.$ U - 3-39 Amendment ho. 2

D-E

d. Every channel is assumed to receive the minimum flow associated with ;

core flow maldistribution. "

e. Every fuel rod in the core is assumed to have a heat input greater than the maximum calculated value. This value is represented by

' engineering hot channel heat input factors, Fq and Fqn, which are greater than 1.0.

f.

Every channel and associated fuel rod has a heat transfer margin above the experimental best-fit limits reflected in DNB ratios greater than 1.0 at maximum overpower conditions.

The statistical core design technique may also be used in a similar manner to evaluate the entire core at the most probable mechanical and nuclear conditions to give an indication of the most probable degree of fuel element jeopardy.

The result of the technique based on the most erobable design conditions leads to a statistical statement which is a corollary to the maximum design statement:

There is at least a 99 per cent confidence that at least 99.992 per cent of the rods in the core are in no jeopardy of experiencing a DNB, even with continuous operation at the design overpower.

The most probable design conditions are assumed to be the same as the maximum design conditions with these exceptions:

a. Every coolant channel is assumed to have the nominal flow area w (FA = 1.0). 1 J

b.

Every fuel rod is assumed to have (1) the maximum calculated value of heat input, and (2) Fq and Fq" are assigned values of 1.0.

c.

The flow in each coolant channel is based on a power analysis with-out flow caldistribution factors, d.

Every fuel rod is assumed to have a nominal value for Fah nuclear.

The full meaning of the maximum and most probable design statements requires additional co==ent. As to the 0.06 per cent or 0.008 per cent of the rods not included in the statements, statistically, it can be said that no more than 0.06 per cent or 0.008 per cent of the rods will be in jeopardy, and that in gen-eral the number in jeopardy vill be fewer than 0.06 per cent or 0.008 per cent.

The statements do not mean to specify a given number of DNB, but only acknowl-edge the possibility that a given number could occur for the 112 per cent overpower conditions assumed.

Analyses for'100 per cent rated power conditions reported in Tables are subject to a DNB.

3-10 and 3-11 show that essentially none of the fuel rods ,

In summary, the calculational procedure outlined here represents a substantially improved design technique in two ways:

, a. It reflects the perfor=ance and safety of the entire core in the re-l sultant power rating by considering the effect of each rod in the ,

power rating. ,

}hO -

k 3-ho

D-B p b. It provides information on the reliability of the calculation and,

\ therefore, the core through the statistical statement.

3.2.3.2.3 Evaluation of the Ther=al and Hydraulic Design

a. Hot Channel Coolant Quality and Void Fraction An evaluation of the hot chanel coolant cenditions provides additional confidence in the thermal design. Sufficient coolant flow has been provided to insure low quality and void fractions. The quality in the hot channel versus reactor power is shovu in Figure 3-17. The sensitivity of channel outlet quality with pressure and power level is shown by the 2,185 and 2,120 psig system pressure conditions examined. These calculations were made for an Fah of 1 75. Addi-tional calculations for a 10 per cent increase in Fah to 1.93 were made at 112 per cent power. The significant results of both calcu-lations are summarized in Table 3-13.

1 Table 3-13

) Hot Channel Coolant Conditions l Exit Exit Void Operating Power, ". Fah Quality, ". Fraction, ", Pressure, nsig 100 1.75 (-)0.6(") 3.h 2,185 r 112 1.75 k.1 23.7 2,185 119 1 75 65 29.8 2,185 112 1.93 11.2 h2.0 2,185 100 1.75 1.9 10.7 2,120 112 1.75 6.6 30.0 2,120 119 1 75 9.6 38.5 2,120 j 112 1.93 14.0 49.4 2,120 l

" Negative indication of quality denotes subcooling.

l l

The conditions of Table 3-13 were determined with all of the hot channel factors applied. Additional calculations were made for unit cell channels without engineering hot channel factors to show the coolant conditions more likely to occur in the reactor core. A value for Fah of 1.75,vas examined with and without fuel assembly flow distribution hot' channel factors at 2,185 psig as shown on Fig-ure 3-25 These results show that the exit qualities from the hottest cells should in general be considerably lower than the maximum de-sign conditions.

b. Core Void Fraction The core void fractions were calculated at 100 per cent rated power for the normal operating pressure of 2,185 psig and for the -inimum

^

operating pressure of 2,120 psig. The influene'e'_Af core fuel essenbly O

3 kl ] 9 j'

D-B flow distribution was checked by determining the total voids for both 100 and 95 per cent total core flow for the tvo pressure conditions. q]

.x The results are as follows:

Flow, % Pressure, usig Core Void Fraction, %

100 2,185 0.098 100 2,120 0.319 95 2,185 0.337 95 2,120 1.250 The most conservative conditioi. of 95 per cent flow at 2,120 psig results in no more than 1.25 per cent void volume in the core. Con-servative maximum design values were used to make the calculation.

The voi egram uses a combination of Bowring's(20) model with Zuber's2ge correlation between void fraction and quality. The Bowring model coneiders three different regions of forced convection boiling. They are: .

1. Highly Subcooled Boiling N

In this region, the bubbles adhere to the wall while moving up-ward through the channel. This region is terminated when the subcooling decreases to a point where the bubbles break through the laminar sublayer and depart from the surface. The highly subcooled region starts when the surface temperature of the clad reaches the surface temperature predicted by the Jens and Lottes equation. The highly subcooled region ends when sat - bulk "

$ = local heat flux, Btu /h-ft n = 1.863 x 10-3 (14 + 0.0068p)

V = velocity of coolant, ft/s p = pressure, psia The void f ac ion in this regicn is computed in the same =anner as Maurer, 22 except that the'end of the region is Gtermined by Equation (A) rather than by a vapor layer thickness. The nonequilibrium quality at the end of the region is computed from 's the void fraction as follows:

~mus 3-h2 -

198

._ n

D-B

^ 1 X*=-

(B) 1+ 1 -1 O

g "d wher X$=nonequilibriumqualityatendofRegion1 a '#" ""

d"* sat

~

bulk p = liquid component density, lb/ft o = vapor component density, lb/ft

2. Slightly Subcooled Boiling In this region, the bubbles depart from the wall and are trans-ported along the channel (condensation of the bubbles is ne-glected). This region transcends to a point where the ther=o-dynamic quality is equal to the apparent quality. In general, this is the region of ma,jor concern in the design of pressurized water reactors.

The nonequilibrium quality in this region is computed from the following for=ula:

P z

' X" = Xd+

ihfg (l + *} -[($-$SP)dz (C) 2

d X* = nonequilibrium quality in Region 2 hp = latent heat of vaporization, Btu /lb 1

, = fraction of the heat flux above the single phase heat flux that actually goes to prodacing voids

$gp = single phase heat flux, Btu /h-ft m = mass flow rate, lb/h Py = heated perimeter, ft z = channel distance, ft The void fraction in this region is computed from X"

38.3 A gg o (}

(, C '

'ogge (o f - og }l/4 2 X*

Og /Pr (1 - X") .

2 m o f ,

3-h3 Amendment No. 2 l

D-B 3

Where g = acceleration due to gravity, ft/s2 g =

constant in Newton's Second Law = 32.17 lbm/ft c -

lbf/s2 C = Zuber's distribution parameter Ag = flow area, in.2 a = surface tension Equation (D) results from rearranging equations found in Reference 26 and assuming bubbly turbulent flow in determining the relative velocity between the vapor and the fluid. Zuber has shown that Equation (D) results in a better prediction of the void fraction than earlier models based on empirical slip ratios.

3. Bulk Boiling In this region, the bulk temperature is equal to the saturation temperature, and all the energy transferred to the fluid results in net vapor generation. Bulk boiling begins when the thermo-dynamic (heat balance) quality, x, is greater than the nonequi-2 librium quality, X*. The void fraction in this region is com-puted using Equation (D) with the thermodynamic quality, X, re-placing X*. '

c. Coolant Channel Hydraulic Stability

]

Flow regime maps of mass flow rate and quality were constructed in order to evaluate channel hydraulic stability. The confidence in the design is based on a review of both analytical evaluations (25-28) and experimental results obtained in multiple rod bundle burnout t stg.

Bubble-to-annular and bubble-to-slug flow limits proposed by Baker 23, are consistent with the BW experimental data in the range of in-terest. The analytical limits and experimental data points have been plotted to obtain the maps of the four different types of cells in the reactor core. These are shown in Figures 3-26 through 3-29 The experimental data points represent the exit conditions in the various types of channels just previous to the burnout condition for a representative sample of the data points obtained at design operat-

, ing conditions in the nine rod burnout test assemblies. In all of the bundle tests, the pressure drop, flow rate, and rod temperature traces were repeatable and steady, and did not exhibit any of the characteristics associated with flow instability.

Values of hot channel mass velocity and quality at 112 per cent and 130 per cent power for oath nominal (=ost probable conditicns) and maximum design conditions are shown on the maps. The potential op-ersting points are within the botinds suggested by Baker. Experi-mental data points for.the' reactor geometry with much higher qualities than the operating cond'itions have not exhibited unstable character-istics.

g~ J, e,

Amendment No.'2 3-kh -

D-B

~

d. Hot Channel DNB Comparisons

{

DNB ratios for the hottest channel have been determined for the W-3 correlation, and the results are shown in Figure 3-16. DNB ratios are shown for the design 150- axial max /avs symmetrical cosine flux shape from 100 to lh0 per cent power. The W-3 DNB ratio at the maxi-mum design power of 112 per cent is 1 50. This compares with the suggested W-3 design value of 1.3. A ratio of 1.3 is reached at 119 per cent power at an exit quality of 9.6 per cent, which is within the prescribed quality limits of the correlation.

The sensitivity of the DNB ratio with F: nuclear was examined from 100 to lLO per cent power. The detailed results are labeled in Fig-ure 3-30. A cosine flux shape with an Fz of 1.80 and an Fah of 1.75 results in a W-3 DNB ratio of 1.30 at 109 per cent power. Corresponding results are shown for a value of Fz of 1.65 and for the design value l2 of 1.5 The influence of a change in Fah was determined by analyzing the hot channel for an Fah of 193. This value is 10 per cent above the maximum design value of 1 75 The resulting W-3 DNB ratio is 1.02 at 112 per cent power. This value is above the correlation best-fit values of 1.0 for the severe conditions assumed.

e. Reactor Flav Effects Another significant variable to be considered in evaluating the de-sign is the total system flow. Conservative values for system and reactor pressure drop have been determined to insure that the re-quired system flow is obtained in the as-built plant. The experi-mental programs previously outlined in 1.5.2 have confirmed the pres-sure drop and related pump head requirements. It is anticipated that the as-built reactor flow will exceed the design value and will lead to increased power capability.

The reactor core flov and power capability were evaluated by deter-mining the steady-state power DNB rates versus flow. Analyses were made for (a) variations of power capability with total reactor flow for a constant DNB ratio of 1.30, (b) DNB ratios for design flow with variations in hot channel mixing coefficients and (c) DNB ratios for gross flow variations of 10 per cent. The results are shown in Figures 3-31 and 3-32, For the analysis shown in Figure 3-31 for design hot channel condition, the flow was determined that would give a DNB ratio of 1.30 for a range of reactor powers. This analysis shows, for example, that a DNB ratio of 1.30 can be maintained in the hot channel at 112 per cent power with a total reactor flow of 119 x .

106 lb/h as compared with the available design flow of 131.3 x 106 lb/h. The results shown by Line 2 in Figure 3-32 are the DNB ratios for rated flow of 131.3 x 106 lb/h versus power. The limiting con-dition is 119 per cent power for a DNB ratio of 1.30. Lines 1 and 3 show the DNB ratios versus pover'where the total system flow has '

been varied by 210 per cent. Adequate DNB ratios can be maintained with a substantial reduction in reactor coolant system flow.

esma 20L,ont2.2

3-h5 _

e U, .* e

- D-B The foregoing. sensitivity analyses were made using a fuel assembly A

)

design mixing coefficient of 0.02. A sensitivity analysis for a '

range of coefficients was made for the rated flow condition. The results are shown by Lines k and 5 of Figure 3-32 and discussed in more detail in 3.2.3.2.3 J .

f. Reactor Inlet Temperature Effects The influence of reactor inlet temperature on power capability at de-sign flow was evaluated. A variation of 1 F in reactor inlet tem-perature vill result in a power capability change of 0.6 per cent at a given DNB ratio.

g. Fuel Temperature

1. Method of Calculation .

A fuel temperature and gas pressure computer code was developed to calculate fuel temperatures , expansion, densification, equiaxed and colu=nar grain growth, center piping of fuel pellets, fission gas release, and fission gas pressure. Program and data compari-sons were made on the basis of the fraction of the fuel diameter within these structural regions: ,

(a) Outer limit of equiaxed grain growth - 2,700 F.

_s (b) Outer limit of columnar grain growth - 3,200 F. '

(c) Outer limit of molten fuel (UO2) - 5,080 F.

Data from References 27 through 30 were used to compare calcu-lated experimentr_1 fractions of the rod in grain growth and cen-tral melting.

The radial expansion of the fuel pellet is computed from the mean fuel temperature and the average coefficient of lir. ear expansion for the fuel over the temperature range considered. This model combined with the model for calculating the heat transfer coeffi-cient was compared with the model developed by Notley, et al.t31) of AECL. The difference in fuel growth for the two calculation models was less than the experimental scatter of data.

The fuel may be divided into as many as 30 radial and 70 axial increments for the analysis. An iterative solution for the tem-perature distribution is obtained, and the thermal conductivity of the fuel is input as a function of temperature. The relative thermal expansion of the fuel and cladding is taken into account when determining the temperatdre drop across the gap between the fuel and cladding surfaces.

The temperature drop across the gap is calculated using a gap con-ductan'.e model based on the methods reported in Reference 34. ,)

l 9

.s . -

3-h6 ,

D-B O The model which has the capability of calculating gap conductance k before fuel sion of the to cladsuggested methods contact by asRoss welland as after Stoute.cent 33 et)is This an exten-fuel to clad heat transfer is a function of gap vidth, gas conductivity, mean conductivity of the interface materials, mean surface rough-ness, caterial hardness and fuel to clad contact pressure. Before total fuel to clad contact is made, a fraction of the fuel, based on fuel OD and gap size, is in contact with the clad. A constant contact pressure is app ed to this fraction to simulate the ef-fects of fuel cracking. 0-h2)

The analytical model co=putes the amount of central void expected whenever the temperature approaches the threshold temperature for -

fuel' migration, and readjusts the density according to the new geometry.

The program uses a polynomial fit relationship for fuel thermal conductivity The B&W reference design (35) curve illustrated in Figure 3-29 4 a modification of the relationship presented in GEAP-h624.t3 I The curve yields a conservative integrated ther-mal conductivity of 93 v/cm vitS relatively little increase in con-ductivity beyond 3,000 F.

2. Fuel Center Temperature Results The results of the analysis for center te=peratures with the methods i

described above are shown in Figure 3-33 and 3-3h for beginning-and end-of-life conditions. The beginning- and end-of-life gas conductivity values are 0.09 and 0.01 Btu /h-ft 2 -F. The design cold initial diametral clearance is 0.0085 inch which is based on the final design specified quality control measurements for fuel clad and UO2 pellet tolerances for a 99 per cent confidence and 95.per cent population relationship. The calculated end-of-life center fuel temperatures are higher than the beginning-of-life values up to 100 per cent power because of the reduction in the conductivity of the gas in the gap. The effect is apparent even though a con-tact condition prevails. The calculation includes the effect of fuel swelling due to irradiation and takes credit for the flux de- ;

pression in the center of the rod because of the self-shielding ef-feet of UO2 (nonuniform power generation).

The most conservative assu=ptions using the B&W design curve with relatively little increase in thermal conductivity above 3,000 F result in a design central fuel melting at about 23 kW/ft, which l

is 3 kW/ft higher than the maximum design value of 19.9 kW/ft at l 112 per cent power as shown in Figures 3-33 and 3-3h. l The transient analysis at accident and normal conditions have

,.been made using the design fuel thermal conductivity curve (Fig-

/ Ore 3-20) to reflect a conservative value for the maximum average temperaturg.and stored energy in the fuel. Use of this curve results in a higher temperature and, therefore, a lover Doppler coefficient, since it decreased with temperature. Thus, the

.k' resultant Doppler effect is also conservative.

3-h7

D-3

,.m 3 Eauilibrium Cycle Average Fuel Temperatures An analysis has been made to show equilibrium average fuel condi-tions in the core. A typical fuel cycle, end-of-life, condition was used to determine the fraction of fuel at a given average condition. The results are shown in Figure 3-35 where the ever-age temperature varies from 1,150 to 3,400 F, and the entire core avefage temperature is about 1,560 F. The bundle powers as shown in Figure 3-36 were used to obtain the fuel rod heat rates. A symmetrical cosine axial power distribution with a 1.50 max / avg value as shown in Figure 3-11 was used to predict the axial heat rate distribution. It was assumed that 97 3 per cent of the power is generated in the fuel. The fuel rods were divided into 14 axial and 10 radial segments to obtain the temperature distribu-tion for this analysis. The heat rate for every fuel rod in the core was increased by a local peaking factor of 1.05 to account for uncertainties in the calculation of local peaks.

The B&W design thermal conductivity was used to provide conserva-tive values for fuel conductivity. The maximum powers occurred in fuel assemblies with one and two cycles of operation as shown in Figure 3-36, and the asserblies with the highest burnup did not exceed 1.0h3 times the average power for the typical case analyzed. The results shown in Figure 3-35 were made by group-ing all seg=ents of fuel by temperature. Typical fuel-to-clad s heat transfer coefficients used were 380 and 640 Btu /h-ft2 for 1 6 and 10-kW/ft heat rates, respectively. The corresponding be- '

ginning-of-life coefficients are about 500 and 700 Stu/h-ft2 -F for 6 and 10-kW/ft heat rates.

h. Fission Gas Release The fission gas release is based on results re orted in GEAP-h596.(36)

Additional data from GEAP-431h,(37) AECL-603,( 8) and CF-60-12-14(39) have been compared with the suggested release rate curve. The re-lease rate curve (36) is representative of the upper limit of release data in the temperature region of most importance. An internal gas pressure of 3,300 psi is used to determine the fuel clad internal design conditions reported in 3.2.4.2, Fuel Assemblies.

The design values for fission gas relasse from the fuel and for the maximum clad 1 ternal pressure were determined by analyzing various operating co:.ditions and assigning suitable margins for possible in-creases in local or average burnup in the fuel. Adequate margins are provided without utilizing the initial porcsity voids present in the UO fuel. A detailed analysis of the design assumptions for fis-2 sion ges release, ana the relationship of burnup, fuel growth, and initial diametral clearance between the fuel and clad, are summarized in the following paragraphs. An evaluation of the effect of having the fuel pellet internal voids available as gas holders is also in-cluded.

2b _/

.m 3-48

D-B

[' l. Design Assumptions (a) Fission Gas Release Rates The fission gas release rate is calculated as a function of fuel temperature at 112 per cent of rated power. The pro-cedures for calculating fuel temperatures are discussed in 3.2.3.2.3 g. The fission gas release curve and the support-ing data are shown in-Figure 3-37 Most of the data are on or below the design release rate curve. A release rate of 51 per cent is used for the portion of the fuel above 3,500 F. The fuel temperatures were calculated using the B&W de-sign, fuel thermal conductivity curve which yields conserva-tively high values for fuel temperatures.

(b) Axial Power and Burnun Assumptions The temperature conditions in the fuel are determined for the most severe axial power peaking expected to occur. Two axial power shapes have been evaluated to determine the maximum release rates. These are 1.50 and 1.70 max / avg shapes as shown in Figure 3-11 and repeated as part of Fig-ure 3-38 of this analysis. The quantity of gas released is found by applying the temperature-related release rates to the quantities of fission gas produced along the length of the hot fuel rod.

The quantity of fission gas produced in a given axial loca-tion is obtained from reactor core axial region equilibrium burnup studies. Three curves showing the axial distribution of burnup as a local-to-average ratio along the fuel rod are shown in Figure 3-38. Values of 100, 300, sr.d 230 days of operation are shown.

The 930-day, or end-of-life condition, is .the condition with the maximum fission gas inventory. The average burnup at the end of life in the hot fuel rod is 38,150 mwd /MTU vhich has been determined as follows:

Calculated Hot Bundle Average Burnup, mwd /MTU 33,000 HotFuelRodBddiuhFactor 1.05 Margin for Calculation Accuracy 1.10 Hot Rod Maximum Average Burnup, mwd /MTU 38,150 s 205 i '

3-h9

D-B The local burnup along the length of the fuel rod is the prod-uct of the hot rod =aximum average value above and the local-

/T'g to-average ratio shown in Figure 3-38. The resulting hot rod local maxi =um burnup for the 930-day, end-of-life con-dition is about h2,000 mwd /MI'U. This is the maximum calcu-lated value. However, local values to 55,000 mwd /MTU have been evaluated to insure adequate local fuel cladding strength for possible increases in average or local burnup over the life of the fuel for various fuel management procedures.

(c) Hot Rod Power Assumptions The maximum hot rod power as a function of life of the fuel has been used to calculate the temperature conditions. A study of the power histories of all of the fuel assemblies (thrcugh 5 cycles) to equilibrium conditions shows that the power in the bundles varies from a =aximum of 1.75 times the average rod power to 1.22 times the average power at the end of life. The peak bundle ratio of 1.65 (1.75 + hot rod ratio) vill only occur during the first two fuel cycles when the fission gas inventory is less than the maximum value.

This results in a maximum linear heat rate of 19 9 kW/ft which corresponds to 112 per cent of the maximum thermal output (17.8 kW/ft) shown in Table 3-1.

)

. .)

(d) Fuel Growth Assumptions The fuel growth was calculated as a function vf burnup as indicated in 3.2.h.2.1. Fuel pellet dimensions in the ther-mal temperature and gas release models were increased to the end-of-life conditions as determined above.

(e) Gas Conductivity and Contact Heat Transfer Assumetions The quantity of fir ~ ton gas released is a function of fuel temperature. The temperatures are influenced by three fac-tors: (a) the conductivity of the fission gas in the gap between the fuel and clad, (b) the diametral clearance beg tween fuel and clad, and (c) the heat transfer conditions when the fuel expands enough to contr.ct the clad.

A gas conductivity of 0.01 Btu /h-ft -F based on h3 per cent release of fissicn gas at the end-of-life conditions was used in the analysis. Diametral clearances of 0.0025 to 0.0065 inch reflecting minimum and maximum clearances after fuel growth were analyzed. The contact heat transfer coef-ficients were calculated as suggested in Reference 36.

J

)

.m.

3-50 m ..

D-B

2. Sumary of Results The fission gas release rates were determined in the first eval-uation. Rates were found for various cold diametral clearances and axial power peaking and burnup shapes. The results are shown in Figure 3-39 The lowest curve is the expe.7ted condition for a 1.70 axial power shape with a 930-day axial burnup distribution as shown in Figure 3-38. The increase in release rate with dia-metral clearance results from the fact that the fuel temperature must be raised to higher values before contact with the fuel clad is made. The release rate at the minimum clearance of 0.00h5 inch is 15 5 per cent. This condition is equivalent to a 0.0025 inch gap after irradiation growth and producer the maximum clad stress (maximum sized pellets with minimum internal diameter clad-ding). The release rate of 30.6 per cent for the maximum dia-metral clearance of 0.0085 inch will not occur with the maximum tensile stress condition due to fuel crowth, since the fuel has more room to grow into the clearance.

Two additional cases were examined to check the s(nsitivity of the calculations to axial power and burnup shapes. The results are shown by the upper two curves in Figure 3-39 The top curve is a plot of the release rates when it is assumed that both the axial power and burnup inventory of fission gas are distributed with a 1 50 max /av6 ratio. Similar results are shown for the 1.70 max / avg power ratio with a 1.50 max / avg burnup ratio. These curves show the release rates expected are not strongly influenced by the various power and burnup shapes.

The second evaluation shows the resulting internal pressures due to the release of fission product gases. Plots of internct clad pressures for the expected 930-day axial burnup distribution and a 1 70 max / avg axial power shape are shown in Figure 3-40. The lover curve is a plot of internal gas pressure assuming that 6.5 per cent of the fuel volume is available to hold the released gas (open porosity). (A corresponding closed pore case is also shown.) The present design condition being used in clad-stress f)U!,(calculationsassumesaclosed pore condition with all released

~

f gas contained outside the f.el pellets in spaces between the ex-panded dished ends of the pellets, the radial gaps (if any), and the void spaces at the ends of the fuel rods. The effects of fuel densification and grain grovth described in 3.2.3.2.3 g are included in the analysis. The calculation of maximum pressure is also relatively insensitive to the axial burnup distribution as shown by the line in Figure 3 ho for a 1 50 =ax/ avg axial power and burnup shape. (This ccrresponds to a local burnup peak of 57,000 mwd /MTU.

The allowable design internal pressure of 3,300 psi is well above the maximum values of internal pressures calculated for open or closed pellet pores, and the maximum internal pressure should only occur with the maximum dismetral clearance condition. An increase in average fuel. bur ternal pressute des,nup can be tolerated within the prescribed in-ign limits.

IIIIIII --

207 3-51

I D-B It has been indicated in Reference 31 and in AECL-1598 that the UO2 fuel is plastic enough to flow under low stresses when the temperature is above 1,800 F. That fraction of the fuel below this temperature may retain a large portion of the original por-osity and act as a fission gas holder. The hottest axial loca-tions producing the highest clad stresses will have little if any fuel below 1,800 F. However, the ends of the fuel rods will have some fuel below thic temperature. The approximate fraction of tne fuel below 1,6% F at overpower for a 1 50 axial power shape is as follows for various cold diametral clea: ances.

Per Cent of Fuel Clearance, in. Below 1,800 F, ".

0.0045 78 0.0070 62 0.0095 32 The retention of fuel porosity in the low temperature and low burnup regions will result in modest reductions in internal gas pressure.

Gas pressures at rated and overpower conditiens are shown in Figure 3-40. The overpower condition is not expected to occur

except for brief periods during operating transients.

i. H f t Channel Factors Evaluation

1. Rod Pitch and Bowing A flow area reduction factor is determined for the as-built fuel assembly by takinr; channel flow area measurements and statisti-cally determining an equivalent hot channel flow area reduction factor. A fuel assembly has been measured, and the results are shown in Table 3-9 Interior channel 'neasurements and measure-ments of the channels formed by the outer =ost fuel rods with adjacent assemblies have been analyzed. Coefficients of varia-tion for each type of channel have been determined. In the analytical solution for a channel flow, each channel flow area is reduced over its entire length by the FA factors shown in Figure 3-23 for the desired population protected + at a 99 per cent confidence. The hot channels have_been analyzed using values fcr 95 per cent population protected, or FA in'the in-terior cells of C.98 and FA in the wall cells of 0.97 as listed in 3.2.3.1.1 j.

Special attention is given to the ' influence of water gap variation between fuel assemblies when determining rod powers. Nuclear analyses have been made for the nominal, maximum, and minimum

, spacing between adjacent fuel assemblies. The nominal and maxi-mur hot assembly fuel rod powers are shown in Figures 3-41 and -

3-h2. The hot channel nuclear power factor (Fah nuclear) or l 1.75 shown in 3.2.3.1. 1 is based on Figure 3 42 for the worst8 l 3-52

D-B vater gap between fuel assemblies. The factor of 1.75 is a prod- { uct of the hot assembly factor of 1.65 times the 1.061 hot rod factor. This power factor is assigned to the hottest unit cell rod which is analyzed for burnout. Peaking factors for other channels are obtained in a similar manner. In all cases, the combined flow spacing and power peaking producing the lovest DNB ratio is used.

2. Fuel Pellet Diameter, Density, and Enrichment Factors Variations in the pellet size, density, and enrichment are re-flected in coefficients of variation Numbers 2 through 7 of Table 3-9 These variations have been obtained from the mea-sured or specified tolerances and combined statistically as de-scribed in 3.2.3.2.2 to give a power factor on the hot rod. For 99 per cent confidence and 95 per cent population conditions,'

this factor, Fq, is 1.011 and is applied as a power increase over the full length of the hot channel fuel rod. The local heat flux factor, F qn , for similar conditions is 1.01h.. These hot channel values are shown in Table 3-1. The corresponding values of Eq and FQ " with 99 99 per cent population protected are 1.025 and 1.03, respectively. A conservative value of F " Q of 1.03 for 99 per cent confidence and 99.99 per c ent population is used for finding the maximum fuel linear heat rates as shown in 3.2.3.1.2. ' These factors are used in the direct solution for channel enthal-pies and are not expressed as factors on enthalpy rise as is aften done. The coefficients of variation vill be under con-tinous review during the final design and development of the fuel assembly.

3. Flow Distribution Effects Inlet Plenum Effects The final inlet plenum effects vill be determined from the 1/6 scale model flow test now in progress. It has been assumed that the flow in the hot bundle position is 5 per c3nt less than aver-age bundle flow under isothermal conditions co: responding to the model flow test conditions. An additional redaction of flow due to hot assembly power is described below.

Redistribution in Adjacent Channels to Dissimilar Coolant f Conditions The hot fuel assembly flow is less than the flow through an aver- " age assembly at the same core pressure drop because of the in-creased pressura drop associated with a higher enthalpy and quality condition. This effect is allowed for by making a direct calculation for the hot assembly flow. The combined effects of upper and lower plenum flow conditions and heat input u the hot 209 3-53

s D-B 7 \ 1i s assemblies have been used to determine hot assembly flows. The vorst flow maldistribution effect has been assumed in the initial

                                                                                    /}
                                                                                       '/

design, and the minimum hot assembly flow has been calculated to be 86 per cent of the average assembly flow at 112 per cent over-power. Actual hot assembly flows are calculated rather than ap-plying an equivalent hot channel enthalpy rise factor. Physical Mixing of Coolant Between Channels The flow distribution within the hot assembly is calculated with a mixing code that allows an interchange of heat between channels. Mixing coefficients have been determined from multirod mixing tests. The fuel assembly, consisting of a 15 x 15 array of fuel rods, is divided into unit, vall, control rod, and corner cells as shown by the heavy lines in Figure 3-41. The mixed enthalpy for every cell is determined simultaneously so that the ratio of cell to average assembly enthalpy rise (Enthalpy Rise Factor) and

the corresponding local enthalpy are obtained for each cell, Typical enthalpy rise factors are shown in Figures 3-h1 and 3 h2 for the hot and surrounding cells. The assu=ptions used to de-scribe the channels for the peaking and enthalpy rise factors shown are given in 3.2.3.2.3 j, which follows.

J. Evaluation of the DNB Ratios in the Unit, Wcll, Control Rod, and Corner Cells T DNB Results at Rated Flow i

.)

The DNE ratios in the hot unit cell at the maximum design condition described in 3.2.3.1 are shown in Figure 3-16. The relationship shown is based on the application of the W-3 correlation. An ad-ditional sensitivity analysis of the assembly corner, vall, and control rod cells has been made for the worst combination of fuel assembly spacing and power peaking. The sensitivity of the assembly design with respect to variations of mass velocity (G), channel spacing, mixing intensity, and local peaking on the DNB ratios in the fuel assembly channels has been evaluated by analyzing the nominal conditions and a postulated worst case condition. The summary results are shown below in Table 3-lh for the nominal case with the most probable conditions and 3-15 cor the maximum design conditions or postulated worst case. The unit cell DNB ratios are repeated for comperison. All of the DNB ratios are for 112 per cent overpever. l l l

4

                                                                                'l0 L'       3-5h

D-B Table 3-lh DNB Ratios in the Fuel Assembly Channels (W-31 Nominal Case (Most Probable Conditions) DNBR (W-3) C , lb /h- ft x 10 -6 Cell Type 2 (112% Pover) Unit 2.h9 1.78 Corner 2.57 1.83 Wall 2.56 1.86 Control Rod 2.39 1.92 Table 3-15 DNB Ratics in the Fuel Assembly Channels (W-3) Postulated Worst Case (Maximum Design Conditions) 2 x 10 -6 DNBR (W-3) Cell Type G , lb /h-ft (112% Pover) Unit 2.23 1.50 Corner 2.13 1.63 Wall 2.18 1.61 Control Rod 2.15 1.65 The DNB ratios in all channele are high enough to insure a confidence-population relationship equal to or better than that outlined in 3 " 'v

 ' ' dan.2.3.1.1 for the hot unit cell channel. All of the vall, corner, d control rod cells have DNB ratiot greater than that of the unit        l2 cell hot channel. This results from a more favorable flow to power ratio in these cells associated with relatively larger flov areas.

The DNB ratios were obtained by comparing the fuel rod local hea'. fluxes and channel coolant conditions with the limitations pred'.cted - by the correlation. The typical resitlts are shown in Figures .5 h3 and 3 h4 for the nominal and vorst c- e conditions in the unit cell. Fuel Rod Power Peaks and Cell Coolant Conditions The noninal case local-to-average rod powers and the local-to-average exit enthalpy rise ratios are shown in Figure 3 kl for the hot corner, 211 my 3-55 Amendment No. 2

D-B vall, control rod, and unit cells in the hot fuel assembly. Values shown are for nominal water gaps between the hot fuel asse;bly and adjacent fuel assemblies with nominal flow te hot fuel assembly, and with a minimum intensity of turbulence, a,(*) equal to 0.02. The postulated vorst case local-to-average rod powers (nuclear peak-ing factor) and exit enthalpy rise factors in the hot fuel assembly are shown in Figure 3 h2. The factors were determined for this case , with the minimum water gap between the hot fuel assembly and adjacent fuel assemblies, with minimum flow to the hot fuel assembly, and with a minimum assumed intensity of turbulence, a, equal to 0.02. An eval-uation of minimum, nominal, and maximum spacing between assemblies showed the minimum to have the lowest DNB ratios. A mixing coefficient of 0.02 was used for both nominal and design vorst case analyses. The influence of mixing coefficients is shown in Figure 3-32, which shows values ranging from 0.01 to 0.06. The value of 0.02 is sufficiently conservative for design evaluation. The conditions analyzed to obtain the DNB ratios for various values of the mixing coefficients shown in Figure 3 '2 were outlined pre-viously in 3.2.3.2.3 1. Fuel Assembly Power and Rated Flow Conditions The nominal and postulated worst cases were run at 122 per cent re- -w actor power with the nominal worst Fah factors shown in 3.2.3.1.1 c. The 150 modified cosine axial power shape of Fi6ure 3-11 was used ] to describe the vorst axial condition. ' The hot assembly flow under nominal conditions vitl'aut a flow mal-distribution effect is 96 per cent of the average casembly flow, and the reduction in flow is due entirely to heat it.put effects. The hot assembly flow under the worst postulated conditions is 86 per cent of the average assembly flow and considers the vorst com-bined effects of heat input and flow maldistribution. (*)The intensity of turbulence, a, is defined as hV{2fy where V{ is the transverse component of the fluctuating turbulent velocity, and V is the coolant velocity in the axial direction. This method of com-puting mixing is described by Sandberg, R. O., and Bishop, A. A., CVTR Thermal-Hydraulic Design for 65 MW Gross Fission Power, CVNA-227 '

                                                     .                              2,12 g

3-56 i

D-B g 3.2.h MECHANICAL DESIGN . 3.2.h.1 Reactor Internals Reactor internal components include the plenum assembly and the core support assembly (consisting of the core support shield, core barrel, lover grid, flow distributor, in-core instrument guide tubes, and thermal shield). Figure 3-h5 shows the reactor vessel, reactor vessel internals arrangement, and the reac-tor coolant flow path. Figure 3-h6 shows a cross section through the reactor vessel, and Figure 3 h7 shows the core flooding arrangement. Reactor internal components do not include fuel assemblies, orifice rod assem-blies, control rod assemblies (CRA), or in-core instrumentation. Fuel assem-blies are described in 3.2.h.2, control rod assemblies and drives in 3.2.h.3, and in-core instrumentation in 7.3.3. The reactor internals are designed to support the core, maintain fuel assembly alignement, limit fuel assembly movement, and maintain PDA guide tube aligraent between fuel assemblies and control rod drives. They also direct the flow of reactor coolant, provide gamma and neutron shielding, and provide guides for in-core instrumentation between the reactor vessel lower head and the fuel as-semblies. All reactor internal components can be removed from the reactor ves-sel to allow inspection of the reactor internals and the reactor vessel inter-nal surface. To minimize lateral deflection of the lower end of the core support asse=bly as a result of horizontal seismic loadings, integral veld-attached, deflec-tion-limiting guide lugs have been placed on the reactor vessel inside vall. In addition, these lugs limit the rotation of the le"'" end of the core sup-port assembly which could conceivably result from f. induced torsional load-ings. The lugs allow free vertical movement of the 1 wer end of the internals for thermal expansion througnout all ranges of reactor operating conditions, but in the unlikely event of a flange, circumferential veld, or bolted joint failure the lugs will limit the possible core drop to 1/2 in. or leas. Pre-liminary calculations indicate the impact loading on the guide lugs for a 1/h in. core drop would be approximately 5 g total. Lug location and geometry will be evaluated and determined to transfer this loading through the versel shell courses to the no::le supports and to the containment building concrete. A significant reduction in impact loading can be achieved through proper guide lug design and detailed analysis. A 1/2 in core drop vill not allow the lower end of the @ neu,tran, absorber rods +' disengage from their respective fuel

       . assembly guMe t'u)edLif the CRA are in the full-out position, since approxi-mately 6-1/2 in. of cod length would remain in the fuel assembly guide tubes.

A core drop of 1/2 in. vill not result in a significant reactivity chan6e. The core cannot rotate and bind the drive lines because rott ion of L..c core support assembly is prevented by the guide lugs. - The failure of the core support shield and core barrel upper flanges, or re-lated flanges and other circumferential joints, is not considered credible on the basis of the conservative design criteria and large safety factors employed in the internals design. The final internals design vill be capable ci with-standing various combinations of forces and loadings resulting frcs the static (

      veight of internals (225,000 lb total, not including'the plenum assembly which veighs 100,000 lb), core with control rod drive line (303,000 lb total),

3-57 m 213

D-B dynamic load from trip (10 g gives 207,000 lb), seismic (0.10 g vertical gives i 53,000 lb), coolant flow hydraulic loading (230,000 lb), and other related loadings. The algebraic sum of this simplified loading case is 559,000 lb. This results in a tensile stress of about 585 psi in the core support shield shell, which is approximately 3 per cent of the material yield strength. Final internals component weights, seismic analysis, dynamic loadings from flow-in-duced vibration, detailed stress analysis with consideration for thermal stress during all transients, and resolution of fabrication details such as shell roll-ing tolerances and weld joint preparation details will increase the stress levels listed above. As a final design criterion, the core support components vill meet the stress requirements of the ASME Code, Section III, during nor=al op-eration and transients. The structural integrity of all core support circum-ferential veld joints in the internals shells vill be insured by compliance with the radiographic inspection requirements in the code above. The seismic analysis will include detailed calculations to determine the maximum structural response of the reactor vessel and internals. This analysis will be performed 1 l as descrioed in 3.1.2.h.l. In the event of a major loss-of-coolant accident, such as a 36-in. diameter re-actor coolant pipe break near the reactor vessel outlet, the fuel assembly and vessel internals would be subjected to dynamic loadings resulting from an os-cillating differential pressure across the core. Some deflection of the inter-nals structures would occur, but internals component failure vill not occur. The occurrence of a loss-of-coolant accident and resulting loadings vill be evaluated during the detailed design period for the fuel assemblies and related internals structural ecmponents. 5

The deflections and movements described above would not prevent CRA insertion because the control rods are guided throughout their travel, and the guide-to-fuel-assembly alignment cannot change regardless of related component deflec-tions. CRA trip could conceivably be delayed momentarily as a result of the oscillating pressure differential. However, the CRA travel time to full inser-tion would remain relatively unaffected as transient pressure oscillations are dampened out in approximately 0.5 s. On this basis, the CRA travel time to 2/3 insertion on a trip co= mand will be approximately 1.6 s instead of the specified 1.h0 s. Also, this possible initial minor delay in trip initiation would not contribute to the severity of the loss-of-coolant accident because at the ini-tiation of CRA trip, the core would be suberitical from voids. Material for the reactor internals bolting vill ba subjected to rigid quality control requirements to insure structural integrity. The bolts will be in-spected for surface flav indications after all fabrication operations have been completed. Torque values vill be specified for the final assembly to develop full-bolting capability. All fasteners will be lock-velded to insure assembly integrity. 3.2.h.l.1 Plenum As wmbly The plenum assembly is located directly above the reactor core and is removed as a single component before refueling. It consists of a plenum cover, upper grid, CRA guide tube assemblies, and a flanged plenum cylinder with openings for reactor coolant outlet flow. The plenum cover is constructed of a series s of parallel flat plates intersecting to form square lattices and has a perfo- ) rated top plate and an integral flange at its periphery. The cover assembly

3-58

D-B is attached to the plenum cylinder top flange. The perforated top plate has (~ matching holes to position the upper end of the CRA guide tubes. Lifting lugs are provided for remote handling of the plenum assembly. These lifting lugs are velded to the cover grid. The CRA guide tubes are velded to the plenum cover Lop plate and bolted to the upper grid. CRA guide assemblies provide CRA guidance, protect the CRA frem the effects of coolant cross-flov, and pro-vide structural attachment of the grid asse=bly to the plenum cover. Each CRA guide assembly consists of an outer tube housing, a mounting flange, 12 perforated slotted tubes and four sets of tube seg=ents which are oriented and attached to a series of castings so as to provide continuous guidance for the CRA full stroke travel. The outer tube housing is velded to a mounting flange, which is bolted to the upper grid. Design clearances in the guide tube accommodate misalignment oetween the CRA guide tubes and the fuel assemblies. Final design clearances were established by tolerance studies and Control Rod Drive Line Facility (CRDL) prototype test results. The test results are de-scribed in 3.3.3.4. 9 The plenum cylinder consists of a large cylindrical section with flanges on both ends to connect the cylinder to the plenum cover and the upper grid. Holes in the plenum cylinder provide a flow path for the coolant water. The upper grid consists of a perforated plate which locates the lower end of the individual CRA guide tube assembly relative to the upper end of a correspond-ing fuel assembly. The grid is bolted to the plenum cylinder lover flange. Locating keyways in the plenum assembly cover flange engage the reactor vessel flange locating keys to align the plenum assembly with the reactor vessel, the reactor closure head control rod drive penetrations, and the core support as-sembly. The bottom of the plenum assembly is guided by the inside surface of the lower flange of the core support shield. 3.2.4.1.2 Core Support Assembly The core support assembly consists of the core support shield, core barrel, lower grid assembly, flow distributor, thermal shield, and in-core instrument guide tubes. Static loads from the assembled components and fuel assemblies, and dynamic loads from CRA trip, hydraulic flow, thermal expansion, seismic dfsturbances, and loss-of-coolant accident loads are all carried by the core support assembly. The core support assembly components are described as follows:

a. Core Suntort Shield The core support shield is a flanged cylinder vnich mates with the reactor vessel opening. The forged top flange rests on a circum- l ferential ledge in the reactor vessel closure flange. The core sup- ;

port shield lower flange is bolted to the core barrel. The inside surface of the lover flange guides and aligns the plenum assembly relative to the core support shield. The cylinder vall has two noz-

le openings for coolant flow. These openings are for=ed by two forged rings, which seal to the reactor vessel outlet nozzles by the differential thermal expansion between the stainJess steel core sup- i l

.                   port shigldjand the carbon steel reactor vessel. The nozzle seal surfaces are finished and fitted to a predetermined cold gap providing 3-59                               -

D-B clearance for core support assembly installation and remeval. At reactor operating temperature, the mating metal surfaces are in con- J)

' /

tact to make a seal without exceeding allowable stresses in either the reactor vessel or internals.

b. Core Barrel The core barrel supports the fuel assemblies, lower grid, flow dis- -

tributor,'and in-core instrument guide tubes. The core barrel con-sists of a flanged cylinder, a series of internal horizontal former plates bolted to the cylinder, and a series of vertical baffle plates bolted to the inner surfaces of the horizontal formers to produce an inner vall enclosing the fuel assemblies. The core barrel cylinder is flanged on both ends. The upper flange of the core barrel cylin-der is bolted to the mating lower flange of the core support shield and the lower flange is bolted to the lower grid assembly. All bolts are lock welded after final assembly. Coolant flow is downward along the outside of the core barrel cylinder and upward through the fuel assemblies contained in the core barrel. A small portion of the cool-ant flows upward through the space between the core barrel outer cyl-inder and the inner baffle plate vall. Coolant pressure in this space is maintained lower than the core coolant pressure to avoid tension loads on the bolts attaching the plates to the horizontal formers.

c. Lower Grid Assembly The lover grid assembly provides alignment and support for the fuel ~s i

assemblies, supports the thermal shield and flow distributor, and / aligns the in-core instru=ent guide tubes with the fuel assembly in-strument tubes. -fhe lower grid consists of two grid structures, separated by short tubular columns, and surrounded by a forged flanged cylinder. The upper structure is a perforated plate, while the lower structure consists of a machined forging. The top flange of the forged cylinder is bolted to the lower flange of the core barrel. A perforated flat plate located midway between the two grid struc-tures aids in distributing coolant flow prior to entrance into the core. Align =ent between fuel assemblies and in-core instruments is provided by pads bolted to the upper perforated plate,

d. Flow Distributor The flev distributor is a perforated dished head with an external flange which is bolted to the bottom flange of the lower grid. The flow distributor supports the in-core instrument guide tubes and distributes the inlet coolant entering the bottcm of the core,

e. Thermal Shielo_

A cylindrical stainless steel thermal shield is installed in the annulus between the core barrel cylinder and reactor vessel inner wall. The thermal shield reduces the incident ga=ma ' absorption in-ternal heat generation in the reactor vessel vall and thereby re-duces the resulting thermal stresses. The thermal shield is a 2 ,

                                                                                 )
                                                ..                         z16 3-60

D-B inch thick cylinder supported by and attached to the lover grid top flange. The thermal shield upper end is positioned by spacers be-tween the thermal shield and the core barrel outer cylinder to mini-mise the possibility of thermal shield vibration. The thermal shield attachment is designed to avoid shear loads on fasteners. All fasten-ers are lock-welded after final assembly.

f. In-Core Instrument Guide Tube Assembly The in-core instrument guide tube assemblies guide the in-core in-strument assemblies between the instru=ent penetrations in the reac-tor vessel bottom head and the instrument tubes in the fuel assem-blies. Minor horizontal misalignment is accommodated between the reactor vessel instrument penetrations' and the instrument guide tubes assembled with the flow distributor. A perforated shroud tube, con-centric with the instrument guide tube, adds rigidity to the assem-bly and reduces the effect of coolant flow forces. Fifty seven in-core instrument guide tubes are provided. The in-core instrument guide tubes are designed so they vill not be affected by the core drop described in 3.2.k.1.

3.2.h.2 Core Components 3.2.4.2.1 Description

a. General Descriction The fuel for the reactor is sintered pellets of low-enriched uranium dioxide clad in Zircaloy L tubing. The clad, fuel pellets, end caps, and the fuel support components form a " Fuel Rod." Two hundred and eight fuel rods, 16 control rod guide tubes, one instrumentat!cn tube, eight spacer grids, and two end fittings make up the basic " Fuel As-sembly" (Figure 3-48). The guide tubes, spacer grids, and end fit-tings form a structural cage which contains the 208 fuel rods in a 15 x 15 array. The center position in the assembly is reserved for instrumentation. The remaining 16 locations in the array are pro-vided for the guide tubes which guide the control rods and provide the vertical support of the assembly.

The complete core has 177 fuel assemblies which are arranged on a square lattice to approximate the shape of a cylinder. All assem-blies are identical in mechanical construction and interchangeable in the core and are designed to accept the control rod assemblies (CRA). The reactivity of the core under operating conditions is con-trolled by 57 control rod assemblies (CRA's), of which 8 are axial power shaping rod assemblies (APSRA's). These axial power shaping < rod assemblies are identical in physical configuration to the full length poison CRA but have poison'in the lover portion of the rcd only. In the fuel assemblies containing no CRA or APSRA, an orifice rod assembly (Figure 3-h9) or a burnable poison rod assembly (Fig-ure 3-50) is inserted into the upper ends of the guide tubes. These assemblies minimize guide tube bypass coolant flow. The lwsped burnable poison, rod assemblies allow a lower boric acid concentra-tion in the'. reactor coolant, thereby lowering the moderator temperature - N . 217 3-61

D-B coefficient. Because of mechanical and geometric identity, the CRA, 73 axial power shaping rods, burnable poison rod assemblies, and orifice rod assemblies are designed to be interchangeable among fuel assem-

blies. Table 3-16 Fuel Assembly Comtonents, Materials, and Dimensicns Item Material Dimensions, in. Fuel UO2 Sintered 0.370 diam Pellets Fuel Clad Zircaloy-b 0.h30 OD x 0.377 ID x 153-1/8 long Fuel Rod Pitch C.568 Fue1 ..ssembly Pitch 8. 5 17 Active Fuel Length 14k Overall Length *165

~'

Control Rod Guide Zirealcy-h 0 530 OD x 0.016 vall Tube Instrumentation Tube Zir:aloy-b 0.h93 OD x 0.441 ID Spacer Grid Inconel-T18 Strips 0.016 thick interior 0.020 thick exterior End Fittings Stainless Steel Casting

b. Fuel The fuel is sintered and ground pellets of uranium dioxide which are fabricated from previously unirradiated material. These slightly en-riched pellets are right circular cylinders with dished ends and a ground diameter. The pellet ends are dished to minimize the differ-ence in axial expansion between the fuel and the cladding. The ncm-inal density of the fuel is 93.5 per cent of theoretical.

Average design burnup of the fue1 is 27,806 mwd /MTU. Peak design burnup is 55,000 !Nd/MTU. At the peak burnup, the fuel growth is

 .              calculated to be 9-1/2 volume per cent by the method given in Refer-enc ( h5   Radial growth of the fuel during burnup is accommodated by pellet porosity, by radial clearance between the pellets and the        't cladding, and by a small amount of permanent strain in the cladding. _,)

218 ents-3-62

D-B Below each fuel column is a thin vall stainless steel spacer which (~ axially locates the bottom of the fuel column and separates the fuel from the lover fuel rod end cap. This pedestal is designed to col-lapse at a predetermined column load to prevent excessive axial strain

  ,      in the cladding.

Above the fuel colu=n is a thin-vall stainless steel spacer tube that separates the fuel from the fuel rod upper end cap. This spacer maintains the fuel column in place during shipping and handling. In operation, the spacer permits axial differential growth and thermal expansion between the fuel and the clad. This spacer also provides radial fuel rod cladding support. Ceramic spacers are located between the fuel pellets and the corru-gated tube spacers, to thermally insulate and separate fuel pellets from tube spacers. Fission gas release from the fuel is vented to voids within the pel-lets, the radial gap between the pellets and the cladding, and to the void spaces at top and bottom ends of the fuel rods.

c. Fuel Assembly

1. General The fuel assembly shown in Figure 3-48 is of the canless type where the eight spacer grids, end fittings, and the guide tubes forn the basic structure. Fuel rods are supported at each spacer grid by contact points integral with the valls of the cell bound-ary. The guide tubes are permanently attached to the upper and lover end fittings. Use of similar material in the guide tubes and fuel rods results in minimum differential thermal expansion.

2. Spacer Grids Spacer grids are constructed from strips which are slotted and fitted together in " egg crate" fashion. Each grid has 32 strips, 16 perpendicular to 16, which forms the 15 x 15 lattice. The square valls formed by the interlaced strips provide support for the fuel rods in two perpendicular directions. Contact points on the valls of each square opening are integrally punched in the I strips. On each of the two end spacer grids, the peripheral I strip is extended and rigidly attached to the respective end fitting.

T 1 3 Lover End Fitting . The lover end fitting positions the assembly when inserted in l the lower core grid plate and supports the fuel assembly weight. l The lower ends of the fuel rods rest on the grid of the lover ! end fitting. Penetrations in the fitting are provided for at- l taching the control rod guide tubes and for access to the instru-mentation tube. p %. 3-63 un

D-B 1 4 Ueper End Pitting

                                                                                       /[[)J' The upper end fitting positions the upp.er end of the fuel assem-bly in the upper core grid plate structure and provides means for coupling the handling equipment. An identifying number on each upper end fitting provides positive identification when handling.

An internal hollow post, velded in the center of the end fitting provides means for retention of the orifice rod assembly and burn-able poison rod assembly. Integral with each upper end fitting is a holddown spring to pro-vide a positive holddown margin to oppose hydraulic forces. Penetrationc in the upper end fitting grid are provided for the guide tuces. 5 Control Rod Guide Tubes The Zircaloy guide tubes provide continuous guidance to the con-trol rods within the fuel assembly during operation and provide structural continuity for the fuel assembly. Welded to each end of a guide tube are flanged and threaded sleeves, which attach the tubes to the end fittings by lock-velded nuts. Transverse location of the guide tubes is provided by the spacer grids.

-~3

6. Instrumentation Tube -

This Zircaloy tube serves as a channel to guide, position, and contain the in-core instrumentation in the center of the fuel assembly. The instrumentation string is guided up through the lower end fitting and through the tube to the desired core eleva-tion. The instrumentation tube provides no structural support of the assembly and is retained axially by the end fittings and radially by the spacer grids. 7 Spacer Sleeves The spacer tube segments fit around the instrument tube between spacer grids and prevent axial move =ent of the spacer grids dur-ing primary coolant flow through the fuel assembly. 3.2.4.2.2 Evaluation

a. Fuel Rod Assembly

l. General The basis for the design of the fuel rod is discussed in 3.1.2.h.

Materials testing and actual operation in reactor service with Zircaloy cladding have demonstrated that Zircaloy h material has sufficient corrocion resistance and mechanical properties to maintain the integrity and serviceability required for design burnup. ")] t: 3-6h

D-B

2. Clad Stress and Strain The cladding of fuel rods is subjected to hydrostatic pressure, gradually increasing internal pressure, thermal stresses, vibra-tion, and to the effects of differential thermal expansion of the fuel and cladding, and by fuel growth due to irradiation ei-fects. In addition, tne properties of the cladding are influ-enced by thermal and irradiation effects, which are analyzed be-lov.

Stress analysis for cladding is based on several conservative as-sumptions that make the actual margins of safety greater than those calculated. For example, it is assumed that the clad with the thinnest vall, the smallest fuel-clad gap, and the greatest ovality permitted by the specification is operating in the region of the core where performance requireme:ns are most severe. Fission gas release - rates, fuel growth, and changes in mechanical properties with irradiation are bast _ on a conservative evaluation of currently available data.

3. Pressure Effects Beginning of Life Power Conditions Clad stresses due to external and internal pressure are consider-ably below the yield strength. Circumferentail stresses due to external pressure, calculated using those combinations of clad dimensions, ovality, and eccentricity that produce the highest stress, are shown in Table 3-17 The maximum compressive stress in the expansion void at the system design pressure is the sum of compressive membrane stress plus compressive bending stress due to ovality at the clad OD. Stress conditions are listed for beginning of life. The maximum stress in the heat-producing zone at the beginning of life at operating pressure is also shown. At this stress, the material may creep enough to allow an increase in ovality until further creep is restrained by sup-port from the fuel. Contact loads between fuel and cladding for this condition are listed.

End of Life Power Conditions At the end of life, fission gas pressure does not exceed operating pressure (3.2.3.2.3), however an internal pressure of 3,300 psi has been selected as the design basis. At this pressure the dif-ferential would result in a circumferential tensil.e stress at nor- s mal operating pressure. This stress value shown in Table 3-17 is j about 1/k of the yield strength and, therefore, is not a potential source of short-time burst. The possibility of stress-rupture burst has been investigated using finite-difference methods to estimate the long-time effects of the increasing design pressure on the clad. The predicted pressure-time relationship produces stresses that are less than 1/3 of the stress levels that vould l .

a g . .

221

D-B produce stress rupture at the end of life. Outpile stress-rup- ) ture data were used, but the greater than 3: 1 margin on stress is more than enough to account for decreased stress rupture strength due to irradiation. End of Life Shutdown Conditions The primary coolant system can be cooled down at a maximum rate of 100 F/h. During this cooldown period it is desirable to main-tain a compressive load on the fuel clad until the clad has been cooled to at least 400 F vhere all significant hydrides have had a chance to precipitate under a favorable stress field. The pres-surizer and primary system vill be cooled down at rates producing a net compressive load on the clad until the clad has reached h00 F. Coolant conditions and stress levels are shown for two

stages of the normal cooldown cycle to the desired temperature 1evel in Table 3-17 Fuel Burnuo, Temrerature and Gas Release Conditions The total production of fission gas in the hottest fuel rod as-sembly is based on the hot rod average burnup of 38,000 mwd /MTU.

The corresponding maximum design burnup at the hot fuel rod mid-point is 55,000 mwd /MTU. The fission gas release is based on temperature versus release fraction as shown in Figure 3-37 Fuel temperatures are calcu-lated for small radial and axial increments. The total fission gas release is calculated by integrating the incremental releases. The maximum release and gas pressure buildups are determined by evaluating the following factcrs for the most conservative condi-tions: (a) Gas conductivity at the end of life with fission gas present. (b) Influence of the pellet-to-clad radial gap and contact heat transfer coefficient on fuel temperature and release rate. (a) Unrestrained radial and : cial thermal growth of the fuel pellets relative to the clad. (d) Hot rod local peaking factors. (e) Radial distribution of fission gas production in the fuel pellets. (f) Fuel temperatures at reactor' design overpower. 222 m ) 3-66

D-B Table 3-17 I- Clad Circumferential Stresses Ultimate Calc. Yield Tensile Stress, Stress, Stress, Operating Condition psi psi osi

1. BOL

Expansion Void Clad "

Conditions at Maximum Overpower Total Stress (Membrane + Bending) Due

to 2,500 psig System Design Pressure Minus 100 psig Fuel Red Internal Pressure. Average Clad Temperature 650 F. Stress (-)33,000 45,000

2. BOL - Fueled Section Clad 2 Conditions at Maximum Overeover Total Stress (Membrane + Bending) Due to System Pressure. Average Clad Temperature 728 F.

Fuel to Clad Contact Load, 20 lb/in. of length. Stress At 2,500 psig System Design Pressure (-)32,000 ho,000 At 2,185 psig System operating (-)27,000 ho,000 Preesure

3. SOL - Fueled Section Clad Conditions at Maximum Overoever 7

Systed bressure - 2,185 psig Fuel Rod Internal Design Pressure - 3,300 psig Average Temperature Through Clad Thickness at Hot Spot - Approxi-mately 702 F. 223

,      Stress Pressure Stress Only                             9,000 3-67                    Amendment lio. 2

D-B Table 3-17 (Cont'd) ,,,; Ultimate

                                                      . Calc.      Yield  Tensile           -

Stress, Stress, Stress, Operating Condition esi psi psi Including k,000 psi Thermal Stress 13,000 38,000 40,000

h. EOL Shutdown Immediately After Shutdown Conditions System Pressure - 2,185 psia Fuel Rod Internal Pressure - 1,k00 .

psig Average Clad Temperature - 593 F Stress (-) 7,100 hh,000 47,000 1-1/3 Hours Later Conditions 'N (50 F/h Pressurizer Cooldown Rate First Hour, up to 100 F/h There-after) Primary System Cooldown, 100 F/h Fuel Rod Internal Pressure - 1,090 psig System Pressure - 1,275 psig Average Clad Temperature h00 F Stress (-) 1,700 52,000 55,000 (* Cladding is specified with h5,000 psi minimum yield strength and 10 per cent minimum elongation, both at 650 F. Minimum room temperature strengths are approximately 75,000 pai yield strength (0.2 per cent offset) and 85,000 psi

  . ultimate tensile strength.

Cladding stresses due to fuel swelling are discussed further in a subsequent paragraph of this section, Effect of Zircaloy Creep. l l

                                                                 " faut              224 u) l                               .

3-68 '

D-3

  ,. The fuel te=peratures used to determine fission gas release and internal gas pressure have been calculated at the reactor over-power condition (112 per cent). Fuel temperatures, total free gas volume, fission gas release, and internal gas pressure have been evaluated for a range of initial diametral clearances. This evalustion shows that the highest internal pressure results when the maximum diametral gap is assumed because of the resulting high average fuel temperature. The release rate increases rap-idly with an increase in fuel temperature, and unrestrained axial growth reduces the relatively cold gas end plenum volumes. A conservative ideal thermal expansion model is used to calculate fuel temperatures as a function of initial cold diametral clear-ance as outlined in 3.2.3.2.3 g(l).

h. Collapse Margins Short-time collapse-tests have demonstrated a clad collapsing pressure in excess of h,000 psi at expansion void maximum tem-perature. Collapse pressure margin is approximately 1.7 Ex-trapolation to hot spot average clad temperature (725 F) indi-cates a collapse pressure of 3,500 psi and a margin of 1.h, which also greatly exceeds requirement. Outpile creep collapse tests have demonstrated that the clad meets the long-time (creep collapse) requirement. Back up radial support has been provided in the urger end void to assure clad dimensional stability in the event that in-pile creep rates are sufficiently high to al-low creep collapse of unsupported cladding. Test results sum-marized in Section 3.3.3.3.1 show the end void spacers are ca-pable of providing backup support. The results of the tests show that creep collapse of the bottom end void vill not occur since .ie clad temperature is about 90 F lower than that in the upper told region. The spacer in the bottom end void is there-fore not required to provide radial support. Its geometry, how-ever, is similar to the upper spacer, and it therefore provides added assurance of clad dimensional stability at the bottom void region.

5 Fuel Irradiation Growth and Fuel-Clad Differentail Thermal Expansion The esults of tests and the operation of 7.ircaloy-clad UO2 fuel rods indicate that the rods can be safely operated to the point where total permanent strain is 1-1/2 per cent, or higher, in the tempersture range applicable to PWR cladding.(h6) The al-lovable design strain is about 1 per cent (3.1.2.h.2 c). Fuei* rod operating conditions pertinent to fuel swelling consid- _ erations are listed belov for end-of-life conditions. Burnup (Design Value), mwd /MTU 55,00C Minimum Fuel-to-Clad Gap 0.0045 (Beginning of Life), in. 9C 't Pellet Nominal Diameter, in. 0.370 }LJ 3-69

D-B Pellet Density (Per Cent of Theoretical), % 93 5 Madding (Zircalcy-4), in. 0.0265 Wall The capability of Zircaloy-clad UO 2 fuel in solid rod form to perform satisfactorily in service has been demonstrated through operation of the SA-1 essembly in the Dresden and Shippingport cores, and through resalts of their supplementary development programs, up to approximately h5,000 mwd /MTU. As outlined belov, existing experimental infor=ation supports the various individual design parr. a rs and operating condi-tions up to and perhaps beyond the maximum design burnup of 55,000 mwd /M'IU, but not in a single experiment. However, the B&W High Burnup Irradiation Program currently in progress does combine the primary items of concern in a single experiment, and the results will be available to contribute to the final design. Application of Experimental Data to Design Adequacy of the Clad-Fuel Initial Gap to Accommodate Clad-Fuel Differential Thermal Exnansion Fanerimental Work Six rabbit capsules, each containing three Zr-2 clad rods of 5-in. fuel leggth, were irradiated in the Westinghouse Test Reactor (h51 at power levels up to 2h kW/ft. The 94 , per cent theoretical density (TD) UO2 pellets (0.h30 OD) had initial clad-fuel diametral gaps of 6, 12, and 25 mils. No dimensional changes were observed. Central melting oc-curred at 2h kW/ft only in the rods that had the 25 mil initial gap. Two additional capsules were tested.(h6) The specimens were similar to those described above except for length and initial gap. Initial gaps of 2, 6, and 12 mils were used in each capsule. In the A-2 capsule, three 38-in.- long rods were irradiated to 3,h50 mwd /MTU at 19 kW/ft maximum. 'In the A-b capsule, four 6-in.-long rods were irradiated to 6,250 mwd /MTU at 22.2 kW/ft maximum. No central melting occurred in any rod, but diameter in-creases up to 3 mils in the A-2 capsule and up to 1 5 mils in the A-4 capsule were found in the rods with the 2 mil initial gap. Anolication

In. addition to demonstrating the adequacy,of Zircaloy-clad UO2 pellet rods to operate successfully at the power levels

of interest (and without central melting),'these experi-ments demonstrate that the design initial clad-fuel gap of k.5 to 8.5 mils is adequate to prevent unacceptable clad diam- ~

eter increase due to differential thermal expansion between the clad and the fuel at beginning of life. A maximum local ,' 3-70 g

D-B diametral increase of less than 0.001 in. is indicated for ( fuel rods having the minimum initial gap, operating at the maximum overpower condition. Adequacy of the Available Voids to Accommodate Differential Expansion of Clad and Fuel, Including the Effects of Fuel Svelling Exterimental Work Zircaloy-clad, UO2 pellet-type rods have performed success-fully in the Shippingport reactor up to approx -at ly 40,000 mwd /MTU. Bettis Atomic Power Laboratory hk has irradiated plate-type UO2 fuel (96-98 per cent TD) up to 127,000 mwd /MTU and a fuel center temperatures between 1,300 and 3,800 F. This work indicates fuel swelling rates of 0.16 per cent AV/1020 f/cc until fuel internal voids are filled, then 0.7 per cent AV/1020 f/cc after internal voids are filled. This point of "breakav..y" appears to be inde-pendent of tenperature over the range s adied and dependent on clad restraint and the void volume a tilable for collec-tion of fission products. The additionui clad restraint and greater fuel plasticity (from higher fuel temperatures) of rod-type elements tend to reduce these swelling effects by providing greater resistance to radial swelling and lover resistance to longitudinal swelling than was present in the plate-type test specimens. This is confirmed in part by the work of Frost, Bradbury, and Griffiths of Harvell(kIl in which 1/h-in. diameter UO 2 pellets clad in 0.020 in. stainless steel with a 2 mil diametral gap were irradiated to 53,300 mwd /MTU at a fuel center temperature of 3,180 F vithout significant dimen-sional change. In other testing (h0) 0.150-in. OD, 82-96 per cent TD oxide pellets (20 per cent Pu, 80 per cent U) clad with 0.016-in. stainless steel with 6-8 mil diametral gaps have been ir-radiated to 77,000 mwd /MTU at fuel temperatures high enough to approach central melting without apparent detrimental results. Comparable results were obtained on rods svaged to 75 per cent TD and irradiated to 100,000 mwd /MTU. Acolication

Nc Based on the BAPL experimental data, swelling of the fuel

          V5bds is estimated as outlined below.                                 "

The fuel is assumed to swell uniformly in all directions, conservatively neglecting axial plastic flov into the end dishes. Thermal expansions are calculated as described in 3.2.3.2.3 g. If the fuel cracks, the crack voids are as- .( sumed to be available to absorb fuel growth, i 227 3-71 .

D-B The external effect of fuel swelling is assumed to occur A at 0.16 per cent AV/1020 f/cc until the as-fabricated void j in the 93.5 per cent pellets is filled. From that time on, swelling is assumed to take place at 0.7 per cent AV/1020 f/cc until the maximum burnup of 13.6 x 1020 f/cc (55,000 mwd /MTU) is reached. Studies of clad strain at various gaps indicate that the rod with the minimum gap experiences the greatest clad strain in spite of its Lnproved gap conductivity. Clad permanent strain reaches a maximum at the end of life, and is 0 7 per cent for nominal density fuel. Clad strain for fuel rods with maxLaum density allowed by the specification vill also meet the design's maximum allovable permanent strain.

6. Effect of Zircaloy Creen The effect of Zircaloy creep on the amount of fuel rod growth due to fuel swelling has been investigated. Clad creep has the effect of producing a nearly constant total pressure on the clad ID by permitting the clad diameter to increase as the fuel diam-eter increases. Based on out-of-pile data,(49) 1 per cent creep will result in 10,000 h (corresponding approximately to the end-of-life diametral swelling rate) from a stress of about 22,000 psi at the 702 F average temperature through the clad at the 'N hot spot. At the start of this high swelling period (roughly ,)

the last 1/3 of the core life), the reactor coolant system pres-sure voi.ld more or less be balanced by the rod internal pressure, so, that total pressure to produce the clad stress of 22,000 psi vould have to come from the fuel. Contact pressure vould be 2,400 psi. At the end of life, the rod internal design pressure ex-ceeds the system pressure by about 1,100 psi, so the clad fuel contact pressure would drop to 1,300 psi. Assuming that irradia-tion produces a 3:1 increase in creep rates, the clad stress for 1 per cent strain in 10,000 h would drop to about 15,000 psi. Contact pressures would be 1,800 psi at the beginning of the high swelling period, 700 psi at the end of life. Since the contact pressure was assumed to be 825 psi in calculating the contact coefficient used to determine the fuel pellet thermal expansion, there is only a short period at the very end of life (assuming tpe 3:1 increase in creep rates due to irradiation) when the pel-let is slightly hotter than calculated. The effect of this would be a slight facrease in pellet thermal expansion and therefore in clad strain,

b. Overall Assembly

1. Assurance of Control Rod Assembly Free Motion The 0.058 in, diametral clearance between the control rod guide tube and the control rod is provided to cool'the control rod and to insure adequate freedom to insert the control rod. As indi- ,

cated below, studies have shown that fuel rods vill not bov 228 3-72 :.- m . 1

D-B sufficiently to touch the guide tube. Thus, the guide tube vill not undergo deformation caused by fuel rod bowing effects. Ini-tial lack of straightness of fuel rod and guide tube, plus other adverse tolerance conditions, conceivably could reduce the 0.088 in. nominal gap between fuel rod and guide tube to a minimum of about 0.038 in. , including a=plification of boving due to axial friction loads from the spacer grid. The maximum expected flux gradient of 1.176 across a fuel red will produce a temperature difference of 12 F, which will result in a thermal bow of less than 0.002 in. Under these conditions, for the fuel rod to touch the guide tube, the thermal gradient across the fuel rod diameter would have to be on the order of 300 F. The effect of a DNB occurring on the side of a fuel rod adjacent to a guide tube vould result in a large temperature difference. In this case, however, investigation has shown that the clad temperature would be so high that insufficient strength.vould be available to generate a force of sufficient magnitude to cause a significant deflection of the guide tube. In addition, the guide tube would experience an opposing gradient that would resist fuel rod bowing, and its internal cooling would maintain temperatures much lower than those in the fuel rod cladding, thus retaining the guide tube strength.

2. Vibration i The semi-empirical expression developed by Burgreen(50) was used to calculate the flow-induced vibratory amplitudes for the fuel assembly and fuel rod. The calculated amplitude is 0.010 in.

for the fuel assembly and less than 0.005 in. for the fuel rod. The fuel rod vibratory amplitude correlates with the measured amplitude obtained from a test on a 3 x 3 fuel rod assembly. In order to substantiate this conserva;ively' calc'ulated amplitude for the fuel assembly, a direct measurement was obtained for a full size prototype fuel assembly during testing of the assembly in the Control Rod Drive Line Facility (CRDL) at the B&W Research Center, Alliance, Ohio. The maximum assembly amplitude was mea-sured to be 0.005 in.

3. Demonstration In addition to the specific items discussed above, the overall mechanical performance of the fuel assembly and its individual co=ponents is being demonstrated in an extensive experimental program in the CRDL.

3.2.k.3 Control Rod-Drive System - 3.2.k.3.1 Descriution - The control rod drive system includes drive mechanisms which actuate control ( rod assemblies and xenon control rod assemblies, drive controls, power sup-(, plies, position indication, operating panels and indicators, safety devices, J . 229 3-73 --

D-B i enclosures, housings, and mountings. Criteria applicable to drive mechanisms for both control rod assemblies and xenon control rod assemblies are given be- '~'} - lov (General Design Criteria)._ Additional requirements for the mechanisms which actuate only control rod, assemblies are given under Additional Design Criterta. General Design Criteria

a. Single Failure No single failure shall inhibit the protective action of the control rod drive system. The effect of a single failure shall be limited to one control rod drive.

b. Uncontrolled Withdrawal No single failure or sequence of dependent failures shall cause un-controlled withdrawal of any control rod assembly (CRA) . ~

c. Ecuiement Removal The disconnection of plug-in connectors, modules, and subassemblies from the protective circuits shall be annunciated or shall cause a reactor trip.

d. Position Indicatien ,

Continuous position indication, as well as an upper and lover posi- ' tion limit indication, shall be provided for each control rod drive. The accuracy of the position indicators shall be consistent with the tolerance set by reactor safety analysis.

e. System Monitoring The control rod drive control system shall include provisions for monitoring conditions that are important to safety and reliability.

These include rod position deviation and power supply voltage.

f. Drive Speed a

The control rod drive control system shall provide for single uni-form speed of the mechanism. The drive controls, or mechanism and motor combination, shall have an IMherent speed limiting feature. The speed of the mechanism shall be 30 in./ min 6 per cent of the predetermined value for both insertion and withdrawal. The with-drawal speed shall be limited so as not to exceed 25 per cent over-speed in the event of speed control fault.

g. Mechanical Stoos Each control rod drive shall have positive mechanical stops at both ends of the stroke or travel. The stops shall be capable of receiving the full operating force of the mechanisms without failure .

qc '3] s-

3-74

D-B c Additional Design Criteria ( The following criteria are applicable only to the mechanisms which actuate con-trol rod assemblies.

a. CRA Positioning The control rod drives shall provide for controlled withdrawal or insertion of the control rod assemblies (CRA) out of, or into, the reactor core to establish and hold the power level required. The drives are also capable of rapid insertion or trip for emergency reactor conditions. Insert command shall have priority over with-draw command. The control rod drive will be capable of overcoming a " stuck rod" condition equivalent to a 400 lb weight.

b. CRA Trip The trip command shall have priority over all other co== ands. Trip -

action shall be positive and nonreversible. Trip circuitry shall provide the final protective action and shall be direct-acting, in-cur minimum delay, and shall not require external power. Circuit-interrupting devices shall not prevent reactor trip. Fuses, where used, shall be provided with blown indicators. Circuit breaker position information shall also be indicated.

c. Group Withdrawal k

The control rod drive system allows only two out of three regulating CRA groups to withdraw at any time subject to the conditions de-scribed in 7.3.2.1.2. 3.2.4.3.2 Control Rod Drive Mechanis=s The control rod drives provide for controlled withdrawal or insertion of the control rod assemblies out of or into the core and are capable of rapid in-sertion or trip. The drives are hermetically sealed, reluct,ance motor-driven screw units. The control rod drive data are listed in Table 3-18. N

   \
                                   .~

s ( , ..__. 231

                       ;g      g 3-75

D-B Table 3-18 O Control Rod Drive Design Data

Item Shim Safety Axial Power Shaning Number of Drives 49 8 Type Roller Nut Roller Hut - Drive Drive Location Top-Mounted Top-Mounted Direction of Trip Down Does Not Trip Velocity of Normal Withdrawal 30 30 - and Insertion, in./ min Maximum Travel Time for Trip 2/3 insertion, s 1.40 Drive has no Trip Function 3/4 insertion, s 1 50 Drive has no Trip Function Length of Stroke, in. 139 139 Design Pressure, psig 2,500 2,500 'h

.)

Design Temperature, F 650 650 Weight of Mechanism 9k0 9h0 (appron); lb Shim Safety Drive Mechanism The shim safety drive mechanism consists of a motor tube which houses a lead screw and its rotor assembly, and a buffer. The top end of the motor tube is closed by a cap and vent assembly. An external motor stator surrounds the motor tube (a pressure housing) and position indication switches are arranged outside the motor tube extension. The control rod drive output element is a non-rotating translating lead screw coupled to the control rod. The screw is driven by separating anti-friction roller nut assemblies attached to seg=ent arms which are rotated magnetically by a motor stator located outside the pressure boundary. Current impressed on the stator causes the separating roller nut assembly halves to close and engage the lead screw. Mechanical springs disengage the roller nut halves from the screw in the absence of a current. For rapid insertion, the nut halves separate to release the screw and control rod, which move into the core by gravity. A hydraulic buffer assembly within the upper housing decelerates s the moving CRA to a lov speed a short distance above the CRA full-in pos4; tion. The final CRA deceleration energy is absorbed by the down-stop buffer spring. '] d

                                        ,_,              'Tnat

D-B 7 .he CRDM is a totally sealed unit with the roller nut assemblies and segment arms magnetically driven by the stator coil through tne motor tube pressure housing wall. The lead screw assembly is connected to the control rod by a bayonet type coupling. An anti-rotation device (torque taker) prevents rota-tion of the lead screv while the drive is in service. A cap and vent assembly is provided at the top of the motor tube housing to permit access to couple and release the lead screw assembly from the control rod. The top end of the lead screw assembly is guided by the buffer piston and its guide. Two of the six phase stator housing vindings are energized to maintain the control rod position when the drive is in the holding mode. The CREM is shown in Figures 3-51 and 3-52. Subassemblies of the CRDM are de-scribed as follows:

a. Motor Tube The motor tube is a three-piece velded assembly designed and manu-factured in accordance with the requirements of the ASME Code, Sec- '

tien III, for Class A nuclear pressure vessel. Materials conform to ASTM or ASME, Section II, Material Specifications. All velding shall be performed by personnel qualified under ASME Code, Section IX, Welding Qualifications. The motor tube vall between the rotor as-sembly and the stator is constructed of magnetic material to present a small air gap to the motor. This region of the motor tube is of lov alloy steel clad on the inside diameter with stainless steel or with Inconel. The upper end of the motor tube functions only as a 4 pressurized enclosure for the withdrawn lead screw and is made of stainless steel transition-velded to the upper end of the lov alloy steel motor section. The lower end of the low alloy steel tube sec-tion is velded to a stainless steel machined forging which is flanged at the face which contacts the vessel control rod no::le. Double gaskets, which are separated by a ported test annulus, seal the flanged connection between the motor tube and the reactor vessel.

b. Motor The motor is a synchrenous reluctance unit with a slip-on stator.

The rcter assembly is described in Paragraph (f). The stator is a 48-slot four-pole arrangement with water cooling coils wound on the outside of its casing. The stator is encapsulated after vinding to establish a hermetically sealed unit. It is six phase star-connected

        ... for operation in a pulse-stepping mode and advances 15 mechanical degrees per step. The stator assembly is mounted over the motor tube housing as shcun in Figure 3-52.

l

c. Cao and Vent Valve i

The upper end of the motor tube is closed by a cap containing a vapor l bleed port and vent valve. The bleed port and vent valve and the l cap-to-motcr tube closures have double seals. The cap is retained l by a bolting ring threaded to the outside of the motor tube. The re-taining bolts are assembled to the bolting ring to prevent inadvertent ( ; .t* less. The bolts are made long so as to be elastic enough to provide l positive seal preload at any assembly temperature frca 20 to 650 F. l The minimum preload is equal to.the 3,750 pais proof pressure force. I m s .; .., .. 2331 3-77 -..

D-B

d. Actuator O

                                                                                .s The actuator consists of the translating lead screw, its rotating nut assembly, and the torque taker assembly on the screw. The actuator lead screw travel is 139 inches.

e. Lead Screw The lead screw has a lead of 0 750 in. The thread is double lead with a single pitch spacing of 0 375 in. Thread lead error is held to 0.0005 in. maximum in any 6 in. for uniform loading with the roller nut assemblies. The thread form is a modified ACME with a flank angle that allo-s the roller nut to disengage without lifting the screw,

f. Rotor Assembly The rotor assembly consists of a ball bearing supported rotor tube carrying and limiting the travel of a pair of segment arms. Each of the two arms carry a pair of ball bearing supported roller (nut) as-semblies which are skewed at the lead screw helix angle for engage-ment with the lead screw. The current in the motor stator (two of a six vinding stator) causes the arms that are pivoted in the rotor tube to move radially toward the motor tube vall to the limit pro-vided thereby engaging the four roller nuts with the centrally lo-cated lead screw. Also, four separating springs mounted in the seg- s nent arms keep the rollers disengaged when the power is removed from i the stator coils. A second radial bearing mounted to the upper end >'

of the rotor tube has its outer race pinned to both sessent arms thereby synchronizing their motion during engagement and disengage-ment. When a three phase rotating magnetic field is applied to the motor stator, the resulting force produces rotor assembly rotation.

g. Torque Extension Tube and' Torque Taker The torque extension tube is a separate tubular assembly containing a keyvay that extends the full length of the lead screw travel. The tube assembly is supported against rotation and in elevation by the upper end of the motor tube extension. The lover end of the tube assembly supports the buffer and is the down stop. A set of index-ing serrations mate to prevent rotation and orient the torq.ue exten-sion tube with the motor tube below the cap and vent valve assembly.

An integral shoulder at the top of the tube rests against a step in

the motor tube inside diameter to provide a vertical support.

The torque taker assembly consists of the position indicator perma-nent magnet, the buffer piston, and a positioning key. The torque - taker key fixed at the top of the lead screw is mated with the tor-que extension tube keyway to provide both radial and tangential po-sitioning of the lead screw. Q v\

                                                 .. y
                     ,           3-T8

D-B

h. Buffer

( The buffer assembly is capable of decelerating the translating mass from the unpressurized terminal velocity to zero velocity without applying greater than ten times the gravitational force on the con-trol rod. The water buffer consists of a piston fixed to the top end of the screw shaft and a cylinder which is fixed to the lower end of the torque extension tube. Twelve inches above the bottom stop, the piston at the top of the screw enters the cylinder. Guid-ing is accomplished because the piston and torque key are in a single part, and the cylinder and keyway are in a single =ating part. As the piston travels into the cylinder, water is driven into the center of the lead screw through holes in the upper section which produce the damping pressure drop. The number of holes presented to the buffer cha=ber is reduced as the rod moves into the core, so that the damping coefficient increases as the velocity reduces, thereby providing an approximately uniform deceleration. A large helical

buffer spring is used to take the kinetic energy of the drive line at the end of the water buffer stroke. The buffer spring accepts a five-foot per second impact velocity of the drive line and control rod with an instantaneous overtravel of one inch past the normal down stop. The inc_usion of this buffer spring permits practical clearances in the water buffer.

1. Lead Screv Guide j The lead screw guide bushing acts as a primary thermal barrier and as a guide for the screw shaft. As a primary thermal barrier, the bushing allows only a small path for free convection of water be-tween the mechanism and the closure head nozzle. Fluid temperature in the mechanism is largely governed by the flow of water up and down through this bushing. The diametral clearance between screw shaft and bushing is Large enough to preclude ja==ing the screv shaft and small enough to hold the free convection to an acceptable value.

In order to obtain trip travel times of acceptably small values, it is necessary to provide an auxiliary flow path around the guide bush-ing. The larger area path is necessary to reduce the pressure dif-ferential required to drive water into the mechanism to equal the screv displacement. The auxiliary flow paths are closed for small pressure differentials (several inches of water) by ball check valves which prevent the convection flow but open fully during trip. J. Position Indications Two methods of position indication are provided; one, an absolute position indicator and the other, a relative position indicator. The absolute position transducer censists of a series of magneti- " cally operated reed switches mountea in a tube parallel to the motor tube extension. Each switch is hermetically sealed. Switch contacts close when a permanent magnet mounted on the upper end of the lead screw extension comes in close proximity. As -he lead screw (and the control rod assembly) moves, switches operate sequentially pro-i ducing an analogue voltage proportional to position. The accuracy oftheanaloguesignal"is!1.(percentandproducesareadoutof aanse . 235 I-' T9

D-B approximately 2.1 per cent accuracy. Additional reed switches are included in the same tube with the absolute position transducer to (h) provide full withdrawal and insertion signals. The relative position indicator consists of a small pulse-stepping motor driving a poten-tiometer th;; generates a signal accuracy of 0.7% producing a posi-tion readou, of 11 7% accuracy.

k. Motor Tube Design Criteria -

The motor tube design complies with Section III of the ASME Boiler and Pressure Vessel Code for a Class A vessel. The operating tran-sient cycles, which are considered for the stress analysis of the reactor pressure vessel, are also considered in the motor tube de-sign. Quality standards relative to material selection, fabrication, and inspection are specified to insure safety function of the housings essential to accident prevention. Materials conform to ASTM or ASME, Section II, Material Specifications. All velding shall be performed by personnel qualified under ASME Code, Section IX, Welding Qualifi-cations. These design and fabrication procedures establish quality assurance of the assemblies to contain the reactor coolant safely at operating temperature and pressure. In the highly unlikely event that a pressure barrier component or the control rod drive assembly does fail catastrophically, i.e., ruptured ~~s completely, the following results would ensue:

1. Control Rod Drite Nozzle The assembly would be ejected upward as a missile until it was stopped by the missile shield over the reactor. This upward mo-tion vould have no adverse effect on adjacent assemblies.

2. Motor Tube The failure of this component anywhere above the lower flange vould result in a missile-like ejection into the missile shield-ing over the reactor. This upward motion would have no adverse effect on adjacent mechanisms.

Axial Power Shaping Rod Drive Mechanisms For actuating the partial length control rods which =aintain their set position during a reactor-trip of the shim safety drive, the CRDM is modified so that the roller nut assembly vill not disengage from the lead screw on a loss of power to the stator. Except for this modification, the shim drives and the axial power shaping rod drives are identical'. 3.2.h.3.3 Control Rod Drive Control System (Control Package)

                          'C'y)ill The control system for the control rod drive is designed to energize and posi-        ?N tion the control rod drive, provide a reactor trip, indicate the control rod        (s,/

assembly (CRA) position in the core, and indicate malfunctions in the system. The control system consists of: 2}h deust 3-80 . t

D-B

   ,-          a. System Control

1. Individual and Group CRA Control (Operator's Panel'

2. Position Indication

3. Automatic Sequencing

h. Position Deviation Monitors

b. Power Supply (Motor Controller)

1. SCR Programmer (CRA Speed Standard)

2. SCR Banks 3 CRA Grouping Panel

h. Transfer Control

c. Trip Figure 3-53 depicts in block diagram form items b and c with ccmmand inputs from item a. I The reactor operator is provided with an operator panel and controls which per-mit manual or automatic group operation, manual single rod eperation, group sequencing and position indication. All manual ecmmands, including operator-initiated trip, are made from the operator's panel.

Position indication is provided on both individual rods and rod groups. In-dividual position meters indicating per cent withdrawn are visible frem the f operator's panel. Four group position meters are provided at the operator's panel. These group meters indicate the position of either the four safety rod groups or the regulating grcups, whichever ir selected by the meter se-lector switch. Automatic sequencing of the regulating groups is provided. The sequencer pro-vides overlapped withdrawal and insertion of Groups 5, 6, and 7 within the 11=- its of T.2.2.1.2. This sequencing is provided in both automatic and manual modes of control. Sequencer logic is derived from position transducer and lie-it switch signals. The system centrol provides the logic to eccmand the proper group power supply. The group power supply provides d-c pover from a three-phase source and applies it as directed by the programmer controlled SCR to the CRA =echanism. The power supply consists of an SCR programmer, SCR gate drivers, SCR banks, transfer re-lays, input pcwer transformers, and rod group patch panels. Input commands of in-hold-out are rec '.ved by the progra=mer, which in turn generates the gating sequences for the SCR banks. The programmer consists of a synchronous motor operating on 60-cycle, a-c power, driving a coded disk through a light beam. The coded light team drives photo detectors, and the photo detectors drive SCR gate driver auplifiers, which in turn gate the SCR. The SCR banks apply a steady voltage to successive motor vindings and are line-commutated by the input a-c pcVer. The motor is 6-phase, star-connected and C produces 15 degrees of mechanical rotation per switching cycle. The group power supplies contain redundant SCR banks, each fed from a different power source but driven from a common but dual-channel progra=mer. Eight groups 3-81 _ 23[

D-B of drives are in the rod drive system., each havin6 its own power supply. A h ninth power supply is provided for single rod control and as an operational (~[l v; spare. Any rod may be operated in the single rod mode by transferring to the ninth power supply. Reactor trip is initiated by de-energizing two circuit breakers supplying con-trol rod drive power or two contactors supplying SCR gate power. Both circuits have two devices in series. The reactor protection system trips the power circuit breakers and control cir-cuit contactors through two out of four logic as shown in the block diagram (Figure 3-53). 3.2.4.3.4 Control Rod Drive System Evaluatinn

a. Design Criteria The system vill be designed, tested, and analyzed for cc=pliance with the design criteria. A preliminary safety analysis of the control rod drive motor control subsystem was conducted to determine _ failures of logic functions. It was concluded that no single failure in any CRA control vould prevent CRA insertion, nor cavre inadvertent CRA withdrawal of another CRA or CRA group.

b. Materials Selection Materials are selected to be compatible with, and operate in, the reactor coolant. Certified mill test reports containing chemical analysis and test data of all materials exposed to the reactor sys-tem fluid vill be provided and maintained for the control rod drives.

Certificates of compliance for other materials and components shall also be provided.

c. Relation to Design Temeerature All parts of the control rod drive exposed to reactor coolant are designed to operate at 650 F, although it is expected that all parts vill operate considerably cooler. Some tests have been completed, and additional tests are planned to determine the operating temper-ature gradients throughout the drive mechanism during all phases of operation. These tests vill also provide an indication of the amount of convection that takes place within the water space of the mecha-nism. The more significant temperature changes will be caused by displacement of reactor coolant in and out of the mechanism water space as the drive line is raised and lowered.

d. Design Life l

The expected life of the control rod drive control system is as fol-lows: s c -, 238 3-82 m

D-B f_ 1. Structural portions, such as flanges and pressure housings, have an expected life of 40 years.

2. Moving parts, such as lead screw and roller nuts have an expected life of 20 years.

3. Electronic control circuitry has an expected life of 20 years.

3.2.4.3 5 Control Rod Assembly (CRA) Each control rod assembly (Figure 3-5k) has 16 control rods, a stainless steel spider, and a femala coupling. The 16 control rods are attached to the spider by means of a nut threaded to the upper shank of each rod. After assembly, all nuts are lock welded. The control rod drive is coupled to the CRA by a bayonet type connection. Full length guidance for the CRA is provided by the control rod guide tube of the. upper plenum assembly and by the fuel assembly guide tubes. The CRA's and guide tubes are designed with adequate flexibility and clearances to permit freedom of motion within the fuel assembly guide tubes throughout the stroke. Each control rod has a section of neutron absorber material. The absorber ma-terial is an alloy of silver-indium-cadmium. It is clad in cold-worked type 30L stainless steel tubing. Stainless steel end pieces are velded to the tub-ing to form a water- and pressure-tight container for the absorber material. The stainless steel tubing provides the structural strength of the control rods and prevents corrosion of the absorber material. A tube spacer sbsilar to that in the fael assembly is used to prevent a v arber motion within the cladding dur-ing shipping and handling, and to persit differential expansion in service. Principal data pertaining to the CRA are shown in Table 3-19 Table 3-19 Control Rod Assembly Data Item Data Number of CRA L9 Number of Control Rods per Assembly 16 Outside Diameter of Control Rod, in. 0.hh0 Cladding Thickness, in. 'O.021 Cladding Material Type 30h SS, Cold-Worked End Plug Material Type 30k SS, Annealed Spider Material SS Grade CF3M Poison Material 80% Ag, 15% In, 5% Cd Female Coupling Material Type 30h SS, Annealed ( Length of Poison Section, in. 134 StrokeofControlf. in. 139 3-83

D-B A CRA prototype of the BW design has been extensively tested at reactor tem-perature, pressure, and flow conditions in the P&W test loop at their Alliance Research. Laboratory. For test program description and results, refer to BAW-10007, " Control Rod Drive System Tee ^ Program." These control rods are designed to withstand all operating loads including those resulting from hydraulic force, thermal gradients, and reactor trip decelera-tion. The ability of the control rod clad to resist collapse has been demon-strated by a test program on cold-worked stainless steel tubing. Because the Ag-In-Cd alloy poison does not yield a gaseous product under irradiation, in-ternal pressure and swelling of the absorber material will not cause excessive stressing or stretching of the clad. Because of their length and the possible lack of straightness over the entire length of the rod, some interference between control rods and the fuel assem-bly guide tubes is expected. However, the parts involved, especially the con-trol rods, are flexible and only small friction drag loads result. Similarly, thermal distortions of the control rods are small because of the low heat gen-eration and adequate cooling. Consequently, control rod assemblies will not encounter.significant frictional resistance to their motion in the guide tubes. 2 The methods and frequency of CRA in-service inspection, as well as the criteria for replacement, will be determined during the detailed design. 3.2.h.3.6 Axial Power Shaping Rod Asse=bly (APSRA) Each axial power shaping rod asse=bly (Figure 3-55) has 16 axial power shaping i rods, a stainless steel spider, and a female coupling. The 16 rods are attached to the spider by means of a nut threaded to the upper shank of each rod. After assembly all nuts are lock welded. The axial power shaping rod drive is coupled to the APSRA by a bayonet connection. The female couplings of the APSRA and CRA have slight dimensional differences to ensure that each type of rod can only be coupled to the correct type of drive mechanism. When the APSRA is inserted into the fuel assembly it is guided by the guide tubes of the fuel assembly. Full length guidance of the APSRA is provided by the control rod guide tube of the upper plenum assembly. At the full out posi-tion of the control rod drive stroke, the lower end of the APSRA remains within the fuel assembly guide tube to maintain the continuity of guidance throughout the rod travel length. The APSRA's are designed to permit maximum conformity with the fuel assembly guide tube throughout travel. Each axial power shaping rod has a section of neutron absorber material. This absorber material is an alley of silver-indium-cadmium and is clad in cold-worked, stainless steel tubing with stainless steel upper and lower end pieces. The end pieces are welded M the clad to form a water- and pressure-tight con-tainer for the absorber mater M . The tubing provides the structural strength of the axial power shaping rods and prevents corrosion of the absorber material. Above the section containing the absorber =aterial is a tubular follower made of cold-worked Zircaloy-h tubing, with Zircaloy-k upper and lower end pieces. The end pieces are welded to the tubing and are vented to permit the coolant-3 moderator to fill the follower. The follower and absorber sections are fitted together, pinned, and lock welded to form a complete axial power shaping rod. Pertinent data on the APS$A is 'shown in Table 3-20. r . ., (h s Amendment No. 2 3-8h

D-B Table 3-20 Axial Power Shaping Rod Assembly Data Item Data Number of Axial Power Shaping Rod Assemblies 8 Number of Axial Power Shaping Rods per Assembly 16 Outside Diameter of Axial Power Shaping Rods, in. 0.hh0 Cladding Thickness, in. 0.021 Cladding Material Type 304 SS, cold-Worked Cita End Plug Material Type 304 CS, Annealed Follower Tube Material Ziresloy-4, Cold-Worked Follower End Plug Material Zircaloy k, Annealed Absorber Section to Follover Pin Material Type 30h SS, Annealed Poison Material 80% Ag, 15% In, 5% Cd Spider Material SS, Grade CF3M Female Coupling Material Type 30k SS, Annealed Length of Poison Section, in. 36 Stroke of Control Rod, in. 139 'k. These axial power shaping rods are designed to withstand all operating loads including those resulting from hydraulic forces and thermal gradients. The ability of the axial power shaping rod clad to resist collapse due to the sys-tem pressure has been demonstrated by an extensive collapse test prcgram on stainless steel tubing. Internal pressure is not generated within the clad since the Ag-In-Cd alloy does not yield gaseous products under irradiation. Swelling .cf the absorber material is negligible, and vill not cause unaccept-able clad strain. Mechanical interference between axial power shaping rods and the fuel assembly guide tubes can be tolerated, since the mechanical interference between axial power shaping rods and the fuel assembly guide tubes must be expected. The parts involved are flexible and result in very small friction drag loads.

     , Thermal distortions of the rods are small because of the lov heat generation and adequate cooling. Consequently, the APSRA's will not encounter signifi-cant frictional resistance to their motion in the guide tubes.

3.2.k.3.T Eurnable Poison Rod Assembly (BPRA) c Each EPRA (Figure 3-50) has 16 burnable poison rods, a stainless steel spider, and a coupling mechanism. The coupling mechanism and the 16 rods are attached to the spider. The EPRA is inserted into the fuel assembly guide tubes through the upper end fitting. The coupling mechanism provides a means for positive coupling between the BPRA and the fuel assembly holddown latch.

241

                           -d                   3-85

D-B Each burnable poison rod has a section of sintered A123 0 -BqC pellets which g pg serve as burnable poison. The burnable poison is clad in cold-worked Zirealoy- ) h tubing and Zircaloy-k upper and lower end pieces. The end pieces are welded ' to the tubing to form a water- and pressure-tight container for the absorber material. The Zircaloy-k tubing provides the structural strength of the burn-able poison rods. In addition to their nuclear function, the BPRA also serve to minimize guide tube bypass coolant flow. Pertinent data on the BPRA is shown in Table 3-21. Table 3-21 Burnable Poison Rod Assembly Data Item Data Number of Burnable Poison Rods per Assembly 16 Outside Diameter of Burnable Poison Rod, in. 0.430 Cladding Thickness, in. 0.035 Cladding Material Zircaloy-h, Cold-Worked End Cap Material Zircaloy-h, Annealed Poi: son Material Al 23 0 -BgC Length of Poison Section, in. :.hk

~')

Spider Material SS, Grade CF3M # Coupling Mechanism Material Type 30k SS, Annealed The burnable poison rods are designed to withstand all operating loads includ-ing those resulting from hydraulic forces and thermal gradients. The ability of the burnable poison rod clad to resist collapse due to the system pressure and internal pressure has been demonstrated by an extensive test program on cold-worked Zircaloy-4 tubing. 3.2.4.3.8 Orifice Rod Assembly (ORA) Each orifice rod assembly (Figure 3-49) has 16 orifice rods and orifice rod springs, a stainless steel spider, and a coupling mechanism. The coupling mechanism provides a means for positive coupling between the ORA and the fuel assembly holddown latch when the orifice reds are inserted into the fuel as-sembly. Springs located between the bottom of the spider and the orifice rods hold the rods firmly in the vertical direction, and permit lateral movement of the rods to facilitate insertion of the orifice rods into the fuel assembly guide tubes. The ORA serves to limit reactor coolant bypass flow through empty guide tubes. Pertinent data on the ORA is shown in Table 3-22.

2 [t 3-86

                                                           %mJ  O

D-B Table 3-22 Orifice Rod Assembly Data

      ,                   Item                                                                             Data Number of Orifice Rod Assemblies                     120 Number of Orifice Rods per Assembly                  16 Outside Diameter of Orifice Rod, in.                 0.k80 Orifice Rod Material                                 Type 30h SS, Annealed Spider Material                                      SS, Grade CF3M Coupling Mechanism Material                          Type 30k SS Annealed, and 17-k PH, Condition H1100 4

( t .i

(E (L ' t.-

m . . amor 243 3-87

                    ,2._'   'O_.,.          , _ _ .            , . . . . . _ _ - - , _ . . _ . _ _ _ . . _    -

D-B 3.3 TESTS AND INSPECTIONS QJ 3.3.1 NUCLEAR TESTS AND INSPECTION 3.3.1.1 Critical Experiments An experimental program (52-5h) to verify the relative reactivity worth of the CRA has recently been completed. Detailed testing established the worth of the CRA under various conditions similar to those for the reference core. These parameters include control rod arrangement in a CRA, fuel enrichments, fuel element geometry, CRA materials, and soluble boron concentration in the moderator. Gross and local power peaking were also studied, and three-dimensional power-peaking data were taken as a function of CRA insertion. Detailed peaking data were also taken between fuel assemblies and around the water holes left by withdrawn CRA. The experimental data are being analyzed and vill become part of the experimental bench mark for the analytical models used in the design. . 3.3.1.2 Zero Power, Approach to Power, and Power Testing Boron worth and CRA vorth (including stuck-CRA vorth) vill be determined by physics tests at the beginning of each core cycle. Recalibration of boron worth and CRA vorth is expected to, be performed at least once during each core cycle. Calculated values of boron vorth and CRA vorth vill be adjusted to the test values as necessary. The boron vorth and CRA vorth at a given time in core life vill be based on CRA position indication and calculated data as adjusted by experi= ental data. 3 The reactor coolant will be analyzed in the laboratory periodically to determine the boron concentration, and the reactivity held in boron vill then be calculated from the concentration and the reactivity worth of boren. The method of maintaining the hot shutdown margin (hence stuck-CRA margin) is related to operational characteristics (load pattern) and to the power-peaking restrictions on CRA patterns at power. The CRA pattern restrictions vill in-sure that sufficient reactivity is always fully withdrawn to provide adequate shutdown with the stuck-CEA margin. Power peaking as related to CRA patterns and shutdown margin vill be monitored by reactivity calculations. Operation under power conditions will normally be monitored by in-core instru-mentation, and the 'resulting data vill be analyzed and compared with multi-dimensional calculations to provide support for further power escalations. 3.3.2 THEPMAL AND HYDRAULIC TESTS AND INSPECTION 3.3.2.1 Reactor Vessel Flow Distribution and Pressure Drop Test A 1/6-scale model of the reactor vessel and " internals has been tested to evaluate:

a. The flow distribution to each fuel assembly of the reactor core and to develop any necessary devices required to produce the desired flow distribution.

N D 244 3-88

a D-B ( b. Fluid mixing between the vessel inlet nozzle and the core inlet, between the inlet and outlet of the core.

c. The overall pressure drop between the vessel inlet and outlet nozzles, and the pressure drop between various points in the reactor vessel flow circuit.

The reactor vessel, thermal shield, flow baffle, core barrel and plenum assem-bly are made of clear plastic to allow use of visual flow study tecuniques. All parts of the model except the core are geometrically similar to those in the prototype reactor. However, the simulated core was designed to maintain dynamic similari. between the model and crototyte. Each of the 177 simulated fuel assemblies contains a calibrated flow nozzle. The test loop is capable of supplying cold water (80 F) to three inlet nozzles and hot water (180 F) to the fourth. Temperature was measured in the inlet and outlet noz-zles of the reactor model and at the inlet and outlet of each of the fuel assemblies. Static pressure taps were located at suitable points along the flev path througn the vessel. This instrumentation provided the data necessary to accomplish the objec-tives set forth for the tests. The results of the test are reported in proprietary B&W Topical Report BAW 10012, " Reactor Vessel Model Flow Tests". 1 3.3.3 FUEL ASSEMBLY, CONTROL ROD ASSEMBLY, AND CONTROL ROD DRIVE MECHANICAL TESTS AND INSPECTION To de=cnstrate the mechanical adequacy and safety of the fuel assembly, control I rod assembly (CRA), and control rod drive, a nc=ber of functional tests have been performed, are in progress, or are in the final stages of preparatien. 3.3.3.1 Prototvne Testing A full-scale prototype fuel assembly, CRA, and control' rod drive have been tested in the Control Rod Drive Line (CRDL) Facility located at the B&W Research Center, Alliance , Ohio. This full-size loop is capable of simulating reactor environ-mental c^nditions cf pressure, temperature, and coolant flow. To verify the meensnical design, operating ec=patibility, and characteristics of the entire control rod drive fuel assembly system, the drive vill be stroked and tripped aptroximately 200 per cent of the expected operating life requirements. A per+. ion of the testir.g was performed with maximum misalignment conditions.

.luipment is available to record and verify data such as fuel assembly pressure iro;, vibration characteristics and hydraulic forces, and to demonstrate con-trei rod drive operation and verify scram times. All prototype components were exanined periodically for signs of material fretting, wear, and vibration /

f atigae to insure that the mechanical design of the equipment meets reactor cpersting requirements. Preliminary test results are noted in 3.2.h.3.5 and - given in BAW-10007, " Control Rod Drive System Test Program."

 , 3,. 3. 3. 2      Model Testing               .

Many functional inprovements have been incorporated in the design of the proto-typc fuel assembly as a result of model tests. Ecr example, the spacer grid to fuel red contact area was fabricated to 10 times: reactor size and tested in a icop simulating coolant flow Reynolds numbers of interest. Thus visually, the ears 245 3-89

D-B shape of the fuel rod support areas was optimized with respec'. to minimizing ,/'?s the severity of flow vortices. A 9-rod (3 x 3) assembly using stainless steel / spacer grid material has been tested at reactor conditions (6h0 F, 2,200 psi, 13 fps coolant flow) for 210 days. Two full-sized canned fuel assemblies with stainless steel spacer grids have been tested at reactor conditions, one for h0 days and t te other for 22 days. A prototype canless fuel assembly using Inconel 718 spacer grids has been tested for approximately 90 days, approxi-mately half of that time at reactor conditions. The principal objectives of these tests were to evaluate fuel assembly and fuel rod vibration and/or fret-ting wear resulting from ficv-induced vibration. Vibratory amplitudes have been found to be very small, and, with the exception of a few isolated instances which are attributed to pretest spacer grid damage, no unacceptable wear has been observed. 3.3.3.3 Commenent and/or Material Testing 3.3.3.3.1 Fuel Rod Cladding Extensive short time collapse testing was performed on Zircaloy k tube specimens as part of the B&W overall creep-collapse testing program. Initial test speci-mens were 0.h36-in. OD vith vall thicknesses of 0.020 in., 0.02h in., and 0.028 in. Ten 8-in. long specimens of each thickness were individually tested at 680 F at slowly increasing pressure until collapse occurred. Collapse pressures for the 0.020-in. vall thickness specimens ranged from 1,800 to 2,200 psig, the 0.00L-in. specimens ranged from 2,800 to 3,200 psig, and the 0.028-in. specimens ranged from h,500 to h,900 psig. The material yield strength of these specimens ranged from 65,000 to 72,000 psi at room temperature, and was 35,800 psi at 'N 650F. / Additional Zircaloy-L short time collapse specimens were prepared with a material yield stress of 78,000 psi at room temperature and h8,500 psi at 615 F. Fifteen specimens having an CD of 0.h10 in. and an ID of 0.365 in. (0.0225-in. nominal vall thickness) were tested at 615 F at increasing pressure until collapse occurred. Collapse pressures ranged frem h,h70 to 4,960 psig. Creep-collapse testing was performed on the 0.h36-in. OD specimens. Twelve specimens of 0.02L-in. vall thickness and 30 specimens of 0.028-in. vall thick-ness were tested in a single autoclave at 680 F and 2,050 psig. During this test, two 0.02h-in. vall thickness specimens collapsed during the first 30 days and two collapsed between 30 and 60 days. None of the 0.028-in. vall thickness specimens had collapsed after 60 days. Creep-collapse testing was then per-formed on thirty 0.h10-in. OD by 0.365-in. ID (0.0225-in nominal vall) speci-mens for 60 days at 615 F and 2,1h0 psig. None of these specimens collapsed, and there were no significant increases in ovality after 60 days. Results of the 60-day, creep-collapse testing on the 0.h10-in. OD specimens showed no indication of incipient collapse. The 60-day period for creep- - collapse testing is used since it exceeds the point of primary creep of the material, yet is sufficiently long to enter the stage when fuel rod pressure begins to build up during reactor operation, i.e. , past the point of maximum differential pressure that the clad would be subjected to in the reactor. These tests were followed by additional creep-collapse tests in which 60 speci-mens of variable vall thickness were subjected to a pressure of 2,085 psi at - e W,. - 3-90

                                                       ,j g 24(J

                                           "D-B 685 F until collapse occurred. The cladding vall thickness was 0.0285, 0.0263,

(' O.0251, and 0.02h0 inch. The cladding thickness included the range of toler-ances for production cladding, and the pressure represented the fuel rod maxi-mum pressure differential at operating conditions. The temperature was selected to conservatively approximate in-pile creep rates. It was found that the 0.02h-in. vall specimens collapsed in less than a month, and several 0.0263-in. vall specimens collapsed in less than 3 months. In view of the unknown increase of in-pile creep rates as compared with out-of-pile creep rates , it was decided to provide backup support for the cladding in the upper end void region where cladding temperatures of 650 F occur in hot channels. A thin-valled stainless steel tube was selected as a backup spacer. This spacer has the desirable property of providing radial support without causing large restraint for the axial expansion of the fuel. Development tests have been performed to select a spacer that can withstand shipping acceleration of fuel pellets and provide the required backup support for the cladding. The tubes were encapsulated in 0.016-in. vall Zircaloy tubing end subjected to a pressure of 2,750 psi at 850 F. This represents a 10 per cent margin on system ' design pressure at the operating temperature for the spacer tube. On section-ing the. specimens , there was no measurable deformation of the tube. 3.3.3.3.2 Fuel Assembly Structural Components The structural characteristice of the fuel assembly which are pertinent to loadings resulting from normal operation, handling, earthquake, and accident conditions vill be investigated experimentally in test facilities such as the CRDL Facility. Structural characteristics such as natural frequency and damp-ing vill be determined at the relatively high amplitude of interest in our seismic and LOCA analysis. Natural frequencies and amplitudes resulting from flow induced vibration vill be measured at various temperatures and flow velo-cities , up to reactor operating conditions. In the mechanical design of the spacer grids, particular attention is given to the ferrule-to-fuel-rod contact points. Sufficient load must be applied to position the fuel rods and to minimize fuel rod vibration, yet allow axial thermal differential expansion, and not produce fretting wear in the fuel rod cladding. Static load and functional testing of the prototype grids will demon-strate their adequacy to perform within the design requirements. 3.3.3.h Control Rod Drive Tests and Insnection 3.3.3.h.1 Control Rod Drive Developmental Tests The testing and development program for the roller nut drive has been completed. The prototype drive was tested at the B&W 7% search Center at A~.liance, Ohio. Wear characteristics of critical componen# 2 have indicated that material com-patibility and structural design of these components would be adequate for the design life of the mechanism. The trip time for the mechanism as determined under test conditions of reactor te=perature, pressure, and flow was well with-in the specification requirements. B&W Topical Report BAW-10007 su=marizes the results of the test program, i 247 3-91 GerP

                           . ca.g

D-B 3.3.3.k.2 Production Tests 'A g Production tests discussed in this section vill be performed either on the

     ' drives installed, or on drives manufactured to the same specifications. The finished control rod drive vill be proof-tested as a complete system, i.e. ,

mechanisms, motor control, and system control working as a system. This proof-

    -testing vill be above and beyond any developmental testing performed in the
    ' product development stages.

Mechanism production tests will include:

    ;       a. Ambient Tests Coupling tests.

Operating speeds. Position indication. - Trip Tests.

b. Operational Tests Operating speeds.

Position indication, 3.3.k fp)

               . INTERNALS TESTS AND INSPECTION The internals upper and lower plenum hydraulic design vill be evaluated and guided by the results from the 1/6-scale model flow test which is described in detail in 3.3.2.1. These test results will indicate areas of gross flev mal-distribution and allow verification of vessel flow-pressure drop computations.

In addition, the test results will provide measured pressure pulses at specific locations to aid in assessing the vibration response characteristics of the in-ternals components. The effects of internals misalignment will be evaluated on the basis of the test results from the ORDL tests described in 3.3.3.4. - These test results, when correlated with the internals guide tube final design, will insure that the CRA vill have the capability for a reactor trip or fast insertion under all modes of reactor operation in the reactor coolant environment. These tests will not include the effects of neutron flux exposure. All internals components can be removed from the reactor vessel to allow in-spection of all vessel interior surfaces (see 4.4.1). Internals components surfaces can be inspe.cted when the internals are removed to the canal storage location. et . .. 248

                                       .i.06J                -

20mm. d 3-92 4

D-B

    ~     3.h       REFERENCES

( (1) Saxton, Large Closed-Cycle Water Research & Development Work Program for the Period July 1 to December 31,196h, WCAP-3269 h. (2) Bohl, H. , Jr. , et al. , P3MG1, A One-Dimensional Multigroup P-3 Program for the Philco-2000 Computer, WAPD-TM-272. (3) Bohl, H., Jr. and Hemphill, A. P., MUFr-5, A Fast Neutron Spectrum Pro-gram for the Philco-2000, WAPD-TM-218. (h) Armster, H. J. and Callahan, J. C. , KATE-1, A Program for Calculating Wigner-Wilkins and Maxwellian-Averaged Thermal Constants on the Philco-2000, WAPD-TM-232. (5) Marlov, O. J. and Suggs, M. C., WANDA-5, A One-Dimensional Neutron Dif-fusion Equation Program for the Philco-2000 Computer, WAPD-TM-2hl. ( 6 ', Honeck, H. C. , THERMOS, A Thermalization Transport Theory Code for Reac-tor Lattices, BNL-5826. (7) Cadwell, W. R., Buerger, P. F., and Pfeifer, C. J., The PDQ-5 and PDQ-6 Programs for the Solution of the Two-Dimensional Neutron Diffusion-Depletion Problem, WAPD-TM-h77. (8) Marlove, O. J. , Nuclear Reactor Depletion Programs for the Philco-2000 ( Computer, WAPD-TM-221. (9) Lathrop, K. P., DTF-IV, A FORTRAN-IV Program for Solving the Multigroup Transport Equation With Anisotropic Scattering, LA-3373. (10) Joanou, G. D. and Dudek, J. S. , GAM-1: A Consistent P1 Multigroup Code for the Calculation of Fast Neutron Spectra and Multigroup Constants, GA-1850. (11) Clark, R. H. and Pitts, T. G., Physics Verification Experiments, Core I, BAW-TM h55. (12) Clark, R. H. and Pitts , T. G. , Physics Verification Experiments , Cores II and III, BAW-TM h58. (13) Spinks, N., "The Extrapolation Distance at the Surface of a Grey Cylin-drical Control Rod," Nuclear Science and Engineering, 22, pp 87-93,1965. (lk) Poncelet, C. G. and Christie, A. M., The Effect of a Finite Time Step Length on Calculated Xenon Ct bi3 ity Characteristics in Large PWR's , ANS Winter Meeting, Noverder 19o7 (15) Tong, L. S. , DNB Prediction for and Axially Nonuniform Heat Flux Distri-bution, WCAP-558h, September 1965 J 249 3-93 w

D-B (16) Tong, L. S. , An Evaluation of the Departure From Nucleate Boiling in O Bundles of Reactor Fuel Rods, Nuclear Science and Engineering, y, 'l' pp T-15,1968. (IT) U.S.-Euratom Joint' R&D Program, Burnout Flow Inside Round Tubes With Nonuniform Heat Fluxes, The Babcock & Wilcox Company, BAW-3238-9, May 1966. (18) Jens , W. H' and Lottes , P. A. , Analysis of Heat Transfer Burncut , Pres-sure Drop, and Density Data for High-Pressure Water, ANL h627, May 1951. (19) Owen, D. B. , Factors for One-Sided Tolerance Limits and for Variable Sampling Plans, SCR-607, March 1963 (20) Bowring, R. W. , Physical Model, Based on Bubble Detachment , and Calcula-tion of Steam Voidage in the Subcooled Region of a Heated Channel, HPR-10, OECD Halden Reactor Project, December 1962. (21) Zuber, N. and Findlay, J. A. , Average Volumetric Concentrations in Two Phase Flov Systems, Presented at the ASME Winter Meeting,196h. To be published in the ASME Transactions. (22) Maurer, G. W. , A Method cf Predicting Steady-State Boiling Vapor Frac-tions in Reactor Coolant Channels, Bettis Technical Review, WAPD-BT-19. (23) Baker, O. Simultaneous Flov of 011 and Gas , Oil and Gas Joumal. Vol 53, pp 185-195, 195h. m s j (24) Rose, S. C. , Jr. , and Griffith, P. , Flow Properties of Bubbly Mixtures , ASME Paper No. 65-HT-38,1965. (25) Haberstroh, R. D. and Griffith, P. , The Transition From the Annular to the Slug Flow Regime in Two-Phase Flow, MIT TR 5003-28, Department of Mechanical Engineering, MIT, June 196h. (26) Bergles, A. E. and Suo, M. , Investigation of Boiling Water Flow Regimes at High Pressure, NYo-330h-8, February 1, 1966. (27) Notley, M. J. F., The Thermal Conductivity of Columnar Grains in Irradi-ated UO2 Fuel Elements, AECL-1822, July 1963. (28) Lyons, M. F., et al., UO2 Fuel Rod Operation With Gross Central Melting, GEAP-h26h, October 1963. (29) Notley, M. J. F., et al., Zircaloy-Sheathed UO 2 Fuel Elements Irradiated at Values of Integral kd8 Between 30 and 83 v/cm, AECL-1676, December 1962. . (30) Bain, A. S. , Melting of UO 2 During Irradiations of Short Duration, AECL-2289, August 1965 c eage P ) L: w . 3-9h A

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 ,- (31)  Notley, M. J. F. , et al. , The Longitudinal and Diametral Expansions of UO2 Fuel Elements, AECL-21h3, November 196h.

(32) Lyons, M. F., et al., UO2 Pellet Thermal Conductivity From Irradiations With Central Melting, GEAP h62h, July 196h. (33) Ross, A. M. and Stoute, R. L. , Heat Transfer Coefficients Between UO2 and Zircaloy-2, AECL-1552, June 1962. (3k) Kjaerheim, G. and Rolstad, E. , Inpile Determination of UO2 Thermal Con-ductivity, Density Effects , and Gap Conductance, HPR-80, December 1967. (35) Lyons, M. F., et al., Reactor Power High Performance UO2 Program - Fuel Design Summary and Program Status , GEAP-5591, January 1968. (36) Hoffman, J. P. and Copling, D. H. , The Release of Fission Gases From Uranium Dioxide Pellet Fuel Operated at High Temperatures, GEAP h596, September 196h. . (37) Spolaris, C. N. and Megerth, F. H., Residual and Fission Gas Release From Uranium Dioxide, GEAP-431h, July 1963. (38) Robertson, J. A. L., et al., Behavior of Uranium Dioxide as a Reactor Fuel, AECL-603, 1958. (39) Parker, G. W. , et al. , Fission Product Release From UO2 by High Tempera- - ture Diffusion and Melting in Helium and Air, CF-60-12-lh , ORNL, February 1961. (h0) Duncombe, E. , Effects of Fuel Cracking, Void Migration, and Clad Collapse in Oxide Fuel Rods, Trans. ANS 11(1), p 132, June 1968. (hl) Bain, A. S. , Microscopic, Autoradiographic and Fuel /Shasth Heat Transfer Studies on UO2 Fuel Elements, AECL-2588, June 1966. (h2) Balfour, M. G. , Post-Irradiation Examinatiot of CVTR Fuel Assemblies, WCAP-3850-2, March 1968. (h3) Daniel, R. C., et al., Effects of High Burnup on Zirealcy-Clad, Bulk UO 2 s Plate Fuel Element Samples, WAPD-263, September 1962. (hh) Fracture of Cylindrical Fuel Rod Cladding Due to Plastic Instability, WAPD-TM-651, April 1967 l (h5) Duncan, R. N. , Rabbit Capsule Irradiation of UO 2, CVNA-lh2, June 1962. (h6) Duncan, R..N., CVTR Fuel Capsule Irradiations, CVNA-153, August 1962. (h7) Frost, Bradbury, and Griffiths ( AERE Harwell), Irradiation Effects in Fissile Oxides and Carbides at Lov and High Burnup Levels, Proceedings of IAEA Symposium on Radiation Damage in Solids .and Reactor Materials, Venice, Italy,,May 1962.

,,j

(

                              ;.( -

e 3-95

                                                         .      auer                251
                .s

D-B (h8) Gerhart, J. M. , The Post-Irradiation Examination of a Pu0 -UO 2 2 Fast Reactor Fuel, GEAP-3833. .

(h9) Physical and Mechanical Properties of Zirecloy-2 and h, WCAP-3269 bl, Figure 18. (50) Burgreen, D. , Byrnes , J. J. , and Benforado , D. M. , " Vibration of Rods Induced by Water in Parallel Flov," Trans. ASME 80, p 991,1958. (51) Large Closed-Cycle Water Reactor R&D Progran, Progress Report- for the Period January 1 to March 31, 1965, WCAP-3269-12. (52) Clark, R. H., Physics Verification Experiments, Cores IV and V, BAW-TM-178, September 1966. (53) Clark, R. H., Physics Verification Experiment, Core VI, BAW-TM-179, December 1966. (Sh) Clark, R. H., Physics Verification Experiment, Axial Power Mapping on Core IV, BAW-TM-255, December 1966. (55) Svenson, H. W., Carver, J. R., and Kakarala, C. R., The Influence of Axial Heat Flux Distribution on the Departure From Nucleate Boiling in a Water Cooled Tube, ASME Pacer 62-WA-297. (56) Burnout for Flow Inside Round Tubes With Nonuniform Heat Fluxes, BAW-

     ~3238-9, May 1966.

(57) O)

    ' Nonuniform Heat Generation Experimental Program, BAW-3238-13, July 1966.

(58) Wilson, R. H. and Ferrell, J. K. , Correlation of Critical Heat Flux for Boiling Water in Forced Circulation at Elevated Pressures, The Babcock & Wilcox Company, BAW-168, November 1961. b

                        'J ~

252 J 3-96 %

D-B 3abcock & Wilcox Topical Report References BAW-10010 Stability Margin for Xenon Oscillations -

         .                               Medal Analysis 4

BAW-1001h Analysis of Sustained Departure From Nucleate Boiling Operation . BAW-10012 Reactor Vessel Model Flow Tests ! l EAW-10007 Control Rod Drive System Test Program i , 1 J t i I ! i ( 253 h aump l Y::,0 3-97 f L ._ ._ . _ - _ _ _ _ . ~ . ._ _

(-. 1200 , , , i i I l 1000 g ,

      =          l
       .                                                                      First Cycle
                                                                                                   ~
     .$                                                   (with burnable poison)

%\, N c 800 g c \

3 _ \ _ E \ 6 00 % 8

     ;                              \

g _ \ _ '( -

E \ ( g 400 g u \ g -

                                                  \                                                 -

t \

      "                                               4 200                                           \                                \
                                                           \

Equilibrium Cycle \

                 - (no burnable poison)                                \                            -

g

I I I I

0 0 94 188 282 376 470 i Core Lif e, Ef f ective Full (Rated) Power. Days 1

0 AVIS-BESSE NUCLEAR POWER STATION BORON CONCENTRATION VERSUS CORE LIFE FIGURE 3-1 ( 84Wir 254

Axial Power Profile for 555 Insertion is Shown on l Figure 3-3 ej i 1

                                                             .I 1.7                                                      -
                                     /
                ~

1.6 > e -

                    /                                                        \

( N 15 33 N

E h w 1.4 E* l.3 ' ' I I I I I 10 20 30 40 50 60 70 80 Rod Insertion, ", DAVIS-BESSE NUCLEAR POWER STATION AXlAL PEAK TO AVERAGE POWER VERSUS XENON OVERRIDE R00 INSERTION FIGURE 3-2 ( 255 8mes-

~.

f 1.8 1.6 / \ 1.4

1. 2 io E

1.0 -

                      /                          N    '

0.8 ) \ E % 0.6 3 h 4 0.4 \ T

0. 2 ._ _

T 0 I"" - 10 20 30 40 50 60 70 80 90 100 110 120 130 140 150 Distance f rom Bottom of Active Fuel, in. DAVIS-BESSE NUCLEAR POWER STATION AXlAL POWER PROFILE XENON OVERRIDE RODS 55 PERCENT INSERTED FIGURE 3-3 ( 256 l l 1M . 1

i l i ( l X X l 1 X X 'X X X ' ' X X X X X X X X X X X X X X X X X ! X X XXX XX X x ! X X X X X X c X X X XX XX X X X X X X X X X X X X X X X X X X X X X X X X X X DAVIS-BESSE NUCLEAR POWER STATION

k. LOCATION OF FUEL ASSEMBLIES CONTAINING BURNABLE POISON ROOS

                'g                  FIGURE 3-4         257 AMENDMENT NO. I

l 110 100 3

                            \
                               \
                                \

t 80 -t i

        *3                           6 is                           '.

E% EO ' a t c E \

        =;                                le 2.41Ak/h
        ~c                                 f t

40 g

5. 45ak/k i g

        '                                         s-20 g_                           -

8 DAVIS-BESSE NUCLEAR POWER STATION 0 i 2 4 5 6 7 8 PER CENT NEUTRON POWER VERSUS TillE FOLLOWING TRIP FIGURE 3-0 iv j cn CD i

2.5 2.6 - 2.4 - 2.2 - 2.0 - 1.8 - Leeer Core s'" 8.5 - f"% g / N /

                                                                       ,r"N g
                                                                              \       /

f

                                                                                          /
                                                                                            #~\     \

f

P/P l.4 -- g

                              / --     N      /              \.    /
                                                                                \   /                 \          j
                         ' '               /              '- O' -- M <'
.          i., -                                                                               -
                                                                                                       \ /__ __Q
                                          \                     \ (              "7 f%          ,e jg ' -

1.o - y N, . . , N\ --

                                                           /
                                                             ,i \                j\ g                 /

i , g / g

             * ~                                  ' "'

Upper Core \ / \ j \ g 6- \/

             .4  -
             .2 -
             .0 O                 I      .

2 3 4 5 Ti.e (T). days NOTES:

i. Pe.e, natie isien 3e in tre. top and noite. er act..e i.ei.

Case I - Ne temperature storateen, fe,,g. 1.400 F. DAVIS-BESSE NUCLEAR POWER STATION Case 2 - T..perature iteration eith fe,si i.4o0 F. Case 3 Is.perature sterateen .ith EFFECT OF FUEL TEMPERATURE (D0PPLER) ON 2. vuel* III I-Oscillation inateated at i . 2 days. XENON OSCILLATIONS - BEGINNING OF LIFE N FIGURE 3-6

    .D so

I 2.4 2.2 -

^ ~

Loser Core

1. 8 - --

1. 6 -

1.4 - pfp I.2 -

1. 0 -

             .8-                                                          l   l
             .6-       /

g~ \ -. -

             .4 -                                                    ~~     y
             .2 -          Upper Core 0

s i O I 2 3 4 isme (i), days DAVIS-BESSE NUCLEAR POWER STATION

1. roser natio taken 36 in. from top and nottom of active tuei.

Case 1 - Temperature Iteration eith I tuel - 3 488 I- EFFECT OF FUEL TEllPERATURE (D0PPLER) ON Case 2 - Temperature Iteration with ifuel . 800 F. XENON OSCILLATIONS - NEAR END OF LIFE

2. Oscillation initiated at T - 388 days. FIGURE 3-7 s

CD i f

I ( 2.5 - 2.4 - , 2.2 _ Upper Core l / [ \ '

2.0 - / g

1.8 - - g

                                                ,-s                                             g
                                                       \                                       I                      \

1.5 - \ 2

                               'l 1.4 -                                                f 1.2 -

A : s g j 1.0 - I f II \# \ f Y , *

                                         .8 - l              \                       s p              \                     s                                       \

s- i

                                                               \                  /

l \

                                                                                                                              \y s                !
                                        .4 -

loser Core g / \

                                         .2-                        g N_'
                                                                             /                                                    \
                                                                                                                                    \s
                                        .0 I                                i 0                           1                               2 se (T ). days NOTES:

1. Case 1 - Divergent oscellation (without temperature iteration).

Case 2 - Poser ratio variation with control (without temperature steration).

2. Oscillation initiated at T . 200 days.

DAVIS-BESSE NUCLEAR POWER STATION CONTROL OF AXIAL OSCILLATION WITH PARTIAL RODS N FIGURE 3-6

AMEN 0 MENT NO. 1

( l l

                                +0.15-                                                                                   '

l

                                + 0.10 6
                      =
                               + 0. 05 3

E Min BOL t D 00 . f i

                               -0.05 Min 80L g Ref EOL Ref BOL M i
                              -0.10 0             25             50                75           100 Flatness, %

DAVIS-BESSE NUCLEAR POWER STATION STABILITY INDEX VERSUS FLATNESS ( BEGINNING AND END OF LIFE (AZIMUTHAL) FIGURE 3-9 2<5 2

(l.P) (P) 1.0 a e e i 0 Nuncer of Data Points . 809 ( Mean Value . 0.996 Standard Deviation . 0.132 0.1 0. 9 94.31 Population Protected 0.01 0.99 99.00!. Population Protected 0.001 \, 0.999 D48R = 1.30 0.0001 0.9999 (DNBR=1.50 \ (Minimum DNBR at Design Conds tions) 0.00001 0.99999 1.0 1. 2 1.4 1. 6 1.6 2. 0 2. 2 2. 4 DNB Ratio (W-3) DAVIS-BESSE NUCLEAR POWER STATION POPULATION PROTECTED. P ANO l-P - ( , VERSUS ONB RATIO (W 3) 265 FIGURE 3-10 O

 .-.     . .   . - - - .              - . - - - . ~ .                                .-

1,8 i i i ig i e i I 1.7 - P/P = l . 70 (Pa r 1.i a l 1.6 Rod Insertion l I j /1 N P/ P - 1. 50 l4 !

7 4 (Rodifind Cosine) 1.3 - / I -

                                                                                                                  \                                          -

1.2 / i

                                                                       /                                                 \

1.1 - I -

                                                                    /
                                                                  /                           I                            \                                                 i 1.0                                                                  -         '                    '
                                                                /                             I
                                                                                                                                \                                            !

l

             'R                                               /
                                                                                                                                 \

l

0. 9 l

                                                           /                                  l                                    \

l 0' 8 0.7 -

                                                        /
                                                         /                                   l                     \                 \

l \ f I \

0. 6 l I '

     <                                                                                                                                     \                                 l

0. 5 -

                                         /                                                                                                   \ %-           -

l 0.4

I 0.3 - ! -
Fuei lidplane i 0.2 -

1 Core l Core 0.1 > Botton l Top - 144" - 0.0 1 i l i l i l i I i i i 1 0 20 40 60 80 100 120 140 10 30 50 70 90 110 130 Distance from Bottom of Active Fuel, in. DAVIS-BESSE NUCLEAR POWER STATION POWER SHAPE REFLECTING INCREASE 0 AXIAL l POWER PEAK FOR 144-INCH CORE l (. FIGURE 3-11

                                                                                                                                  . .~ .          __

f 1.80 N 1,60 \ - 1 40 \ h

d i.20 2_

       .53 1. 00 h  '
       %u fb  0.80                                   ,

33 2 0~60 2 0.40 , \, 0.00 0 10 20 30 40 50 60 70 80 90 100 Percentage of Fuel Rods with Higher Peaking Factors Than Point Values,% , 4 I DAVIS-BESSE NUCLEAR POWER STATION DISTRIBUTION OF FUEL ROD PEAKING l l g FIGURE 3-12 265

    .~ ._                                                   .          -      _          .. _ _ .

(e 100

            ,                                             Line i E                                             95% Flow                     ,

e w M C l.3 DNBR Limit a f. 60 z 3 l12% Overoower , (- , o

,\

S b

          $    20 5           A Line 2 100% Flow 0

100 110 12D l Rated Power (2633 MWt) % DAVIS-BESSE NUCLEAR POWER STATION POSSIBLE FUEL R00 DNB'S FGR MAXIMUM - DESIGN CONDITIONS - 36,816 R00 CORE FIGURE 3-13 O 2bb

20 m E

   =    16 o                                     I 1                                                            F  -
    ;; 12 a.

Maximum Overpower +

g (l12%) E

   ,      8 C

2 g

0 E \ 0 1 20 i 105 l10 115 100 Rated Power (2633 MWt), % DAVIS BESSE NUCLEAR POWER STATION POSSIBLE FUEL'R00 DNB'S FOR MOST PROBABLE CON 0lT10NS - 3 6,816 - R00 CORE FIGURE 3-14 _ (- 1 m 267

(1-P) (P)

             .I                                                                                   '9

[

          .01                                                                                     '99 k

001 . '999 i

                      \
                       \

g ! L I

        .0001                    ,                                                                 9999 X    _

N A A,

     .00001                                                                                        99999 0         10       20   30      tiO   50      60       70      80      90     100 Percentage of Rods With A Lower Value of P, 7 DAVIS-BESSE NUCLEAR POWER STATION DISTRIBUTION OF POPULATION PROTECTED P, AND 1-P VERSUS NUMBER OF RODS FOR MOST PROBABLE CONDITIONS A.T THE DESIGN l[

i OVERPOWER OF 112", FIGURE 3-15

                                                              .' M              AMEN 0 MENT NO. 3        268

2.0 6 1.8 m E 99% Confidence Basi s bo 1 .6 b I.50

      ,.E        _ _. _        -

U ! l.4 i I t.v I

\

j l.2 l l f 5

     =                           l f    1.0

j o I Design h

Overpower (l12%) E I 0.8 l l l I O. 6 ! 100 l10 120 130 1I40 150 Rated Power (2633 MWt), % DAVIS-BESSE NUCLEAR POWER STATION , DNB RATIOS (W-3) IN HOT UNIT CELL l VERSUS REACTOR P0nER l ( FIGURE 3-16 l M~~ AMENDMENT NO. 3 2(39 l l

1 [, 20 85 18 - 87 2120 PSIG 16 > 90 e gg _ 2185 PSIG - 93 $2

                                                                                                    $5 12                                  ,                                           96       4*

_7 E%

u. c 10 -

99 y S-

,?

8 I 103 EC

                        /       f                                                                   u   N 4               /                                                                            Ie b    6   -
                   /
                                                                                       -   109 s e.

i I 4 - 5*" 116 g. E' s 2 - 1 27 y w Qual i ty ggg 8 I l Subcooled

         -2   -

Design Overpower -

         -4
         -6 100 110        120           1 30            140              !50       I60 RatedPower(2633MWt),7 i

I l DAVIS-BESSE NUCLEAR POWER STATION

                                                '4AXIMUM HOT CHANNEL EXIT QUALITY iI
    '                                                  VERSUS REACTOR POWER S                           FIGURE 3-17

3.0

2. 5 ,

2 Pumps- 3 Pumps 7 g 2 Loop 6

g 2.0 2 ~ E 5 1.5

1. 0 50 60 70 80 90 100 110 Reactor Power (2633 MWt), 7 DAVIS BESSE NUCLEAR POWER STATION ll0T CHANNEL DNB RATIO (W-3) VERSUS POWER -

FOR PARTIAL PUMP OPERATION FIGURE 3-18 ( - 271 amt

I 20.0

c 5 6

I 15.0

E e 2 PUMP - 2 LOOP

       .e
       .5 x

g 10.0 c ( - m 5 3 j 5.0 5- 3 PUMP T* 8 v 0.0 50 60 70 80 90 100 Reactor Power (2633 MWt), % DAVIS-BESSE NUCLEAR POWER STATION i HOT CHANNEL QUAllT) AT POINT OF MIN. DNBR VERSUS POWER FOR PARTIAL l PUMP OPERATION FIGURE 3-19 m 272

4. 00 g

c'

                    .                                                           I
                                                                              .i U02 ge l ts     c'i l

l 1 l-

      't to-                                                                     i 3

J. 3. 00 E, l Data Based On M00, I 7 EEAP 4624

l, 1 l l

       .?                                           N'L k de = 93 w/ cm ]

t I i I ( 0 2 2. m l e 1 5 1 d I j 1

1. 00 0 10 2 2000 3000 4000 5000 !

Temperature, F - 1

DAVIS-BESSE NUCLEAR POWER STATION THERMAL CONDUCTIVl!Y FIGURE 3-20 2,73 3

l l 6000 , , l ( 002 Helting l Temperatu re 5000 1 I l H)00 # u. J b , [ 3000 l l 5 ! c I l 0 1007. Power - I

  ,        I

l 2000 i 1127. Power i

l l 1000 ) l l l l 0 5 10 15 20 25 I l J Linear Heat Rate, kw/ f t

l l l l l 7' DAVIS-BESSE NUCLEAR POWER STATION FUEL CENTER TEMPERATURE AT THE HOT

                                            -      O                    SPOT VER$US LINEAR POWER         7 FIGURE 3-21

70 - - Gaussian Distribution 80 - /

      . 50 -                                ~

2-

     "  40 -

o

     $  30 -
     =                                               .

20 . ~ ( 10 0 - 0.6 cd, 4(1 0.8 i.0 i.2 i.4 i .'s i .'s

+l+

EC DAVIS-BESSE NUCLEAR POWER STATION NUMBER OF DATA POINTS VERSUSE D .OC - FIGURE 3-22 1 275

[ l.025 - l 1.020 -

    .                                                                        l 1.015    -

1.010 - l Fg 1.005 - g 1.000 60 70 80 90 100 0 995 t 0.99 0 - a h 0.985 - FA (Interior Bundle Cell 0.980 - 0.975 - 0.970 - FA (Wall Bundle Cell) 0.965 - 0.960 - Population Protected, % DAVIS-BESSE NUCLEAR POWER STATION HOT CHANNEL FACTORS VERSUS PER CENT POPULATION PROTECTED FIGURE 3-23 6 276

l 100 l Infinite Sample 100% Confidence 4 r-0 j _

                                       - F i n i te S amp le -

o 90% Confidence a. o F i n i t e S amp l e-79 997. Confidence 5 ( -

60 50 1.0 1. I 1.2 1. 3 1.4 1.5 1.6 1.7 Burnout Factor, DNB Ratio (W-3) DAVIS BESSE NUCLEAR POWER STATION BURNOUT FACTOR (W-3) VERSUS - POPULATION FOR VARIOUS CONFIDENCE LEVELS FIGURE 3-24 b 277

D 1 (_ 16 Hot Channel With 5% 14 - Flow Distribution - Facto r 12 10 - 8

                                                                 - Nominal Channel 6    -          l                                            Without Flow  _

j f

                                         /                           Distribution Factor i   2   -
                      /
                        / /                                                       -

0 9"* ' # [ Subcooled

      -2    -
      -4 l
      -6    -
                                                                                  -                        I Design Overpower (ll2%)
      -8 I

l 100 110 120, I30 140 150 Rated Power .(2633 MWt) % j DAVIS-BESSE NUCLEAR POWER STATION DESIGN HOT CHANNEL AND NOMINAL CHANNEL EXIT QUALITIES VERSUS REACTOR POWER (WITHOUT e ENGINEERING HOT CHANNEL FACTORS) ! C. FIGLIRE 3-25 h

                                           %-                                         278                  ,

l l

Sundle Burnout Test Conditions here Staale Operations u

( Were Observed A Worst Condi tions, l12% Power B Worst Conditions,130% Power ! O lioninal Conditions, ll2% Power 4 Ilominal Conditions,130% Power

3. 0 15 0 .

A

e 2.0 '

M l 4.s b ,. y . . . 7 . . . t .

       $      l.5 o

l t 3

z . g, , . **

l 1.0 Bubble To Annular 1 ( Baker) ,

                                                                   .                *. .      ,*            .   .j      a
               .5                                                                                               .Y
                                                                                                            )
                                                                                ~
                                                           ':.:::- v                                                                 .

0 5 10 15 20 25 30 Quality (lb vapor / total Ib), % ( DAVIS-BESSE NUCLEAR POWER STATION FLOW REGIME U P FOR THE HOT UNIT CELL

            .                                                                                        FIGURE 3-26 K['              ,.

279

l + Bundle Burnout Test Conditions where Stable Operations ( i were Observed. A Worst Conditjons. I12% Power a Worst Conditions 1301 Power g Nominal Conditions,112% Power 4 Nominal Conditions, 130% Power t ++ t

2. 5 +

9 y +

                                 +   ++++,+',       +e       .
                                                                      +
                                                                         ++
                                     +

o. M 2.0 m+ 5 3 i ,+ + tF + + +** + , i + o ( 3 1.5 h .. >%+++ , ,j + Bubble To 2

Annular

                                                                         .         ( Baker) 1.0
                                                                                                     )
                                                                                      +          +
                                                         +        .        ...      .                .  :+
             .5                                        Bubble To
                                                                                                     ).

Slug (Baker)

                -5 0               5            10             15             20            25             30 Quality (lb vapor / total I b) . Y, DAVIS-BESSE NUCLEAR POWER STATION

( ~ FLOW REGINE WAP FOR THE HOT CONTROL R00 CELL FIGURE 3 27

._

Bundle Burnout Test Conditions where Stable Operations

[' Were Observed. A Worst Conditions, ll2% Power a Worst Conditions. 130% Power g Mominal Condi tions.112% Power 4 Moeinal Conditions,130% Power 2.5 db o - y . o ..

       =                            A w                                                                .

2.0 E a 7 j . . ". . . =. . . . Bubble To ( 2 1. 5 Annular . E

I . ( Baker)

                                               ,    y      ,        , ..               ..

1.0

                                                                                                                    \
                                                                          .                   .                    .~ ~

Bubble To )

            .5                                                                                         -

Slug (Baker) d

                -5
                                                                                                                                   ~

0 5 10 15 20 25 30 Quality (It vapor / total Ib), % DAVIS'BESSE NUCLEAR POWER STATION

O : LOW REGIME MAP FOR THE HOT WALL CELL FIGURE 3-28 281 I

e- Sundle Burnout Test Conditions Where Stable Ooerations Were Observed [ A Worst Conditiens,112% Power a Worst Conditio.es, 130% Power ) g Nominal Conditions,ll2% Power 4' Nominal Conditions,130% Power 3.0

                                                                                                            ..             \

2. 5 a 4 ,

i T .. E A . i

2. 0 ::

                 't                                                                                                                                        !

Tw . g . . e . . *. . e . ( j l.5

Bubble To y Annular ( Baker)

                                                                                   .. .             ..                        **                           1
                                                         . go .         .                                                                              !

l.0 l

                                                                     .                 o                 .* .        ,
                       *5 Bubble To Srug (Baker)                                       2               .}
                          -5     0                5              10                       15                      20                  25              30

Quality (Ib vapor / total Ib), T.

( DAVil-BESSE NUCLEAR POWER STATION 1 Fl.0W M GIME MAP FOR HOT CORNER CELL l l FIGURE 3-29 d l 1 w _ _ _ . , ,

l l

2. 0 l.8 2-T Design Overpower k l.6 4 --

E

               .2
              =

1. 5 Cosine am g 1.4 '

              -                                        \

E r 1.65 Cosine C 3 g I.2 \l l g .8 Cosine 1 I .

l.0 I l I 0.8 < 100 110 120 ~ l 30 140 150 Rated Power (2633 MWt), 7 DAVIS-BESSE NilCLEAR POWER STATION HOT CHANNEL DNB RATIO (W-3) VERSUS POWER FOR VARIOUS AXIAL FLUX SHAPES FIGURE 3-30 gg3 iMEUB.

150 1140 Design Flowrate-y ~~ ~ ( 131. 32 X 10 6 LB/HR) x 130

     .5 2

i

     -   120                                                           I/

r u

     .3 2                                                                l "E  I10                                                r

( - l l Design Overpower (l127.X2633MWt) 100

                                                /
                                                                     \

l DNBR (W-3) . l. 30 90 l l l l 2I40 0 2600 2600 3000 3200 3%0 Reactor Core Power, NWt DAVIS-BESSE NUCLEAR POWER STATION REACTOR COOLANT SYSTEM FLOW VERSUS POWER g FIGURE 3-31

}g4

(

2. 0

                                                                  ,               i MIXING LINE                FLOW   COEFFICI ENT            l 1                 l10%      .02 a

1. 8 - 2 100% .02 a ~

i g 3 90% .02 a ! 4 100% .06 a 7 \\ 5 100% .01 a b 1.6 i I

      ,o i'
                                    \
      .                           \

E g\ j l.4 -

                                         \                             1.30 (W-3) o                                     \i

(

1. 2 A '

io 3 0.8 100 llo 120 130 140 150 Rated Power (2633 MWt), % DAVIS-BESSE NUCLEAR POWER STATION HOT CHANNEL DNB RATIO (W-3) VERSUS POWER WITH REACTOR SYSTEM FLOW AND ENERGY MIXlNG AS PARAMETERS FIGURE 3 ( M' ~~

                                                    ~
                                                                                          . 285

6000 " 5500 U02 Melting Temperature 5000 u g J

      %                                                   )
       ; 4500
       =

v 112% Overpower h 1005 Power

( 3500 3000

7 2500 6 10 14 18 22 26 30 Linear Heat Rate, kw/ft DAVIS-BESSE NUCLEAR POWER STATION

FUEL CENTER TEMPERATURE FOR BEGINNING ( OF LIFE CONulTIONS Figure 3-33 N . 286

( 6000 1 5500 UO2 Melting l Temperature 5000 d !

                % 4500                                      2 Lesign
                ,                                                                     Overpower 2 4000                                                                                 i

(

3500 1005 Power 3000 7 2500 - 6 10 14 18 22 26 30 ) Linear Heat Rate,kw/ft DAVIS-BESSE NUCLEAR POWER STATION FUEL CENTER TEMPERATURE FOR END (' OF LIFE CON 01TIONS Figure 3 34

                                           .aust                                                287
                                                      . c.:. . . , ,.

1 l 3t00 l l 3000 l l I 2600

l

         .                                                                               l 8
        "                                                                                l D%         ,

E e e

%=

u. l800

(. I1400 ' 1000 ' O 20 16 0 60 80 100 Volume Fraction of Total Fuel , At or Above Fuel Temperature, f. DAVIS-BESSE NUCLEAR POWER STATION i FUEL TEMPE,RATURE VERSUS TOTAL FUEL VOLUME FRACTION FOR EQUILIBRIUM CYCLE AT END OF ( LIFE l g FIGURE 3-35 288 l

s Ne - 3 3- -- -- 2 --- 3--- ---3--- --I a 0.780\ 0.830 0.839 0.836 1.043 0.986 0.995 1.008

                               \r                                                          -

1 3 3 2 3 1 1 0.937 \ 0.738 0.794 1.054 1.005 1.311 0.976 m 3 2 2 3 2 1 ON4\ 1.097 1.182 1.025 1.082 0.781 Nn 2 2 2 1 J 1.202\ 1.148 1. 091 1.034 ('~ s Xe 3 1 1 a

1. 043 \ 1.218 0.798 g Number Cycles Burned B

O.911\ Assem0ly P/P DAVIS-BESSE NUCLEAR POWER STATION TYPICAL REACTOR FUEL ASSEMBLY POWER DISTRIBUTION AT ENO 0F LIFE EQUILIBRIUM CYCLE CONDITIONS FOR 1/8 CORE - FIGURE 3 36 1( l M 289 l

          ,O
                                                                                                                                                       .]

i00.00 So.co *

.~

_n -- u. 6 U 7 -A sn r '.

                                           / s.

io.oo - l3 ,

5"

                  .                 4
                                     /

a _m

5 j BO ^

u & 3 '* . E o.5o

                           ,"                            1                :

f

                     .    >                +         +

f a O Gear - 4596 0 'O

                      /a                                         4             . star - uis
                                                                              + arct - 603 3                                                           A CF-60-12 it (cant) 0 05         "

l l l l DAVIS-BESSE NUCLEAR POWER STATION i2co avoo isoo isoo 2000 220o 2 00 2600 2 o0 sooo 3200 asoo sioo PERCENT FISSION GAS RELEASE 0 AS A FUNCTION volumetric Averase Temperature, r 0F THE AVERAGE TEMPERATURE OF THE UO FUEL 2 FIGURE 3-37 N

        %O CD i

l.8 i i :l i i i I

   "'                                       ~
                                       /          N            P/P - 1.70 (Parti al Rod
                                                    \               l 1.6                         '

Irsertion)

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20. 40 60 80 100 120 140 Distance from Bottom of Active Fuel,.in.

DAVIS-BESSE NUCLEAR POWER STATION AXIAL LOCAL-TO AVERAGE BURNUP (' AND INSTANTANEOUS POWER g C0llPARISONS FIGURE 3-38 2, 9 ]

50 uo

1. 5 Axi al Power and 1. 5 Bu rnup Shape 1.7 Axi al Power and I. 5 g Burnup Shape 9 30 - #

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0 1.7 Axial Power and w 930 Day Bo rnup Snape 10 0 0 2' 4 6 8 10 Ini ti al Col d Di anetral C1earance. (i n x 10-3) .

  '                                                       DAVIS-BESSE NUCLEAR POWER STATION FISSION GAS RELEASE FOR 1.5 AND pg]
                                                            .7 MAX / AVG AXIAL POWER SHAPES FIGURE 3-39

(~ 3500 Design Limit 3000

                         - - - -         112%, Overpower 100% Power
   ._  250 0         -

E Closed Pores j i / 5 / o 4 200 0 ,

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500 - k_ Open Pores I I O 2 4 6 8 !0 Ini ti al Col d Diamet ral Clearance.(i n x 10-3) t DAVIS-BESSE NUCLEAR POWER STATION ( GAS PRESSURE INSIDE THE CLAD FOR VARIOUS AXlAL POWER AND BURNUP SHAPES g FIGURE 3-40 2g3

A - - - -

4 I i l 4, _ INST' 1.023 1.0 994 0.9 0.98 -0.98 __ t.043 1.050 1.058 1.060 1.048 1.029 1.007 975 1.053 1. 044 1.046 1. 05 2 1.034 1.015 1.002 0.97: 0 1.058 1.045 1.042 1.045 1.025 1.010 1.000 0.971 ( 1.058 1. 051 1. 045 1.034 1.019 1.017 0.999 0.967 04 7 1.034 1.025 1.020 1.014 1.013 0.992 . 953 I 1.030 1.015 1.011 1. 01 8 1.014 1.002 0.977 .943 hh" 1.110 8 1.1104 1.002 b'3 'h 1.001 0.99 0.978 k' 0.951

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                                                                                                       "              "           8"     '
                      @        HOT UNIT CELL DAVIS-BESSE NUCLEAR POWER STATION
     .                @        HOT CO RNER CELL NOMINAL FUEL R00 POWER PEAKS AND l
                      @        HOT CONTROL R00 CELL                                                                       CELL EXIT ENTHALPY RISE RATIOS                                 ,

g FIGURE 3-41 . d94

i , k (- l i (_.- INST .035 1.0 007 1.00 0.99 0.9 A9 __ l.025 1.035 1. 04 1 1.042 1. 031 1.020 1.014 1 33

1. 03 4 1.027 1.029 1.036 1. 019 1 006 1.009 1. 01D 1.039 1.02 1.028 1.035 1.01 1.002 1.1307 1. 000 1.040 1.034 1.031 1.025 1.007 1.009 1.005 0.991

( R0 , 1.031 1.018 1.011 1.007 1. 00 1 1 .0 05 1.000 0. 9 91 1.021 1.008 1,005 1.011 1.007 1.002 0.f 93 0 964 0 0 0 0. 1 94 9 9 0.983 - 0.988 .984 0.974 0.964 0 58 0.958 1.,011 1.010 011 1,.,004 0.,996 0.g86 0.}78 0.973

                                                               ^ b uclN  ear Peaking Factor                                    l HOT UNIT CELL                       En thal py Ri se Facto r l @
               @          HOT WALL CELL                                                                                        l
               @          HOT CORNER CELL DAVIS-BESSE NUCLEAR POWER STATION
               @          HOT CONTROL R00 CELL                        MAXIMUM FUEL R00 POWER PEAKS AND CELL EXIT ENTHALPY RISE RATIOS

! FIGURE 3-42 i 295 l

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1.6 g l G = 2.49 x 106 l b/ hr-f t2 1.4 '

W-3 DMB Heat Flux (DesignLimit)

1. 2

1.0 b 3= 0.8 L i C Minimum DNBR = 1.78 0.6 - 3o 0.4 7 N Cal culated ~ Surf ace Heat Flux 0.2 0.0 _ 540 580 62D 660 700 Local Enthalpy. Stu/lb ' OAVIS-BESSE NUCLEAR POWER STATION CALCULATED AND DESIGN LIMIT LOCAL l(- HEAT FLUX VERSUS ENTHALPY IN THE HOT UNIT' CELL AT NOMINAL CONDITIONS l Figure 3-43 2,96

l.6 ( G - 2. 2 3 x 106 l b/ n r- f t2 1.4 N l.2 W-3 DNB Heat Flux (Design Limit) x 1.0 "O L - 0.8 ' d C Minimum DNBR = I.50 E x

        ;  0.6

( S r ' 0.4 g Calculated Surface Heat Flux

0. 2 O.0 540 580 620 660 700 Local Enthalpy. Btu / l b DAVIS-BESSE NUCLEAR POWER STATION l

' CALCULATED AND DESIGN LIMIT LOCAL HEAT FLUX VERSUS ENTHALPY IN THE HOT UNIT CELL AT DESIGN CONDITION ( , Figure 3:44 2o7 918it- - e

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DAVIS-BESSE NUCLEAR POWER STATION REACTOR VESSEL AND INTERNALS GENERAL ARRANGEMENT FIGURE 3-45 AMENDMENT NO. 3 f 298 tiudis

FUEL ASSEMBLY

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SURVEILLANCE SPECIMEN HOLDER TUBE DAVIS-BESSE NUCLEAR POWER STATION REACTOR VESSEL AND INTERNALS-CROSS SECTION FIGURE 3-46 AMENDMENT NO. 3'

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BURNABLE PolSON MATERIAL {,'. s. b b b DAVIS-BESSE NUCLEAR POWER STATION

                                  %                        BURNABLE POISON R0D ASSEMBLY FIGURE 3-50 303

i I li l , POSmON INDICATOR ASSEM8LY N r WOTOR TU8E t I I l i I i 7-STATOR ASSEWBLY id. f -REACTOR VESSEL HEAD L.51 /

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Pt e P.St s 85.et. s est f. DAYlS-BESSE NUCLEAR P0WER STATION CONTROL R00 DRIVE SYSTEM AND TRIP BLOCK DIAGRAM FIGURE 3-53 g AMEN 0 MENT NO. 2

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TOP VIEW M NEUTRON ASSORBING MATERIAL 4

                                                                                           +
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a v _ _ _ _ _ t'i G U V UU t' g DAV,lS-BESSE NUCLEAR POWER STATION AX1AL POWER SHAPING ROD ASSEMBLY j}Q FIGURE 3-55 l 1

          -     _ - _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _                       . _ _ . ___________________________-_.__-________-____O

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