ML19329E136

Chapter 3 of AR Nuclear 1 PSAR, Reactor. Includes Revisions 1-18

ML19329E136
Person / Time
Site:	Arkansas Nuclear
Issue date:	11/24/1967
From:	ARKANSAS POWER & LIGHT CO.
To:
References
NUDOCS 8005300715
Download: ML19329E136 (200)

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TABLE OF COUTE'!TS Section Page 3 REACTOR 3-1

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3.1 DESIGN BASES 3-1 3.1.1 PERFORMANCE OBJECTIVES 3-1 3.1.2 LIMITS 3-1 3.1.2.1 NuclearLi$its 3-1 3.1.2.2 Reactivity Control Limits 3-2 3.1.2.3 Thormal and Hydraulic Linits 3-2 3.1.2.h Mechanical Linits 3-3 3.2 REACTOR DESIGN 3-5 3.2.1 GENERAL SU: GARY 3-5 3.2.2 NUCLEAR DESIGN AND EVALUATION 3-7 3.2.2.1 Nuclear Characteristics of the Desien 3-7 3.2.2.2 Nuclear Evaluation 3-20 3.2.3 THERMAL AND HYDRAULIC DESIGN AND E7ALUATION 3-31 3.2.3.1 Thermal and Hydraulic Characteristics 3-31 3.2.3.2 Thermal and Hydraulic Evaluatien 3-ho 3.2.h MECEMiICAL DESIGN LAYOUT 3-68 3.2.h.1 Internal Layout 3-68 3.2.h.2 Fuel Assemblies 3-72h 3.2.h.3 Centrol Rod, Drive System 3-85 3.3 TESTS AND INSPECTIONS 3-95 3.3.1 NUCLEAR TESTS MID INSFECTIONS 3-95

3. 3.1.1 ; Critical Experiments -

3-95

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j 3.3.1.2 Zero Pover, Actreach to Pover, and . Power Testinc: 3-95 ,

3.3.2 THERMAL AND HYDRAULIC TESTS AND INSFECTION -6 M 3-1

CONTENTS (Cont'd)

Section Page 3.3.2.1 Reactor Vessel Flow Distribution and Pressure Drop Test 3-96 3.3.2.2 Fuel Assembly Heat Transfer and Fluid Flow Tests 3-96 3.3.2.3 Preoperational Testing and Postoperational Testinst 3-98 3.3.3 FUEL ASSEMBLY, CONTROL ROD ASSEMBLY, AND CONTROL ROD DRIVE MECHANICAL TESTS AND INSPECTION 3-98 .

3.3.3.1 Prototype Testing 3-98a l3 3.3.3.2 Model Testing 3-99 3.3.3.3 Component and/or Material Testing 3-99 3.3.3.h Control Rod Drive Tests and Inspection 3-100 3.3.h INTERNALS TESTS AND INSPECTIONS 3-103

3.4 REFERENCES

3-105 35 NUCLEAR STFAM SUPPLY SYSTEM REVISIONS 3-110 13 0099 y

3-11 10-31-69 Supplement No. 13

LIST OF TABLE Table No. Title Page 3-1 Core Design, Thermal, and Hydraulic Data 3-6 3-2 Nuclear Design Data 3-8 3-3 Excess Reactivity Conditions 3-9 3-k First Cycle Reactivity Control Distribution 3-9 3-5 Shutdown Reactivity Analysis 3-13 3-6 Soluble Boron Levels and Worth 3-14 3-7 Exterior Neutron Levels and Spectra 3-17 3-8 calculated and Experimental Rod and Rod Assembly Comparison 3-22 3-9 Reference Core Parameters 3-28 3-10 First Mode Threshold Dimensions and Flatness 3-28 3-11 Threshold Ratio and Power Flatness 3-29 3-12 coerricients or variation 3-35 3-13 DNB Results - Maximum Design Condition 3-37 3-14 DNB Results - Most Probable Condition 3-37 3-15 Heat Transfer Test Data 3-47 3-16 comparison or Heat Transfer Test Data 3-50 3-17 Hot Channel Coolant Conditions 3-51 3-18 DNB Ratios in the Fuel Assembly Channels 3-65 3-19 clad circumferential Stresses 3-77 3-20 LRD Fuel Swelling Irradiation Program 3-83 3-21 control Rod Drive Design Data 3-87 3-22 control Rod Assembly Design Data 3-93 Ol.M 3-111

LIST OF FIGURES (At rear of Section)

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 Bods 55 Per Cent Inserted 3-4 Moderator Temperature Coefficient versus Baron Concentration 3-5 Moderator Temperature Coefficient versus Moderator Temperature and Various Boron Levels 3-6 Per Cent Neutron Power versus Time Following Trip 3-7 Effect of Fuel Te=perature (Doppler) on Xenon Oscillations -

Beginning of Life 3-8 Effect of Fuel Temperature (Doppler) on Xenon Oscillations -

Near End of Life 3-9 Control of Axial Oscillation with Partial Rods 3-10 Population Included in the Statistical Statement versus DNB Ratio 3-11 Power Shape Reflecting Increased Axial Power Peak for 144-Inch Core 3-12 Distribution of Fuel Rod Peaking 3-13 Possible Fuel Rod DNB's for Maximum Design Conditions -

36,816-Rod Core 3-14 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 (BAW-168) versus Reactor Power 3-17 Maximum Hot Channel Exit Quality versus Reactor Overpower 3-18 Thermal Conductivity of 00 2 3-19 Fuel Center Temperature at the Hot Spot versus Linear Power

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3-20 Number of Data Points versus $g4 C 3-21 Hot Channel Factor versus Per Cent Population Prot +g 4%f

FIGURES (Cont'd) ,-

Figure N% Title 3-22 Burnout Factor versus Population for Various Confidence Levels

?-23 Rods in Jeopardy versus Power 3-24 Ratio of Experimental to Calculated Burnout Heat Flux 3-25 Ratio of Experimental to Calculated Burnout Heat Flux 3-26 Ratio of Experimental to Calculated Burnout Heat Flux 3-27 Ratio of Experimental to Calculated Burnout Heat Flux 3-28 ' Ratio of Experimental to Calculated Burnout Heat Flux 3-29 Ratio of Experimental to Calculated Burnout Heat Flux l

l 3-30 Ratio of Experimental to Calculated Burnout Heat Flux 3-31. Ratio of Experimental to Calculated Burnout Heat Flux

'3-32 Ratio of Experimental to Calculated Burnout Heat Flux

! 3-33 Ratio of Experimental to Calculated Burnout Heat Flux 3-34 Ratio of Experimental to Calculated Burnout Heat Flux l

l 3-35 Ratio of Experimental to Calculated Burnout Heat Flux 1

3-36 Ratio of Experimental to Calculated Burnout Heat Flux 3-37 Ratio of Experimental to Calculated Burnout Heat Flux 3-38 Ratio of Experimental to Calculated Burnout Heat Flux 3-39 Ratio of Experimental to Calculated Burnout Heat Flux 3-40' Maximum Hot Channel Exit Quality versus Reactor Power 3-41 Hottest Design and Nominal Channel Exit Quality versus Reactor Power (without Engineering Hot Channel Factors) l

! 3-42 Flow Regime Map for Unit Cell Channel at 2,120 psig '

3-43 Flow Regime Map for Unit Cell Channel 3-44 Flow Regime Map for Corner Channel 3-45 Flow Regime Map for Wall Channel 0103 3-v m- . . .

FIGURES (Cont'd)

Figure No. Title 3-46 Hot Channel DUB Ratio Comparison 3-47 Reactor Coolant Flow versus Power 3-48 Thermal Conductivity of 95 Per Cent Dense Sintered UO Pellets 2

3-49 Fuel Center Temperature for Beginning-of-Life Conditions 3-50 Fuel Center Temperature for End-of-Life Conditions 3-5] Per Cent Fission Gas Released as a Function of the Average Temperature of the UO Fuel 2

3-52 Axial Incal to Average Burnup and Instantaneous Power Compar-isons 3-53 FissionGasReleasefor150and170 Max /AvgAxialPower Shapes 3 -54 Gas Pressure inside the Fuel Clad for Various Axial Burnup and Power Shapes 3-55 Nominal Fuel Rod Power Peaks and Cell Exit Enthalpy Rise Ratios 3-56 Maximus Fuel Rod Power Peaks and Cell Exit Enthalpy Rise Ratios 3-57 Calculated and Design Limit Local Heat Flux versus Enthalpy in the Hot Corner Cell at the Nominal Condition 3-58 Calculated and Design Limit Incal Heat Flux versus Enthalpy in the Hot Corner Cell at the Postulated Worst Condition 3-59 Reactor Vessel and Internals - General Arrangement 3-60 Reactor vessel and Internals - Cross Section 3-61 Core Flooding Arrangement 3-61a Internals vent Valve 3-62 Fuel Assembly 3-63 Orifice Rod Assembly 3-64 Control Rod Drive - General Arrangement 3-65 Control Rod Drive - Vertical Section

'3 Control Rod Drive Control System Block Diagram 3-67 Limit Signal and Position Indication Syst,em 3-68 Reactor Trip Circuit g

3-69 Control Rod Assembly M 3-vi REVISED, : 2-8-68 o

3 REACTOR 37 DESIGN BASES The reactor is designed to meet the performance cbjectives specified in 31.1 without exceeding the limits of design and operation specified in 31.2.

3 1.1 PERFORMANCE OBJECTIVES The reactor is designed to operate initially at 2,452 E t(*) with sufficient design margins to accommodate transient operation and instrument error without damage to the core and without exceeding the pressure at the safety valve set-tings in the reactor coolant system. The ultimate operating pcwer level of the reactor core is expected to be 2,568 Wt, but additional operating infomation will be required to justify operation at this higher power level. Thus, 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 anticipated core life. The effects of gas release, fuel dimensional changes, and corrosion-or irradiation-induced changes in the mechanical properties of cladding are con-sidered in the design of fuel assemblies.

Reactivity is controlled by control rod assemblies (CRA's) and soluble boron in 3 the coolant. Sufficient CRA vorth is available to shut the reactor down (keff n 0 99) in the hot condition at any time during the life cycle with the most reac-tive CRA stuck in the fully withdrawn position. Redundant 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's, and the rate at which reactivity can be added, is limited to insure that credible reactivity accidents cannot cause a transient capable of damaging the reactor coolant system or causing significant fuel fail-ure.

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 an average burnup of 28,200 WD/MEU and for a maximum burnup of 55,000 WD/MI'U.

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.

c. Control systems will be available to handle core xenon instabili-ties should they occur during operation, without Jeopardizing the safety conditions of the system.

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d. The core win have sufficient excess reactivity to produce the design

,, power level and lifetime without exceeding the control capacity or shutdown margin.

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e. Controlled reactivity insertion rates have been limited to 5.8 x lg-5 L see for a single regulating CRA group withdrawal, and 7 x 10-4 see 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.

S e 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 win provide adequate control of the core reactivity and power distribution. The following control limits will be met:

a. Sufficient control will be available to produce a shutdown margin of atleast1%ZA/k.

b. The shutdown margin wL11 be maintained with the CRA of highest worth stuck out of the core.

c. CRA withdrawal limits the reactivity insertion to 5.8 x 10-5 Ak/k/see on a single regulating group. ron dilution is also limited to a reactivity insertion of 7 x 10- .ik/k/sec. ,

3 1.2.3 Thermal and Hydraulic Limits m e reactor core is designed to meet the following limiting thermal and hydrau-lic conditions:

a. No central melting in the fuel at the design overpower (114 percent). 3

b. a 99 percent confidence that at least 99 5 percent of the fuel rods in the core are in no jeopardy of experiencing a departure from nucleate boiling (DNB) during continuous operation at the design overpower.

c. Essentially 100 percent confidence that at least 99 96 percent of the fuel rods in the core ape in no jeopardy of experiencing a DNB during continuous operation'at rated power.

d. The generation of net steam in the hottest core channels is permis-sible, but steam voids will be low enough to prevent flow instabilities.

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 2,452 Et, 3-2 O.!.05 1

5-3-68 Supplement No. 3 l 1

3 1.2.4 Mechanical Limit's 3.1.2.4.1~ 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 3 coolant pumps running; and shutdown conditions. No damage to the reactor in-ternals will occur as a result of loss of pumping power.

Reactor internals will be fabricated from SA-240 (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 welds will be assured by compliance with ASME Code Sections III and IX, radiographic inspection acceptance standards, and welding qualifications.

Se 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-7024,

" Nuclear Reactors and Earthquakes".

Iateral deflection and torsional rotation 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's).

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

S e structural internals will be designed to maintain their functional integrity in the event of a major' loss-of-coolant accident as described in 3 2.4.1. The dynamic loading resulting from the pressure oscillations because of a loss-of-coolant accident will not prevent CRA insertion.

Internals vent valves are provided to relieve pressure generated by steaming in the core following a postulated reactor coolant inlet pipe rupture so that the core will remain sufficiently covered with coolant.

3.1.2.4.2 Fuel Assemblies The fuel assemblies are designed to operate satisfactorily to design burnup and to retain adequate integrity at the end of life to permit safe removal from the Core.-

,The assemblies are designed to operate safely 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. Re cold-worked Zircaloy-4 cladding is designed to be free-standing.~ Fuel rods are held in place by mechanical. spacer grids that are de-

~s igned to maintain dimensional control of the fuel rod spacing throughout the design life without impairing cladding integrity. Contact loads are limited to prevent 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 F

structure are both indexed to.the grid plate above the fuel assemblies, thus insuring. continuous alignment of the guide channels for the CRA's. The control rod travel is designed so that the rods are always engaged in the fuel assembly 0106 i; 3-3 M 5-3-68 L -

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s guide tubes, thus insuring that CRA's can always be inserted. The assembly

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structure is also designed to withstand handling loads, shipping loads, and earthquake loads. .

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

a. Stresses that are not relieved by small deformations of the material vill be prevented from leading to failure by not permitting these stresses 'to exceed the yield strength of the material nor to exceed levels that would use in excess of 75 per cent of the stress rupture life of the material. An example of this type of stress is the cir-cumferential membrane stress in the clad due to internal or external pressure.

b. Stresses that are relieved by small deformations of the material, and the single occurrence of which will not make a significant contribu-tion to the possibility of a failure, will be permitted ta exceed the yield strength of the material. Where such stresses exceed the mate-rial yield strength, strain limits will be set, based on low-cycle fatigue techniques, using no more than 90 per cent of the material fatigue life. Evaluations of cyclic loadings vill be based on con-servative estimates of the number of cycles to be experienced. An example of this type of stress is the thermal stress resulting from the thermal gradient across the clad thickness,

c. Combinations of these two types of stresses, in addition to the in-dividual treatment outlined above, will be evaluated on the low-cycle fatigue basis of Item b. Also, clad plastic strain due to diameter increases resulting from thermal ratcheting and/or creep, including the effects of internal gas pressure and fuel swelling, will be lim-ited 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, en 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.

(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.h.3 Control Rod Assembly (CRA)

The control rod clad is designed to the same criteria as the fuel clad, as ap-plicable. Adequat e clearance vill be provided between the control rods and the guide tubes, which position them within the fuel assembly, so that control rod overheating vill be avoided and unacceptable mechanical interference between the control rod and the guide tube vill not occur under any operating condition, in-cluding earthquake.

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Overstressing of the CRA components during a trip will be prevented by minimiz-ing the shock loads by snubbing and by providing adequate strength.

.3.1.2.4.h . Control Rod Drive Each control rod drive is provided with a pressure breakdown seal to allow a con-trolled leakage of reactor coolant water. All precsure-containing components are designed to meet the requirements of the ASME Code,Section III, Nuclear

-Vessels, for Class A vessels.

The control rod drives provide control rod assembly (CRA) insertion and v'ith-drawal rates consistent with the required reactivity changes for reactor opera-tional load changes. This rate is based on the worths of the various rod groups, which have been established to limit power-peaking flux patterns to design val-ues. The maximum reactivity addition rate is specified to limit the magnitude of a possible nuclear excursion resulting from a control system or operator mal-function. The normal insertion and withdrawal velocity has been established as 25 in./ min.

The control rod drives provide a " trip" of the CRA's which results in a rapid shutdown of the reactor for conditions that cannot be handled by the reactor control system. The trip is based on the results of various reactor emergency analyses, including instrument and control delay times and the amount of reac-tivity that must be inserted before deceleration of the CRA occurs. The maxi-mum travel time for a 2/3 insertion on a trip conmand of a CRA has been estab-lished as 1.h sec.

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

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

Structural integrity and adherence to allovable stress limits of the control rod drive and related parts during a trip will be achieved by establishing a limit on impact loads through snubbing.

3.2 REACTOR DESIGN 3.2.1 GENERAL SUM ERY The important core design, thermal, and hydraulic characteristics are tabulated in' Table 3-1.

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Table 3-1 3 Core Design, Thermal, and Hydraulic Data J Reactor Type ,

Pressurized Water Rated Heat Output, MWt 2,h52 Vessel Coolant Inlet Temperature, F 555 Vessel Coolant Outlet Temperature, F 602.8 Core Outlet Temperature, F 60h.3 Operating Pressure, psig 2,185 Core and Fuel Assemblies Total Number of Fuel Assemblies in Core 177 Humber of Fuel Rods per Fuel Assembly 208 Humber of Control Rods per Control Rod Assembly 16 Number of Incore Instrumentation Positions per Fuel Assembly 1 Fuel Rod Outside Diameter, in. 0.h20 Clad Thickness, in. 0.026 Fuel Rod Pitch, in. 0 558 Fuel Assembly Pitch Spacing, in. 8.587 Unit Cell Metal / Water Ratio 0.80 Clad Material Zirealoy-b (cold-worked)

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Material UO2 Form Dished-End, Cylindrical Pellets Diameter, in. 0.362 Active Length, in. 1hh Density, % of theoretical 95 Heat Transfer and Fluid Flow at Rated Power Total Heat Transfer Surface in. Core, ft2 h8,578 Avert.ge Heat Flux, Btu /hr-ft2 167,620 Maximum Heat Flux, Btu /hr-ft2 Sh3,000 Average Power Density in Core, kv/l 79.60 Average Thermal Output, kw/ft of fuel rod . 5.h Maximum Thermal Output, kv/ft of fuel rod 17.h9 Maximum Clad Surface Temperature, F 65h Average Core Fuel Temperature, F 1,385 Maximum Fuel Central Temperature at Hot Spot, F h,160 Total Reactor Coolant Flow, lb/hr 13^..32 x 106 Core Flow Area (effective for heat transfer), ft2 h7.75

. Core Coolant Average Velocity, fps 15.7 Coolant Outlet Temperature at Hot Channel, F 6hh.h Power Distribution Maximum / Average Power Ratio, radial x local (Fdh nuclear) 1.85 0109

Tablo 3-1 (Cont'd)

Mavimm/AveragePowerRatio, axial (F2 nuclear) 1 70 Overall Power Ratio (Fa nuclear) 3.U Power Generated in Fuel and Cladding, % 97 3 Hot Channel Factors Power Peaking Factor (Fo) 1.008 Flow Area Reduction Factor:(F ) 0 992 Local Heat Flux Factor (Fq") A 1.013 Hot ~ Spot Mavi==/ Average Heat Flux Ratio (Fqnuc. and mech.) 3 24 DNB Data Design Overpower Ratio 1.14 DNB Ratio at Design Overpower (BAW-168) 1 38 DNB Ratio at Rated Power (BAW-168) 1.60 3.2.2 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 level over the specified fuel cycle.

b. Sufficient reactivity control is provided to permit safe reactor operation and shutdevn at all times during core 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 reactivi-ties associated with various core conditions are tabulated in Table 3-3 The core will operate for 410 full power days for the first cycle and will have a 310 full-power day equilibrium cycle. Design limits will be held with respect to reactivity control and power distribution. Incore instrumentation will be used to indicate power peaking levels. Single fuel assembly reactivity in- 3 formation is also included in Table 3-3 t

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Nuclear Design Data ]

Fuel Assembly Volume h actions Fuel'- 0.285 Moderator 0 590 Zircaloy 0.099 Stainless Steel 0.011 Void 0.015 1.000 Total UO 2, metric tons 91.61 Core Dimensions, In.

Equivalent Diameter 129 9 Active Height. 144.0 Unit Cell Hg0 to U Atomic Ratio (fuel assembly)

Cold 2 97 Hot 2.13 Full Power Lifetime, days )

First Cycle 410 Each succeeding Cycle 310 Fuel Irradiation, ETD/ML'U First Cycle Average 12,460 Succeeding Cycle Average 9,410 FeedEnrichments,w/oU-235 First Cycle 2.29/2.64/2 90 (by zone)

Equilibrium Cycle 2 94(a)

Control Data Control Rod Material Ag.-In-Cd 3 Number of Control Rod Assemblies 69 Total Rod Worth (ak/k), y, 10.0 Control Rod Cladding Material Type 304 SS (a)Averagefeedenrichment.

3-8 Olli 5-3-68.

Supplement No. 3

Table 3-3 Excess Reactivity Conditions Effective Multiplication - BOL(a)

Cold, Zero Power 1.302 Hot, Zero Pcuer 1.2h7 Hot, Rated Power 1.229 Hot, Equilibrium Xe, Rated Power 1.192 Hot, Equilibrium Xe and Sm, Rater Power (b) 1,15g Single Fuel Assembly (c)

Hot 0.77 Cold (d) 0.87

" BOL - Beginning-of-Life (b) Includes burnup until equilibrium samarium is reached.

Based on highest probable enrichment of 3.5 weight per cent.

(d)A center-to-center assembly pitch of 21 in.

is required for this k rr e in cold, nonborated water with no xenon or samarium.

The minimum critical mass, with and without xenon and samarium poisoning, may be specified in a variety of forms, i.e., single assembly, multiple assemblies in various geometric arrays, damaged or crushed assemblies, etc. The unit fuel assembly has been investigated for comparative purposes. A single cold, clean assembly containing a maximum probable enrichment of 3 5 vt % is suberitical.

Two assemblies side-by-side are supercritical except when both equilibrium xe-non and samarium are present. .Three assemblies side-by-side are supercritical with both equilibrium xenon and camarium present.

3.2.2.1.2 Reactivity Control Distribution Control of excess reactivity is shown in Table 3-4.

Table 3-4 First Cycle Reactivity Control Distribution

% Ak/k

1. Controlled'by Soluble Boron

a. , Moderator. Temperature Deficit (70 to 520 F) 3.h c, - '

b. . Equilibrium Xenon and-Samarium 25

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Table 3-h (Cont'd) .-

% Ak/k

c. Fuel Burnup and Fission Product Buildup 16.0

2. Controlled by Inserted Control Rod Assemblies Transient Xenon (normally inserted) 1.h

3. Controlled by Movable Control Rod Assemblies

a. Doppler Deficit (0 to 100% rated power) 1.2

b. Equilibrium Xenon 1.0

c. Moderator Temperature Deficit (0 to 15% power at end of life) 0.6

d. Dilution Control 0.2

e. Shutdown Margin 1.0 Total Movable Control Worth Required h.0

h. Available Control Rod Assembly Worths

a. Total CRA Worth 10.0

b. Stuck Rod Worth (rod of highest reactivity value) (-) 3.0

c. Minimum Available CRA Worth 7.0

d. Minimum Movable CRA Worth Available 5.6 Explanation of Items Above

1. Soluble Boron Boron in solution is used to control the following relatively slow-moving reactivity changes:

a. The. moderator deficit in going from ambient to operating te=peratures.

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

b. Equilibrium samarium and a part (approximately 1.h% Ak/k) of the equili-brium xenon.

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c. The- excess reactivity required for fuel burnup and _ fission product build-up throughout' cycle life.

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

Control rod assemblies (CRA's) will be used to control the reactivity changes associated with the following:

2. Inserted Control Sufficient rod worth remains inserted in the core during normal operation to overcome the peak xenon transient following a power reduction. This override capability facilitates the return to normal operating conditions without extended delays. The presence of these rods in the core during op-eration does not produce power peaks above the design value, and the shut-down margin of the core is not adversely affected. Axial power peak varia-tion, resulting from partial or full insertion of xenon override rods, is described fully in Figures 3-2 and 3-3. The loss of movable reactivity con-trol due to the insertion of this group produces no shutdown difficulties and is reflected in Table 3-5

3. Movable control

a. Power level changes (Doppler) and regulation.

b. The portion of the equilibrium xenon not controlled by soluble boron, approximately 15 Ak/k, is held by movable CRA's.

c. Patween zero and 15 per cent of rated power, reactivity compensation by CMA's may be required as a result of the linear increase of reactor cool-ant temperature from 520 F to the normal operating value,

d. Additional reactivity is held by a group of partially inserted CRA's (25 per cent insertion maximum) to allow periodic rather than continuous soluble-boron dilution. The CRA's_are inserted to the 25 per cent limit as the boron is diluted. Automatic withdrawal of these CRA's during op-eration is allowed to the 5 per cent insertion limit where the dilution procedure is again initiated and this group of CRA's is reinserted.

e. A shutdown margin of 1% Ak/k to the hot critical condition is also re-quired as part of the reactivity controlled by CRA's.

h .' Rod Worth- -

Atotalof4.ONAk/k(a)isrequiredinmovablecontrol. Analysis of the 69 CRA's under the reference fuel arrangement predicts a total CRA worth of at least 10.0% Ak/k. Th9 qtuck-out CRA worth was also evaluated at a value no larger than 3.0% Ak/ktbl. This evaluation included selection of the highest (ah -

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-(b)Do,es not_ include transient control.

See Table 3-h. ()}.1 k First cycle. See. Table 3-4. -

3-11

4 1 1 worth CRA unde the first CRA-out condition. The minimum available CRA worth

-of 5.6% ok/k a is sufficient to meet movable control requirements. -

3.2.2.1.3 Reactivity Shutdown Analysis The ability to shut down the corci under both hot and cold conditions is illus-trated 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 shut-down capability.

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Table 3-5 Shutdown Reactivity Analysis First Cycle Eauilibrium Reactivity Effects, % Ak/k BOL EOL BOL EOL

1. Maximum Shutdown CRA Requirement Doppler (100 to 0% Power) 1.2 15 1.2 15 Equilibrium Xenon 1.0 1.0 1.0 1.0 Moderator Deficit (15 to 0% Power) 0.0 0.8 0.0 0.8 Total 2.2 3.3 2.2 3.3

2. Maximum Available CRA Worth ("} -10.0 -10.0 -10.0 -10.0 Transient Xe Insertion Worth 1.h 1.4 1.h 0.0 Possible Dilution Insertion 0.2 0.2 0.2 0.2

3. Minimum Available CRA Worth All CRA's In -8.h -8.h -8.4 -9.8 One CRA Stuck-Out(b) -5.h -5.k -5.h -6.8

h. Minimum Hot Shutdown Margin All CRA's In -6.2 -5.1 -6.2 -6.5 One CRA Stuck-Out -3.2 -2.1 -3.2 - 3. 5 5 Hot-to-Cold Reactivity Changes (c)

All CRA's In 0.0 +6.h +3.0 +8.0 One CRA Stuck-Out -0 9 +5 5 +2.1 +7.1

6. Cold Reactivity Condition (d)

All CRA's In -6.2 +1.3 -3.2 +1 5 One CRA Stuck-Out -h.1 +3.h -1.1 +3.6 7 PPM Boron Addition Required for keff

= 0.99 (cold)

All CRA's In 0 170 0 190 One CRA Stuck-Out 0 330 0 350 (a) Total worth of 69 CRA's.

(b)CRA of highest reactivity value.

(c) Includes changes in CRA vorth, moderator deficit, and y, equilibrium Xe held by soluble boron. .

(d)No. boron addition.

3-13

m; . ~ -

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 1% Ak/k at_the end-of-life.

Under conditions where a cooldown to reactor building ambient temperature is required, concentrated soluble boron vill be added to the reactor coolant to produce a' shutdown margin of at least 1% Ak/k. The reactivity changes that take place between the hot zero power to cold conditions are tabulated, and the cor-responding increases in soluble boron are listed. Beginning-of-life boron lev-els for several core conditions are listed in Table 3-6 along with boron worth values. Additional soluble-boron could be added for situations involving more than a single stuck CRA. The conditions shown with no CRA's illustrate the highest requirements.

Table 3-6 Soluble Boron Levels and Worth ,

BOL Boron Levels,

.., Core Conditions ppm

1. Cold, keff = 0.99 No CRA's In 1,820 All CRA's In 1,290 One-Stuck CRA 1,h50 /

2. Hot, Zero Power, Keff = 0.99 No.CRA's In 2,080 All CRA's In 1,080 One Stuck CRA 1,380

3. Hot, Rated Power No CRA's In 1,860

h. Hot, Equilibrium Xe and Sm, Eated Power No CRA's In 1,360 Core Condition Boron Worth, (% Ak/k)/ ppm Hot 1/100 Cold 1/75 0 ?..'19

3 2.2.1.4 Reactivity coefficients

-Reactivity coefficients form the basis for analog studies involving nor=al and abnormal reactor operating conditions. These coefficients have been investi-gated as part of the analysis of this core and are described below as to fune-tion 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-duction-in neutron production. The range for the Doppler coefficient under operating conditions is expected to be -1.1 x 10-> to -1 7 x 10-2 (ak/k)/F.

b. Moderator Void Coefficient The moderator void coefficient relates the change in neutron multi-plication to the presence of voids in the =oderator. Cores controlled by appreciable amounts of soluble boron may exhibit a small positive coefficient for very small void levels (several per cent void), while highervoidlevelsproduceincreasinglynegativecoeffgeients. The expected range for the void coefficient is + 1. J x 10- to -3 0 x 10-3 (ak/k)5 void.

c. Moderator Pressure Coefficient The moderator' pressure coefficient relates the change in moderator density, resulting from a reactor coolant pressure change, to the corresponding effect on neutron production. This coefficient is opposite 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 -1 x 10-6 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-tors using soluble boron as a reactivity control have fever negative moderator temperature coefficients than do cores controlled solely by movable or fixed CRA's. The major te=perature 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 reac-tivity held by soluble baron.

The moderator te=perature coefficient has been parameterized for the L reference core in terms of boron concentration and reactor coolant

'j . . , , . temperature. The results of the study are shown in Figures 3-4 and ,

0118 W

\

1 3-15

3-5 Figura 3-4 shows th2 coefficient variation for ambisnt and operating temperatures as a function of soluble boroj concentration. The operating value ranges from a proximately +1.0 x 10 at the beginning of the first cycle to -3 0 x 10 (ak/k)/Fattheendoftheequilibriumcycle. Figure ]

3-5 shows the moderstor temperature coefficient as a function of tempera-ture for various poison concentrations for the first cycle. The coeffi-cients of the equilibrium cycle vill be more negative than those of the first cycle since the boron concentration levels are considerably lower.

The positive temperature coefficient occurs durin6 the initial portion of 3 the first cycle only and vill not constitute an operational problem. The Doppler deficit represents a much larger reactivity effect in the negative direction and, together with the CRA system response, vill provide adequate control. Should detailed analysis result in a requirement that the moder-ator temperature coefficient be made less positive, fixed shims vill be used in the unrodded fuel elements to reduce the boron level and consequent-ly the moderator temperature coefficient.

e. pH Coefficient Currently, there is no definite correlation to predict pH reactivity effects between various operating reactors, pH effects versus reactor operating time at power, and changes in effects with various clad, temperature, and water chemistry. Yankee (Rowe, Mass.), Saxton, and Con Edison Indian Point Station No. 1 have experienced reactivity changes at the time of pH changes, but there is no clear-cut evidence that pH is the direct influencing variable without considering other items such as clad materials, fuel assembly crud deposition, system average temperature, and prior system water chemistry.

5 Saxton experiments have indicated a pH reactivity effect of 0.16 per cent reactivity per pH unit change with and without local boiling in the core.

Operating reactor data and the results of applying Saxton observations to the reference reactor are as follows:

(1) The proposed system pH will vary from a cold measured value of approx-imately 5 5 to a hot calculated value of 7.8 with 1,400 ppm boron and 3 pin KOH in solution at the beginning of life. Lifetime bleed dilu-tion to 20 ppm boron vill reduce pH by approximately 0.8 pH units to a hot calculated pH value of 7 0.

(2) _Considering the maximum system makeup rate of 70 gps, the corresponding j changes in pH are 0.071 pH units per hour for boron dilution and 0.231 '

pH units per hour for KOH dilution. Applying pH vorth values of 0.16%  :

Ak/

10}perpHunit,asobservedatSaxton )

(%Ak/k)/seeand1.03x10-5(%Ak/kinsertion rates are These sec, respectively. 3 16 xin- 1 sertion rates correspond to 1.03% power urand3.4% power / hour, respectively, which are easily compensated by the operator or the automatic control system.

1 3 2.2.1 5 Reactivity Insertion Rates '

Figure 7-7 displays the integrated rod worth of four overlapping rod banks as a

function of distance withdrawn. The indicated groups are those used in the core l during power operation. Using approximately 1.2%dk/k CRA groups y.nd a 25 in./ min l

l 3-16 g3 ,

4 I

5-3-68 4

Supplement No. 3

drive speed in conjunction with the reactivity response given in Figure T-T yields a maximum reactivity insertion rate of 5.8 x 10-5 ( Ak/k)/sec. The maxi-mum reactivity insertion rate for soluble boron removal is T x 10-6 (ak/k)sec.

3.2.2.1.6 Power Decay Curves Figure 3-6 displays the beginning-of-life power decay curves for the two least effective CRA worths as outlined in Table 3-5, Item No. 3. The power decay is initiated by the trip release of the CRA's with a 300 msee delay from initiation to start of CRA motion. The time required for 2/3 rod insertion is 1.4 sec.

3.2.2.1 7 Neutron Flux Distribution and Spectrum The neutron flux levels at the core edge and the pressure vessel wall are given h in Table 3-7 At both locations the valves shown include an axial peaking factor of 1.3, a scaling factor of 2, and a safety margin of 1 9 Table 3-7 Exterior Neutron Levels and Spectra Neutron Flux Levels, n/cm /see Interior Wall of Flux Core Edge Pressure Vessel Group (x 1013) (x 1010) 1 0.821 Mev to 10 Mev 6.0 3.4 2 1.230 Key to 0.821 Mev 9.0 T.5 3 0.hlh ev to 1.230 Kev 6.2 57 3 h Less than 0.hlh ev T.1 2.1 The calculations were performed using The Babcock & Wilcox Company's LIFE code (BAW-293, Section 3.6.3) to generate input data for the transport code, TOPIC.

A h-group edit is obtained from the LIFE output which includes diffusion coef-ficients, absorption, removal, and fission cross sections, and the zeroth and first moments of the scattering cross section. TOPIC is an S code designed to n

solve the 1-dimensional transport equation in cylindrical coordinates for ve, to six groups of neutrons. For the radial and aximuthal variables, a linear approx-imation to the transport equation is used; for the polar angle, Gauss quadrature is used. Scattering functions are represented by a Legendre Series. The azi-

~methal angle can be partitioned into 4 to 10 intervals on the half-space between 0 and w. The number of mesh points in the radial direction is restricted by the number of these intervals. For the core exterior flux calculations, four inter-vals on the azimuthal were used. This allows the maximum number of mesh points (2h0) in the "r" direction to describe the shield complex. An option is avail-able to use either equal intervals on the aximuthal angle or equal intervals on 0120

, W .

6-5-68 3-17 Supplement No. 4

the cosine of the angle. Equal intervals on the cosine were chosen since this s provides more detail in the forward direction of the flux (toward the vessel). .o )

Five Gaus,s quadrature points were used on the cosine of the polar angle in the half-space between 0 and tr.

Results from the above method of calculation have been compared with thermal flux measurements through an array of iron and water slabs in the LIDO pool reactor.(2)

Although this is not a direct comparison with fast neutron measurements, it does provide a degree of confidence in the method since the magnitude of the thermal flux in shield regions is governed by fast neutron penetration.

Results of the comparison showed that fluxes predicted by the LIFE-TOPIC calcula-tion were lover, in general, by a'uout a factor of 2. Results of the fast flux calculations are, consequently, increased by a factor of 2 to predict the nyt in the reactor vessel.

The following conservatisms were also incorporated in the calculations:

a. Neutron fluxes outside the core are based on a maximum power density of kl vatts/cc at the outer edge of the core rather than an estimated average of 28 vatts/cc over life,.resulting in a safety margin of about 45 per cent.

b. A maximum axial power peaking factor of 1.7 was used. This is about 30 per cent greater than the 1.3 expected over life.

Uncertainties in the calculations include the following: s

/

1. The use of only four neutron groups to describe the neutron energy spectrum.

2. Use of the LIFE code to generate the k-group cross sections. In the LIFE program, the h-group data in all regions are computed from a fis-sion spectrum rather than a leakage spectrum.

3 Having only four intervals, i.e., n = h in the Sn calculation, to de-scribe the angular segmentation of the flux.

It is expected that the combination of 1 and 2 above vill conservatively predict a high fast neutron flux at the vessel vall because it underestimates the effee-tiveness of the thermal shield in reducing the fast flux. In penetration through water, the average energy of the neutrons in the group above 1 Mev increases above that of a fission spectrum, i.e., the spectrum in this group hardens. For neutrons above 1 Mev, the nonelastic cross section of iron increases rapidly with energy. Therefore, the assumption of a fission spectrum to compute cross sec-tions in the thermal shield, and the use of a fev-group model to cover the neu-tron energy spectrum, would underestimate the neutron energy loss in the thermal shield mal and the shield. .Thesubsequent results fromattenuation by the(vqter 3h-group P3MG1 between the 31 calc.uations vessel show thatand ther-reduction of the flux above 1 Mev by the thermal shield is about a factor of h greater than that computed from the b-group calculations.

l The'effect of'3 above is expected to underestimate the flux at the vessel vall.

In calculations at ORNL using the Sn technique, a comparison between an S4 and 3-18

an S12 ' calculation was made in penetration through hydrogen. The results for a variety of energies over a penetration range of lh0 cm shoved the Sh calculation to be lower than the S12 by about a factor of 2 at maximum. Good agreement was obtained between the S12 and moments method calculations.

The above uncertainties indicate that the calculation technique should over-estimate the fast flux at the reactor vessel vall. However, the comparison with thermal flux data indicates a possible underestimate. Until a better comparison with data can be made, we have assumed that the underestimate is correct and ac-cordingly have increased the flux calculations by a factor of 2 to predict the nvt in the reactor vessel.

The reactor utilizes a larger water gap and thinner thermal shield between the core and the reactor vessel vall when compared to currently licensed plants.

The effect of this steel-water configuration on (a) the neutron irradiation, and (b) the thermal stresses in the reactor vessel vall, were evaluated as follows:

a. Heutron Irradiation Calculations were performed in connection with the reactor vessel de-sign to determine the relative effects of varying the baffle and ther-mal shield thicknesses on the neutron flux (>l Mev) at the vessel vall.

These calculations were performed with the P1 option of the P3MG1~ code (3) using 3h fast neutron groups. The results showed that the neutron flux level at the vessel vall is dependent, for the most part, on the total metal and water thickness between the core and the vessel. However, there was some variation in fluxes depending upon the particular con-figuration of steel-vater laminations. Also, the gain in neutron at-tenuation by replacing water with steel diminishes somewhat with in-creasing steel thickness.

In general however, the results showed that for total steel thicknesses in the range of 3 to 6 in., 1 in. of steel in place of 1 in. of water would reduce the neutron flux above 1 Mev by about 30 per cent. In pure water the calculations showed that the neutron flux would be re-duced, on the average, by a factor of 6 in 6 in. of water.

Based on the above analysis a comparison has been made of the neutron attenuation in this reactor vessel with those in San Onofre, Turkey Point 3 and h, Indian Point 2, and Ginna. The total distance between this core and the reactor vessel.is 21 in. This provides from 1.5 to as much as 5 75 in. more distance between the core and the vessel than in the other reactors. For neutrons above 1 Mev it was found that this additional distance would provide additional attenuation ranging from a factor of 1.1 to 5 times greater than that in the other PWR's con-sidered.

b. Thermal Stresses The gamma heating in the reactor vessel is produced by primary ga= mas-from the core and by secondary gammas originating in the core liner,

' barrel, thermal shield, and the vessel itself. In this reactor design the major portion of the heat is generated by gamma rays from the core v/ and by. secondary gamma rays from the core liner and barrel.

0122

[ '3-19

• er

.g - -.

p/u

-Since$ the gammas 'from each of these sources must penetrate the thermal shield to reach the vessel, the vessel heating rate is dependent on -)

the thermal. shield thickness.

For designs which employ thicker thermal shields, or in which internals are to be exposed to higher neutron fluxes, gamma rays originating in the thermal shield or in-the vessel itself may govern the vessel heat-

-ing rates. Since gamma rays from these sources would have to penetrate

~

only portions or none of the thermal shield to reach the vessel, the vessel heating in such cases would be less dependent on thermal shield

. thickness than in this reactor design.

A comparison was made between the gamma attenuation provided by the water and metal in this reactor vessel and that in other PWR's by as-

. suming that, in each design, the vessel heating was dependent on the gamma ray attenuation provided by the thermal shield. This approach would be conservative since, as noted above for some designs, gamma sources other than those attenuated by the thermal ~ shield may contrib-ute appreciably to the vessel heating. The results of the comparison showed that the difference in gamma attenuation between this reactor and other PWR's ranged from negligible difference to a factor of 5.3 less for this_resetor design.

The maximum steady-state stress resulting from gamma heating in the vessel has been calculated to.be 3,190 psi (tension). This is a rel-atively low value, and no problems are anticipated from thermal stresses in the reactor vessel vall.

'3.2.2.2L ' Nuclear Evaluation Analytical models and the application of these models are discussed in this sec-tion. Core _ instabilities associated with xenon oscillation are also mentioned, with threshold data evaluated under reference conditions.

3.2.2.2.1 Analytical >Models Reactor design calculations are made with a large number of computer codes. The choice of which code set or sets to use depends en which phase of the design is being analyzed. A list of codes used in core analysis with a brief discussion follows in 3.2.2.2.2.

2. Reactivity Calculations Calculation of the reactivity of a pressurized water reactor core is performed in one, two, or three dimensions. The geometric choice de-pends 'cn1 the type of calculations to be made. In a clean type of cal-culation where there are no strong, localized absorbers of a type dif-

~

fering from the rest of the lattice, 1-dimensional analysis 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.'h),.KATE (Ref. 5), RIP, WANDA (Ref. 6), and a depletion routine.

Normally the MUFT portion is used with 3h-energy groups , an exact treat-ment of hydrogen, the Greuling-Goertzel approximation for elements of mass less than 10, and Fermi age for all heavier elements. The KATE

=#

j,QWf 3-20 m

A portion normally uses a Wigner-Wilkins spectrum. In WANDA, k-enerEy groups are utilized._ Disadvantage factors for input to the thermal group.are' calculated with the THERMOS (Ref. 7) 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%-ak/k.

A 1-dimensional analysis of a geometric arrangement, where there are localized strong absorbers such as CRA's, 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 problem 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 reactiv-ity coefficients, fuel cycle enrichments, lifetimes, and soluble poi-son concentrations can.be found to improve the initial conditions spec-ified for 2-dimensional analysis.

Two-dimensional reactivity calculations are done with either the PDQ (Ref. 8) or TURBO (Ref. 9) diffusion and/or depletion codes. These 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 de-tail 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-dimen-sional calculations. This.is particularly true when local power peak-ing as a function of power _ history.is of interest. It is necessary to study this type of problem vith at least a 2-dimensional code, and in somefeases 3-dimensional calculations are necessary.

Use of the 2-dimensional programs requires the generation of group con-stants as a' function of material composition, power history, and geo-metry. For regions where diffusion theory is valid, MUFT and KATE with THERMOS disadvantage factors are used to generate epithermal and-thermal coefficients. This would apply at a distance of a few mean

. paths from boundaries or discontinuities in the fuel rod lattice. Dis-continuities . refer 'to ~ fuel assembly can, water channels , instrumenta-tion ports, and CRA guide tubes. The interfaces between regions of different' 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 represent-ed well in slab geometry. This. problem is analyzed by P3MG (Ref. 3) 1which is~ effective in slab geometry. _The coefficients so generated

-are utilized in the epithermal energy range. Coefficients.for the thermal energy range are generated by a slab THERMOS calculation. The N/ regions adjacent to an interface of material of different enrichment 3 ,.are also

~

s-well represented with the P3MG code. () bqh'- '

%.h a i' g 3-21 $

7 - ,.

i

. p_

, l) :

The arrangement of instrumentation ports and control rod guide tubes

. lends _itself to cylindrical geometry. DTF-IV (Ref. 10) is quite ef-fective in the analysis of this arrangement. Input to DTF-IV_is from '}

-GAM (Ref. 11) and THERMOS or KATE. Iteration is required between 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 transi-tion from fuel to reflector and baffle is also represented by the DTF-IV code. The 3-dimensional analysis is acco=plished 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 was made by the comparative analysis of 14 e itiegl experiments with varying rod and rod assembly configurations.13,14) 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 as part of the lattice study. The resulting comparison of the analyt-ical and experimental vorths are shown in Table 3-8. Details of the critical configurations are given in References 13 and 14.

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 Worth, % A k/k 5-B 4 4 252 2.00 1 98 4-F 4 9 0 3 38 3 34 5-c 2 12 276 2 38 2 35 4-D 1 16 0 1.43 1.42 5-D 2 16 284 2.80 2.82 4-E 1 20 0 1 54 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 experimen-tal 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-veighting. 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-

, quired eignvalues. ,

- &.h,,

~""

6diis

GAKER~and LIBPM are used to prepare data for THERMOS. GAKER generates scattering cross sections for hydrogen by the Helkin 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-weighted disadvantage factor. This is used in the homogenization of fuel cells and gives a first order cor-rection 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 sec-oud 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-veighted con-trol rod cell coefficients as a function of boron concentration. As a cheak on the validity of the THERMOS approach, extrapolation dis-tanc were compared to those given by the Spinks method.(15) The agrt2 ment was within 2.2 per cent for a set of cases wherein the num-ber 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 whe . 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, void-ing, and pressure, and for fuel temperature. These are calculated 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 Calculations' ,

This section contains a brief description of codes mentioned in the preceding sections.

THERMOS (Ref. 7) - This code solves the integral form of the Boltzmann Transport Equation for the neutron spectrum as a function of po-sition. A diagonalized connection to the isotropic transfer ma-trix has been incorporated allowing a degree of anisotropic scat-tering.

t MUFT (Ref. h) - This program solves the F1 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. l

, Coefficients are generated with MUFT for the epithermal energy i range.

, , Ol?J I

~3-23

KATE (Ref. 5) - The code solves the Wigner-Wilkins differential equation for a homogeneous medium moderated by chemically unbound hydro- ,

gen atoms in thermal equilibriun. Coefficients for the thennal 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 peaks, and computes L-factors for input to MUFT, PDU, and P3MG.

. WANDA (Ref. 6) - This code provides numerical solutions of the 1-dimen-sional fev-group neutron diffusion equations.

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

GAM (Ref. 11) - This code is a multigroup coefficient generation program that solves the Pl equations and includes anisotropic scattering.

Inelastic scattering and resonance parameters are also treated by GAM.

P3MG (Ref. 3) - The code solves the multienergy transport equation in various geometries. The code is pri=arily used for epithermal coefficient generations.

IYrF (Ref.10) - This code solves the multigroup,1-dimensional Boltzmann .

transport equaticn by the method of discrete ordinates. DIF al- /

lovs cultigroup anisotropic scattering as well as up and down scattering.

PDQ (Ref. 8) - This program solves tne 2-dimensional neutron diffusion-depletion problem with up to five groups. It nas a flexible rep-resentation of time-dependent cross sections by means of fit op-tions.

TURBO (Ref. 9) - This code is -similar in application to the PDQ depletion program. It, however, lacks the great flexibility of the PDQ fit options.

CANDLE (Ref. 9) - This code is similar to TURBO, but solves the diffusion equations in one dimension.

~ TNT (Ref. 9) - This code is similar in application to TURBO, but is a 3-dimensional code extended from DRACO.

3.2.2.2.3 Xenon Stability Analysis Initial studies of the initial and equilibrium cores, where realistic fuel tem-peratures are generated by thermal-nuclear iteration, indicate no instability at any time during the life cycle. These results are encouraging, but until more detailed analyses are completed, it vill be assumed that axial xenon os-cillations are possible. Animuthal oscillations are unlikely, and radial os-cillations vill not occur.

4 3-24 -E L

-An extensive investigation must be ecmpleted before th~e stability of a core can be ascertained. An adequate solution can be found by first using analytical Ltechniques in the manner of Randall and St. John to predict problematic areas, and then by analyzing these with diffusion theory programs that are coupled with

~

heat transfer equations.

The results of the. stability analysis of the reference core are presented be-low, followed by the methods section containing the details of the threshold and diffusion theory calculations employed. The closing section outlines an overall approach to the solution of the stability problem in regard to addi-tional detailed calculative programs as well as a method for the correction of unbalanced power distributions.

a. Summary of Results (1) Threshold Analysis In the threshold analysis axial, azimuthal, and radial oscilla-tions were investigated for beginnin~-of-life, flattened, and slightly dished power distributions. 17) The results are as follows:

(a) 'For a fixed dimension, the tendency toward spatial xenon oscillation is increased as the flux increases.

(b) For a fixed flux, the tendency toward spctial oscillation is increased as the dimension of the core increases.

(c) The large size of current FWR designs permits an adequate xenon description using 1-group theory.

(d) Flattened power distributions are more unstable than nor-mal beginning-of-life distributions. Dished power distri-butions are even vorse.

(e) In a modal analysis of the reference core, modal coupling can be ignored. In addition, the core is not large enough to permit second-harmonic instability.

(f) A large, negative power coefficient tends to dampen oscil-lations. If this coefficient is sufficiently large, os-cillations cannot occur regardless of core size or flux level. Current PWR designs have a substantial negative power coefficient.

(g) The critical diameter for azimuthal oscillations is larger than the critical height for axial oscillations.

(h) The reference core design is not large enough to excite radial oscillations. )

(1) Examination of the diameter, height, and power coefficient for this reference design indicates that oscillations

~~/ should not occur at the beginning of life with unflattened J N 3-25 1

u _

'2h, 5 ,

s n;,

')

_ power distributions. However, there exists a finite prob- ~)

i ability of oscillations at some later time, since core de-pletion tends to flatten the power distribution.

(j) The period of oscillation (25 to'30 hours3.472222e-4 days <br />0.00833 hours <br />4.960317e-5 weeks <br />1.1415e-5 months <br />) is long enough to permit easy control of the oscillations.

(k) The modal analysis of this core toward the end of the ini-tial cycle (with about 80 per cent flatness) showed that axial oscillations are possible, azimuti.al oscillations are unlikely, and radial oscillations will not occur.

(2) Depletion Analysis Diffusion-depletion calculations coupled with heat transfer equations were employed to investigate further the axial sta-bility of the core since the analytical study indicated that this was the most probable mode of oscillation. The results follow:

(a) Axial instability did not occur at any time during the ini-tial cycle. An avera6e fuel temperature of 1,400 F was maintained during the cycle.

(b) The threshold for axial instability near the end of tne initial cycle was found to coincide with a core average fuel te=perature of 900 F.

Diffusion theory was also used to examine the problem of can-trolling the system with rods if the stabilizing power Doppler was not present. The following was concluded:

(a) Partial control rods are quite adequate in controlling axial oscillations. These rods have 3-ft-long poison see-tions which are moved up and down about the midplane of the core to offset oscillatory power shifts.

(b) Detailed power profiles will be available to the reactor operator as output from the instrumentation. The large period of the oscillation vill allow partial rod movement such that axial power peaks are held well within allowable limits.

b. Methods (1) Threshold Analysis The method used in the threshold analysis is an extension of the 1-group treatment including power coefficient introduced by Randall and St. John. One- and 2-group treatments have been compared, and the conclusion drawn is that a 1-group model is satisfactory for large cores. For all three geometries, data

-were generated as a function of:

P 0129 3-26

(a) Core size.

(b) Flux level.

(c) Degree of flatness in the power distribution.

(d) Power coefficient.

(e) Reactivity held by saturation xenon.

In addition, slightly dished power distributions were investi-

. gated to show that any dishing resultin6 from high depletion is not sufficient to require correction to data based on replacing the dished segment with a flat power distribution.

The effect of modal coupling has been exa=ined and shown to be of no consequence for cores similar to the reference reactor design. Values of the critical dimension varied no more than 1 to 2.8 per cent for the same core with and without modal cou-pling. The lover value was computed with a zero power coeffi-cient and was not conservative without modal coupling. The higher value was co=puted with the reference power coefficient and was conservative without modal coupling.

Table 3-9 summarizes those parameters for the reference core which affect the xenon stability threshold. The parameters were calculated at two substantially different times in core life.

Reference physical di=ensions are also shown for comparison pur-poses in the following discussion.

Table 3-10 shows the threshold dimensions for first mode in-st~ ability as a function of flux flattening. The percentage of flattening is defined as 100 per cent times the ratio of the flattened power distribution to the total physical dimension under consideration. The parameters of Table 3-9 at two full power days were used~ since they are virtually the same as those at 150 days but are more conservative. Axial depletion studies show that power distributions are flattened by 0, 63, and 73 per cent at 2, 150, and 35h full power days, respectively. A maximum flatness of approximately 80 per cent may be expected for long core life.

An examination of the data in Table 3-10 shows that--with the maximum flatness--axial oscillations are possible, azimuthal oscillations are unlikely, and radial oscillations will not oc-cur.

Threshold dimensions for second mode oscillations were 50 per cent larger.in magnitude than those shown in Table 3-10 for the first mode. Oscillations in the second mode vill not occur in the reference core.

0130

.I +

3 hkhkh 3-27

Table 3-9 Reference Core Parameters \

Two Full (Rated) 150 Full (Rated)

Power Days Power Days M2 , em2 57.0 57.0 ith, n/cm 2-sec 3.9 x 1013 3.8 x 1013 ax (reactivity held by saturation-xenon),Ak/k O.03h 0.033 Doppler Coefficient, (Ak/k)/F -1.1 x 10-5 -1.1 x 10-5 Moderator Temperature Coefficient Positive but Small Negative aT (power doppler coeff.),(Ak/k)/ unit flux = -2.2 x 10-16 = -2.3 x 10-16 Equivalent Dimensions, ft Height 12.00 Diameter 10 74 Radius 5 37

't Table 3-10 First Mode Threshold Dimensions and Flatness Flatness, %

Threshold Dimensions , ft O 50 80 Threshold height (axial oscillations) 18.5 lh.1 11.8 Threshold diameter (azimuthal oscillation) 20.h 16.5 1h.0 Threshold radius (radial oscillation) 16.8 16.7 1h.5 Table 3-11 shows the values of H/D versus power flatness for equal likelihood of axial, azimuthal, and radial first harmonic

' oscillations, i.e. , if the core is just at the axial threshold for axial oscillations, it can also be expected that there vill be azimuthal and radial oscillations provided the value of H/D in Table 3-11 is satisfied. H/D for this reference reactor is 1.12.

g 0131 t

3-28 i

E

Table 3-11 Threshold Ratio and Power Flatness Flatness, %

Ratio 0 20 50 80 100 H/D (axial versus asimuthal) 0.91 0.87 0.86 0.86 0.85 H/D (axial versus~ radial) 0 55 0.k9 0.h2 0.hl 0.hl The model methods used to examine the xenon oscillation problem made use of core-averaged quantities such as flux, power coeffi-cient, and reactivity held by saturation xenon. In addition, flux distributions were limited to (a) Geometric distributions.

(b) Partially or totally flat.

(c) Slightly dished.

The power distribution during early life is such that no xenon instabilities will occur. The power flattening effect of fuel burnup with time renders the core more susceptible to xenon os-cillations.

(2) Depletion Analysis Core-averaged quantities were used in the analytical analysis.

For a more comprehensive investigation, it is desirable to study xenon' oscillations with diffusion-depletion programs including heat transfer. Such calculations, which include the important local temperature effects, allow the designer to look for xenon oscillations under actual operating conditions. For these rea-sons, the B&W LIFE depletion program was modified to include axial heat transfer. The equations and iteration scheme are outlined below:

(a) The average fluid temperature for each axial core region is computed from a previously known power density distribu-tion as follows:

ATi = (Tout - Tin)1=C)Z PD(Z)dZ (A) in where ATt= temperature change in re61on "i" '

PD (Z) = power density in Z direction v' '

l1 . +,

Zin' Zout = region "1" boundaries 03.32 3-29

4 and' ~

)

C= (B)

PD (Z) dZ vhere H = active fuel height.

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

Tout + Tin T

• fluidi 2 (b) The newly computed region-averaged fluid temperatures are used to compute new fluid densities. These fluid densities

. are then used to adjust the number densities for water and soluble poison. Iocal or bulk boiling is not permitted.

(c) The average fuel temperature for each axial core re5 1on is then computed from the average fluid temperatures and power densities:

l Tfueri = K PDi+Tfluid i e t

where 5 1= coverage power density of region "1" and K is defined by fuel ~ fluid core K= __ (E)

PDcore (d) After the new fluid temperatures, moderator densities, and fuel temperatures are obtained, these quantities are used as new LIFE input to obtain a'new power distribution until either a convergence criterion is met or a specified num-ber of iterations is made.

This analysis used.an exact solution in that the spectrum was recaluelated for e'ach zone (11 axial zones described the reae-

! ' tor) for each iteration at every time step. This included the effects of the moderator coefficient.

This LIFE package was used to determine the effects of the un-certainty in the power Doppler on the stability of the core.

The uncertainty in the Doppler was more than compensated witn a reduction in fuel temperature of 500 degrees. The reference core was analyzed with core average fuel temperatures of 1,h00 0133 L.

4 F and 900 F. Figure 3-7 compares the cyclic response of these two cases following the 3-ft insertion and removal (after two hours) of a 1.2% Ak/k rod bank near the beginning of life.

These studies were made at beginning-of-life boron levels of ap-proximately 1900 ppm. This level is approximately 200 ppm above

.the predicted beginning-of-life level and, consequently, reflects a more positive moderator coefficient than would be expected.

Case 1 on Figure 3-7 depicts the behavior of the core if the heat transfer equations were not included in the calculation.

Figure 3-8 shows the effect 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 oscillation in this core. On the basis of the information presented, it can be said that for a realistic fuel temperature this core does not exhibit axial instability at any time during the initial cycle.

The 1-D 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 oscil-lation as shown in Figure 3-9 A study was completed wherein a 1% Ak/k rod bank with a 3-ft-long section of regular control rod material was successfully maneuvered to control the core after a perturbation of the power shape at a point about 3/L of the way through Cycle 1. The controlled results are also shown in Figure 3-9 The minimum rod motion was one foot, and the time step employed was h.8 hours9.259259e-5 days <br />0.00222 hours <br />1.322751e-5 weeks <br />3.044e-6 months <br />. More precise rod movement over shorter time periods would produce a much smoother power ratio curve. This control mechanism appears quite adequate,

c. Conclusions Instability in the radial or azimuthal mode is not expected since the diffusion theory study showed that the core is stable throughout life-time and the L/D ratio is 1.1. The results are encouraging, but un-til additional analyses are completed, it will be assumed that axial xenon' oscillations are possible. Consequently, rod motion vill be used to compensate for unbalanced power distribution as indicated by the instrumentation.

Work is underway to provide a 2-dimensional depletion program which allows nuclear-thermal iterations. A detailed quantitative analysis of core stability and control procedures is to be undertaken with the new program.

3.2.3 THERMAL AND HYDRAULIC DESIGN AND EVALUATION 1

3.2.'3.1 Thermal and Hydraulic Characteristics

)

3.2.3.1.1 Fuel Assembly Heat Transfer Design a.- Design Criteria i 1

-' i The criterion for heat transfer design is to be safely below Departure l

'from Nucleate Boiling (DNB) at the design overpower (llh per cent of

~'

t",',..

N '

03M 3-31 1 c

rated power). A detailed description of the analysis is given in ~.

3.2.3.2.2,. Statistical Core Design Technique. l

.The input information for the statistical core design technique and for the evaluation of individual hot channels consists of the follow-ing:

(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, and calculations as outlined below,

b. Heat Transfer Equation and Data Correlation The' conditionsheat transfer relationship is presented used to) predict in BAW-168.(10 limiting heat transfer The equation is as follows:

_q " =,(1.83 - 0.000h15 P) x 90,000

-G 0.3987 + 0.001036 ATese -

. 7x -

(ATesc}

240 .

q" = critical heat flux as predicted by the best fit form, Btu /hr-ft2 P = core operating pressure, psia G = channel mass velocity, lb/hr-ft2 S = channel equivalent diameter, ft L = length up the channel to the point of interest,.ft AT,3c = inlet subcooling Msat ~

inlet '

T = coolant saturation temperature corre-sponding to P, F

-This equation was derived from experimental heat transfer data. An analysis of heat transfer data for this and other relationships is described in detail in 3.~2.3.2.3, correlation of Heat Transfer Data.

Individual channels are analyzed to determine a DNB ratio, i.e. , the ratio of the heat flux at which a DNB is predicted to occur to the 5e 32 ,

01.%

a

heat . flux'in the channel being investigated. This DNB ratio is re-

. lated to_the data correlation as in Figure 3-10. A confidence and population value is associated with every DNB ratio as described in the Statistical Core Design Technique. The plot of DNB versus P shown is for confidence of 99 per cent.

The DNB and population relationships shown are also the values asso-ciated with the single hot channel analysis for the hottest unit cell where a 1.38 DNB ratio corresponds to a 99 per cent confidence that at least 94.5 per cent of the population of all such hot channels are in no jeopardy of experiencing a DNB. This statement is a corollary to the total core statistical statement given in 3.1.2.3, Thermal and Rydraulic Limits. The criterion for evaluating the thermal design margin for individual channels or the total core is the confidence-population relationship. The DNB ratios required to meet the basic criteria or limits are a function of the experimental data and heat transfer correlation used, and vary with the quantity and quality of data.

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 comprising a coolant channel is determined from a nuclear analysis of the core and fuel assemblies. The results of this analysis are as follows:

(1) The nominal nuclear peaking factors for the worst time in core life are FAh = 1.79 Fz = 1 70 ,

Fq = 3.04 (2) The design nuclear peaking factors for the worst time in core life are FAh = 1.85 Fz = 1 70 Fq = 3.15 where FAh = max / avg total power ratio (radial x local nuclear)

Fz = max / avg axial power ratio (nuclear)

Fq = FAh-x Fz (nuclear total)

The nominal values are the maximum calculated values. The design

. . values are obtained by increasing -the maximum calculated total power ratio, FAh, from 1.79 to 1.85 to obtain a more conservative design.

..  : 013G g

=t-  :,

3-33

J The axial nuclear factor, Fz, i; illustrated in Figure 3-11. The dis-tribution of power expressed as P/P is shown for two conditions of re- ]

actor 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 re-sults 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 withdrawn. 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 15 cosine power distribution. The in-let peak shape has a larger maximum value. However, the position of the 1.5 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.(19) The 1.5 cosine axial shape has been used to determine individual channel DNB limits and 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. Line B shows the distribution of the maximum calcu-lated values of Fah for nominal conditions with a maximum value of 1.79 The distribution of peaking factors for the design condition is obtained by increasing the maximum calculated value for all rods in the core by the ratio of 1.85/1.79 or 1.033 to provide conserva- s tive results. Determination of the peaking distribution for the de- /

sign condition in this manner has the effect of increasing reactor power by about 3 per cent. This assu=ption is conservative since the distribution with a maximum peak Fah of 1.85 vill follow a line simi-lar to Line C where the average power of all rods in the core is repre-sented by an Fah of I.0. The actual shape of the distribution curve is dependent upon statistical peaking relationships, CRA positions, moderator conditions , emd operating history. The shape of the distri-bution curve vill be more accurately described during the detailed core design.

d. Engineering 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 deter-mined' statistically from fuel assembly as-built or specified data where Fq is a heat input factor, QF " 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 3-12. The factor, the coefficient of variation, the standard deviation, and the mean value are tabulated.

~

3-3h u

Table 3-12 Coefficients of Variation CV No. Description a 5I CV 1 Flow Area 0.00075 0.17625 0.00h26 2 Local Rod Diameter 0.000485 0.k20 0.00116 3 Average Rod Diameter 0.000h85 0.k20 0.00116 (Die-drawn, local and average same) h Local Fuel Loading 0.00687 Subdensity 0.006h7 0 95 0.00681 Subfuel area 0.000092 0.1029 0.00089 (Diameter effect) 5 Average Fuel Loading 0.00370 Subdensity 0.0032h 0.95 0.003h1 Sublength 0.16181 1h4 0.00112 Subfuel area 0.000092 0.1029 0.00089 (Diameter effect) 6 Local Enrichment 0.00323 2.2k 0.001hh 7 Average Enrichment 0.00323 2.24 0.001hh CV Coefficient of Variation o/x o Standard Deviation of Variable x Mean Value of Variable (Enrichment values are for worst case normal assay batch; maximum variation occurs for minimum enrichment.)

e. Core Flow Distribution Hot Channel Factors The physical arrangement of the reactor vessel internals and nozzles results in a nonuniform distribution of 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 is being performed to demonstrate the adequacy of the internal arrangements. The final variations in flow will be determined when the tests are completed.

Interim factors for flow distribution effects have been calculated from test data on reactor vessel models for previous pressurized water reactor designs.

A flow distribution factor is determined for cach 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 01.38

],-

g 3-35 9

j -

ratio may be greater'or less than 1.0 depending upon the position of the assembly being evaluated. The flow in the central fuel asserblies

-is in general larger than the flow in the outermost 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.hi. 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 hot-test 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 flow factors 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 Desien Overpower Core performance is assessed at the maximum design overpower. The selection of the design overpower is based on an analysis of the re-actor protective system as described in Section 7 The reactor trip point is 107.5 per cent rated power, and the maximum overpower, which is 114 per cent, vill not be exceeded under any conditions.

g. Maximum Design Conditions Analysis Summary The Statistical Core Design Technique described in 3.2.3.2.2 was used to.analy::e the reactor at the maximum design conditions described previously. The total number of fuel rods in the core that have a possibility of reaching DNB is shown in Figure 3-13 for 100 to 118 per cent overpower. Point.A on Line 1 is the maximum design point for 114 per cent power with the design FAh nuclear of 1.85 Line 2 was calculated using the maximum calculated value for Feb nuclear of 1 79 to show the margin between maximum calculated and design condi-tions. It is anticipated that detailed core nuclear analyses vill permit a lowering of the maximum design value for Fah.

The number of fuel rods that may possibly reach a DNB at the maximum design condition with an Fah of 1.85 and at 11h per cent overpower, represented by point A on Figure 3-13, forms the basis for this sta-tistical statement:

There is a 99 per cent confidence that at least 99.5 per cent of the fuel rods in the core are in no ,)eopardy of experiencing 3-36 "

()].39

a departure from nucleate boiling (DNB) during continuous oper-ation at the design overpower of 114 per cent.

- Statistical results for the maximum design condition calculation shown by Figure 3-13 may be summarized as follows in Table 3-13 Table 3-13 DNB Results - Maximu= Design Condition (99 per cent confidence Level)

Power, Possible Population Point  % of 2,h52 MWt FAh DNB's Protected, ".

A 11h 1.85 184 99 50

- B 11h 1.79 100 99.73 C 100 1.85 17 99.95 -

D 100 1.79 10 99.98 E 118 1.79 184 99.50

h. Most Probable Desien Condition Analysis Summary The previous maximum design calculation indicates the total number of rods that are 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 described in 3.2.3.2.2. For example, all channels are analyzed with FA (flow area factor) less than 1.0, Fq (heat input factor) greater than 1.0, and with minimum fuel assembly flow. It is physically impossible for all channels to have hot channel character-istics. A 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 3-1h. The analysis may be su=marized as follows in Table 3-14.

Table 3-lh DNB Results - Most Probable Ccndition Power, Possible Population Point  %.of 2,h52 MWt Fah DNB's Protected, %

F 100 1.79 2 99.99h G 114 1.79 32 99.913 ,

H 118 1.79 TO 99.815 '

/ 1 s -

0140 l 3-37 s

n . _

7 . - .

The analysis was made from Point F at 100 per cent power to Point H 'T at 118 per cent power to show the sensitivity of the analysis with '

power. The vorst condition expected is indicated by Point G at 11h per cent power where it is shown that there is a small possibility that 32 fuel rods may be subject to a departure from nucleate boiling (DNB). This result forms the basis for the following statistical statement for the most probable design conditions:

There is at least a 99 per cent confidence that at least 99 9 per cent of the rods in the core are in no jeopardy of experi-encing a DNB, even with continuous operation at the design over-power of 114 per cent.

i. Distribution of the Fraction of Fuel Rods Protected The distribution of the fraction (P) of fuel rods that have been shown statistically to be in no jeopardy of a DNB has been calculated for the maximum design and most probable design conditions. The computer programs used provide an output of (N) number of rods and (P) fraction of rods that vill not experience a DNB grouped for ranges of (P). The results for the most probable design condition are shown in Figure 3-15 The population protected, (P), and the population in jeopardy, (1-P),

are both plotted. The integral of (1-P) and the number of fuel rods gives the number of rods that are in jeopardy for given conditions as s shown in Figures 3-13 and 3-lk. The number of rods is obtained from ).

the product of the percentage times the total number of rods being considered (36,816). The two distributions shown in Figure 3-15 are for the most probable condition analysis of Points F and G on Figure 3-lh. The lower line of Figure 3-15 shows P and (1-P) at the 100 per cent power condition represented by Point F of Figure 3-lh. The upper l

curve shows P and (1-P) at the lik per cent power condition represen-ted by Point G of Figure 3-14. The integral of N and (1-P) of the upper curve forms the basis for the statistical statement at the most probable design condition described in paragraph h above.

j. Hot Channel Performance Summary The hottest unit cell with all surfaces heated has been examined for l 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 1.85 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 refle -

ted as an increase in AT of the coolant.

Error bands of 65 psi operating pressure and 2 F are reflected in the total core and hot channel thermal margin calculations in the direction producing the lowest DNB ratios or highest qualities.

W -

05lW 3-38

The DNB ratio versus power is shown in Figure 3-16. The DNB ratio in the hot channel at the maximum overpower of 114 per cent is 1.38 which corresponds to a 99 per cent confidence that at least 94.5 per cent of the fuel channels of this type are in no jeopardy of experiencing a DNB. The engineering hot channel factors corresponding to the above confidence-population relationship are described in 3.2.3.2.2 and listed below:

Fq = 1.008 Fqn = 1.013 FA = 0 992 1

The hot channel exit quality for various powers is shown in Figure 3-17 The combined results may be summarized as follows:

Reactor Power. 5 DNB Ratio (BAW-168) Exit Quality, %

100 1.60 0 107.5 (trip setting) 1.47 2.6 114'(maximum power) 1.38 5.h

' 1h9 1.00 23.0 3.2.3.1.2 Fuel and Cladding Thermal Conditions

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

The program uses uniform volumetric heat generation across the fuel diameter, and external coolant conditions and heat transfer coeffi-cients determined for thermal-hydraulic channel solutions. The fuel thermal conductivity is varied in a radial direction as a function of the temperature variation. Values for fuel conductivity were used as shown in Figure 3-18, a plot of fuel conductivity versus temperature.

The heat transfer from the fuel to the clad is' calculated with a fuel and clad expansion model proportional to temperatures. The temper-ature drop is calculated using gas conductivity at the beginnin lifeconditionswhenthegasconductivityis0.1 Btu-ft/hr-F-ftg-of- .

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 coefficient is de-pendent.upon pressure and gas. conductivity.

A plot of fuel center temperature versus linear heat rate in kw/ft is shown in Figure 3-19 for beginning-of-life conditions. The linear-heat rate at the maximum overpower of 114 per cent is 19.9 kv/ft.

The corresponding center fuel temperature shown in Table 1-2 is h,h00 F. . The center and average temperatures at 100 per cent power are

.4,160 and 1,385 F as shown in Table 3-1.

();jg;3

, jb(%few .3-39 ,

9 p , .N1 -

- .s The peaking factors used in the calculation are Fah = 1.85 ' ')

Fz = 1.70 Fqn = 1.03 Fq (nue. and mech.) = 3.2h A conservative value of 1.03 was assumed for the heat flux peaking factor, Fqa. The assigned value corresponds to a 99 per cent confi-dence and 99.99 per cent population-protected relationship as describ-ed in the statistical technique.

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

Boiling cond tigns prevail at the hot spot, and the Jens and Lottes relationship 20/ 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 psig.

3.2.3.2 Thermal and Hydraulic Evaluation 3.2.3.2.1 Introduction .

I Summary 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 hat channel DNB data were presented to relate the nev

. method with previous criteria. A comprehensive description of the new tech-nique is included in this section to permit a rapid evaluation of the methods used.

The BAW-168 correlation is a B&W design equation. An extensive review of data available in the field was undertaken to derive the correlation and to deter-mine the confidence, population, and DIIB relationships included in this section.

A comparison of the BAW-168 correlation with other correlations in use is also in'cluded.

A' detailed evaluation and sensitivity analysis of the design has been made by examining the hottest channel. in the reactor for DNB ratio, quality, and fuel temperatures. BAW-168 DNB ratios have been compared with W-3 DNB ratios to facilitate a comparison of the design with PWR reactor core designs previously reviewed.

1 0143 p

M 3-ho t _U

3.2.3.2.2 Statistical Core Design Technique The core thermal design is based on a Statistical Core Design Technique devel-

.oped 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 resultant 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 will provide a good measure of the core safety and re-liability since the method provides a statistical statement for the total core.

This statement also reflects the conservatism or design margin in the calcula-tion.

A reactor safe operating power has always been determined by the ability of the coolant to remove heat from the fuel material. The criterion that best mea-sures this ability is the DNB, which involves the individual parameters of heat flux, coolant temperature rise, and flow area, and their intereffects. The DNB criterion is commonly applied through the use of the departure from nu-cleate boiling 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 input and the flow rate. It is possible to separate these two effects; the statisti-cal hot channel factors required are a heat -input factor, FQ, and a flow area factor, F A. In addition, a statistical heat flux factor, Fq", is required; the heat flux factor statistically describes the variation in surface heat flux.

The DNBR is most. limiting when the burnout heat flax is based on minimum flow area (small F A) and maximum heat input (large Fq), and when 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 correlation is based.

To afford protection against DNB, the DNB heat flux computed by the best-fit correlation is divided by a DNB factor (B.F.) greater than 1.0 to yield the design DNB surface heat flux. The basic relationship DNBR = B.F. x f(F A , FQ) x Q" xFnq f,

involves as parameters statistical hot channel and DNB factors. The DNB fac-ter (B.F.) above is usually assigned a value of unity when reporting DNB ra- ,

tios so that the margin at.a.given condition is shown direc;tly by a DNBR 1 greater than 1.0, i.e., 1.38 in the hot channel. 'v.

3.

'Q

~

0144

+

• j

.3-h1

~

.To find the DNB correlation, selected correlations are compared with DNB data obtained in the B&W burnout loop and with published data. The comparison is ']

facilitated by preparing histograms of the ratio of the experimentally deter-mined-DNB heat: flux- E($ ) to the~ calculated value of the burnout heat flux (4C). A typical histogram'is shown in Figure 3-20.

A histogram is obtained for each DNB correlation considered. The histograms indicate the ability of the correlations to describe the data. They indicate, qualitatively, the . dispersion of the data about the mean value--the smaller the. dispersion, the better the correlation. Since thermal and hydraulic data generally are well represented with a Gaussian (normal) distribution (Figure 3-20), mathematical parameters that quantitatively rate the correlation can be easily obtained for the histogram. -These same mathematical paramete'rs are the basis for the statistical burnout factor (B.F.).

In analyzing a-reactor core, the statistical information 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-factors can be shown graphically (Figures 3-21 and 3-22).

All th'e 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 statisti-cal hot channel and burnout factor is a function of the percentage of confi-dence desired in the result, and the portion of all possibilities desired, as

. well as the amount of data used in determining the statistical factor. A fre- )

quently used assumption in statistical analyses is that the data available rep-resent an infinite sample of that data. The implications of this assumption should be noted. For instance, if limited data are available, such an assump-tion leads to a somewhat optimistic result. The assumption also implies that more information 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 pro-cedure 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 influence of the amount of data for instance can be illustrated easily as follows: . Consider the heat flux factor which has the form Fn q =1+K F p

where. pn .is the statistical hot channel factor for heat q

flux K is a statistical multiplying factor

a. 'is the-standard deviation of the heat flux fac-Q"' tor, including the effects of all the subfactors

>If opqn = 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 is

Fqn = 1 + (2.608 x 0.050) = 1.1304 g _.

• P 03.45 l

W 3-4 I .

'~

L

and will 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 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 Fq , = 1 + (2.326 x 0.050) = 1.1163 which implies 100 per cent confidence in th9 calculation. The values of the K factor used above are taken from SCR-607.t21) The same basic techniques can be 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 factors 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 ra-tio that is substantially greater than 1.0, even at the design overpower, and that theoretical 13 no rod can have a statistical population factor of 100 per cent, no matter how large its DNB ratio.

L 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 summation. The calculation is made as a function of power, and the plot of rods in jeopardy versus reactor overpower is obtained (Figure 3-23).

The summation of the fraction of rods in jeopardy at each peaking factor summed over all peaking factors can be made in a statistically rigorous man-ner only if the confidence for all populations is identical. If an infinite sample is not assumed, the confidence varies with population. To form this '

summation then, a conservative assumption is required. B&W's total core model assumes that the confidence for all rods is equal to that for the least-pro-tected rod, i.e., the minimum possible confidence factor is associated with the entire calculation.

The result of the foregoing technique, based on the maximum design conditions (114 per cent' power), is this statistical statement:

~

There is a least a 99 per cent confidence that at least 99.5 per cent of the rods in the core are in no jeopardy of experiencing a DNB, even with continuousEoperation at the design overpower.

The maximum design conditions are-represented by these assumptions:

a. The maximum design values of Fah (nuclear max / avg total fuel rod heat input) are obtained by increasing the maximum calculated value of Fah by a factor of 1.033 to provide additional design margin.

b. . The maximum value of F (ndelearmax/avgaxialfuelrodheatinput) is determined for the limiting transient o'r steady state condition.

03/W

^

. i "MT ,

3-43 m .

J

c. Every coolant channel in the core is assumed to have less than the nominal flow area represented by engineering hot channel area factors, -]/

FA, less than 1.0.

d. Every channel is assumed to receive the minimum flow associated with core flow maldistributi3n.

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 Fqa, 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 D:!B 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 The result of the technique based on the most probable design conditions leads to a statistical statement which is a corollary to the maximum design state-ment:

There is at least a 99 per cent confidence that at least 99 9 per cent of the rods in the core are in no jeopardy of experiencing a DNB, even with continuous operation at the design overpower. 1 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 (Fg = 1.0).

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

c. The flow in each coolant channel is based on core flow and power distributions.

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 comment. As to the 0 5 per cent or 0.1 per cent of the rods not included in the statements, statistically, it can be said that no more than 0 5 per; cent or 0.1 per cent of the rods vill be in jeopardy, and that in general the number in jeopardy vill be fewer than 0 5 per cent or 0.1 per cent. The statements do' not mean to specify a given number of DNB's, but only acknowledge the possibility that a given number could occur for the con-ditions assumed. '

In summary', the calculational procedure outlined here represents a substan-tially improved design technique in two ways:

k. * * #

(,47, 9 7 hh h 6* a

4

a. It . reflects the' performance and safety of the entire core in the re-sultant power rating by considering the effect of each rod on the

. power rating.

b. It provides information on the reliability of the calculation and, therefore, the core through the statistical statement.

3.2.3.2.3 Correlation of Heat Transfer Data The BAW-168 ' report (Ref.18) serves as a reference - for the "best-fit" form of the design relationship used by B&W. This heat transfer correlation has been found to be the most satisfactory in the representation of both uniform and nonuniform heat flux. test data. The BAW-168 correlation is used by comparing the integrated average heat flux along a fuel rod to a DNB heat flux limit predicted by.the correlation. For uniform heat flux the integrated average heat flux is equal to the local heat flux. .The comparison is carried out over the entire channel length. The point--at which the ratio of the DNB heat flux to the integrated average heat flux is a minimum--is selected as the DNB' point, and that value of the ratio at that point is the DNB ratio (DNBR) for that channel.

This particular discussion deals with the comparison of DNB data to three par-ticular correlations. The correlations selected were the B&W correlation in the case of BAW-168,(10) a correlation with which the industry is familiar in the case of WAPD-188,(22) and a correlation recently proposed for use in the design of pressurized water reactors in the case of W-3.(23) j The data considered for the purpose of these comparisons were taken from the following sources:

1

a. WAPD-188 (Ref. 22).

4

b. AEEW-R213 (Ref. 2h).

c. Columbia = University Data (Ref. 25, 26, and 27).

d. .Argonne National Laboratory Data, ANL (Ref. 28).

e. The Babcock & Wilcox Company-Data, B&W (Ref. 29).

f. . The Babcock & Wilcox Company Euratom Data (Ref. 30).

The' comparison of. data to the BAW-168 correlation is presented as histograms of the ratio of the experimental DNB heat flux (4E) to the calculated heat flux'(4C). -The data.from each source were grouped by pressure and analyzed as a group; batches.were then prepared including common pressure groups from

.ill sources. Altogether there are 41 different data groups and batches con-sidered.- Histograms for only' the BAW-168 correlation are presented to mini-mize-the graphical material. The information required for the generation of histograms _of the other two correlations was also prepared.

The comparison of the various correlations to each other is facilitated qf through the use of tabulations of : pertinent statistical parameters. The stan-

-f..da.rd deviation and mean value.were obtained from the., computed. values of Mt A' 01.4s

'. a 3-45

($E/4 0-

) for each group or batch. A comparison of standard deviations is some- -~

what indicative of the ability of the correlation to represent the data. ).

However, differences -in mean values from group to group and correlation to cor-relation tend to complicate this type comparison. A relatively simple method may be used to compare the correlations for various data; this method uses the coefficient of . variation (Ref. 31) which is the ratio of the standard devia-

~

tion (c) to the mean T. The coefficient of variation may be thought of as the standard deviation given in per cent; it essentially normalizes the various standard deviations to a common mean value of 1.0.

Table 3-15 is a tabulation of the data source, heat flux type, and correspond-ing histogram numbers. The histograms are shown on Figures 3-24 through 3-39

.)

e 9

W 01.49 e ', .- . : -

3-46

Table 3-15 Heat Transfer Test Data Histogram Figure Source Heat Flux Type Number Nu=ber WAPD-188 Unifom 1-9 3-24 3-25 3-26 AEEW-R-213 Unifom 10-14 3-26 3-27 3-28 Colu=bia Unifom 15-19 3-28 3-29 3-30 ANL Uniform 20 3-30 B&W Unifom 21 3-31 B&W-Euratom Unifom 22-24 3-31 3-32 Ccebined Data (500-720 psia) Uniform 25 3-32 Combined Data (1,000 psia) Uniform 26 3-33 Combined Data (1,500 psia) Unifom 27 3-34 Combined Data (2,000 psia) Uniform 28 3-35 Combined Data (1,750-2,750 psia) Uniform 29 3-36 B&W-Euratom Chopped Cosine Nonunifom 30-32 3-37 B&W-Euratom and B&W Inlet Peak Nonunifom 33-35 3-37 3-38 Euratom and B&W Outlet Peak Nonunifom 36-38 3-38 3-39 Combined Nonunifom '(1,000 psia) Nonuniform 39 3-39 Combined Nonunifo m (1,500 psia) Nonunifom 40 3-39

~

Combined Nonuniform (2,000 psia) Nonuniform 41 3-39 4

3-47 ,

The histograms graphically demonstrate the distribution of (&E/&C) for. ecch data group. The Gaussian type distribution of ($E/4C) about the mean for the ')

group is apparent in the large data groups. Some data groups are too small to provide meaningful histograms, but they are presented in order to completc this survey.

~

The data were used as presented in the source for the calculation of (&E/4C).

no points were discarded for any reason. A good correlation should be capable of representing DNB data for a full range of all pertinent parameters. The result of the comparison on this basis is demonstrated in Table 3-16. The data source, pressure, histogram figure number, heat flux type, and number of data points in the group are tabulated. For each of the three correlations the following data'are indicated:

o/E The coefficient of variation based on all available data in the group.

n The number of data points rejected using Chauvenet's criterion (Ref.

R 32). This criterion is statistical in nature and is applied to the values of ($E/ 4C). Data points that fall outside certain limits with respect to the main body of data are rejected.

(c/E)' The coefficient of variation based on the original data sample less those points rejected by Chauvenet's criterion, i.e. , based on n-nR values of ($E/4C)*

It is unfortunate that Chauvenet's criterion must be applied to the values of

($E/ fC) rather than to the original data, since application to (&E/&C) leads )

to the rejection of points for either of two reasons:

a. Bad data points.

b. Inability of the correlation to represent a particular data point.

It is not desirable to reject points for the second reason, and yet one might expect to encounter some bad data. The logical choice then is to present data both ways, i.e., with and without Chauvenet's criterion applied. Of the L1 groups and batches analyzed the following is observed from Table 3-16:

Groups and Batches of Data Groups and Batches of Data With Smallest c/E Without With Smallest c/I With Correlation Chauvenet's Criterion Chauvenet's Criterion Baw-168 38 36 WAPD-188 2 3 W-3 1 2 Chauvenet's criterion rejected the following number of points for each corre-lation 01h1 m~ ' ,

WJT 3-48 s

Uniform Nonuniform Total BAW-168 (Groups Only) 32 1 33 BAW-168 (Batches only) 39 0 39 WAPD-188 (Groups Only) 3h 2 36 WAPD-188 (Batches Only) 33 0 33 W-3 (Groups only) 59 12 71 W-3 (Batches Only) 50 9 59 Several notable peculiarities exist in the tabulation of Table 3-16. The Col-umbia data 500 psia group contained only five data points; four were rejected by Chauvenet's criterion,' leaving one point. A standard deviation cannot be computed for one point; therefore all three values of (a/X)' are shown as not available (H.A.). Neither the BAW-168 nor the WAPD-188 predicted any negative DNB heat fluxes; the W-3 predicted 93 negative values for uniform data. The fact that only 59 were rejected for this correlation indicates that the re-maining 3h uniform points which were negative (93 - 59 = 3h) were close enough to the body of the data to be considered statistically significant. Table 3-16 may be consolidated somewhat as below by tabulating the number of groups and batches of data having coefficients of variation within a specified inter-val for each correlation.

(a/I)

Interval BAW-168 BAW-168'("} WAPD-188 WAPD-188'("} W-3 W-3'

• Negative 0 0 0 0 2 0 0-0.1 6 8 0 0 0 1 0.1-0.2 2h 2h 13 13 1 5 0.2-0.3 8 8 7 8 3 1 0.3-0.4 1 0 3 4 1 2 0.4-0.5 1- 0 5 7 5 6 0.5-0.6 0 0 6 5 3 h 0.6-0 7 0 0 3 2 1 1 0 7-0.8 0 0 2 1 7 8 0.8-0 9 1 0 0 -0 1 5 0 9-1.0 0 0 0 0 1 0

>1.0 0 0 2 0 16 7 Total kl 40 41 40 41 h0

" Chauvenet's criterion applied.

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As is seen from the tabulation the column for BAW-168 with Chauvenet's crite-rien applied indicates a grouping of 0,1 to 0.2, and a maximum value of

 -0.28780 is'noted from Table 3-16. For WAPD-188 the spread is greater with a maximum value of 0 7kO18. For W-3 the spread is still greater, and a maximum value of 1.Th83 is noted. The negative values of DNB heat flux predicted by
 .the W-3 correlation are in part responsible for the large spread in (c/i).
                         ~

The ability of the BAW-168 correlation to fit both uniform and nonuniform heat flux _ data over a wide range of pertinent variables leads us to believe that it is the best DNB correlation available. 3.2.3.2.h Evalustion of the Thermal and Hydraulic Design

a. Hot Channel Coolant Quality and Void Fraction An evaluation of the hot channel coolant conditions provides addi-tional 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 shown in Figure 3 h0. 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 1.85. Additional calculations for a 10 per cent increase in Fah to 2.035 were made at 114 per cent power. The significant results of both calculations are su=marized in Table 3-17. The effects of us-ing an Fah of 1 79 are shown in Figure 3 h0.

Table 3-17 Hot Channel Coolant  :.ditions Exit Exit Void Operating Power, 5 Fah Quality, % Fraction, % Pressure, psig 100 1.85 (-)2.h(D) 0.5(*) 2,185 llh 1.85 2.8 13 5 2,185 130 1.85 9.h 36.9 2,185 llh 2.035 8.7 2,185 100 1.85 0 35.0(*) 3.8 2,120 114 1.85 5.h 25.2 2,120 130 1.85 12.1 h5.2 2,120 11h 2.035 11.3 h3.h 2,120 (" Subcooled voids. Negative indication of quality denotes subcooling of 10.2 Btu /lb. / Th9 conditions' of Table 3-17 vere determined with all of the hot

1. channel factors applied. Additional calculations were made for unit 0154 3-51

p W y '

           - cell channels without engineering' hot channel factors to show the       ">

i

           ' coolant conditions more likely to occur in the reactor core. Values for Fah of,~1 79 and 1.85 were examined with and without fuel assem-          I
           -bly flow distribution hot channel factors at 2,185 psig as shown on Figure 3-41. These results show that the exit qualities from the hottest cells should in general be considerably lower than the maxi-mum design 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 minimum operating pressure of 2,120 psig. The influence of core fuel assem-bly flow distribution was checked by determining the total voids for both 100 and 95 per cent total core flow for the two' pressure condi-tions.

The results are as follows: Flow, % Pressure, psig Core Void Fraction, % 100 2,185 0.007 100 2,120 0.033

                       -95             2,185                   0.0h1 95             2,120                   0.127 The most-conservative _ condition of 95 per cent flow at 2,120 psig          )

results in no more than 0.13 per cent void volume in the core-Conservative. maximum design values for Fah nuclear described by Line A of Figure 3-12 were used to make the calculation. The void ogram uses a combination of Bowring's(33) model with Zuber's(3 correlation between void fraction and quality. The Bowring model considers three different regions of forced convec-tion boiling. They are: (1)'- Hichly Subcooled Boiling In this region the bubbles adhere to the wall while moving up-vard 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 fuel reaches the-surface temperature predicted by the Jens and Lottes equation. The highly subcooled region ends when T sat

                                                 -T bulk
                                                         =AV                      (A) where-
                               , local heat flux, Btu /hr-ft n =-1.863 -x 105 (1h + 0.0068p)

V = velocity of coolant, ft/see

                            -p =~ pressure, psia Qj,b,D
                                         .3-52                       -
                                                              .a y

i The void fraction in this region is computed in the same manner as Maurer,(35) except that the end of the region is determined by Equation (A) rather than by a vapor layer thickness. The nonequilibrium quality at the end of the region is computed from the void fraction as follows: 1 x*d = (B) y , E 1, y Pg a d where x" = nonequilibrium quality at end of Region 1 ad = v id fraction at T, -T = Pr = liquid component density, lb/ft 3 Pg = 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 point where the thermo-dynamic quality is zero. In general, this is the region of major concern in the design of pressurized water reactors. The nonequilibrium quality in this region is computed from the following formula: P z x" = x* + g ,g ($ - $gp)dz M d where x, = nonequilibrium quality in Region 2 h f

                               = latent heat of vaporization, Btu /lb 1

1** = fraction of the heat flux above the single phase heat flux that actually goes to producing voids

                          $gp = single phase heat h, Mub M s = mass flow rate, lb/hr P = heated perimeter, ft h

z = channel distance, ft n The void fraction in this region is computed from i

 /

0356 3-53 ',

J i 38.3 A P Fogg (P - P )'l/4 f Cg +P g f ~ + 6 p2

                                                                    .      f      .

wher g = acceleration due to gravity, ft/see

                                                                                      * #t g = constant in Newton's Second Law = 32.17 2

lb f see C = Zuber's distribution parameter A = flow area, ft f o = surface tension Equation (D) results from rearranging equations found in Ref-erence (34) 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-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 A flow regime map was constructed to evaluate channel hydraulic sta-bility. The transition from bubbly to annular flow at high mass ve-locities was determined using Baker's(36) correlation, and the tran-flow which occurs at low mass velocities

               -sition was'det'from  bubbly ermined withto slug (37) correlation. The transition'from Rose's slug flow to annular flow was determined by Haberstroh's(38) corre-lation. Bergles(39) found that these correlations, which were de-veloped from adiabatic data, are adequate for locating flow regime transitions with heat addition, and that they adequately preiict the effects of pressure. Figure 3-42 shows the flow regime map on which has'been plotted a point representing operating conditions in the hot channel at 114 per cent overpower. To aid in assessing the conserva-tism of the design, an additional point is plotted at 130 per cent overpower. Inspection shows that both points lie well within the
 .,            . bubbly flow regime. Since the bubbly flow regime is hytraulically stable, no flow instabilities should occur. This flow regime map was prepared for the hot unit cell at the maximum design condition
               > characteristics outlined in 3.2.3.1.1.

yL

The confidence in the design is based on both experimental results obtained in multiple rod bundle burnout tests and analytical evalua-tions. Three additional flow regime maps were constructed for nom-inal and postulated worst case conditions to show the sensitivity of the analysis with respect to mass flow rate, channel dimensions and mixing intensity in unit, corner, and wall-type cells. The results are shown in Figures 3 43, 3 kh, and 3 h5 The mass velocity and quality in each type of channel for the two cases are plotted on the figures. The conditions assumed for the nominal and postulated worst case are given in 3.2.3.2.h J. Data from the burnout tests performed by B&W on a 9-rod bundle simu-lating the core gecmetry are also plotted on the maps. The open data points on the maps represent the exit conditions in the various type channels just previous to the burnout condition for a representative sample of the data points obtained at the design operating pressure of 2,200 psia. In all of the bunale tests the pressure drop, flow rate, and rod temperature traces were steady and did not exhibit any of the characteristics associated with flow instability. Inspection of these maps shows that the nominal conditions are far removed from unstable flow regimes. The evaluation also shows that I under the worst conditions that have been postulated the reactor will be operating in the hydrodynamically stable, bubbly flow regime.

d. Hot Channel DNB Comparisons DNB ratios for the hottest channel have been deter =ined for the BAW-168 and W-3 correlations. The results are shown in Figure 3-46. DNB ratios for both correlations are shown for the 1.50 axial max / avg symmetrical cosine flux shape from 100 to 150 per cent power. The BAW-168 DNB ratio at the maximum design power of 114 per cent is 1.38; the corresponding W-3 value is 1 72. This compares with the suggested W-3 design value of 1.3. It is interesting to note that the calculat-ed DNB ratio reaches a value of 1.0 at about 150 per cent power with the BAW-168 equation which adequately describes DNB at the high qual-ity condition of 20 per cent. The W-3 calculation is accurate to about 130 per cent power, but because of quality 1Laitations it can-not be used to examine the channel at the 150 per cent power condi-tion.

The sensitivity of DNB ratio with FAh and F: nucler.r was examined from 100 to 114 per cent power. The detailed results are labeled in Figure 3-46. A cosine flux shape with an Fz of 1.80 and an Fah of 1.85 results in a W-3 DNB ratio of 1.45 and a BAW-168 ratio of 1.33. The W-3 value is well above suggested design values, and the BAW-168 value of 1.33 corresponds to a hot channel confidence of 99 per cent that about 93 per cent of the population is in no jeopardy as shown in the Population-DNB ratio plot in 3.2.3.2.2, Statistical Core Design Technique. The influence of a change in FAh was determined by analyzing the hot channel for an FAh of 2.035 This value is 14 per cent above the maximum calculated value of 1.79 and 10 per cent above the maximum 3-55 '

a a design value of 1.85 The resulting BAW-168 DNB ratio is 1.22 and ) l the W-3 value is 1.26. Both of these values are well above the cor- ' relation best-fit values of 1.0 for the severe conditions assumed,

e. Reactor Flow Effects Another significant variable to be considered in the evaluation of the design is the total system flow. Conservative values for system and reactor pressure drop-have been determined to insure that the

j. required system flow is obtained in the as-built plant. The experi-mental programs previously outlined in Section 1 will confirm the pressure 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.

     -             An evaluation of reactor core flow and power capability was made by determining the maximum steady state power rating versus flow. The analysis was made by evaluating the hot channel at the overpower con-ditions while maintaining (a) a DNB ratio of 1.38 (BAW-168), and (b) the statistical core design criteria. The results of the analysis are shown in Figure 3 h7 The power shown is the 100 per cent rat-ing, and the limiting condition is llh per cent of the rated power.

An examination of the slope of the curve indicates stable character-istics, and a 1 per cent change in flow changes the power capability by only about 1/2 per cent. I Reactor Inlet Temperature Effects / f. The influence of reactor inlet temperature on power espability at a given flow was evaluated in a similar manner. A variation of 1 F in reactor inlet temperature will result in a power capability change of slightly less than 1/2 per cent,

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

(1) Outer limit of equiaxed grain growth - 2,700 F. (2) Outer limit of columnar grain growth - 3,200 F (3) Outer limit of molten fuel (UO2) - 5,000 F. l Data from References h0 through h3 were used to compare calculated and. experimental fractions.of the rod in grain growth and central melting. The radial' expansion .of the fuel pellet is computed from the mean

                     . fuel temperature and the average coefficient of linear expansion for
       . s< .

e

                       ~

the fuel over the temperature range considered. This model combined with the model for calculating the heat transfer coefficient was ccm-pared with the model developed by Notley et al(kk) of AECL. The dif-ference 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 in-crements for the analysis. An iterative solution for the temperature 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 temperature drop across the gap between the fuel and cladding' surfaces. The temperature drop across the gap is a function of vidth, mean tem-perature, and gas conductivity. The conductivity of the gas in the gap is deter =ined as a function of burnup and subsequent release of fission product gases. In the event of fuel clad contact, contact coefficieng8).are and Stoute determined The contacton the basis ofis coefficient methods suggested determined by Ross as a function of the mean conductivity of the interface materials, the contact pressure, the mean surface roughness, the material hardness, and the conductivity of the gas in the gap. The analytical model computes the amount of central void expected whenever the temperature approaches the threshold temperature for fuel migration, and readjusts the density according to the new ge-cmetry. The program uses a polynominal fit relationship for fuel thermal con-ductivity. Three relationships were used to evaluate the effects of conductivity. A comparison of these conductivity relationships with the reference design CV7 The values suggested in GEAP-h62h( D 4-)lh2(h5) is shown in Figure 3 h8.and CV 3,000 F, and the former values are more conservative above 3,000 F. McGraththT) concludes that the CVNA-2h6 values are lower limits for the high temperature conditions. Fuel center temperatures for all three of the conductivity relationships at the peaking factors given in 3.2.3.1.2 have been calculated to evaluate the margin to central melting at the maximum overpower and to show the sensitivity of the calculation with respect to thermal conductivity. Since the power peaks will be burned off with irradiation, the peaking factors used are conservative at end-of-life. Results The results of the analysis with the methods described above are shown in Figures 3 h9 and 3-50 for beginning and end-of-life condi-tions. The beginning and end-of-life gas conductivity values are 0.1 and 0.01 Btu /hr-ft 2 -F respectively. The calculated end-of-life center fuel temperatures are higher than the beginning-of-life val-ues because of the reduction in the conductivity of the gas in the gap. The effect is apparent even though a contact condition pre-vails. The calculation does not include the effects of fuel swell-

    ,         ing due to irradiation. -The calculated contact pressures are con-g   ,

servatively lower than those expected at end-of-life conditions in

                                                      .t 0160 v J+Mf"                         3                                                              -

1

the hottest fuel rods, and the fuel temperatures shown in the above figures are conservatively higher. } The B&W model gives very good results when ec= pared to the results of

                                                                   ~

others in the field as is shown in Figure 3-50. In the linear heat range of most interest, i.e., approximately 20 kv/ft, there is only about 300 F difference between the maximum and minimum values calcu-lated. Also the small differences between the'B&W curve and the other curves' indicate the relative insensitivity of the results to the shape of the conductivity at the elevated temperatures. The most conservative assumptions, using GEAP-h62h d'ata with relatively little increase in thermal conductivity above 3,000 F, result in cen-tral fuel melting at about 22 kv/ft, which is 2 kv/ft higher than the maximum design value of 19 9 kv/ft at llh per cent power. Further evaluation of the two figures shows that central fuel melting is pre-dicted to occur between 22 and 26 kv/ft depending on the time-in-life and conductivity assumptions. The transient analyses at accident and normal conditions have been made using the GEAP k62h fuel thermal conductivity curve to reflect a conservative value for the maximum average temperature and stored energy in the fuel. Use of this curve results in a higher tempera-ture and therefore a lower Doppler coefficient, since it decreases with temperature. Thus the resultant Doppler effect is also con-servative.

h. Fission Gas Release The fission gas release is bas d on results re AdditionaldatafromGEAP-h31450),AECL-603(5$ortedinGEAPh596(k9

                                                           ),andCF-60-12-lhl52) with the suggested release rate curve. The re-havebeenecmparg9)isrepresentativeoftheupperlimitofrelease lease rate curve data in the temperature region of most importance. A design release rate of h3 per cent and an internal gas pressure of 3,300 psi are used to determine the fuel clad internal design conditions reported in 3.2.k.2, Fuel Assemblies.

The design values for fission gas release from the fuel and for the maximum clad internal pressure were determined by analyzing various operating conditions and assigning suitable margins for possible in-creases in local or average burnup in the fuel. Adequate margins are provided without utilizing the initial porosity voids present ir. the UO2 fuel. A detailed analysis of the design assumptions for fission gas release, and 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 included. W 3-58 , , (3 3.43 l.

(1) Design Assumptions (a) Fission Gas Release Rates The fission gas release rate is calculated as a function of fuel temperature at the design overpower of lik per cent. The procedures for calculating fuel temperatures are dis-cussed in 3.2.3.2.4 g. The fission gas release curve and the supporting data are shown in Figure 3-51. Most of the data is on or below the design release rate curve. A re-lease rate of 51 per cent is used for the portion of the fuel above 3,500 F. The fuel temperatures were calculated using the OEAP-462h fuel thermal conductivity curve to ob-tain conservatively high values for fuel temperatures. (b) Axial Power and Burnup 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 max-imum release rates. These are 1.50 and 1.70 max /avs shapes as shown in Figure 3-11 and repeated as part of Figure 3-52 of this analysis. The quantity of gas released is found by applying the temperature-related release rates to the quan-tities 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 burnup studies. Three curves showing the axial distribution of burnup as a local to average ratio along the fuel rod are shown in Figure 3-52. Values of 100, 300, and 930 days of operation are shown. The 930-day, or end-of-life condition, is the condition with the maximum fission gas inventory. The average burn-up at the end of life in the hot fuel rod is 38,150 MWD /MTU which has been determined as follows: Calculated Hot Bundle Average Burnup, MWD /MTU 33,000 Hot Fuel Rod Burnup Factor 1.05 Margin for Calculation Accuracy 1.10 Hot Rod Maximum Average Burnup, MWD /MTU 38,150 The local burnup along the length of the fuel rod is the product of the hot rod maximum average value above and the local to average ratio shown in Figure 3-52. The resulting hot rod local maximum burnup for the 930-day, end-of-life condition is about 42,000 MWD /MTU. This is the maximum ~' ' calculated value. However, local values to 55,000 MWD /MTU O have been evaluated to insure adequate local fuel cladding g 3, 0162 3-5F

strength for possible increases in average or local burnup over the life of the fuel for various fuel management pro- '.} cedures. (c) Hot Rod Power Assumptions The maximum hot rod total power occurring at any time in the life of the fuel has been used to calculate the overpower temperature conditions. A hot rod power of 1.85 times the average rod power has been applied. This results in a max-imum linear heat rate of 19 9 kv/ft which corresponds to 114 per cent of the maximum thermal output (17.49 kv/ft) shown in Table 3-1. This is a conservative assumption when cou-pled with the end-of-life fission gas inventory since bundle and individual fuel rod power is expected to decrease with fuel burnup. A study of the power histories of all of the fuel assemblies to equilibrium conditions shows that the powers in the bundles during the last 300 days of operation are not more than 1.3 times the average bundle power. The peak bundle ratio of 1.69 (1.85 + hot rod ratio) vill only occur during the first two fuel cycles when the fission gas inventory is less than the maximum value. (d) Fuel Growth Assumptions l The fuel growth was calculated as a function of 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 Assumptions The quantity of fission gas released is a function of fuel l ' temperature. The temperatures are influenced by three factors: (a) the conductivity of the fission gas in the gap between the fuel and clad, (b) the diametral clearance between fuel and clad, and (c) the heat transfer conditions l vhen the fuel expands enough to contact the clad. 2 A gas conductivity of 0.01 Btu /hr-ft -F based on h3 per cent release of fission gas at the end-of-life condition was used l in the analysis. Diametral clearances of 0.0025 to 0.0075 in, reflecting minimum and maximum clearances after fuel growth were analyzed. The contact heat transfer coefficients were calculated as suggested in Reference h8. (2) Summary of'Results

The fission gas release rates were determined in the-first eval-~

uation. Rates were founu for various cold diametral clearances and axial power peaking and burnup shapes. The results are shown in Figure 3-53. The lowest curve is the expected condition for a 1.70 axihl power shape with a 930-day e.xial burnup distribution as shown in Figure 3-52. The increase in release rate with r, 01.63 g 3-60 (..

diametral clearance results from the fact that the fuel temper-ature must be raised to higher values before contact with the fuel clad is made. The release rate at the minimum clearance of 0.0025 in. is 19 per cent. This is the condition that pro-duces the maximum clad stress due to fuel growth with irradia-tion. The assembly of maximum size pellets with minimum inter-nal diameter cladding will produce this condition after fuel growth. In the event a few hot pellets have the maximum diam-eter and the remainder have the minimum diameter, then the aver-age cold gap would be 0.0035 in. producing a slightly larger release rate. The release rate of 33 per cent for the maximum diametral clearance will not occur with the maximum stress con-dition due to fuel growth, since the fuel can grow into the clearance. Two additional cases were examined to check the sensitivity of the calculations to axial power and burnup shapes. The results are shown by the upper two curves in Figure 3-53. 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 1.70 max / avg ratio as shown on Figure 3-52. Similar results are shown for the 1 50 max / avg ratio. These curves show the release rates expected are not strongly influ-enced 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 pressures for the expected 930-day-axial burnup distribution and a 170 max / avg axial power shape are shown in Figure 3-54 The lower curve is a plot of internal gas pressure with open pores (5 per cent of the fuel volume is available to hold the released gas). The upper data band is for a closed pore condition with all re-leased gas contained outside the fuel pellets in spaces between the expanded dished ends of the pellets, the radial gaps (if any), and the void spaces at the ends of the fuel rods. The band of data shown reflects the effect of fuel densification and grain growth described in 3.2.3.2.k. The upper limit is for an ideal thermal model without, grain growth or densifica-tion; the lower limits are for the design model. The calcula-tion of the maximum pressure is also relatively insensitive to the axial burnup distribution as shown by the dashed line in Figure 3-54 for a 1 50 maximum to average axial power and burn-up shape. (This corresponds to a local burnup peak of 57,000 MWD /MEU.) 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 diametral clearance condi-tion. A' modest increase in average fuel burnup can be toler-ated within the prescribed internal pressure design limits.

 < '   .It has been indicated in Reference kh and in AECL-1598 that the
     ,   UO 2 fuel is plastic enough to flow under lo stresses when the 3- 4                           0164

temperature is above 1,800 F. That. fraction of the fuel below this temperature may retain a large portion of the original po-rosity 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 this temperature. The approximate frac-tion of the fuel below 1,800 F at overpower for a 1.70 axial power shape is as follows for various cold diametral clearances. Clearance, Per Cent of Fuel in, below 1,800 F, % 0.0025 h0 0.005 20 0.0075 5 The retention of fuel porosity in the low temperature and low burnup regions will result in modest reductions in internal gas pressure.

1. Hot Channel Factors Evaluation (1) Rod Pitch and Bowing A flow area reduction factor is determined for the as-built fuel T assembly by taking channel flow area measurements and statisti- 1 cally determining an equivalent hot channel flow area reduction factor. A fuel assembly has been measured with the results l shown in Table 3-12. In the analytical solution for a channel flow, each channel flow area is reduced over its entire length by the FA factor shown in Figure 3-21 for 99 per cent confi-dence. With a 99 per cent confidence and 94.5 per cent popula-.

tion relationship described in 3.2.3.1.1 for the hot channel, the area reduction factor is 0.992. The approximate limit of this factor is obtained by examining the value in Figure 3-21 i ' as the population protected approaches 100 per cent. FA at 99.99 per cent of the population protected is 0.983 The hot channel value is shown in Table 3-1. Special attention is given to the influence of water gap varia- , tion between fuel assemblies when determining rod powers. Nu-clear analyses have been made for the nominal and maximum spac-ing between adjacent fuel assemblies. The nominal and maximum hot assembly fuel rod powers are shown in Figures 3-55 and 3-56 respectively. The hot channel nuclear power factor (Fah nuclear) of 1.85 shown in 3 2.3.1.1 is based on Figure 3-56 for the maxi-mum water gap between fuel assemblies. The factor of 1.85 is a product of the hot assembly factor of 1.69 times the 1.096 hot rod factor. This power factor is assigned to the hottest fuel rod which is analyzed for burnout.under unit cell, wall cell, and corner cell flow conditions.

                                                   ~

W ;3-62 0165 e

(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 3-12. These variations have been obtained from the mea-sured or specified tolerances and ccabined statistically as de-scribed in 3.2.3.2.2 to give a power factor on the hot rod. For the hot channel confidence and population conditions, this factor, Fq, is 1.008 and is applied as a power increase over the full length of the hot fuel rod. The local heat flux fac-tor, Fq", for 99 per cent confidence and 94.5 per cent popula-tion is 1.013. These hot channel values are shown in Table 3-1. The corresponding values of Fq and Fqn with 99.99 per cent pop-ulation protected are 1.017 and 1.03 respectively. A conserva-tive value of FQ " of 1.03 for 99 per cent confidence and 99.99 per cent population is used for finding the maximum fuel line:ar 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 often done. The coefficients of variaticn will be under contin-uous review during the final design and development of the fuel assembly. (3) Flow Distribution Effects Inlet Plenum Effects The final inlet plenum effects will be determined from the 1/6 scale model flow test now in progress. The initial runs indi-cate satisfactory flow distribution. Although the final nuclear analysis and flow test data may show that the hot bundle posi-tions receive average or better flow, it has been assumed that the flow in the hot bundle position is 5 per cent less than aver-age bundle flow under isothermal conditions corresponding to the model flow test conditions. An additional reduction of flow due to hot assembly power is described below. Redistribution in Adjacent Channels of Dicsimilar Coolant 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 pressure drop associated with a higher enthalpy and qual-ity 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 to the hot assemblies will result in a hot assembly flow of about 85 to 95 per cent of the average assembly flow depending on the final ple-num effects and assembly power peaks. The worst combination of effects has been assumed in the initial design, and the hot as-sembly flow has been calculated to be about 85 per cent of the average assembly flow at llh per cent overpower. Actual hot U - 0166 3-63

z ._ l assembly fisvs are calculated rather than applying an equivalent ~ hot ~channe" 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, wall, and corner cells as shown by the heavy lines in Figure 3-55 The mixed enthalpy for every cell is determined sLmultaneously so that the ratio of cell to average assembly enthalpy rise (Enthalpy Rise Factor) and the corresponding local enthalpy are obtained for each cell. Typi-cal enthalpy rise factors are shown in Figures 3-55 and 3-56 for cells surrounding the hottest fuel rod located in the cor-ner of the assembly. The assumptions used to describe the chan-nels for the peaking and enthalpy rise factors shown are given in Wall and Corner Channels Evaluation, 3.2.3.2.h j, which fol-lows.

j. Evaluation of the DNB Ratios in the Unit, Wall, and Corner Cells DNB Results The DNB ratios in the hot unit cell at the maximum design condition i described in 3.2.3.1 are shown in Figure 3 b6. The relationships shown data inare thebased on the a pplication of single channel heat transfer BAW-168(18) and W-3(23,68) correlations. An additional sensitivity analysis of the assembly has been made utilizing 9-rod assembly heat transfer DNB test data that is more representative of the actual vall and corner cells geometry effects than single chan-nel data.

The sensitivity of the assembly design with respect to variations of mass flo" rate (G), channel spacing, mixing intensity, and local peaki- 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 3-18. g OMA

                                                                    ^

3-6h .,

Table 3-18 DNB Patios in the Fuel Assembly Channels Nominal Case Cell Tyre G, lb/hr-ft x 10' DNBR Corner 1.59 2.20 Wall 1 90 2.11 Unit 2.52 2.01 Postulated Worst Case

                                                     ~

Cell Tyne G, lb/hr-ft x 10 DNBR Corner 1.32 1.70 Wall 1.6h 1.65 Unit 2.29 1 73 The DNBR's above'are ratios of the limiting heat flux to the 1ccal flux along the length of the channels. The limiting heat fluxes have been determined from the 9-rod assembly DNB test data. The DNB ratios in all channels are high enough to insure a confidence-population relationship equal to or better than that outlined in 3.2.3.1.1 for the hot anit cell channel. The postulated worst case conditions are more severe than the required maximum design condi-tions. The results of th,e assembly tests and this evaluation show that the performance of the vall and corner cells is more sensitive to local enthalpy than to the local mass velocities. Although the mass flow rates in the corner and vall cells are lower than in the unit cell,

      -the total flow in these cells is relatively higher than the mass flow rates imply because of the increased space between the outer rods and the perforated can. This results in more favorable power-to-flow ratios than the mass flow rates indicate.

The DNB ratios were obtained by comparing the local heat fluxes and coolant conditions with heat transfer data points from 9-rod fuel assembly heat transfer tests for uniform heat flux with an appropri-ate correction for a nonuniform axial power shape. Typical results are shown in Figures 3-57 and 3-58 for the nominal and worst case conditions in the corner cell. The line defined by a best fit of the data is shown on each figure as a solid line. A design limit

     'line, shown as dotted, has been determined by lowering the best-fit line to account for the effects of nonuniform flux shapes. The mag -

nitude of the reduction was determined by comparison with the results of the Euratom nonuniform test data (19) and the results of more recent

   ,   nonuniform tests conducted by B&W.

y q. 0168

                                   .3-65

The limiting best-fit lines were derived from a 9-rod fuel assembly e 's test section 72 in. long with rod diameter, pitch spacing, and spacer ) grids of the type to be used in the reference design. A total of 513 data points between 1,000 psi and.2,h50 psi has been obtained. One

                                                     ~

hundred and sixty-two of these points were used for the limiting lines in the PWR pressure and mass flow ranget,. The ranges of test variables for the 162 data points used_ vere: Pressure - 1,800 to 2,450 psi 6 2 Mass Flow Rate - 1.0 to 3.5 x 10 lb/hr-ft Quality - -5 to +20 per cent All of the cell conditions of interest in this analysis fall within this range of parameters. Fuel Rod Power Peaks and Cell Coolant Conditions The nominal case local-to-average rod powers and the local-to-average exit enthalpy rise ratios are shown in Figure 3-55 for the hot cor-ner, hot vall, and hot unit cells in the hot fuel assembly. Values shown are for nominal water gaps between the hot fuel assembly and adjacent fuel assemblies with nominal rod-to-vall spacing, with nom-inal flow to the hot fuel assembly, and with a nominal intensity of turbulence, a,(*) equal to 0.03 Additional tests are being run to determine the maximum values of in-3 tensity of turbulence associated with the fuel assembly. The expect- j ed value is greater than 0.03 since this value is obtained in smooth tubes, and the spacers and can panel perforations should induce more turbulence. The postulated worst case local-to-average rod powers and exit en-thalpy rise ratios in the hot fuel assembly are shown in Figure 3-56. The factors were determined for this case with twice the nom-inal water gaps between the hot fuel assembly and adjacent fuel as-semblies with minimum rod-to-vall spacing, with minimum flow to the hot fuel assembly, and with a minimum assumed intensity of turbu-lence, a, equal to 0.01. I In neither the nominal nor the postulated worst case analysis has any credit been taken for the coolant which is flowing in the water (*)The intensity of turbulence, a, is defined as V' /V whereV[isthetransversecomponentoftheiluctuatingturbulentvelocity, and V is the coolant velocity in the axial direction. Thh method of ecm-puting mixing is described by Sandberg, R. O., and Bishop, A. A., CVTR Thermal-Hydraulic Design for 65 MW Gross Fission Power, CVNA-227_. I

• 3-66 .

                         ~

gaps between the fuel assemblies and which serves to reduce enthal-pies in the peripheral cells of the hot fuel assembly by mixing with the coolsnt in those cells through the can panel perforations. In both cases, hcwever, the effective roughness of the can panel perfo-rations and its effect on reducing the flow in the peripheral cells of the fuel assembly has been accounted for. The magnitude of the effective roughness was obtained from the results of a series of flow tests performed on a mockup of the outer two rows of fuel rods and the can panels of two adjacent fuel assemblies. The rod-to-vall spacing in the peripheral cells of the fuel assembly has been increased to compensate for the-effects of the can panel in reduc-ing the flow in the peripheral cells. The nominal distance from the center of the outside rods to the can panel is 0.324 in. The corresponding postulated worst case dimension was assumed to be 0.310 in. Fuel Assembly Power and Flow Conditions The nominal and postulated worst cases were run at 114 per cent re-actor power with the nominal and worst Fah factors shown in 3.2.3.1.1 c. The 1.50 modified cosine axial power shape of Figure 3-11 was used to describe the worst axial conditicn. The hot assembly flow under nominal conditions without a flow mal-distribution effect is 93 per cent of the average assembly flow, and the reduction in flow is due entirely to heat input effects. The hot assembly flow under the worst postulated conditions is 85

      .per cent of the average assembly flow and considers the worst com-bined effects of heat input and flow maldistribution.

Summary Analysis of all B&W bundle data to date indicates that the B&W meth-od cill correlate data with less deviation than previous methods. Indications are that this is also true when considering nonuniform axial power distributions. Additional bundle tests will be conduct-ed with nonuniform axial power distribution to confirm that the use of a power shape correction factor based on single channel and an-nular specimens is conservative. Completion of the test programs outlined in this report and evalua-tion of the experimental data vill provide final design correlations and flow relationships that will give complete confidence in the con-servatism of the design and the B&W analytical procedures. It.should be noted that the postulated worst case is worse than the hot channel permitted by our specifications. Even with this postu-lated worst case, the design is still conservative, and there is very little difference in the performance of the various channels. This indicates that the outside cell geometries have been compen-sated correctly to account for wall effects. t n .gs. , ono-

   -i                                                  

3-67

3 .'2 .4 L 18CHANICAL.DESIG9 LAYOUT 3 2.4.1 . Internal Layout Reat. tor internal componente include the plenum assembly and the core support acsembly (consisting of the c e a support shield, vent valves, core barrel, lower grid flow distributor, incore instrument guide tubes, thermal shield, and sur-veillance holder tubes). Figure 3-59 shows the reactor vessel, reactor vessel internals arrangement, and the reactor coolant flow path. Figure 3-60 shows a cross section throu8h the reactor vessel, and Figure 3-61 shows the core flood-ing crrangement. Reactor internal components do not include fuel assemblies, control rod assem-

  ~blies (CRA's), surveillance specimen assemblies, or incore instrumentation. Fuel
  -assemblies are~ described in 3 2.4.2, control rod assemblies and drives in 3 2.4 3,
                      ~
  . surveillance specimen assemblies in 4.4 3, and incore instrumentation in 7 3 3 The reactor internals are designed to support the core, maintain fuel assembly alignment, limit fuel assembly movement, and maintain CRA guide tube alignment between fuel assemblies.and control rod drives. They also direct the flow of reactor coolant, provide gnmm9 and neutron shielding, provide guides for incore instrumentation between the reactor vessel lower head and the fuel assemblies, support the surveillance specimen assemblies in the annulus between the themal shield and the reactor vessel vall, and support the internals vent valves.

These vent. valves are provided' to relieve pressure generated by steaming in the core following a reactor coolant inlet pipe rupture so that the core vill remain sufficiently covered'with coolant. All reactor internal components can , be removed from the reactor vessel to allov inspection of the reactor inter- I. nals and the reactor vessel internal surface. A shop fitup and checkout of all internal components in an as-built reactor ves-sel mockup vill insure proper alignment of mating parts before shipnent. Dummy , fuel assemblies and control rod assemblies vill be used to check fuel assembly clearances and CRA free movement. In anticipation of lateral deflection of the lower end of the core support assembly as a result of horizontal seismic loadings, integral veld-attached, deflection-l biting spacer blocks have been placed on the reactor vessel inside vall. .In addition, these blocks limit the rotation of the lower end of the core support assembly which could conceivably result from flow-induced torsion-al loadings. The blocks allow free vertical movement of the lower end of the internals for thermal expansion throughout all ranges of reactor operating con-ditions, but in the unlikely event of a flange, circumferential veld, or bolted jointfailuretheblocksvilllimitthepossiblecoredropto1/2in.orless. The final elevation plane of these blocks will be established near the same elevation as the-vessel support skirt attachment to minimize dynamic loading effects on the vessel shell or bottom head. Preliminary calculations indicate

                                      ~

the. impact loading on the stop bloch for a 1/4 in.' core drop would be approx-imately 5 g's total. Block location and geometry will be evaluated and de-

 .termined to transfer this loading through the vessel support skirt to the reactor building conert.te. A significant reduction in impact loading can be achieved through proper c N block design and detailed analysis. A1/2in.

3-68 REVISED, 2-8-68

                   ,r g                                                 Oiqo-L

core drop vill not allow the lover end of the CBA neutron absorber rods to disengage from their respective fuel assembly guide tubes if the CRA's are in the full-out position, since approximately 6-1/2 in of rod length would remain in the fuel assembly guide tubes. Acoredropof1/2in.villnotresultina significant reactivity change. The core cannot rotate and bind the drive lines because rotation of the core support assembly is prevented by the stop blocks. The failure of the core support shield and core barrel upper flanges, or related flanges and other trcumferential joints, is not considered credible on the basis of the conse cative design criteria and large safety factors employed in the internals des 1 6n. The final internals design vill be capable of withstand-ing various combinations of forces and loadings resulting from the static ve16 ht of internals (225,000 lb total, not including the plenum assembly which weighs 100,000 lb), core with control rod drive line (303,000 lb total), dynsmic load from trip (10 g's gives 207,000 lb), seismic (0.10 g vertical gives 53,000 lb), coolant flow hydraulic loading (230,000 lb), and other related loadings. The al Sebraic sum of thf s 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 com-ponent weights, seismic analysis, dynamic loadings from flow-induced vibration, detailed stress analysis with consideration for thermal stress during all transients, and resolution of fabrication details such as shell rolling toler-ances and weld joint preparation details vill increase the stress levels listed above. As a final design criterion, the core support :omponents vill meet the stress requirements of the ASME Code, Section III, during normal operation and transients. The structural integrity of all core support circumferential veld joints in the internals shells vill be insured by compliance with the radio-graphic inspection requirements in the code above. The seismic analysis vill include detailed calculations to determine the mw % structural response of the reactor vessel and internals. This analysis will be perfomed as described in 3 1.2.4.1. In the event of a major loss-of-coolant accident, such as a 36-in. diameter reactor coolant pipe break near the reactor vessel outlet, the fuel assembly and vessel internals would be subjected to dynamic loadings resulting from an oscillating (approximately sinusoidal) differential pressure across the core. A preliminary analysis of this postulated accident indicates that the fuel assemblieswouldmoveupwardlessthan3/8in. Some deflection of the inter-nals structures vould 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 components. 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 vould remain relatively unaffected as transient pressure oscillations are dampened out in approximately 0 5 sec. On this basis, the CRA travel time to 2/3insertiononatripcommandwillbeapproximately155seeinsteadofthe specified 1.40 sec. Also, this possible initial minor delay in trip initiation lM "4 ~ 3-69 REVISED, 2-8-68 TTE . m os.an.

vould not contribute to the severity of the loss-of-coolant accident because at the initiation of CRA trip, the core would be suberitical from voids. Material for the reactor internals bolting vill be subjected to rigid quality control requirements to insure structural integrity. The bolts will be dye-penetrant inspected for surface flaw indications after all fabrication opera-

 ' tions have been completed. Torque values vill be specified for the final as-sembly to develop full-bolting capability. All fasteners will be lock-velded to insure assembly integrity.

3:

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L s I 3-69a asyIszD, 2-8-68 p +M osa 1 c

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3.2.4.1.1 1 Plenum Assembly The plenum assembly is located directly above the reactor core and is removed 3 cs 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 a series of parallel flat plates intersecting to form square lattices with a perforated top plate and flange, and is attached to the plenum cylinder top flange. Three lifting lugs are pro-vided for the plenum assembly handling. 7he CRA guide tubes are welded to the plenum cover top plate and bolted to the upper grid. CRA guide assemblies provide CRA guidance and protect the CRA from the effects of coolant cross-flow, and provide structural attachment of the grid assembly 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 segments which are properly

oriented and attached to a series of castings to provide continuous guidance for the CRA full stroke travel. Design clearances in the guide tube will accom-modate some degree of misalignment between the CRA guide tubes and the fuel assemblies. Final design clearances will be stablished by tolerance studies and by the results of the Control Rod Drive Line Facility (CRDL) prototype tests.

, Prelimina7 test results are described in 3.2.4.3 5. The upper grid assembly consists of parallel flat bars intersecting to form square lattices. - The bars are attached to a flange which is bolted to the plenum cylinder lower flange. The upper grid assembly locates the lower end of the individual CRA guide tube assembly relative to the upper end of the corresponding fuel assembly. Iocating keyways in the plenum assembly cover flange engage the reactor vessel top flange locating keys to align the plenum assembly with the reactor vessel, ' rcactor closure head control rod drive penetrations, and the core support assembly. 7he 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, incore instrument guide tubes, surveillance specimen holder tubes, and internals vent valves.

Static loads from the assembled components and fuel assemblies, and dynamic loads from CRA trip, hydraulic flow, thermal expansion, seismic disturbances, and loss-of-coolant accident considerations, are all carried by the core sup-port assembly. The core support assembly components are described as follows:

a. Core Support Shield 7he . core support shield is a large flanged cylinder which mates with the reactor vessel opening. The top flange rests on a circumferential ledge in the reactor vessel top closure flange. The core support 3-70

() D .1 W J 5-3-68 Smf@mrs9Jhm_3

shield lower flange is bolted to the core barrel. The cylinder wall 3 has two nozzle openings for reactor coolant outlet flow. Se inside surface of the lower flu.ge guides and aligns the plenum assembly relative to the core support shield. The core support shield outlet nozzles are sealed to the reactor ves-sel outlet nozzles by the differential thermal expansion between the stainless steel core support shield and the carbon steel reactor ves-sel. The nozzle seal surfaces are finished and fitted to a predeter-mined cold gap providing clearance during core support assembly instal-lation and removal. At reactor operating temperature the mating metal surfaces are in contact to make a seal without exceeding allowable stresses in either the reactor vessel or internals. Internals vent valves are installed in the core support shield cylinder wall to re-lieve the pressure generated by steaming in the core following a pos-tulated cold leg (reactor coolant inlet) pipe rupture (see 3 2.4.1;.

b. Core Barrel The core barrel supports the fuel assemblies, lower grid, flow dis-tributor, and incore instrument guide tubes. The core barrel consists of-a flanged cylinder, a series of internal horizontal spacers bolted to the cylinder, and a series of vertical plates bolted to the inner surfaces of the horizontal spacers to form an inner wall enclosing the fuel assemblies. Construction of the core barrel will be similar to that of the reactor internals component developed by B&W for the Indian Point Station Unit No.1.

Coolant flow is downward along the outside of the core barrel cylin-der and upward through the fuel assemblies contained in the core bar-rel. A mall portion of the coolant flows upward through the space between the core barrel outer cylinder and the inner plate wall. Coolant pressure in this space is maintained slightly lower than the core coolant pressure to avoid tension loads on the bolts attaching the plates to the horizontal spacers. The vertical plate inner wall will be carefully fitted together to reduce reactor coolant leakage to an acceptable rate. The upper flange of the core barrel cylinder is bolted to the mating lower flange of the core support shield assembly, and the icwer flange is bolted to the mating flange of the lower grid assembly. All bolts will be inspected and installed as described in 3 2.4.1, and will be lock-welded after final assembly. Lifting lugs attached to the core

       - barrel are provided for core barrel and core support assembly handling.

c. Iower Grid Assembly The lower grid assembly provides alignment and support for the fuel assemblies, supports the thermal shield and flow distributor, and aligns the incore instrument guide tubes with the fuel assembly in-strument tubes._ The lower grid consists of two flat plate and bar lattice structures separated by short tubular columns surrounded by e-

       - a flanged cylinder. -The top flange is bolted to the lower flange of the core barrel. A perforated flat plate located midway between the two' lattice structures aids in distributing coolant flow.                     '

1: Wr ' 3-71 om 5-3-68 .-

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Supplement No. 3

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d. Flow Distributor .l3

          !s         The flow distributor is a perforated, dished head with an external         1 flange which is bolted to the bottom flange of the lower grid. The M                flow distributor supports the incore instrument guide tubes and dis-tributes the reactor coolant entering the bottom of the core.
    ~

e. Thermal Shield A cylindrical stainless steel thermal shield is installed in the an-nulus between the core barrel cylinder and the reactor vessel inner wall. The thermal shield reduces the neutron and gamma internal heat generation in the reactor vessel wall and thereby reduces the result-ing thermal stresses.

The thermal shield is supported on, positioned by, and attached to the lower grid top flange. The thermal shield upper end is positioned by spacers between the themal shield and the core barrel outer cylin-der to minimize the possibility of thermal shield vibration. 'Ihe thermal shield attachment is designed to avoid shear loads on fasten-ers. All fasteners are lock-welded after final assembly.

f. Surveillance Specimen Holder Tubes Surveillance specimen holder tubes are installed on the core support assembly outer wall to contain the surveillance specimen assemblies.

The tubes extend from the top flange of the core support shield to the lower end of the thermal shield. 'lhe tubes will be rigidly at-tached to prevent flow-induced vibration. Slip joints at the inter- 3 mediate supports and top end of the assemblies permit axial motion for differential thermal expansion,

g. Incore Instrument Guide Tube Assembly The incore instrument guide tube assemblies guide the incore instru-ment assemblies between the instrument penetrations in the reactor vessel bt,ct,om head and the instrument tubes in the fuel assemblies.

Minor horizontal misalignment clearance between the reactor vessel 3 instrument penetrations and the instrument guide tubes assembled with the flow distributor is provided. A perforated shroud tube, concentric with the instrument guide tube, adds rigidity to the assembly and reduces the effect of coolant flow forces. Fifty-two incore instrument guide tubes are provided. The incore instrument guide tubes are designed so they will not be affected by the core drop described in 3 2.4.1.

h. Internals Vent Valves Internals vent valves are installed in the core support shield to prevent a pressure unbalance which might interfere with core cooling following a loss-of-coolant accident. In its natural state and under all normal operating conditions, the vent valve will be closed. In the event of a loss-of-coolant accident in the cold leg of the reactor N 3-72 01.% 5-3-68 i

Supplement No. 3

loop, the valve will open to permit steam generated in the core to flow directly to the leak and will prevent the core from becoming morethan1/2uncoveredafteremergencycorecoolanthasbeensup-plied to the reactor vessel. The pre 14minn7 design of the internals vent valve is shown in Figure 3-61a. Each valve assembly consicts of a hinged disc, valve body with seal-ing surfaces, split-retaining ring, and fasteners. Each valve assem-bly is installed into a machined mounting ring, integrally welded in the core support shield wall. The mounting ring contains the neces-sary features to retain and seal the perimeter of the valve assembly. Also, the mounting ring includes an n14gnment device to maintain the correct orientation of the valve assembly for hinged-disc operation. Each valve assembly will be remotely handled as a unit for removal or installation. Valve component parts, including the disc, will be of captured-design to minimize the possibility of part loss to the coolant system, and all fasteners will include a positive locking device. The hinged-disc will include an integral arm hook, eye, or other device for remote inspection of disc function. The preliminnvy arrangement consists of 14" diameter vent valve assem- 3 blies installed in the cylindrical wall of the internals core support shield (refer to Figure 3-59). The valve centers are coplanar and are 42" above the plane of the reactor vessel coolant nozzle centers. In cross section, the valves are spaced around the circumference of the core support shield wall. The hinge design will consist of a shaft, two valve body journal re-ceptacles, two valve disc journal receptacles, and four f1snged shaft journals (bushings). Loose clearances will be used between the shaft and journal inside diameters, and between the journal outside diameters and their receptacles. This feature provides eight loose rotational clearances to minimize any possibility of impairment of disc-free motion in service. In the event that one rotational clearance should bind in service, seven loose rotational clearances would remain to allow unhampered dise-free motion. In the worst case, at least four clearances must bind or seize solid to adversely affect valve disc-free motion. In addition, the valve disc will contain a self-alignment feature so that the external differential pressure will adjust the disc seal face to the valve body seal face. This feature minimizes the possi-bility of increased leakage and pressure-induced deflection loadings on the hinge parts in service. The external side of the disc will be contoured to absorb the impact load of the disc on the reactor vessel inside wall without transmit-tin 8 excessive impact loads to the hinge parts as a result of a loss-of-coolant accident. 3-72a OJ.W 1 , 5-3-68 Supplement No. 3

 .3 2.4.2       Fuel Assemblies 3 2.4.2.1       Description

a. General Description The fuel for the reactor is sintered pellets of low enrichment ura-nium dioxide clad in Zircaloy-4 tubing. The clad, fuel pellets, end supports, holddown spring, and end caps form a " Fuel Rod". Two hun-dred and eight fuel rods are mechanically , joined in a 15 x 15 array to form a " Fuel Assembly" (Figure 3-62). The center position in the assembly is reserved for instrumentation. The remaining 16 positions in the array are provided with " Guide Tubes" for use as control rod locations. The complete core has 17( fuel assemblies. All assemblies are identical in mechanical construction, i.e., all are designed to

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• l I

3-72b , REVIsra, 2-8-48 i

accept the control rod assemblies (CRA). However, only 69 have CRA's to control the reactivity of the core under operating conditions. In the 108 fuel' assemblies containing no CRA during a given core cy-cle, the guide tubes are partially filled at the top by an " Orifice Rod Assembly" (Figure 3-63) in order to minimize bypass coolant flow. These orifice rod assemblies also tend to equalize coolant flow be-tween fuel assemblies with CRA's and those with orifice rod assem-blies. Fuel assembly components, materials , and dimensions are listed below. Item Material Dimensions, in. Fuel UO2 Sintered 0.362 dian Pellets Fuel Clad Zircalcy h 0.h20 OD x 0.368 ID x 152-7/8 long Fuel Rod Pitch 0.558 Fuel Assembly Pitch 8.587 Active Fuel Length 1hh Overall Length =165 Control Rod Guide Zircaloy L 0.530 OD x 0.015 vall Tube 3 Incore Instrument Zircaloy k 0.530 OD x 0.064 vall Guide Extension Spacer Grid Stainless Steel, Spaced at 21-7/16 in. Tp-30h can Panel Stainless Steel, 0.031 thick Tp-30h End Fitting Stainless Steel, Tp-30k

b. Fuel The fuel is in the form of sintered and ground pellets of uranium dioxide. The pellets are dished on each end face to minimize the difference in axial thermal expansion between the fuel and cladding.

The density of the fuel is 95 per cent of theoretical. Average design burnup of the fuel is 28,200 MWD /MTU. Peak burnup is 55,000 MWD /MTU. At the peak burnup, the fuel growth is calculated to be 9-1/2 volume per cent by the method given in Reference 53. This grovth _is accommodated by pellet porosity, by the radial clear-ance provided between the pellets and the cl'adding, and by a small

      )
        . amount of plastic strain in the cladding.

0179 N , 3-73 .. 5-3-68 Supplement No. 3

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Each fuel column is located, at the bottom, by a thin-vall stainless [$(. je e i

                         '      ' steel pedestal and is held in place during handling by a spring at the top. The spring allows axial differential thermal expansion be-
                                                                                                                   . .)

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                    .p           tween fuel and cladding, and axial fuel growth.         The bottom pedestal is also collapsible, thus providing a secondary buffer to prevent ex-

_ cess cladding axial strain. r

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in, , < Fission gas release from the fuel is accommodated by voids within the fuel, by the radial gap between the pellets and cladding, and by void space at the top and bottom ends of the fuel rod.

       ~

c. Fuel Assembly Structure (1) General The fuel assembly shown in-Figure 3-62 is the canned type.

Eight spacer grids and four perforated can panels form the basic structure. The panels are welded together at the corners for the entire length. The spacer grids are velded to the panels, and the lower and upper end fittings are velded to the panels to

    ^

complete the structure. The upper end fitting is not attached until the fuel rods, guide tubes, and instrumentation tube have been installed. At each spacer grid rissembly each fuel rod is supported on four sides by integral leaf-type springs. These springs are designed to provide a radial load on the fuel rod sufficient to restrain it so that flow-induced vibrational am-plitudes are minimal. However, to avoid undesirable bowing of the fuel rods, the spring leads are designed small enough to permit the relative axial motion required to accommodate the differential thermal expansion between the Zircaloy fuel rod and the~ stainless steel structure. (2) Spacer Grid These grids are composed of ferrules =ade of square tubing. The ferrule has a portion of each side formed into spring sections i which have hydrodynamically shaped " dimples" that contact the l fuel rods. The ferrules are joined together by brazing to form the spacer grids. The grids, which provide the desired pitch spacing between fuel rods, are spot-welded at intervals to the perforated stainless steel can panels. (3) _ Lower End Fitting The lover end fitting is constructed from Type 30h stainl'ess steel members which when joined together form a box structure. Four deep cross members serve as the positioning surfaces for the fuel assembly when it is inserted into the lower core sup- ! port structure. The assembly includes a grid structure which provides a support base for fuel rods while maintaining a maxi-mum inlet flow area for the coolant. OL80 3-Th , , L _ __

A (h) Upper End Fitting The upper _end fitting is similar to the lower end fitting. It positions the upper end of the fuel asse=bly and provides coup-ling between the fuel assembly and the handling equipment. A hollow post, velded in the center of the assembly, is designed to provide a means of uncoupling the CRA-to-drive connection and to retain the orifice rod assembly. In order to identify a fuel assembly under water, a serial number is milled into a flat, chrome-plated surface which is welded to the box frame. (5) Control Rod Guide Tubes The Zircaloy guide tubes serve to guide the control rods within the fuel assembly during operation. The tubes are restrained axially by the upper and lower end fittings in the fuel asse=bly and radially by the spacer grids in the sa=e manner as the fuel rods. 3.2.h.2.2 Evaluation

a. Fuel Rod Assembly (1) 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 has demonstrated that Zircaloy h material has ample corrosion resistance and sufficient mechanical properties to maintain the integrity and serviceability required for design burnup. (2) Clad Stress Stress analysis for cladding is based on several conservative assumptions that make the actual margins of safety greater than calculated. For example, it is assumed that the clad with the thinnest vall and the greatest ovality permitted by the specifi-cation is operating in the region of the core where performance requirements are most severe. Fission gas release rates, fuel growth, and changes in mechanical properties with irradiation are based on a conservative evaluation of currently available data. Thus, it.is unlikely that significant failure of the clad- - ding will result during operation. The actual clad stresses are considerably below the yield strength. Circumferential stresses due to external pressure, calculated using those combinations of clad dimensions, ovality, and eccen-tricity that produce the highest stresses, are shown in Table 3-19 The maximum stress of 33,000 psi compression, at the de-sign pressure of 2,500 psi, is the sum of 22,000 psi compressive membrane stress plus 11,000 psi compressive bending stress due

      .           to ovality at the clad OD in the expansion void, and at the be-ginning-of-life. The maximum stress in the heat-producing zone fa -

3-75 0181

is 32,000 psi at design pressure, 27,000 psi at operating pres-sure. .At this stress, the material may creep sufficiently.to allow an-increase in ovality until further creep is rectrained ~} by support from the fuel. Contact loads on the order of 20 lb/ in. of length are sufficient to counteract the bending stress. Creep collapse tests have indicated a long time collapse resis-tance in excess of the requirement to prevent collapse in the end void. As the fuel rod internal pressure builds up with time, these stresses are reduced. Late in life, the fuel rod internal pressure exceeds the system pressure, up to a maximum difference of 1,110 psi. The resul-tant circumferential pressure stress of 9,000 psi is about 1/h 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 esti-mate the long time effects of the increasing pressure on the clad. The predicted pressure-time relationship produces stresses that are less than 1/3 of the stress levels that vould produce stress rupture at the end-of-life. Outpile stress-rupture 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. Clad circumferential stresses are listed in Table 3-19 The free gas content of the fuel rod is calculated by consider-ing (1) initial helium fill gas, (2) initial water vapor and atmospheric gases adsorbed on the fuel, and (3) fission product  ; gases. The water vapor present initially is expected to disso-ciate _over the life of the fuel and enter into hydriding and oxidizing reactions. The gas remaining at the end-of-life, when the maximum internal pressures exist, consists of the atmospheric gases and helium present initially plus the released fission gases. The fission gas production is evaluated for a range of neutron fluxes and the fissionable material present over the life of the fuel.(54) A design value for gas production has been determined as 0.29 atoms of gas per fission. 0182 3-76

Table 3-19 Clad Circumferential Stresses Ultimate Calc. Yield Tensile Stress, Stress, Stress, Operating Condition psi usi psi

1. BOL(") - Operating at Design Pressure Total Stress (membrane + bending) Due to 2,500 psig System Design Pressure Minus 100 psig Fuel Rod Internal Pressure Average Clad Temperature - Approxi-mately 625 F (expansion void) -33,000 46,000

2. EOL - Maximum Overpower System Pressure - 2,185 psig Fuel Rod Internal Pressure -

3,300 psig Average Temperature Through Clad Thickness at Hot Spot - Approxi-mately 725 F Pressure Stress Only(b) 9,000 Including 4,000. psi Thermal Stress 13,000 36,000 38,000

3. EOL - Shutdown Immediately After Shutdown System Pressure - 2,200 psig Fuel Rod Internal Pressure -

1,750 psig

   -Average Clad Temperature - Approxi-mately 575 F                                  4,000     45,000      48,000 (a) Cladding is being ordered with 45,000 psi minimum yield strength and 10 per cent minimum elongation, both at 650 F. Minimum room temper-ature strengths will be approximately 75,000 psi yield strength (0.2 per cent offset) and 85,000 psi ultimate tensile strength.

(b) Cladding stresses due to fuel swelling are discussed further on another page of 3.2.4.2.2. t 3-77 s

s

                                                                                      -) '

Table 3-19 (Cont'd) Ultimate Cale. Yield' Tensile Stress, Stress, Stress, Operating Condition psi psi psi 3 Hours Later (50 F/hr Pressurizer Cooldown Rate) Fuel Rod Internal Pressure - 1,050 psig System Pressure - 680 psig Average Clad Temperature - Approxi-mately h25 F 3,300 52,000 55,000 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 h midpoint is 55,000 MWD /MTU. This fission gas release is ased on temperature versus release  ; fraction experimental data.(49) Fuel temperatures are calculated ' for small radial and axial increments. The total fission gas re-lease is calculated by integrating the incremental releases. The maximum release and gas pressure buildups are determined by evaluating the following factors 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.

             -(c) Unrestrained radial and axial 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.

The fuel temperatures used to determine fission gas release and internal gas pressure have been calculated at the reactor over-power condition. Fuel temperatures, total free gas volume, fis-sion gas release, and internal gas pressure have been evaluated

                               """                               6-5-68 e      ,

()j g34," Supplement No. 4 3-78 l 1_

for a range of initial diametral clearances. This evaluation shows that the highest internal pressure results when the maxi-mum diametral gap is assumed because of the resulting high aver-age fuel temperature. The release rate increases rapidly with an increase in fuel temperature, and unrestrained axial growth reduces the relatively cold gas end plenum volumes. A ccuser-vative ideal thermal expansion model is used to calculate fuel temperatures as a function of initial cold diametral clearance. Considerably lower resistance to heat transfer between the fuel and clad is anticipated at the end-of-life due to fuel fracture, swelling, and densification. The resulting maximum fission gas release rate is h3 per cent. (3) Collause Margins Short time collapse tests have demonstrated a clad collapsing pressure in excess of 4,000 psi at expansion void maximum tem-perature. Collapse pressure margin is approximately 1 7 Ex-trapolation to hot spot average clad te=perature (: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. (4) Fuel Swelling Fuel rod average and hot spot operating conditions and design parameters at 100 per cent power, pertinent to fuel swelling considerations, are listed below. Average Maximum Heat Flux, Btu /ft2 -hr 167,620 543,000 Linear Heat Rate, kw/ft 5.4 17.5 Fuel Temperature, F 1,385 4,160 Burnup (MWD /MTU) at Equilibrium 28,200 55,000 Nominal Values Pellet OD, in. 0.362 Pellet Density, % of Theoretical 95 Pellet-Clad Diametral Gap at Assy., in. 0.00h - 0.008 Clad Material Cold-Worked Zr h Clad Thickness, in. 0.026 The capability of Zircaloy-clad UO2 fuel in solid rod form to perform satisfactorily in PWR service has been amply demonstrated through operation of the CVTR and Shippingport cores, and through results of their supplementary development programs, up to ap-proximately 40,000 MWD /MTU. 0185

                 \

3-79 -

   .y.

4 4 h + As outlined below, existing experimental information supports the various individual design parameters and operating conditions up ") ' to and perhaps beyond the maximum burnup of 55,000 WD/MfU, but not in a single experiment. However, the Babcock & Wilcox irra- 3 diation test program, currently in progress, does combine the items of concern in a single experiment, and the results will be available to contribute the final design confirmation.

             .(5) Application of Experimental Data to Design Adequacy of the Clad-Fuel Initial Gap to Accommodate Clad-Fuel Differential Thermal Expansion Experimental Work Six fuel rabbit length,capsules, each containing were irradiated            three Zr-2 Test in the Westinghouse  clad Reactor rods of 5-(in.

45) at power levels up to 2h kv/ft. The 9h per cent theoretical den-sity (T.D.) UO2 pellets (0.430 OD) had initial clad-fuel diametral gaps of 6, 12, and 25 mils. No dimensional changes were observed. Central melting occurred at 2h kw/ft only in the rods that had the 25 mil initial gap. Two additional capsules were tested. (55) The specimens were sim-ilar 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 3,450 WD/MI'U at 19 kv/ft maximum. In the A-h capsule, four 6-in.-long rods were irradiated to 6,250 WD/MIU at 22.2 kv/ft maximum. No central melting occurred in any rod, but diameter increases up to 3 mils in the A-2 capsule and up to 1.5 mils in the A-h capsule were found in the rods with the 2 mil initial gap. Application In addition to demonstrating the adequacy of Zircaloy-clad UO 2 pellet rods to operate successfully at the power levels of in-terest (and without central melting), these experiments demon-strate that the design initial clad fuel gap of h to 8 mils is adequate -to prevent unacceptable clad diameter increase due to differential thermal expansion between the clad and the fuel. A maximum local diametral increase of less than 0.001 in. is in-dicated for fuel rods having the minimum initial gap, operating at the maximum overpower condition. (6) Adequacy of the Available Voids to Accommodate Differential Expansion of Clad and Fuel, (

                  . Including the Effects of Fuel Swelling Experimental Work Zircaloy-clad, UO2 pellet-type rods have performed successfully

An the Shippingport reactor up to approximately h0,000 WD/MTU.

3-80 5-3-68 Supplemcnt No. 3 m .

Bettis Atomic Power Laboratory (53) has irradiated plate-type UO2 fuel (96-98 per cent T.D.) up to 127,000 WD/MTU and at fuel cen-ter temperatures between 1,300 and 3,800 F. This work indicates fuel swelling rates of 0.16% AV/1020 f/cc until fuel internal voids are filled, then 0 7% AV/1020 f/cc after internal voids are filled. This point of " breakaway" appears to be independent of temperature over the range studied and dependent on clad re-straint and the void volume available for collection of fission products. The additional clad restraint and greater fuel plas-ticity (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 Harwell >61 in which 1/4 in, diameter UO2 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 dimensional change. In other testing (57) 0.150 in. OD, 82-96 per cent T.D. 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 irradiated to 77,000 MWD /MTU at fuel temperatures high enough to approach central melting without apparent detrimental results. Compar-able results were obtained on rods svaged to 75 per cent T.D. and irradiated to 100,000 MWD /MTU. Applicatien Based on the BAPL experimental data, swelling of.the fuel rods is estimated as outlined below.  ; Fuel is assumed to swell uniformly in all directions. Clad-pellet differential thermal expansion is calculated to be about 0.004 in, at the maximum linear heat rate, so that all of the minimum initial gap of 0.004 in. is filled up by thermal expan~ sion. If the initial gap exceeds the minimum, the additional gap volu=e is assumed available to acec=modate swelling. This additional veld volume may initially tend to be filled by pellet ther=al expansion because of the low contact pressure and re-sultant low contact coefficient, but as the fuel swells , the contact pressure must increase if the clad is to be stretched. Where fuel cracking tends to fill the radial gap, it is assumed that the crack voids are available to absorb swelling. The external effect of fuel swelling is assumed to occur at 0.16% ^ AV/1020 f/cc until the 5 per cent initial void in the 95 per cent T.D. pellets-is filled at about 9 x 1020 r/cc. From that time on, swelling is assumed to take place at 0 7% AV/1020 f/cc until the  ; 2 maximum burnup of 13.6 x 10 0 r/cc (55,000 MWD /MTU) is reached. Total fuel volume increase is h-1/2 per cent, which results in a 121/2 per cent diameter increase in a rod with the 0.004 in. mini-mum initial gap. Clad stress is estimated at 22,000 psi, so t at 5;t

                            '3-01 7

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the elastic strain in about 0.2 per cent. Net plastic strain is ' l.3 per cent. Similar calculations indicate that fuel rods with maximum burnup and the nominal clad-fuel gap (0.006 in, at as-sembly) vill have clad plastic strains of about 0.6 per cent at the end-of-life. Based on 'outpile data, stress rupture should not be a problem at these strains. Qualitative information from ISBRI30) suggests that swelling rates for this design may exceed those indicated by the BAPL data because of the higher fuel temperatures. However, the A.E.R.E. tests (56) and the General Electric tests (57) do not support more than a small increa.se in post " breakaway" swelling rates at tem-peratures of interest. Fuel Svelling Studies 3 Dimensional stability of UO2 under inpile conditions simulating large reactor environments is under investigation. Parameters contributing to swelling are burnup, heat rating, fuel density and grain size, and clad restraint. These are systemati-cally being studied by irradiating a series of capsules contain-ing fuel rods. Test variables are shown in Table 3-20 and the program schedule is given in Table 3-20a. f k L f 93.38

                       . M **
                        !'                 3 .82 5-3-68
                                                                             ; Supplement No. 3

Table 3-20 3 8

           ,   '{                                                                                                           BW High Burnup Program - Capsule Fuel Test Burnup                                        Heat Rate Identification Irradiation Diametral        Clad 14fD/     Fissions /cc                        Initial,    Final      Gap,      Thickness C^p mle ;'. Fuel Rod                                                                Facility                 x10]-     x 10-23(1)   Irradiation Calendar    Tim 9d )

Months W/Ft IGi/Ft mils mils B-1 B-1 RS-3 10 25 4 18 B-2 17 5 4-5 25 25 4 17 5 7-8 25 B-3 25 4 B-2 B-4 20 17 5 7-8 15 RS-5 4.46 7 18 16.9 Powder B-5 25 50 7 16 9 4-5 25 B-3 B-7 RS-6 30 75 n 18 16.1 B-8 4-5 25 75 13 16.1 7-8 25 B-9 -7 5 n 16.1 B-4 B-10 RS-6 7-8 15 45 10.05 17 18 14.9 Bowder 25 B-11 11.25 17 m 14 9 4-5 25 "do B-5 B-13 RS-5 55 13 75 21 18 14.1 4-5 25 B-14 13.75 21 14.1 B-15 7-8 25 13 75 21 14.1 7-8 15 B-6 B-16 RS-3 65 14.50 24 18 13.3 Powder 25 B-17 16.24 24 13 3 7-8 25 B-7[ B-19 RS-4 75 16.73 28 '18 12 5 Powder 25 B-20 18.75 28 12 5 7-8 25 C H B-8 -B-22 RL-2 25 6.25 10 21} 20 5 4-5 25

@                   B-23                                                                                                         6.25           10 CD                  B-24                                                                                                                                                    20 5       7-8            25 6.25           10                          20 5       7-8            15 B-9        B-25                                                                             RL-2           50         12.5             18 B-26                                                                                                                                        21i         19.3       4-5            25 12.5             18                          19 3       7-8            25 B-27                                                                                                       12 5             18                          19 3 B-lo       B-28                                                                            RL-2 7-8            15 75         18.75            28              21          17.8       7-8            25 B-29                                                                                                       18.75            28                          17.8 B-30                                                                                                                                                               7-8          ~ 25 18.75            28                          17.8      7-8             15 (1) Based on 200 Mev per fission.

(2) Based on 80 per cent reactor efficiency. 5-3-68 Supplement No. 3

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Effect of Zircaloy Creen The effect of Zircaloy creep on the a=ount 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,(60) 1 per cent creep vill result in 10,000 hr (corresponding approximately to the end-of-life diametral swelling rate) from a stress of about 22,000 psi at the 2720 F average temperature through the clad at the 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 would more or less be balanced by the rod internal pres-sure, so the total pressure to produce the clad stress of 22,000 psi vould have to come from the fuel. Contact pressure would be 2,h00 psi. At the end-of-life, the rod internal pressure ex-ceeds the system pressure by about 1,100 psi, so the clad-fuel contact pressure would drop to 1,300 psi. Assu=ing that irradia-tion produces a 3:1 increase in creep rates, the clad stress for 1 per cent strain in 10,000 hr would drop to about 15,00 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 the 3:1 increase in creep rates due to irradiation) when the pellet is slightly hotter than calculated. The effect of this would be a slight increase in pellet thermal expansion and therefore in clad strain. Considering the improbability that irradiation will actually increase creep rates by 3:1, no change is anticipated.

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 free ;m to insert the control rod. As indi-cated belov, studies have shown that fuel rods vill not bow sur-ficiently to touch the guide tube. Thus, the guide tube vill not andergo 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 0.083-in, nominal gap between fuel rod and guide tube to a minimum of about 0.045 in., including amplification of bowing due to axial friction loads from the s gradient of 1.176 across~ apacer grid.vill The fuel rod maximum expected flux 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.

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The effect of a DNB occurring on the side of a fuel rod adjacent to a guide tube would result in a large temperature difference. In this case, however, investigation has shown that the clad temperature would be so high that insufficient strength would 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 re-sist 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 The semiempirical expression developed by Burgreen(61) 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 what is believed to be a conservatively calculated amplitude for the fuel assembly, a direct measure-ment vill be 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. (3) Demonstration In addition to the specific items discussed above, the overall mechanical performance of the fuel assembly and its individual components is being demonstrated in an extensive experimental program in the CRDL. 3.2.4.3 Control Rod Drive System 3.2.4.3.1 Control Rod Drive System Design Criteria The control rod drive system shall be designed to meet the following perfor-mance 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 chain of failures. shall cause uncontrolled with-drawal of any control red assembly (CRA).

c. Equipment Removal The disconnection of plug-in type connectors , modules, and subassem-

           ; blies from the protective circuits shall be annunciated or shall cause a. reactor' trip.                         g 3-85 019?-

s- . cs

d. Control Rod Assembly (CRA) Trip The trip command shall have priority over all other commands. 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. Puses, where used, shall be provided with blown indicators. Circuit breaker po-sition information shall also be indicated.

e. CRA Insertion Insert co= mand shall have priority over withdraw command. The con-trol rod drive vill be capable of overcoming a " stuck-rod" condition equivalent to a k00 lb weight.

f. Withdrawal The control rod drive system allows only two out of four regulating CRA groups to withdraw at any time subject to the conditions de-scribed in 7.2.2.1.2.

g. Position Indication 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.

h. 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.

1. Drive Steed The control rod drive control system shall provide for single uniform speed of the mechanism. The drive controls, or mechanism and motor combination, shall have an inherent speed-limiting feature. The speed of the mechanism shall be 30 in./ min plus or minus 10 per cent of the predetermined value for both insertion and withdrawal. The withdrawal speed shall be limited so as not to exceed 25 per cent overspeed in the event of speed control fault.

J. Mechanical Stops Each control rod drive shall be provided with positive mechanical stops at both ends of the stroke or travel. The stops shall be cap-able of receiving the full operating force of the mechanisms without failure. z

                                                                            ~

l " 3-86 5-3-68 Supplement No. 3

3.;2.k.3.2 Control Rod Drive The control rod drives 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. The control rod drives are buffer seal, rack-and-pinion type drives under development by Diamond Power Specialty Corporation. The control rod drive data are listed in Table 3-21. A control rod drive consists of a rack housing,

• snubber bottoming spring assem-bly,. rack, rack pinion, coupling assembly, drive shaft housing, miter gear set, drive shaft assembly, buffer seal assembly, magnetic clutch, gear reducer, drive motor, position indication transmitters , and limit switch system. The spool piece serves to join the drive assembly to the reactor closure head nozzle as shown in Figure 3-6h.

The drive motor supplies torque through the magnetic clutch to the drive shaft-gear system to provide vertical positioning of the rack. Table 3-21 Control Rod Drive Design Data Item Data Number of Drives 69 Type Buffer Seal, Rack and Pinion Location Top-Mounted Direction of Trip Down Velocity of Normal Withdrawal 3 and Insertion, in./ min 30 Maximum Travel Time for 2/3 Trip Insertion, see 1.4 Length of Stroke, in. 139 Design Pressure, psig 2,500 Design Temperature, F 650 The control rod drive is shown on_ Figures 3-64 and 3-65 Subassemblies of the control rod drive are described as follows: t, .. 01.94 8 a 3-87 5-3-68

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    - sn :

a. - Rack Housing The rack housing contains the hydraulic snubber, the bottoming spring assembly,' the rack, rack pinion assembly, and a rack guide bushing.

The lower guide tube is attached to the lower end of the rack hous-ing, and the cap and drive line vent assembly is mounted on the upper end of the rack housing. The hydraulic snubber decelerates the moving elements of the drive at the end of. travel by controlled orificing of reactor coolant water. The bottoming spring assembly absorbs the bottoming impact in a stack of spring washers. The rack is guided by an upper shoe attached to the upper end of the rack, a rack guide bushing located at the pin-ion, and a lower guide tube bushing located at the lower end of the lower guide tube. The rack pinion is carried by two ball bearings. The valve on the cap and drive line vent assembly is used to bleed air or gases from the rack housing during reactor startup. The re-moval of this assembly provides the access for CRA coupling and un-coupling, and for securing the racks in the retracted position when - the reactor closure head or individual drives are to be removed.

b. Drive Shaft Housing The drive shaft housing consists of the miter gear set, the drive shafts, and their supporting ball bearings. The drive shaft assem-bly is made up of two shafts with an intermediate bearing to increase their critical speed.

The drive shaft housing is attached to the rack housing by four through bolts. All gasketed joints are of the double Conoseal-type with a pressure testing tap between the seals,

c. Buffer Seal A pressure breakdown-type seal is employed to seal the drive shaft penetration in the reacter coolant pressure container. Seal system vater is injected between the eighth and ninth stages of a nine stage seal to provide a controlled leakage of approximatelv 5 gal /hr into the reactor coolant system and 20 gal /hr to the makeup tank. The seal water is cooled below 120 F, and specially filtered before injection into the seal. A conventional rotary seal is employed to prevent seal water from entering the drive package,

d. _ Drive Package l

The drive package is a synchronous type containing a self-locking vorm gear reducer, a magnetic clutch, position indication transmitters, and a limit switch system. In conjunction with the magnetic clutch is a unidirectional mechanical clutch which vill allow the motor to drive m

                            . n t-==                            '

v 3288 0195

       'the rod down to the full-in position should a " stuck-rod" condition develop in the course of a trip action.      The motor has inherent brak-ing so no separate brake is required. The self-locking worm gear re-ducer prevents torque feedback to the motor.

The unidirectional feature of the magnetic clutch assembly, which is located between the drive motor and the buffer seal, will function as follows: (1) With the clutch de-energized, the clutch will allow the control rods to fall into the reactor by gravity. However, the unidiree-tional feature will allow the motor to drive through the clutch only in the direction of inserting the control rods, thus allow-ing backup drive-in of control rods following a trip. (2) With the clutch de-energized, the control rods will be held in position in the core even with a net upward force on the control rod because the drive shaft vill drive through the clutch to the motor gear asse=bly which cannot be driven from the reverse di-rection. (3) With the clutch energized, the motor can drive the control rods in both directions, outward or inward.

e. Position Transmitters and Limit Switches The position transmitters and limit switches are located between the buffer seal and the gear motar in the power package and supply re-dundant position signals and limit switch contacts.

There are three separate devices included in the position and limit switch transmitter assembly. A potentiemeter generates an analog position signal, a linear variable differential transformer (LVDT) generates both an analog position signal and limit contacts, and the limit switch mechanism pro" ides limit contacts. Refer to Figure 3-67 The potentiometer is geared directly to the drive shaft and gives a continuous d-c signal proportional to the CRA position. The LVDT transmitter has a core that is moved by means of a ball screw mech-anism geared.to the drive shaft. A demodulator located within the control' cabinet contains the necessary electronic circuitry to gen-erate the analog d-e signal. This demodulator also has relays with adjustable set points for position contacts. The limit switch assem-bly consists of switches operated by linear ca=s that are moved by a ball screw. This is also geared directly to the drive shaft. By using these three transmitters, it is possible I to get both redun-dant position and redundant li=it signals.

f. Housing Design Criteria The control rod. drive assembly housings are designed to the same de-sign criteria as is the' reactor pressure vessel. Accordingly, the drive shaft and rack housings comply with Sec_t(on III of the ASME

                 ~

4 Boiler and Pressure Vessel Code under classification as Class A vessels. The operating transient cycles, whf ch are considered for )' the stress analysis of the reactor pressure vessel, are also con-sidered in the housing designs. Quality standards relative to material selection, fabrication, and irspection are specified to insure safety function of the housings essential to accident prevention. Materials conform to ASTM or ASE, Section II, Material Specifications. All velding shall be performed by personnel qualified under ASE Code, Section IX, Weld-ing Qualifications. These design and fabrication procedures estab-lish quality assurance of the assemblies to contain the reactor coolant safely at operating temperature and pressure. For vibratory and seismic loadings, the assemblies are restrained with a series Of contoured plates that are bolted to the main sup-port structure. These plates are contoured to restrain the upper flange outside diameters of the drive shaft and rack housings as shown in Section CC, Figure 3-65 The main support structure is bolted to the reactor closure head. These plates vill provide lateral support only. Vertical motion of the housings resulting from thermal expansion vill not be restricted. In the highly unlikely event that a pressure barrier component or the control rod drive assembly did fail catastrophically, i.e., a complete rupture, the following results vould ensue: (1) Control Rod Drive Nozzle For the failure of this component, the assembly would be ejected upward as a missile until it was stopped by the reactor building missile shield. This upward motion would have no adverse effect on adjacent assemblies. (2) Rack Housing The failure of this component anywhere above the lover flange would result in a missile-type ejection into the missile shielding of_the reactor building. There would be no adverse effect on adjacent mechanisms. (3) Drive Shaft Housing  ! The failure of this component could result in contact with adjacent assemblies only if the bolting by which the drive i shaft housing is attached to the rack housing failed com-pletely. However, it should be noted that this bolting vill be designed to Section III of the ASME: Boiler and Pressure Vessel Code, and therefore has the same integrity as the housing itself. A general design criterion--that no single bolt failure vill lead to subsequent failure of the remaining bolts--has been imposed upon the design. 3-90 REVISED, 2-8-68

                 , { {7 s                         g.-                ~

L

In addition, the distance between adjacent drive shaft housin6s is small, approximately 1/4 in., and should a failure be postulated, the short travel distance avail-able will not pemit the failed drive to obtain a high velocity. The design criterion for the rack housing is to accept this impact load without failure. If a rod ejection accident were assumed to occur, even though design precautions have been taken to insure the integrity of the control rod drive assembly, no further damage to the reactor internals or the reactor coolant system vill occur, nor would gross fuel failures result. This analysis is presented in 14.2.2.2. Since the control rod drive assembly is designed to minimize the probability of an accident leading to control rod ejection, and since the consequences of such an accident (if it did occur) do not lead to a serious potential safety problem, a holddown mechanism is not re-quired. 3 2.4 3 3 Control Rod Drive Control System (Control Package) The control system for the control rod drive is designed to energize and position the control rod drive, indicate the control rod assembly (CRA) position in the core, and indicate malfunctions in the system. As shown on Figure 3-66, the con-trol system consists of: Power supplies and monitors. Clock (CRA speed standard). Control rod drive grouping panel. Individual CRA control logic. Position indicator system. Travel limit system. 4 3-90a -- _ EEVISED, 2-8-68 hh 1

Automatic Sequ;nca Iogic Trip System Position Deviation 2 nitors The control rod drive control system provides the reactor operators with the flexibility of CRA grouping, manual or automatic group operation, automatic CRA group sequencing, and information of CRA position in the core. A total of 8 CRA groups is available through facilities of a control rod drive 3 grouping panel which enables up to 12 CRA's to be assigned to each group. In-dividual position indicators are provided for all 69 CRA's and are visible to the operator. The operating control panel includes four group position indicators. The , indicators are assigned to groups 1 through 1+ during startup when the safety rods are being withdrawn, and to groups 5 through 8 during operation. CRA groups are progranned so that the power peaking values listed in Table 3-1 are never exceeded. Automatic sequencing (group overlap) of groups 5 through 8 is provided and is 3 available for automatic or manual operator CRA motion requirements. It allows a limited overlap of operation of any two groups in a fixed sequence, but no more than two. Inputs from CRA position and travel limits feed this system. Automatic and manual control is provided. In " automatic", the selected control rod drive group receives an automatic cnmmand signal from the integrated control 3 system. In " manual", provision is made for operation of any individual CRA or group of CRA's. Manual or automatic operation of four CRA groups in a preset sequence is provided as described above. Grouping is determined at the control rod drive grouping panel prior to reactor operation. The drive gate is part of the individual CRA control logic circuitry which per-forms the function of selection and gating. It receives inputs from the clock, the IN and OUT control buses, motion " Enable", and travel limits. The drive gate sends pulses to the translator upon receiving (a) clock pulses, (b) "En-able" input, and (c) and IN or OUT control signal. End travel limits and the driver monitor provide inputs to stop CRA motion. Output signals of the drive gate feed into the translator. This unit produces the pre er signals for the drive motor. Direction is determined by the IN and OUT commands, and speed is determined by the fixed clock frequency. The position indication and travel limit systems consist of three different types of~ transmitters and produce two independent analog position signals and two independent limit signals. One of the devices, the LVDT, produces both position and limit signals. Either source of signals can be used for the posi-tion and for the limit signals. Position output jacks are provided for a precision meter and for computer moni-toring. Calibration of the potentiometer and the LVDT is accompidshed by initial adjustments prior to installing the power package and also by making adjustments within the-control cabinet. The limit switches are adjusted prior to installation of the drive package. 3-91 ~ lN r. 5-3-68 Supplement No. 3 l i

A fault detection circuit monitors signals to provide extra protection against

   -unwanted withdrawal and intertion motion. See Figure 3-66.

T The rod drive control system has two speed-limiting features. First, the motor speed is limited by the frequency of the input power set by a clock or pulse generator. Second, this limit is followed by a speed-saturating circuit which has the inherent property of not responding to a frequency greater than 125 per cent of rated frequency. These features vill prevent an over-frequency and overspeed of the drive. In addition to speed limitation, the rod groups have independent " Enable", sig-nals and gates such that no more than two groups can be enabled simultaneously for withdrawal motion in accordance with the description in 7.2.2.1. These two features, frequency limit and group " Enable" limits, hold the maximum with-drawal rate well below that analyzed in 14.1.2.3. Trip is initiated by deenergiz_ng either of two series circuit breakers in each of two power sources (Figure 3-68). Each loss-of-voltage trip coil is fed by a sepa-rate two-out-of-four relay circuit powered by four inputs from the reactor pro-tection system. Failure of any two inputs causes trip. The manual trip push-button opens all trip circuit breakers. Test pushbuttons are provided to test each circuit breaker action. 3.2.k.3.4 Control Rod Drive System Evaluation

a. Design Criteria The system vill be designed, tested, and analyzed for compliance with s 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 cause inadvertent CRA withdrawal of another CRA or CRA group.

b. Materials Selection Materials are selected to be compatible with, and operate in, the re-actor coolant. Certified mill test reports containing chemical anal-ysis and test data of all materials exposed to the reactor system fluid shall 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 Temperature All parts of the control rod drive exposed to reactor coolant are de-signed to operate at 650 F, although it is expected that all parts will operate considerably cooler. Some tests have been completed, and ad-ditional tests are planned, to closely 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 mechanism.

It is expected that the more significant temperature changes vill be t , 200 6-5-68 3-92 Supplement No. 4 e

caused by displacement of reactor coolant in and out of the mechanism vater space as the drive line is raised and lowered.

d. Design Life The expected life of the control rod drive control system is as fol-icvs:

(1) Structural portions, such as flanges and pressure housings, have an expected. life of 40 years. (2) Moving parts , such as rack, pinions , and other gears, 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 is made up of 16 control rods which are coupled to a single Type 304 stainless steel spider (Figure 3-69). Each control rod con-sists of an absorber section of silver-indium-cadmium clad with cold-worked, Type 30k stain 1 css steel tubing and Type 30k stainless steel upper and lover end pieces. The end pieces are velded to the clad to form a water and pressure-

   - tight container for the absorber. The control. rods are loosely coupled to the spider to permit maximum conformity with the channels provided by the guide            ;

tubes. The CRA is inserted through the upper end fitting of the fuel assembly,  ! each control rod being guided by an incore guide tube. Guide tubes are also provided in the upper plenum assembly above the core so that full length guidance of the control rods'is provided throughout the stroke. With the reactor assem- , bled, the CRA cannot be withdrawn far enough to cause disengagement of the con-trol rods from the incere guide tubes. Pertinent design data are shown in Table 3-22. Table 3-22 Control Rod Assembly Design Data Item Data Number of Rod Assemblies 69 Humber of Control Rods per Assembly 16 Outside-Diameter of Centrol Rod, in. 0.440 Cladding Thickness, in. 0.018 Cladding Material- Type 30k SS, cold-worked

              -Absorber Material                           80% Ag, 15% In, 55 Cd Length of Absorber Section, in.              134 Stroke of Control Rod, in.                   139 This type of CRA has been developed under the USAEC Large Reactor Development
   . Program and offers the following significant advantages:

V 201 3-93 C.

a. More uniform distribution of absorber throughout the core volume.

b. sy Shorter reactor _ vessel and shorter internals owing to elimination of -

control rod followers.

c. Lower reactor building requirements owing to reduction of reactor coolant inventory.

d. Better core power distribution for a given CRA vorth.

A CRA prototype similar to the B&W design has been extensively tested (62) at reactor temperature, pressure, and flow conditions under the LRD program. The silver-indium-cadmium absorber material is enclosed in stainless steel tubes to provide structural strength of the control rod assemblies. These rods are designed to withstand all' operating loads including those resulting from hy-draulic forces, thermal gradients, and reactor trip deceleration. The cladding of the absorber section also prevents corrosion and eliminates possible silver contamination of the reactor coolant. The ability of the absorber clad to resist collapse due to the system pressure-has been demonstrated by an extensive collapse test program on cold-worked stainless steel rods. The actual collapse margins are higher than the require-ments. Internal pressure and absorber swelling are not expected to cause stressing or stretching of the clad because the Ag-In-Cd alloy does not yjeld a gaseous prod-uct under irradiation. Because of their great length and unavoidable lack of straightness, some slight mechanical interference between control rods and guide tubes must be expected. However, the parts involved, especially the control rods, are so flexible that only very small friction drags will result. Similarly, thermal distortions of the control rods are expected to be small because of the low heat generation and adequate cooling. Consequently, it is not anticipated that the control rod assemblies vill encounter significant frictional resistance to their motion in the guide tubes. Lifetime tests are being performed on a prototype CRA in the CRDL Facility de-scribed in 3.3.3.1 and in accordance with the program outlined in 3.3.3.k.1. Approximately 2,200 full-stroke cycles and 250 full-stroke trips have been com-pleted with the reference design CRA at reactor operating conditions of pres-sure, temperature, flow, and water chemistry. This is approximately equivalent to 20 years of operation on the CRA. Evidence of contact was noticed on the lead-in tip of the control rod assembly, but no measurable amount of metal had been removed. Visual inspection of the spider shows an insignificant amount of wear. At the end of h10 full-stroke qycles and 50 full-stroke. trips (the equivalent of three years operation in one assembly), the incore guide tubes in the fuel assembly were examined. ' Wear marks were noted at the entrance of the guide tubes ~, and these marks extended into .the guide tubes approximately 5 in. Ap-proximately 7 mils of metal had been removed longitudir. ally from the guide tubes at the upper end. Since no change in the time required for two-thirds {

~~

      . insertion was noted over the ' duration of the testing performed to date, it-is concluded that wear of the guide tubes and the CRA's will not be of concern.

These tests will be continued to cc=pletion. The methods and frequency of CRA in-service inspection as well as the criteria for replacement will be. determined during the detailed design. 3.3 TESTS AND INSPECTIONS 3.3.1 NUCLEAR TESTS AND INSPECTION 3.3.1.1 Critical Experiments An experimental program (65-67) 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 ele-ment gecmetry, CRA materials, and soluble boron concentration in the moderator. Gross and local power peaking were also studied, and three-dimensional power-peeking data were taken as a function of CRA insertion. Detailed peaking data vere drawnalso taken between fuel assemblies and around the water holes left by with-CRA's. 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 worth (including stuck-CRA worth) vill be determined by physics tests at the beginning of each core cycle. Becalibration of boron worth cycle. and CRA worth is expected to be performed at least once during each core

             . Calculated values of boron worth and CRA worth will be adjusted to the test values as necessary. The boron worth and CRA worth at a given time in core life justed  byvill be based on experimental   data.CRA position indication and calculated data as ad-Th'e reactor coolant will be analyzed in the laboratory periodically to deter-mine the boron concentration, and the reactivity held in boron vill then be calculated from the concentration and the reactivity worth of boron.

The method of maintaining the hot shutdown margin (hence stuck-CRA margin) is related to_ operational characteristics (load patterns) and to the power-peak-ing restrictions on CRA patterns at power. The CRA pattern restrictions vill insure that sufficient reactivity is always fully withdrawn to provide adequate shutdown with the stuck-CRA margin. Power peaking as related to CRA patterns and shutdown margin will be monitored by reactivity calculations, and inter-locks will.be provided to prevent CRA patterns that produce excessive power peaking and/or reduction of shutdown margin. Operation under all power conditions will be monitored by incore instrumenta-tion, and the resulting data vill be analyzed and ccmpared with multidimen-sional calculations

  ;further power            in a continuing effort to provide sufficient support for escalations..

203: 3-95 - <

3.3.2 THERMAL AND HYDRAULIC TESTS AND INSPECTION 3 3.2.1 Reactor Vessel Flow Distribution and Pressure Drop Test A1/6-scalemodelofthereactorvesselandinternalswillbetestedtomeasure

a. 'Ihe flow distribution to each fuel assembly of the reactor core and to develop, if necessary, devices required to produce the desired flow distribution.

b. Fluid mixing between the vessel inlet nozzle and the core inlet, and 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 3 flow circuit.

d. The internals' vent valves will be evaluated for closing behavior and for the effect on core flow with valves in an open position.

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 techniques. A u 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 similarity between the model and prototype. Each of the 177 s h lated fuel assemblies contains a calibrated flow nozzle at its inlet and outlet. The test loop is capable of supplying cold water (80 F) to three inlet nozzles and hot water (180 F) to the fourth. Temperature will , be measured in the inlet and outlet nozzles of the reactor model and at the in-let and outlet of each of the fuel assemblies. Static pressure taps will be located at suitable points along the flow path through the vessel. This instru-mentation will provide the data necessary to accomplish the objectives set fourth ) for the tests, p 3.3.2.2 Fuel Assembly Heat Transfer and Fluid Flow Tests B&W is conducting a continuous research and development program for fuel assem-bly heat transfer and fluid flow applicable to the design of the reference re-actor. Single-channel tubular and annular test sections and multiple rod as-semblies have been tested at the B&W Research Center. The reactor thermal design is based upon burnout heat transfer experiments with (a) multiple rod, heated assemblies with uniform heat flux, and (b) single rod, annular heaters with nonuniform axial heat flux, at design conditions of pres-sure and mass velocity. 'Ihese experiments are being extended to test nonuniform multiple rod heater assemblies as described in 15.4. The results of these tests will be applied to the final thermal design of the reactor and the specification of operating limits. 3.3.2.2.1 Single-Channel Heat Transfer Tests A large quantity of uniform flux, single-channel, critical heat flux data has been obtained. References to uniform flux data are given in BAW-168 and 3.2 3 2.3 of this report. The effect on the critical heat flux caused by non-uniform axial power generation in a tubular test section at 2,000 psi pressure 2U4 3-96 =-3-68 Supplement no. 3 l:

vas investigated as early as 1961.(29) Thisprogramwasextendedtoincludg pressures of 1,000, 1,500, and 2,000 psi and mass velocities up to 2.5 x 10 lb/hr-ft .(63) 2 The effect on the critical heat flux caused by differences in the radial and axial power distribution in an annular test section was recently investigated at reactor design conditions.(64) Data were obtained at pressures of 1,000 1,500, 2,000, and 2,200 psi and at mass velocities up to 2 5 x 106 lb/hr-ft The tubular tests included the following axial heat flux shapes where P/F is local to average power: ,

                                            .s

a. . Uniform Heat Flux (P/F) = 1.000 constant

b. Sine Heat Flux (P/F)m, = 1.396 @ 50% L

c. Inlet Peak Heat Flux (P/F)=,x = 1.930 @ 25% L

d. Outlet Peak Heat Flux (P/F) = 1.930 @ 75% L Tests of two additional, nonuniform, 72-in. heated length, tubular tests were undertaken to obtain data for peaking conditions more closely related to the reference design. The additional flux shapes being tested are

a. Inlet Peak Heat Flux (P/F)g g = 1.65 @ 28% L

b. Outlet Peak Heat Flux (P/F), = 1.65 e 72% L These tests, still in progress, will cover approximately the same range of pressure, mass flov, and AT as the multiple rod fuel assembly tests.

3.3.2.2.2 Multiple Rod Fuel Assembly Heat Transfer Tests Critical heat flux data are being obtained from 6-ft-long, 9-rod fuel assem-blies in a 3 x 3 square array. A total of 513 data points were obtained cover-ing the following conditions: O s ATg5 250 1,000 5 P 5 2 h00 0.2 x 1065 G 5 3.5 x 10 6 where ATg = inlet subcooling, F P = pressure, psia G = mass velocity, Ib/hr-ft 2 The geometry of this section consisted of nine rods of 0.k20-in. diameter on a 0 558-in. square pitch. Analysis of the last data of this set is in process. 203r 3 'g .

3 3 2.2 3 FUEL ASSEMBLY FIOR DISTRIBUTION, MIXING AND PRESSURE DROP TESTS 3

                                                                                       )

Flow visualization and pressure drop data have been obtained from a 10-times-full-scale (10X) model of a sin 6 1e rod in a square flow channel. These data have been used to refine the spacer ferrule designs with respect to mixing turbulence and pressure drop. Additional pressure drop testing has been con-ducted using 4-pin (5X), 4-pin (1X),1-pin (2X), and 9-pin (1X) models. Testing to detemine the extent of interchannel mixing and flow distribution also has been conducted. Flow distribution in a square four rod test assem-bly has been measured. A salt solution injection technique was used to deter-mine the average flow rates in the simulated reactor assembly corner cells, vall cells, and unit cells. Interchannel mixing data was obtained for the same assembly. These data have been used to confim the flow distribution and mixin6 relationships employed in the core themal and hydraulic design. Flow tests on a mockup of two adjacent fuel assemblies have been conducted to detemine the friction effects at the perforated vall boundary. Additional mixing, flow distribution, and pressure drop data vill be obtained to improve the core power capability. The following fuel assembly Geometries will be tested to provide additional data:

a. A 9-pin (3 x 3 array) mixing test assembly, of the same bundle geometry as the DNB bundle described previously, has been con-structed to detemine flow pressure drop, flow distribution and degree of mixing present during the DNB investi 6 ations. Testing with this assembly is in progress.

b. A 16-rod assembly simulating the junction of four fuel assemblies at the corner is under construction. This assembly vill be tested to detemine the degree of mixing which occurs between fuel assemblies.

c. Several 64-rod assemblies simulating larger regions and various mechanical arrangements within a 15 x 15 fuel assembly and be-tween adjacent fuel assemblies vill be flow tested. The hydrau-lic facility for the tests is now under construction.

3323 Preoperational Testing and Postoperational Testing Themocouples are included as part of the incore monitoring system and will enable postoperational temperature measurements to be made at the entrance and exit of all 52 instrumented fuel assemblies. The results of these tests vill be compared to the results of the model tests used for design calcula-tions. 333 FUEL ASSEMBLY, CONTROL ROD ASSEMBLY, AND CONTROL RCD DRIVE MECHANICAL TESTS AND INSPECTION To demonstrate the mechanical adequacy and safety of the fuel assembly, con-trol rod: assembly (CRA), and control rod drive, a number of functional tests have been perfomed, are in pro 6ress, or are in the final stages of prepar-ation. 3-98 206 5-3-68 Supplement No. 3

3331 Prototype Testing A full scale prototype fuel assembly, CRA, and control roi drive is presently being tested in the Control Rod Drive Line (CRDL) Facility located at the B&W Research Center, Alliance, Ohio. This full-si::e loop is capable of simu-lating reactor environmental conditions of pressure, temperature, and coolant flow. . To verify the mechanical desi6 n, operating compatibility, and charac-teristics of the entire control rod drive fuel assembly system, the drive vill be stroked and tripped in excess of expected operating life requirements. A portion of the testing vill be performed with maximum misalignment condi-tionc. Equipnent is available to record and verify data such as fuel assembly pressure drop, vibration characteristics, hydraulic forces, etc, and to demon-strate control . 3-98a :i l

                                                           .<*I

5-3-68

                                                                  ' Supplement No. 3

rod drive operation and verify scram times. All prototype components vill be examined periodically for signs of material fretting, vear, and vibration / fatigue to insure that the mechanical design of the equipment meet reactor operating requirements. Preliminary test results are given in 3 2.4 3 5 3 3332 Model Testing Many functional improvements have been incorporated in the design of the prototype fuel assembly as a result of model tests run to date. For example, the spacer grid to fuel rod contact area was fabricated to 10 times reactor size and tested in a loop simulating coolant flow Reynolds numbers of interest. Thus, visually, the shape of the fuel rod support areas was optimi::ed with respect to minimizing the severity of flow vortices. Also, a 9-rod (3 x 3) actual size model was fabricated (using production fuel assembly materials) and tested at 650 F, 2,200 psi, and 13 fps coolant flow. Principal objectives of this test were to evaluate fuel rod cladding to spacer grid contact wear, and/or fretting corrosion resulting from flow-induced vibration. A vide range of contact loads (including mall clearances) was present in this specimen. No significant wear or other flow-induced damage was observed after 210 days of loop operation. 3333 Component and/or Material Testing 33331 Fuel Rod Cladding Extensive short time collapse testing was performed on Zircaloy-4 tube spec-imens as part of the B&W overall creep-collapse testing program. Initial test specimens were 0.436-in. OD with wall thicknesses of 0.020 in., 0.024 in., and 0.028 in. Ten 8-in.-long specimens of each thickness were indi-vidually tested at 680 F at slowly increasing pressure until collapse oc-curred. Collapse pressures for the 0.020-in. vall thickness specimens ranged from 1,800 to 2,200 psig, the 0.024-in. specimens ranged from 2,800 to 3,200 psig, and the 0.028-in. specimens ranged from 4,500 to 4,900 psig. The caterial yield strength of these specimens ranged from 65,000 to 72,000 psi at room temperature, and was 35,800 psi at 680 F. Additional Zircalcy-4 short time collapse specimens. vere prepared with a ma-terial yield stress of 78,000 psi at room temperature and 48,500 psi at 615 F. Fifteen specimens having an OD of 0.410 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 from 4,470 to 4,960 psig. Creep-collapse testing was perfomed on the 0.436-in. OD specimens. Twelve specimens of 0.024-in. vall thickness and 30 specimens of 0.028-in. vall thickness were tested in a single autoclave at 680 F and 2,050 psig. During this test, two 0.024-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 thickne3s specimens had collapsed after 60 days. Creep-collapse test-ing was then performed s

208 -

5-3-68 Supplement No. 3 u d

s on thirty 0.kl0-in. OD by 0.365-in. ID (0.0225-in. nominal vall) specimens T' for 60 days at 615 F and 2,140 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.kl0-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. In order to help optimize the final clad thickness, additional clad-collapse 3 testing is scheduled for 1969 using specimens fabricated to the reference de-sign fuel clad dimensions, =aterial specifications, and operating conditions. 3.3.3.3.2 Puel Assembly Structural Components The mechanical design of the prototype can panel assembly is the result of an extensive can panel design and structural evaluation program. The full-size, simulated panel designloop, functional criteria. testing noted in 3.3.3.1 is expected to verify can Prototype static and dynamic load testing is underway to andverify seismiccanloads. panel structural adequacy for vibration, handling, operation, In the mechanical design of the spacer grids, particular attention is given to the ferrule-to-fuel-rod contact points. Sufficient lead 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 vill dem-onstrate their adequacy to perform within the design requirements. 3.3.3.k Control Rod Drive Tests and Inspection 3.3.3.L.1 Control Rod Drive Developmental Tests The prototype rack and pinion, buffer seal drive is under development at the B&W Research Center, Alliance, Ohio.

                                 ^

Wear characteristics of critical ccmponents, such as sleeve bearings, pinien and rack teeth, snubber piston and sleeve, etc., during tests to date indicate that material compatibility and structural design of these co=ponents vill be adequate for the life of the mechanism. Subsequent to completion of the development program, the complete prototype control rod drive vill be subjected to environmental testing under simulated reactor conditions (cxcept radiation) in the Control Rod Drive Line (CRDL) Facility at Alliance. be limited to, the following: Environmental tests vill include, but not necessarily Operational Tests Operating speeds. Temperature profiles. 209 3-100 5-3-68 Supplement No. 3

                                                                                                                                 ^

l l Trip times for full and partially withdrawn control rod assemblies (CRA) for various flow-induced pressure drops across the CRA. Life Tests I (With internals assembled to mHmm nisalignment permitted by drawing dimensions and tolerances.) No. of Partial Stroke Span of Control Rod from " Full In" 3 , Stroke Cycles Length Position to be Tested 1550 83" From 56" to 139" 5400 50" From 71" to 121" 8500 25" From 114" to 139" 8500 13" From 126" to 139" No trip cycles 500 139" From 0" to 139" Misalignment Tests 100 full strokes and 100 full stroke trips with internals tolerances altered to 1 5 times = H m m allowable misalignment. Coupling Tests Complete check of coupling operations after testing. The cycles above meet the total test requirements of 5,000 full strokes and 500 trips. 'Ihe assembly will be completely disassembled and inspected at various B&W facilities after completion of environmental tests. 3.3.3.4.2 Control Rod Drive Control System Developnental Tests A control rod drive motor control unit has been built in breadboard form. Following the testing of the breadboard version, prototype circuits for plug-in modules will be designed and tested. Testing will consist of bench testing, life testing, and determining the effects of simulated failures. The s % 1ated-failure testing will be designed to verify the safety analysis. The control rod drive control system will be tested in conjunction with the con-trol rod drive motor control to insure proper operation. Simulated failure test-ing will also be performed on the combined system to insure that protective re-quirements are being met. The position indicator' ard limit switch subsystem has been built in prototype form and life-tested mechanically under expected environmental conditions. Further testing, both mechanical and electrical, will be done under expected environmental conditions at the B&W Research Center. Characteristics to be determined will include accuracy, repeatability, linearity, short term stability, and long term stability. 3-101 2!0 , S-3-68 Supplement No. 3

t 3.3.3.4.3 Production Tests 'i 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 perfor=ed 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. Partial and full stroke cycles. Partial and full stroke trip cycles. Control system production tests will be performed as described in the follow-ing paragraphs. The finished hardware vill be systematically operated through all of its oper-ating modes, checked over the full range of all set points, and checked for proper operation of all patch plugs. This will check completeness and proper functioning of viring and components. The operating modes to be checked vill include such things as automatic opera-tion, manual grvup operation, trim or single CRA operation, position indication of all CRA's, travel limit on all CRA's , trip circuit operations , IN command, OUT command, etc. The trip circuit or circuits will be tested by repeated operation. The over-all trip time vill be measured. The accuracy and repeatability of the position indication and limit switch sys-tems will be tested. Power supply tests will be performed to deter =ine the upper and lower operating voltage and to prove immunity to switching transients. Fault conditions vill be simulated to prove that no unsafe action results frcm defective components, circuits, or viring. Ability to detect unsafe fault con-ditions at the operating console vill be determined. Typical of faults to be simulated are 211 3-102

a. Defective limit switch or circuit, b.- Improper CRA grcup patch,

c. Defective patch plugs.

d. Defective group sequencer.

e. Defective clock.

f. Defective automatic control' signal.

g. Defective command line.

h. Defective fuses..

i. Defective single CRA control circuit or switch.

j. _ Defective pcVer supply.

k. Defective motor translator.

1. Defective motor cable.

m. Defective position transmitter.

The finished hardware vill be visually inspected for quality of workmanship. This inspection vill include an examination of the enclosure, cable entrances, dust-tightness, maintenance features, drawers and cable retractors, fasteners, stiffeners , module mounts, wire harnesses, and other similar details. 3.3.h

                               ~

INTERNALS TESTS AND INSPECTIONS The internals upper and lover 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 vill indicate areas of gross flow mal-distribution and allow verification of vessel flow-pressure drop computations. In addition, the test results vill 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 CRDL tests described in 3.3.3.k. These test results, when correlated with the. internals-guide tube final design, vill 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 vill not include the effects of neutron flux exposure.

     ' After completion of shop fabrication, all internals components vill be she,-

fitted and assembled to final design requirements. The assembled internals components will be installed in a mockup.of the as-built reactor vessel for final shop fitting'and align =ent of the internals for the mating fit with the reactor vessel. Dum=y fuel and CRA's will be used to check out and insure that smple clearances exist between the fuel and internals structures guide tubes to allow free movement of the CRA throughout its full stroke length in various core locations. Fuel assembly mating fit will be enecked at all core

    ' locations. The dummy fuel and CRA's vill be identical to the production.com-ponents except that they will be manufactured to the most adverse tolerance space envelope; even though the assembly weights will be representative of the production materials.

units, the du=my components vill not contain fissionable or poison s Internals shop fabrication quality control tests , inspection, procedures, and methods vill be similar to the pressure vessel tests described in detail in x k.l.k. 2f2 3-103- ?

            ,                c         ,    -

A During refueling outages after the reactor, vessel head and the. internals. plenum assembly have been removed, the vent valves will be accessible for

       . visual.and mechanical inspection. . A remote inspection tool will be pro-vided to engage with the previously mentioned valve disc hook or eye. .With

the aid of this tool, the valve disc can be manually exercised to' evaluate the disc freedom. The hinge design will incorporate special features, as described in'3.2.f.l.2(h), to minimize the possibility of valve disc motion

                                  ~

impairment during_its service life.. Remote : installation and' removal of the vent valve assemblies will be per-

        ' formed with the aidtof another tool which will include unlocking and operating features for the wedge ring. -This handling tool design will be functionally
        ' developed' and tested on a full-size mockup of the vent valve installation         y configuration ~ prior to valve manufacture.
       ' With the aid 1of the. above _ described inspection tool, a visual inspection' of the valve. body and disc sealing faces can be performed for evaluation of ob-served-surface irregularities.
                                      ~

The valve. disc, hinge shaft, shaft journals (bushings), disc journal receptacles, and valve body journal receptacles will be designed to withstand without failure the internal and external differential pressure loadings resulting from a loss-of-coolant accident. These valve materials will be nondestructively tested and accepted in;accordance with the ASME Code III requirements for Class A pressure vessels. The hinge materials will be selected on the basis of their corrosion resistance, surface hardness, antigalling. characteristics, and compatibility with mating

       -materials in the reactor coolant environment.

A remote inspection.of hinge parts is not planned until such time as a valve-

                               ~
     ~
       . assembly is removed because its free-disc motion has been imparied.      In the'un-1ikely. event that a' hinge part should fail during normal operation, the most significant indication of such a fialure would be a change in the free-disc
     - motion as a result of alteredl rotational clearances.
                           ^
                                    ^ h f][                       ~6-5-68

3-10ka. supplement No. 4

3.k REFERENCES (1) Putnam, G. E. , TOPIC - A Fortran Program for Calculating Transport of Particles in Cylinders, IDO-16968, April 1964. (2) Avery, A. F., The Prediction of Neutron Attenuation in Iron-Water Shields, AEEW-R125, April 1962. (3). Bohl, H., Jr., et al., P3MG1, A One-Dimensional Multigroup P-3 Program for the Philco-2000 Computer, WAPD-TM-272. (h) Bohl, H., Jr. and Hemphill, A. P., MUFT-5, A Fast Neutron Spectrum Pro-gram for the Philco-2000, WAPD-TM-218. (5) Armster, H. J. and Callaghan, J. C., KATE-1, A Program for Calculating Wigner-Wilkins and Maxwellian-Averaged Thermal. Constants on the Philco-2000, WAPD-TM-232. (6) Marleve, O. J. and Suggs, M. C., WANDA-5, A One-Dimensional Neutron Dif-fusion Equation Program for the Philco-2000 Computer, WAPD-TM-2hl. (7) Honeck, H. C., THERM 0S, A Ther=alization Transport Theory Code for Reac-tor Lattices, BNL-5826. (8) 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. (9) Marlove, O. J., Nuclear Reactor Depletion Programs for the Philco-2000 Computer, WAPD-TM-221. (10) Lathrop, K. P., DTF-IV, A FORTRAN-IV Program for Solving the Multigroup Transport Equation With Anisotropic Scattering, LA-3373. (11) Joanou, G. D. and Dudek, J. S., GAM-1: A Consistent P 1 Multigroup Code for the Calculation of Fast Neutron Spectra and Multigroup Constants, GA-1850. (12) Baldwin, M. N., Physics Verification Experiments, CORE I, p28 and Initial Conversion Ratio Measurements, BAW-TM 45h. (13) Clark, R. H. and Pitts, T. G., Physics Verification Experiments, Core I, BAW-TM h55. (14) Clark, R. H. and Pitts, T. G., Physics Verification Experiments,. Cores II and III, BAW-TM h58. (15) Spinks, N., "The Extrapolation Distance at the Surface of a Grey Cylin-drical Control Rod", Nuclear Science and Engineering 22,, pp 87-93, 1965 (16) _ Clark, R. H., Batch, M. L., and Pitts, T. G., Lumped Burnable Poison Program - Final Report, BAW-3h92-1. -(17) Neuhold, R. J., Xenon Oscillation, BAW-305, 1966.

                    -ll 3-105

(32) Worthing, A. G. and Geffner, J., Treatment of Experimental Data, John Wiley & Sons, Inc., New York, 1943. (33) 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. (34) Zuber, N. and Findlay, J. A. , Average Volumetric Concentrations in Two Phase Flow Systems, Presented at the ASME Winter Meeting, 1964 To be published in the ASME Transactions. (35) Maurer, G. W. , A Method of Predicting Steady-State Boiling Vapor Frac-tions in Reactor Coolant Channels, Bettis Technical Review, WAPD-BT-19. (36) Baker, O., Simultaneous Flow of 011 and Gas, 011 and Gas Journal, Vol, y,3, 3 pp 185-195, 1954. (37) Rose, S. C., Jr., and Griffith, P., Flow Properties of Bubbly Mixtures, ASME Paper No. 65-HT-38, 1965 (38) 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. (39) Bergles, A. E. and Suo, M., Investigation of Boiling Water Flow Regimes at High Pressure, NYO-330h-8. February 1, 1966. - (k0) Notley, N. J. F., The Thermal Conductivity of Columnar Grains in Irradi-ated UO2 Fuel Elements, AECL-1822, July 1963. (k1) Lyons, M. F., et al., UO 2 Fuel Rod Operation With Gross Central Melting, GEAP h264, October 1963. (42) Notley, M. J. F., et al., Zircaloy-Sheathed UO 2 Fuel Elements Irradiated at Values of Integral kde Between 30 and 83 w/cm. AECL-1676, December 1962. (k3) Bain, A. S. , Melting of UO 2During Irradiations of Short Duration, AECL-2289, August 1965 (4k) Notley, M. J. F., g g,., The Longitudinal and Diametral Expansions of UO2 Fuel Elements, AECL-21L3, November 1964. (45) Duncan, R. N., Rabbit Capsule Irradiation of UO ,2 CVNA-lh2, June 1962. (46) Lyons, M. F., el al_., UO 2 Pellet Thermal Conductivity From Irradiations With Central Melting, GEAP-462h. July 1964. (47) McGrath, R. G., Carolinas-Virginia Nuclear Power Associates, Inc., Re-search and Development Program, Quarterly Progress Report for the Period April - May - June 1965, CVNA-2h6. 1 215  ! 3-107

I (64) Nonuniform Heat Generation Experimental Program, BAW-3238-13, July 1966. (65) Clark, R. H., Physics Verification Experiments, Cores IV and V, BAW-TM-178, September 1966. (66) Clark, R. H., Physics Verification Experiment, Core VI, BAW-TM-179, December 1966. f (67) Clark, R. H. , Physics Verification Experiment, Axial Pcver Mapping on Core IV, BAW-TM-255, December 1966. (68) Larsen, P. S. , et al. , DNB Measurements for Upwards Flcv of Water in an Unheated Square Channel with a Single Uniformly Heated Rod at 1600-2300 psia, Proceedings of the Third International Heat Transfer Conference, August 1966. 216 3-109

                                                                                      - - _ - ~

35 NUCLEAR STEAM SUPPLY SYSTEM REVISIONS Arkansas Power & Light Comparrf is planning to adcpt the following revisions to the Nuclear Steam Supply System:

a. Roller nut drive rod mechanism. This has been submitted with Three Mile Island No. 2 PSAR.

b. MLrk B. Fuel. This has been submitted with the Duke FSAR.

c. Lump Burnable Poison Assemblies. This has been submitted with 13 the Duke FSAR. -

d. Axial Power Shaping Rods. 'Ihis has been submitted with the Duke FSAR.

Information pertinent to these changes will be included in cur FSAR. l 3-110 C 10-31-69 C# 217 Supplement No. 13

2000 I 1800 1

         $ 1600    t N      8 CD   Y   IH)0 5

c \ x l200 N O N .

g First Cycle 1000 N O 800 \ \ 3 N N

       -                        N   \-                                     N 5    600                                                                  N
    . 2                                   N                                       '

N

    =  a                                          N
    =  m S

w>0 - N \ n g g Equilibrium Cycle

                                                             )N                              N R
    =

a 200 ' A N 0 E O 50 100 150 200 250 300 350 400 Core Life, Effective Full (Rated) Power Days 5 n

3 . E i  ?

Axial Power Profile for 55% insertion is shown on i Figure 3-3. ;I i 1

                                           /~

O A / - 1. 6 e s.Na. j / l5 == s It g 1.4

1. 3 10 20 30 40 50 60 70 80 Rod insertion, 7.

AXlAL PEAK TO AVERAGE POWER VERSUS XENON OVERRIDE R00 INSERTION figure 3-2 219

1.8

                   /       N i4           >

l.2 i - 1 .0 / N ' Y

            /

A \ \ g 0.8 g E 0.6 I 'G

• 04 3

0. 2 \3 0

144" h 10 20 30 40 59 60 70 80 90 100 110 120 130 140 150 Distance from Bottom of Active Fuel. in.

AXIAL POWER PROFILE. XENON OVERRIDE RODS 55 PER CENT INSERTED Figure 3-3  ; 220

       +2 S
       .I                                                     /

x 580 F j / u s e E. 7 8 1 #

                                                                   /

0 S ~ -

                                   ,  7              -
     "                                          68 F
                          /

S L -l i /

                /
           /
       -2 0    2   4     6      8    10    12    14    16   18    20    22
                                                                ~

Moderator Boron Concentration, ppm boron x 10 i l l l H0DERATOR TEMPERATURE COEFFICIENT VERSUS BORON CONCENTRATION 2 Figure 3-4

        +2
   ,    +1                                                  #
    -o                                              7 8                  2000 ppm              p   #        1500 ppm 5:                           ,
                                   -'               ,            -=

7 / p # j-1000 ppm b Mb' f- 500 'p7% 3  :  :  % N 2 E N s N \ E -I , g Mo Boron

         -2 h

100 200 30 0 400 500 600 Moderator Temperature, F MODERATOR TEMPERATURE COEFFICIENT VERSUS MODERATOR TEMPERATURE AND VARIOUS BORON LEVELS , Figure 3-5

I l i i 100 N

                      \

. \

                        \

I't u 7 4 u a \ f _ g a \

          . H                t g s

V ^ 5. u ak/k

    $j                          \
         ~                       \
             %                     \
                                    \

1

                                      \

8.M ak/k .\

                                         \

0 0 1 2 3 4 5 6 7 Time, see PCR CENT NEUTRON POWER VERSUS TIME TOLLOWING TRIP Figure 3-6 223 .

2.8 2.6 - 2.4 - 2.2 -

2. 0 -

l.S - N # N Lower Core fs

                                                                                                                                     's                  '
                                                                                                                                ,'         g           j
                                                                              ' s N 1.6 -                                                                    '

A f s / \ PN ~ fs j / ', / \

                                                                                                                                                     /

l'" N ' f g / j -

                                                                                                                         /              _ . .            _ _ _ _ _

l.2 -

                        ;                                                                     y                   ,-   "j \ %                            

o \ e N

                                                                                                                                           ' I\

1. 0 - ,, / \ / #

N - "/ s \ 3 [ y \ / \ /

                        '                               's /                 \

f \ / \ / \

                                 .8 -             Up,er Core                                                                 '

c s,/ s s, E 6- s. k g 4-

                     =O EE
                     , rn
                                 .2 -
                     =

o o ig e n a e u op 0 1 2 3 4 5 Cg Time (T), days

                     ;g                  Notes:
                        ,                l. Power. Ratio taken 36 in. f rom top and bottom of active f uel.

m g Case I - No temperature eteration. I 1.aio0 F. o Case 2 - Temperature iteration with T = l.400 F. e fuel C Case 3 - Temperature iteration with 7 - 900 F. m p fuel io g 2. Oscillation initiated at T 2 days. o o u t .

                                                                                                            ~

o I

2. %

2.2 - 2.0 - Lower Core

1. 8 - --

j

l. 6 - \  !

i.s - l i.2 - P/P I.o -

      .8 -                                                                      I
      .6-         ,I
      .4-                     s,                                          ,,

2

        ,_           Upper Core
      .0                      ,                  ,                 ,

0 1 2 3 Time (T). days Notes:

1. Power Ratio taken 36 in. from top and bottom of active fuel.

Case I - Temperature iteration with 7,g l.400 F. Case 2 - Temperature lieration with 7 - 900 F.

2. Oscillation initiateo at T 300 days.

l I l l EFFECT OF FUEL TEMPERATURE (DOPPLER) ON XENON OSC '.LATIONS . MEAR END 0F LIFE 225 Figure 3-8

2. 6 -

2. 4 - j eN 24 -

Upper Core l

                                                                             /-
                                                                                    \ \

2. 0 -

       , . . - ,,                                                         ,/            \

y

                         \                                                                                                             *

1. 6 - \ 2 i l.4-

                 \           \                                         I g

P/P \ \ , I

                   \                              ~                  l

1. 2 - j l \.

                    \e ,     -                      s
                                                           /                                                 t

1. 0 - I ' v, '/

                                \                       Y        I                                                  \
         .s -                    \                              l                                                 \
                                  \                            l

{

         .6 -                                                 /
                                   '\

J s / \ g_ \ / \

         .2-Lower Core         N
                                          \           /
                                                        /                                                                    g
                                            %d                                                                                 N s,
         .0                                                                      ,

O I 2 Time (T), days Notes: 1 Case I - Divergent oscillation (without temperature iteration). Case 2 - Power ratio variation with control (without temperature iteration).

2. Oscillation initiated at T . 200 days.

CONTROL 0F AIIAL OSCILLATION WITH PARTIAL RODS Figure 3-9

100

 "    90 -

E Finite Sample - 80 - OI' " Id'"** 2 2 a. 8

 ;    70 -

2 E E 60 - 50 - , , , , 1.0 1.2 1.4 1. 6 1.8 2.0 DNB Ratio POPULATION INCLUDED IN THE STATISTICAL STATEMENT VERSUS DN8 RATIO Figure 3-10 227 I

1.6 1.7 l'  : I I I l C P/P = 1.70 (Partial Rod 1.6  ! , insertion) 1.5 , f - "T- ' P/ P = 1. 50

1. 4 ) l s. (Modified Consine)

    ..,            I                   / \              l            ;~
    ..,           /                   /         \                      i 1.1                             '
                                                                          \

1.0 / I ,

 $ 0.9                        f I

0.8 /

                                                                                 \

I

                          /

K \ O.7 / I ( 0.6 1 0.5 1 \ 0.4 / f' l I T

                                                                                       \
                                                                                            's. .,

0.3 I Fuel Midplane 0.2 Core  : Core 3 0.l Boitom Top , c I44" " 0.0 0 20 40 60 80 100 120 140 10 30 50 70 90 110 130 Distance from Bottom of Active Fuel, in. POWER SHAPE REFLECTING INCREASED AXIAL POWER PEAK FOR 144-INCH C' ORE Figure 3-11 2}h.

1.90 1.85 . l.70 . l.60 _

                       \                     l'neA(Design)

N 1.50 ., N Line B (Nominal - Maximum Calculated) 1.40 s Line C (Troical True Distribution)

1. 30

l. 20

     '                                             N 1.10   -

p% N I I 1. 00 . N ES 0.90 - (1) Line A . i.79 x Line B ie s

     "_ d   0.80     (2) Line C is for illustration only              's I              (3) Line 5 is based on detailed data                  N w                                                                         N 0.70 -         from a rod by rod printout of a                                               i g                       !

PDQ ( two dimensional power and 0.60 - g flux calculation for worst time 0.50 _. encorelife). 0.40 . N N

0. 30

                                                                                                 \

0.20 . 0.l0 , 0 10 20 30 40 50 60 70 80 90 100 Percentage of Fuel Rods with Higher Peaking Factors Than Point Values. '. DISTRIBUTION OF FUEL ROD PEAKING Figure 3-12 229

l E 180 E 160 [ Line 1 o Fah Nucl ear = l.85 x / , 5 I

                                                                /

2 Y) / Fan N ea 3 1,79

/

3 [ Maximum 60 j / Overpower oc C o go - 3 1 2

                 /       /
                               /                          l x   M                   '

l A o i l 100 102 104 106 108 110 112 114 116 118 120 Rated Power (2.tl52 MWT),f. I POSSIBLE FUEL ROD DN8'S FOR MAXIMUM DESIGN CONDITIONS - 36,816-ROD CORE Figure 3-13 I l 2.30

10 0 I g" 80 I a l H o 70 ' '

   =

8 5 2 60 1- / I

• so I

o l [

 ~
                                                           /

Fah Nuclear a 1,79 30 ' s S Maximum 20 2 b verpower i O T l 10 -

                      /                               I O                                               I 100 102    10 4 106     108   ll0    112   114   l16     118   120 Rated Power (2.452 MWt). 7 POSSIBLE FUEL R0D DNB'S FOR MOST PROBABLE CONDITIONS - 36,816-ROD CORE Figure 3-14 231

(I-P) (P) O.i 0.9 0.01 , 0.99 i

               \
                \
                 \
                  \
                    \

0.001 ' 0.999

                                   ' l i tt f. '

k

           \                  \
            \                  \
               \                 (
                 \                \
                   \

0.0001 , s 0.9999

                           's              i F

100%[j \ q kg

                                   \             \
                                         \            \

0.00001 ~ 0.99999 0 10 20 30 50 60 14 0 70 80 90 100 Percentage of Rods wi th a Lower Value of P, % DISTRIBUTION OF POPULATION PROTECTED P, AND l-P VERSUS NUMBER OF RODS FOR MOST PROBABLE CONDITIONS Figure 3-15

1.6 1 Design Overpower (l147.) 1.5 o.

                                  /
          =
         ,    I.4 2             1. 38 ,,_

E E o. E t 1. 3 a. E 5 1. 2 3

• t o

m e 1. I

          ?

cc E l.0 _ _ _ _ _ _ _ _ 100 110 120 130 140 150 Rated Power (2.452 MWt). T. DNS RATIOS (BAW-168) VERSUS REACTOR POWER Figure 3-16 2,

     +N

• 79

     + 22                                                               81
     + 20                                                  7            83
     + 18                                                               85
     . ,.                                        >                                  =

Design 88 e " { Overpower (114*,)

     + 14                                  '

93 2* o5

     + 12                                                               94       3g

s

   , + 10                                                               98       8$

a - t ge d ' 0 ' 102 fa EO

     +6             '                                                    108     $,

r

                                                                                    ~
     . 4         F                                                      116        g.

5

     *2       '

129 j Quality  % 0 144 Subcooleo

       -2
       -4
       -6 100   110     120          130          140           :50 Rated Power (2.452 MWt), t MAXIMUM HOT CHANNEL EXIT QUAllTV VERSUS REACTOR OVERPOWER Figure 3-17
                }'

4.00 l I UO2 Melts r u I i i I 5 I e

      %   3.00 l

_i T 1 3 4 I

      $                                                                                                                                                                 )
      ~

2.00

      ~                                                                                                      /                                    I I

w - 1 I Data Based on CVMA - 142 (June 1962) l 1.00 l A 0 1000 2000 3000 4000 5000 Temperature. F l l THERMAL CONOUCTIVITY OF UO 2 235 Figure 3-18

6000 l UO2 Melting Temperature l 50

• _ _ _ _. _ _ _ ____. _ _ p__

l 1 l w 4000 ' 5 l

   $                                                                    143%

D Power  ! 3000 a l 7 / I s l 1 l 5 l I g 100f, Power i I a m x i I l

        .                                                          !            l l

114% Power l l 1000  ! lh I I I I O  ! O 5 10 15 20 25 Linear Heat Rate, kw/ f t FUEL CENTER TEMPERATURE AT THE HOT SPOT VERSUS LINEAR POWER Figure 3-19 236-l

70 . Gaussian Distribution 60 .

                                        ~
     .5                                        -

E - 2 40 . u - 30 . 4 - N -

                            ~

10 . o

              - c-
                     /
                        -r-                                 s i   i     i    i      i       i    i      i
                                                                - -e r

0.6 4 i 0.8 1.0 1.2 1.4 1.6 i 1.8 d+ E C NUMBER OF DATA POINTS VERSUS.p /Ys EC Figure 3-20

   ~                              '

237

1. 025 - F ,,

1.020 - a 1.015 - t 2 1.010 _ j l.005 - F Q d q 1.000 60 70 80 90 100 -

c

       .995 -
       .990 -
       .985 -                                        A 1

Population Protected. % I H0T CHANNEL FACTOR VERSUS PER CENT POPULATION PROTECTED Figure 3-21 238

 ~

l I 100 Infinite Sample

         -100% Confidence 80 -

w Fini te Sample - f 90% Confidence B E 80 - 2  : e ' .9 Finite Sample - 70 - 99% Confidence 5 a. 60 - 50 e i

• i 1.0 l.2 1. 4 1.6 1.8 2.0 Burnout Factor, DNS Ratio I

BURMOUT FACTOR VERSUS POPULATION FOR VARIOUS CONFIDEMCE LEVELS I Figure 3-22 239 l

                                                                                                                                                                   ...m-w ns

200 180 ' I 160 I g l l  ; 140 g r l l l d 120 I E {O 100 I

                                                   )            I 4

I

   .5     80

• 4 l E 60 l

( I w /  ! 20 s / i I O l.00 1.05 1.10 1.15 1.20 Fraction of Rated Power (2.452 MWt)~ RODS IN JEOPARDY VERSUS POWER / Figure 3-23 2t;0

10 - -

                                                                     ~

0 i e i i i i

                             -r                          Tr i i i Dn i i i         e    i i      i i      e    i i
                  .2     .4           .6           .8           1. 0.          1.2     1.4       1.6      1.8     2.0>2
        ,)0
                                                                   +E l 'C WAPD-188 500 psia Data and BAW-168 20 -

l 10 - _ _ 0 i i i i i i e i iii g iiii7i i

    , ( 2) 0     .2     .4          .6             .8          1.0           1. 2     1. 4      1.6     1.8      2.0>2 7

j sg/ +C g WAPD-188 600 psia Data and BAW-168 g 20- ~ T 10- _ O i i i i a i i a a i i i < a T1- a m fk i i i < i i 0 .2 .4 .6 .8 1.0 1. 2 1.4 1.6 1.8 2.0>0 (3) s.g hC WAPD-188 1000 psia Data and BAW-168 10 - _ 0 i i i e i i i r-i , i i i i i n Ih i . i i i i, 0 .2 .4 .6 .8 1.0 1. 2 1.4 1.6 1.8 2.3>2 (4) 4E/4C ., WAPD-188 1500 psia Data and BAW-168 RATIO 0F EXPERIMENTAL TO

 '~

CALCULATED BURMOUT HEAT FLUX 241 rise,e 3-24

     .                                                                                                                                      i l

l J l 10 - 0 F

                                                                                     . .i . .         . i i       -

i i i i i i i i i i i i i i 0 .2 .4 .6 .8 1.0 1. 2 1. 4 1.6 1.8 2. 0> 2 (5)

                                                            + l+

E C WAPD-188 1750 psia Data and BAW-168 70 _ l l 60 _ SC - 5 - O o

       ; 40        _

n 5 -

       =

30 - 20 . 10 - _

                                                   -                            -                                                           1
                                                 ~

0 r, - r, r, I i i i i i . I. . . . i i i i i i i i i i i 0 .2 .4 .6 .8- 1.0 1. 2 1.4 1.6 1.8 2.022

                                                          .t. l+

( S) E C WAPD-188 2000 psia Data and BAW-168 RATIO OF EXPERIMENTAL TO CALCULATED BURMOUT HEAT FLUX Figure 3-25 u 242 l l _ . _ . ~ . .

10 . O 0

                        . 2__
                                   .4
                                           .6
                                                      .8 1.0

1. 2

                                                                                   . M .4M.1.6 1.8

2. 0> 2 (7) E C WAPD-188 2250 psia Data and BAW-168 10 _

0 0 i

                       .2 i     i i
                                 .4 i
                                           .6
                                              . i i
                                                      .8
                                                              .      .     . i         .      .l l O.. r-f.l .                     . ,
   ,,,                                                            1.0         1.2           1. 4        1.6           1.8        2.0>2 e

(8) E C o WAPD-188 2500 psia Data and BAW-i68

   )*  10    -

E o 0 i

                       .2 i    i     i
                                   .4 i
                                           .6 i     i i
                                                      .8 i

1.0 i i 1.2

                                                                                  .     . .          . i 1.6 d.1.8  .    .  %.        .

1. 4 2.0>2 (9)

                                                                       *E # C WAPD-188 2750 psia Data and BAW-168 20 _
                                                                   ~                        ~

10 - _ _

                                                          ~

0 4 - - I i i i a i e i iie i . i i i i i e i i 0 .2 .4 .6 .8 1.0 1. 2 1. 4 1.6 1.3 2.0>2 E C (10) AEEW-R-213 560 psia Data and BAW-168 ' RArl0 0F EXPERIMENTAL TO l 2h3 CALCULATED BURNOUT HEAT FLUX l Figure 3-26

10 - 0 i i i i i i . . J . i i n i i n n ,,n i i i i i i i i 0 .2 .4 .6 .8 1.0 1.2 1. 4 1. 6 1.8 2. 0> 2 (ll) *E # C AEEW-R-213 720 psia Data and BAW-168 50 - 40 - c 30 -

     ,5                                     _              _

o . 5* m- -

                                                                   ~
     =

10 - __ _ 1 n r, n 1 0 i i i i i i i i i i . . . i i e i i i i . 0 .2 .4 .6 .8 1.0 1. 2 1.4 1.6 1.8 2. 0 > 2 (12) *E #C l i I psia a a and B W -IS8 10 - 1 ,a 0 e i i i i i i i iie i e i e i i i i i . 0 .2 .4 .6 .8 1. 0 1. 2 1. 4 1. 6 1.8 2. 0> 2 (13) *E # C AEEW-R-213 1300 psia Data and BAW-168 RATIO OF EXPERIMENTAL TO '~- CALCULATED BURNOUT HEAT FLUX 244 Fio <e 3-27

         %~                                        _

20 ' 10 - O . , , , . , , , , , , , , . . . . . . . . 0 .2 .4 .6 .8 1.0 1. 2 1.4 1.6 1.8 2.0>2 4 h (14) EC 5

      '                         AEEW-R-213 1500 psia Data and BAW-168 o

m 10 - E

      =

r, r,, r, n 0 " ' ' ' ' 8 5 I g I a 3 i s  : i n , 0 .2 .4 .6 .8 1.0 1. 2 1.4 1. 6 1. 8 2. 0> 2 ( l'5) 4EYC Columbia 500 psia Data and BAW-168 10 - O i i . . . . . h, h, , i r, 0 .2 .4 .6 .8 1.0 1. 2 1.4 1.6 1.8 2. 0> 2 (16) 4 / g, Columbia 720 psia Data and BAW-168 RATIO OF EXPERIMENTAL TO CALCULATED BURMOUT HEAT FLUX Figure 3-28

                 .         245

50 - 40 - - 30 - - 20 - 10 - 2 - 5 - o i E m o 0 i i i i i i i e i i i i i i i i e i i

   ;     O       .2     .4         .6       .8         1.0       1.2        1.4     1. 6   1.8     2. 0> 2 n

3 (l7) x />E 'C Columbia 1000 psia Data and BAW-168 10 - O i i i e i i i i i i e i  : i i i i i e i i 0 .2 .4 .6 .8 1.3 1. 2 1.4 1. 6 1.8 2.0>2 (18) 'E 'C Columbia 1200psiabataandBAW-168 RATIO 0F EXPERIMENTAL TO CALCULATED BURNOUT HEAT FLUX

 \

_ Figure 3-29

20 10 - -

                                                              ~

0 . , , , . . . . , , .?. , , , , , . , , 0 .2 .4 .6 .8 1.0 1.2 1.4 1.6 1.8 2.0>2 ( 19) (E C Columbia 1500 psia Data and BAW-168 60 - a _ E o

 $       %-                                                         ~

s

 =

M-20 - 10 - 0 . . . r n.

                               . .       n. . . .

1 . , , . . . . . . . 0 .2 .4 .6 .8 1.0 1. 2 1.4 1. 6 1.8 2. 0 > 2 (20) c EC Argonne 2000 psia Data and BAW-168 3 RATIO OF EXPERIMENTAL TO i CALCULATED BURNOUT HEAT FLUX Figure 3-30 v 247

10 - 0 . . . . . .E . '. . . . . . . . . . . . . 0 .2 .4 .6 .8 1.0 1.2 1.4 1. 6 1.8 2. 0> 2

                                                           *E        C (21)        ,

B&W 2000 psia Data and BAW-168 10 -

   .5                                              _

O - _ I o 0 . . . . . .R . . .

                                                              . T.l    .   .     .   ,    ,    .    .   . .         i g          0    .2      .4       .6       .8          1.0          1. 2     1. 4      1.6       1.8     2. 0> 2 1                                                        4E/4C 5       (22)

Euratom 1000 psia Data and BAW-168 10 ~ 0 F i i i i e iii ie i i i . . i i , ,

 -           0       .2       .4      .6        .8        1.0          1. 2      1. 4     1. 6      1.8     2.0>2

( 23) E C Euratorm 1500. psia Data and BAW-168

                                                     /

RATIO OF EXPERIMENTAL TO CALClILATED BURNOUT HEAT FLUX [ Figure 3-31

20 - 10 - 0 * ' > > > > iii ,,,,,,,.,,,, 0 .2 .4 .6 .8 1.0 1. 2 1.4 1. 6 1.8 2.0>2 (24I f E'#C

• Euratom 2000 psia Data and BAW-168 50 -

   .5 2

o 40 t n 1 -

   =                                               -

30 - - 20 - _ 10 - '- 0 . . . . . .. . . . T ii... 0 .2 .4 .6 .8 1.0 1.2 1. 4 1. 6 1.8 2. 0> 2 (25) .p EC All 500-720 psia Data and BAW-168 RATIO 0F EXPERIMENTAL TO CALCULATED BURNOUT HEAT FLUX ../ 2kh Figure 3-32

80 - 70 -

                                                       ~     ~

60 - 50 -

     .5 E

o 40 - _ _ 2 5 x 30 -

g. _ _

                                                                                 ~-

10 - _ 0 i i i i i i i

                                                                                         -f" ~l fTh i    i               8     3   i    i      i i    i   i   i    i     i   5 0   .2     .4        .6        .8         1.0        1.2      1.4      1.6-     1.8       2.0>2
                                                            #E #C (26)

All 1000 psia Data and BAW-168 RATIO 0F EXPERIMENTAL TO

                      ,                                                             CALC'Ji!TED BURNOUT HEAT FLUX l

250 Figure 3-33 1

80 - 70 - 60 - 50 - h n. i 1 40 - i t i, 1 x 30 - '

                                              ~

20 - 10 - ~

                                      -                                                                        1 0          i i   i   i i   i    e i      e   i     e mbi i i       i   ,    r', O, l

0 , ,

                     .2   .4    .6         8      1.0       1.2 i

1.4 1.6 1.8 2.0>2 (27)

                                                    +E #C All 1500 psia Data and BAW-168 RATIO 0F EXPERIMENTAL 70 CALCULATED BURN 0UT HEAT FLUX Figure 3-34 a

             '221 100                                        %

90 - 80 - l 70 - 0 60 -

        .E E

g 50 - n e 40 - 30 - 20 - I 10 - _ O i i , i e i M i , , b, ,r, ,r, i i , , , , , , , 0 .2 .4 .6 .8 1. 0 1. 2 1. 4 1. 6 1.8 2.032 (28) #E#C All 2000 psia Data and BAW-168 RATIO OF EXPERIMENTAL TO CALCULATED BURNOUT HEAT FLUX

   's -

2}} Figure 3-35

126 - -

               ~
                                                     " ~

100 90 - 4! 70 - _

   . 60 -

c 8 o m 50 - 2 5 -

   =

g- _ M-20 - 10 ~ _ 0 i e iiiiiie M i i i i i i T1-f M. . i i i i 0 .2 .4 .6 .8 1.0 1. 2 1. 4 1.6 1.8 2. 0 > 2 (29) PE#C All 1750-2750 psia Data and BAW-168 RATIO 0F EXPERIMENTAL TO CALCULATED BURNOUT HEAT FLUX Figure 3-36

                        }

10 - 0 , , . , , , i , 0 .2 ;4 .6 .8 1.0 1. 2 1. 4 1.6 1.8 2. 0 > 2 ( 30) 4E#C Euratom Chopped-Cosine 1000 psia Data and BAW-168 10 - 0

                  . . .         . . i i ,              ,     . .        . .          . i i i         .   .   . .

0 .2 .4 .6 .8 1.0 1. 2 1. 4 1.6 1.8 2.0 >2 O ( 31) 5 / o +E C a. g Euratom Chopped-Cosine 1500 psia Data U 10 - and BAW-168 2 0 . . i i i i i i e s i i e i e a i 4 i i i 0 .2 .4 .6 .8 1.0 1. 2 1. 4 1.6 1.8 2.0 > 2 ( 32) dE *C Euratom Chopped-Cosine 2000 psia Data 10 - and B W-168 0

                  . . .        . i i i .              . i i          i . .               . . i i i i        1 0     .2       .4         .6     .8       1.0        1.2           1.4       l.6    1.8     2.0 >2 (33) og/*C                                                                          ,

Euratom and B&W Inlet Peak 1000 psia Data and BAW-168 RATIO 0F EXPERIMENTAL TO CALCULATED BURN 0UT HEAT FLUX

(./ '" '- 7 54

10 O r, i . . i i , , i , , , , i i i i . i e i i 0 .2 .4 .6 .8 1.0 1.2 1.4 1.6 1.8 2.0 > 2 (34) e E *C Euratom and B&W Inlet Peak 1500 psia Data and BAW-168 10 - O i i . . i i . . i i i i i i i i , i e i i 0 .2 4 .6 .8 1.0 1.2 1.4 1.6 1.8 2.0>2 e

                                                      /

2 (35) +E .C Euratom and B&W Inlet Peak 2000 psia Data 2 10 - and BAW-168

 =

0 i e i i i i e i i e i i i i i i i i i i i 0 .2 4 .6 .8 1.0 1.2 1.4 1.6 1.8 2.0>2 (36)

                                                   *g/*C Euratom and B&W Outlet Peak 1000 psia Data and bad.168 10 -

0 i i i e i i a i i i i i e i i i i i i 0 .2 .4 .6 .8 1.0 1.2 1.4 1.6 1.8 2.0 >2 (37)

                                                   *E     C Euratom and B&W Outlet Peak 1500 psia Data and BAW-168 RATIO 0F EXPERIMENTAL TO CALCULATED BURNOUT HEAT FLUX 255                                                    F igure 3-38 a                     .--           e     m       ,.m   ,          - ga

10 - __ 0 r, F i , , , ,,iii,,.......ii 0 .2 4 .6 .8 1.0 1.2 1. 4 1.6 1.8 2.0>2 (38)

                                                           ,E 'C Euratom and B&W Outlet Peak 2000 psia Data and BAW-168 10 -

_- ~

                                             ~

0 , i i i i i i i

                                                  -f  ii               .iii             ..         i

( 0 .2 .4 .6 .8 1.0 1.2 1. 4 1.5 1.8 2.0> 2 (39) *E # C All 1000 psia Nonuniform Data and BAW-168 U l 4 10 - _ l l i 0 1 i e i i i iiiiii .iiii iii . 0 .2 .4 .6 .8 1.0 1.2 1.4 1.6 1.8 2.02 2 (40) #E 'C All 1500 psia Nonuniform Data and BAW-168 20-

                                    -        -     ~

10- - 0 n 1 i i i i i iiiii ie i i i i i , 0 .2 .4 .6 .8 1.0 1.2 1. 4 1.6 1.8 2.02 2 (41)

                                                            *g/*C All 2000 psia Nonuniform Data and BAW-168 RATIO OF EXPERIMENTAL TO CALCULATED BURNOUT HEAT FLUX-   l 256                                                                          I Figure 3-39  I

18 88 16

                                                               #          # 90             en Design                    /                             -

Overpower (i14%) 14 ,

                                                         /           j       93           y 12
                                                                 ,.!         96            =

10 7 l // g g 100

                                                                                    -1 y      g 8                                                                   104
                                          , e, u -

6

                                  /    /    ,/                               109 b

n

     ;                        g                                                     g     g g             "
                         /                                                   i,0 f                                              5 7 2      r'
                   / j[/                                                     126 e

E E e

              /       f/ r

• 0

                       /                                    Ouality 144 g                                                                       y Sube oled                    -
                            - - - 2.120 os i g      (Fah. l.85)
       ,g                              2.185 psig   (Fah- 1.85)
                                  - 2.185 psig      (Fah. l.79)                         g     1
       -6 100       110            120          130           140        150 Rated Power (2.452 MWt) .f.

l l

                                                                                              )

i I l 257 MAXIMUM HOT CHANNEL EXIT QUALITY r VERSUS REACTOR POWER

 ~.)

Figure 3-40

      + I4               I               l        l           l ll                                       l St below Average Assembly Flow.
      + 12        - ~ ~ Average Assembl y Flow.

ll l l i FAh . l .85 - l llFah.1.79

      +8 4 6
   +  , 4 i        a
                                                      /

b l / 2 2 l /

   $                                       /,
                       )   l    >    l      /
                                                   /
                /

faf e

                                      /( Fah , 1.79
         -2            e        ,,

7 j Fah . l.85 v/

             /                       Design Overpower (il47.)
         -8                                                                                                                            ,

I 100 110 120 130 140 . Rated Power ( 2.452 MWt). :1 HOTTEST DESIGN AND NOMINAL CHANNEL EXIT QUALITY VERSUS REACTOR POWER (WITHOUT ENGINEERING HOT CHANNEL FACTORS) Figure 3-41  !

 /

258 . l l

3.0

                                                           \
                                                             \

ll4% Power \

                                                                \

l , 2.5 130% Power \,

   . 2.0
    'S M

ed 7 ' a 6 3 i.s  ;

     $      Bubble to                                $

h

     'G     Slug S                                               }

(Gri f fith and y Rose) \ 2 l Bubble to E I.0 ' Annular g (Baker)

                                                          \

Slug to Annular (Haberstroh) \

         .5                     ,'            ',

Bubble to Slug \ - (Baker) N

3 0 10 20 30 40 l

Quality (lb vapor / total Ib), s FLOW REGINE MAP FOR UNIT CELL CHANNEL AT 2.120 PSIG

                                                                                                                                    ]

Figure 3-42 259 l l

i g Bundle Burnout Test Conditions where stable operations were observed. A Hot Unit Cell Worst Conditions

                  + Wot Unit Cell Nominal Conditions 3.0
                                                               \
                $          a ma                     a           \
                                                                  \
                                                                   \
      .                        .                                     \
       ' o_ 2.5 a
      't                           A                  A

• ua A

        $         a        O 1

2.0 - 4 _o a a E 1.5 Bubble to Slug (Griffith and

              \                                   Ro w) 1.0 Bubble to Annular (Baker)
                          /
             .5 Subble to Slug (Saker) 0         10               20           30        40 Quality (Ib vapor / total Ib), 7

,. / FLOW REGIME MAP FOR UNIT CELL CHANNEL Figure 3-43

g Bundle Burnout Test Conditions where stable operations were observed. 9 Hot Corner Channel Worst Conditions

            + Hot Corner Channel Nominal Conditions 3.0 0
                                                          \
                                                            \
                                                              \

2. 5 g

O O T

   .o

2. 0 G

  %            g     ,g    g                      g i

2 2._ l.5 N 7 3 6 O l.0 g -- Bubble to Slug Bubble to Annular + (Gri f fi th ( Bak er)

        *5                            andRose)                            >
                   /,

Bubble to Slug (Baker) I C e 0 10 20 30 14 0 Quality (Ib vapor / total Ib). f. ~J 26\ FLOW REGIME MAP FOR CORNER CHANNEL Figure 3 tM

g Bundle Burnout Test Conditions where stable operations were observed. 3 Hot Cell Worst Conditions

              +  Hot Cell Channel Nominal Conditions 3.0
                                                    \

s S

                                                      \
                                                        \
                                                          \

2. 5 g

g 8 8 e

  'S M

w, 2. 0 g * . 8 2 b 8 5 5 1.5

    ;                    g       . Bubble
   ;                                  to Slug (Griffith and                                        '

Rose) 1.0 di Bubble to Annular (Baker) f

        .5 Bubble to Slug (Baker) t 0           10            20            30         40 Quality (Ib vapor / total Ib). f.

2II FLOW REGIME MAP FOR WALL CHANNEL Figure 3-45 I I

h , Design Overpower (114%)' g 2.0 -N \

           \ \ \                    I.65 Cosine (W-3) 1.8        \        '
               \   \                1.80 Cosine (W-3) 1.6          \           \
  .9                              \              BAW-168 Design
  ~
                         \             g 1.4                     -- %-- DNBR ( l .38)
  $                                        & - - W-3 Design 1,2  -

DNBR ( l.30)

   ,!                   1.65 Cosine (BAW-168) \

0 1*0 - \

  ,                1.80 Cosine ( BAW-168)                \

E \ 0.8 - \ N l.50 Cosine (BAW-168) 0.6 -

                                                                    \    \

1.50 Cosine (W-3) 0.4 - 0.2 - 0 100 110 120 130 140 150 Rated Power (2.452 MWt). % H0T CHANNEL ON8 RATIO COMPARISON Figure 3-4 w-263

150 l g Design Power <o (2,452MWt) Kl 140

I 7

    . 130
                                   )  I b                                   I

u. l

  .t                                  i M

• 120 0 l 3 1 1

110 I y l l 2300 2400 2500 2600 Reactor Core Power, MWt REACTOR COOLANT FLOW VERSUS POWER Figure 3-47 264

4.00 I I UO Melts 2 1 i n ' 1 - L I a O 3.00 y g i

                        '\

1

          ~.

i \ l

                           'h        84W Design Value (CVMA- 142) l 0                                                   CVMA - 246             I M                                                                       #

g I

2.m k, 7 f i i

        !                           \                                      /

3 A i

                                                                      /             '

i A- / /: g I p GEAP-4624 l 1.00 ' 0 1000 2000 3000 4000 5000 Temperature. F THERMAL CONDUCTIVITY OF 95 PER. CENT DENSE SINTERED U0 PELLETS ' 2

 ~

265 Figure 3-48

                                                                                                                                                           . 1

6000

                                                                                          /

5500 - Design Overpower (114%) ,e f 100% Power 5000 UO Welting Temserature \

                                                  ---        ----             p-
                                                                                       ,/

u

                                                                         /
                                                                        / i
                                                                           /
                                                                                /

4500

                                                              /,//
 ?                                                               I 8                                                         !!

t fjh 2 4000 e f-I w 3500 Y 3000 B&W Design Value (CVNA-142)

                                 - ~ ~- Ref. 46 (GEAP-4624)
                        /
                                 ---- Ref. 47 ( CVN A-2 46)
                                        '       '          '         '      I 2500 6      8    10       12      14      16         18        20     22      24    26     28      30 linear Heat Rate, kw/ft FUEL CENTER TEMPERATURE FOR BEGINNING-OF-LIFE CONDITIONS
                                                                                                 Figure 3-49

6000 l

            -- Design Overpower (114%)
                                                                               )             /

5500 / /' l 100#. Power /

                                                                                   /
                                                                                     /

U0 Melting Tempe ature l  ! 5000 -

                                                              }    -                ----
   '.                                                  /   /       /

i / u j p /y

                                                     /

y 4 4000 A/,// 8

                            /

3500 y J 88WDesignValue(CVMA-142) [

                           - --- Ref. 46 (GEAP-462 4)
                           ---- Ref. 47 (CVN A-246) 3000 2500 6     8     10      12    14      16   18    20      22    24       26       28      30 Linear Heat Rate. kw/ f t FUEL CENTER TEMPERATURE FOR END-OF-LIFE CONDIT10NS

' Figure 3-50 267 l

100.00 i i 50.00 i l v rf.

                                                                  , n   -A                             O
            -                                                   /

7 . e 7 *

/ 8 ."

10.00 "- W i /n ,

                                            /~                                                                      E j    5.00                                ; p'     ;
                      ~

e a .i

-                                      ,/

E a b O}-

• . +

A

• l . 00 a'

~ -- p l -n  ;  :. 1 1

0. 5 '

                             ./                .                  l
                        . p                 :                        i
           -                                        .             L            !
                           /                                      la    I      i                          !

O GEAP - 4596 0.10 o . e GEAP 4314 l + AECL - 603 J 4 A CF-60-12-14(0RNL) 0 05 l l t , I l l l IN)0 l@ 1600 1800 2000 2200 2400 2600 2800 3000 3200 3400 3600 Valuestric Average Temperature. F PER CENT F13310N GAS RELEASED As A FUNCTION OF THE AVERAGE TEMPER ATURE OF THE UO FUEL 2

                                                              ,                                                               Figure 3-51 268

1.8 l

                                 ~

l

                                /       N                 P/P = 1.70 (Partial Rod
     ,                       /            \                      ,         insertioni
                           /                    ,

i

1. 5 7 g , , -T'N j- P/ P = 1. 50

                       /

f( l \ / (Modi fied Cosine)

1. 3

[ /

                                              ,       \
                                                        \

l 100 Days

                                                      %                                 ;i
     '.2
                   ,             /g/                      \                               300 Days l.l j         fp,                                                            ,

j , ,, ,,, ,

                                %_                             l \                \% /i
                  ![/             /                            l     \                   \

i' W/ / l \ \ \T 2'

               ///           /                                !           \               \\ \\

N/ / i x \\ \\

       "   '/// /                                             l                    N          \\ \\
       "     //    /                                         l s     % \\

0.4 V i s Y# 0.3

           /                                                 i                                  N      \,0 i

g F Fuel Midplane \ 0.2 Core i Bottom 1 Core \ ) 0.1 , Top \ g

                                                                                                  ;          3 I     '       '                     '               '    '

0'0  !  ! ' ' I 'I ' 20 40 60 80 100 (20 140 10 30 50 70 90 llo 130 Distance from Bottom of Active Fuel, in. AXIAL LOCAL TO AVERAGE SURNUP AND INSTANTAMEOUS POWER COMPARISONS Figure 3-52 269

                                        .c

l

                                                                                                     \

50 Design Limit 40

                                     /

i 30 1.70 BU and /// = Axial Shape. / / l.50 BU and .9 a 20 Axial Shape j 930-Day BU y [f[ and 1.70 Axial c Shape. 10 0 0 1 2 3 4 5 7 6 8 Cold Diametral Clearance, in. x 10 FISSION GAS RELEASE FOR I 50 270 ANo i.70 wax / AVG AXI AL POWER SHAPES Figure 3-53

3500 Design Limit 3000 Closed Pores l l .b 2500 1a

  . 2000        1.5 Axial Power g               and Burnup Shape.

2 --

 .E 1500    --
                              /

a-l.7 Axial Power And 930-Day Burnup Shape. l Open Pores o

                                /
        .00                              II 0         1      2    3 11 4      5     6     7      8 Cold Diametral Clearance, in. x 10 GAS PRESSURE INSIDE THE FUEL CLAD FOR i

VARIOUS AXlAL BURNUP AND POWER SHAPES

                                    }Jj                                        Figure 3-54

t l I l (___ ___

           +Mev>M 4444%v>M
          +4xH44 M M4+@4 $4                                                                 0.970         0,998 0.987    1.012         0.996          0.97               0.970       0. 976        1. 0 10
          +                          1.004 Nuclear Peakin9 Factor
                                                   /

0.998 0.998 1.015 1.010 1.0 18 1 One 1.019 c o n. i . , . . .. , x t., i l l NOMINAL FUEL RCD POWER PEAKS AND CELL EXIT ENTHALPY RISE RATIOS l ' -- Figure 3-55 1

I 1

    ~~~
  -                                                                         -~-
          %%9%+

9%%%+ 0.978 1.022 0.978 0.933 f.04 O.0 22 f.044 1.09 Buclear Peating Factor Enthaler Rise Factor MAXIMUM FUEL R0D POWER PEAKS AND CELL EXIT ENTHALPY RISE RATIOS s_ _ Figure 3-56 273

1.4 g l.3 G = 1.59 x 106 l b/ h r- f t2 - N \\

    ,2              -
                       \                 Best Fit I.1
                            \

1.0 , o Design Limit

 = 0.9 5
 ; 0.8
                                        \    \

Y Minimum DNBR .2.20

N \

0.7 , 3 i \ 0.6 \ '

 %                                                       \

E N f 0.4 g

                                                                \

g Calculated 0.2 S

                                         \

0.1 0 580 600 620 640 660 680 700 720 7u0 760 780 800 Local Enthalpy. Stu/lb CALCULATED AND DESIGN LIMIT LOCAL HEAT FLUX VERSUS ENTHALPY IN THE HOT CORNER CELL AT THE MOMINAL CONDITION Figure 3-57 274

I. 4 i

1 i i 1.3 G - 1.32 x 106 lb/ hr-f t2 _

l.2 ' l .1

                                  \          -- Best Fi t
 ,     1.0                           \     T x  0.9   -

Design Limit g a \g% 5m 0.8 \s \ e s

                                                    \

0.7 g Hinimum DNBR . l.70 O.6 3 1

   -                                                          Nh
                        -             w 3

3 / [ \

          "                                         \               \\
             /                                        \                  NNs s 0.3                                                .

s Calculated -

                                                                                 \

0.2 g y 0.1 0 580 600 620 640 660 680 700 720 740 760 780 800 Local Enthalpy. Btu /lb CALCULATED AND DESIGN LIMIT LOCAL HEAT FLUX VERSUS ENTHALPY IN THE NOT CORNER CELL AT THE POSTULATED WORST CONDITION l Figure 3-58 275

h, /- %" " " I i i e,

      ""~'

m ch m 03 a y m

                                   ==                        -
                                                                      = = = .                     .=     -

2 s'\s ,

                                                                    ; GI7 !%!                                                            s,..:

i  ?% l N ' T i i , i

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                     -    i
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                                                                                                                                         .......a....
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l _

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      ......                   \);

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no..........

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(0=418 "'"" { [ d[ B( BJ.14 s , k  ! s . w \ J  % l l

                         ,.; _                'N                  ;         hg                 i     i                           ,

u u REACTOR VESSEL AND INTERNALS - (, _ GENERAL ARRANGEllENT

                                                                                                    /

Figure 3-59 Supplement 3

FUEL ASSEMBLY l

                                      ,f i^f's , ', , ,
                 /[ '

0 o SURVEILLANCE

                                   .o                                  ' ',, [            SPECIMEN s

0; l j ,j i l O, f, HOLDER TUBE ql xl isI*!al 4l 10 g '<7' O anel 1 10(*j el LO [ J l la, l 121 l 3 l @ ! 10lale f I l @ldl0I 8! a$2l 101. e ; [O

                                                                                    /          CONTROL ROD ASSEMBLY LOCATION f
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             *l1 0l s1

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• G l // / LOCATICN p

Ol G Ol 'I@ ll@ @ G O

• O l O  :* @l @ O o O *i G 91@l
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                                                                        /
                                                                ,/,                       THERMAL SHIELD
                            ' -/
                               ' , , , , ," ,,, /

CORE BARREL REACTOR VESSEL AND INTERNALS - q CROSS SECTION. 277 Figure 3-60 Supplement 3

ca t.. .a CD CD

                     ,                  'O f}

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9 ' N N V' N i::, N w# ,  ; ;x CORE Fl.00DlHG ARR ANGEllENT 278 i Figure.3-61 l S'upplement 3 I

                                                                                        'S E                                  .

3

        /                                          /      l s,

A  % Di y x f t. T"d"kyus l _H2 e o 5 I-(

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u l 's y

                          +                            .

x xc sd

                     's u    _  J-   _P
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(- 2~/9 FIGURE 3-61-a REVISED 2 8-68

                                                                                            }

1 I tmn Em r  : :f Il g) . ,, I e I

                                                         /        '/                                 l
                                                          ~~

_43_ _ _ p,..J

                                                                                 ~~~

SPACER GRIO # d[ ~ d'J L*J i f8R8g8s8R8s8# m 9 sh r l I k= = = A 5 ',_ _ 7.---,

               ;M,_e . g-              CAN PIMIEL 1
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          == == == ==

h5 , I

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                      /
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1 cross sEcTion l 1

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1

a. ~_ .

t ., . .- e_  ?~ E'"N 1 f 1 i'- FUEL ASSEMBLY Figure 3-62

                                           }gQ

coun m I V in d.$ ' W,. /

                                                    $h  nr "
                                  ~-               Q//) W
                                                        -  ^

y,r-onmics noo -WD n

                            \                ?  ! 11 1

I ,

         ?

b  : - l UV U U VV Top VIEW l lg ORIFICE-ROD. 28l ASSEMBLY t i; Figure 3-63 (

i 1 Dervt uoTJt GEAR REDUCER m

   ~m ll u 24
    =u;
      ~7 l                 RACM MOU$1NG
      !' ((

t i! -parve SMAFT MoussNG

           ' I,
           ; /

a

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c'l 1

                      -$90CL PECE o

9 .-nE ACTOR CLOSURE HE AD j.N Qi R rn iji i 1 i CouPUNG ASSEMBLT CONTROL R00 DRIVE -

GENERAL ARRANGEMENT 202 .

Figure 3-64 Supplement 3

1

        \

I tjr: 70ft rostos [ A T r

                    .            f fQY

_. .. ~_C t _ g.{

                                                                                ; * : QlO                a p-D y    ,z inki&r!E                                 +                     -

b~ - . p; i lli i g M M 'O 't ' t. e T_ P - = x. = _z; := --- - - = ~1 - 5 rk! .. +. 7p "me.% bpd,y;*g vs

                                                                                                                ~a v                          ,j 7 -+
                             ,7             __

normen .. sumes AA In ' 'a-

                                                                                                                         ~

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