ML19319D680

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Chapter 3 to Crystal River 3 & 4 PSAR, Reactor. Includes Revisions 1-10
ML19319D680
Person / Time
Site: Crystal River, 05000303  Duke Energy icon.png
Issue date: 08/10/1967
From:
FLORIDA POWER CORP.
To:
References
NUDOCS 8003240654
Download: ML19319D680 (200)


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TABLE OF CONTERS Section Page 3 RFAC'IOR 3-1 31 DESIGN BASES 3-1 3 1.1 PERFORMANCE OBJECTIVES 3-1 3 1.2 LIMITS 3-1 3 1.2.1 Nuclear Limits 3-1 3 1.2.2 Reactivity Control Limits 3-2 3 1.2 3 Themal and Hydraulie Limits 3-2 3 1.2.4 Mechanical Limits 3-3 32 REACTOR DESIGN 3-6 3 2.1 GENERAL

SUMMARY

3-6 3 2.2 NUCLEAR DESIGN AND EVALUATION 3-7 3 2.2.1 Nuclear Characteristics of the Design 3-7 3 2.2.2 Nuclear Evaluation 3-20 323 THERMAL AND HYDRAULIC DESIGN AND EVALUATION 3-32 3 2..'.1 Themal and Hydraulic Characteristics 3-32 3232 Thermal and Hydraulic Evaluation 3-41 3 2.4 MEtJANICAL DESIGN IAYOUT 3-68 3 2.4.1 Internal Layout 3-68 3 2.4.2 Fuel Assemblies 3-73 3 2.4 3 Control Rod Drive System 3-86 33 TESTS AND INSPECTIONS 3-95 331 NUCLEAR TESTS AND INSPECTION 3-95 3 3 1.1 Critical Experiments 3-95 , n 3 3 1.2 zero Power, Approach to Power, and Power Testing 3-95 U 332 THERMAL AND HYDRAULIC TESTS AND INSPECTION 3-95 00000iO 3-1

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                                                                   ' CONTENTS (Cont'd)

Section g 3 3 2.1 Reactor Vessel Flow Distribution and Pressure Drop Test 3-95 3 3 2.2 ruel Assembly Heat Transfer and Fluid Flow Tests 3-96 3323 Preoperational Testing and Postoperational Testing 3-98 333 FUEL ASSEMBLY, CONTROL ROD ASSEMBLY, AND CONTROL ROD DRIVE MECHANICAL TESTS AND INSPECTION 3-98 3331 Prototype Testing 3-98 3 3 3,2 Model Testing 3-98 [ 3333 Component and/or Material Testing 3-99 3 3 3.4 Control Rod Drive Tests and Inspection 3-100 O t 334 INTERNAIS TESTS AND INSPECTIONS 3-103 34 REFERENCES 3 104 l 00000142 r V ! r 3-11 1

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e

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   '(j                                      LIST OF TABLES' Table No.                            Title                       h-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 Analysic                           3-13 3-6     Soluble Boron' Levels and Worth                        3-lh 3-7     Exterior Neutron Levels and Spectra                    3-17 3-8     Calculated and Experimental Rod and Rod Assembly Comparison                                             3-22
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3-9 Reference Core Parameters 3-28 3-10 First Mode Thr.shold Dimensione and Flatness 3-28 3-11 Threshold Ratio and Power Flatness' 3-29

              -3 12    Coefficients of variation                              3-35~

3-13 DNB Fesults - Maximum Design Condition 3-57 3-lh DNB Results - Most Drobable Condition 3- W 1 3-15 Heat Transfer Test Data 3-3-16 Comparison of Heat Transfer Test Data 3-50 3-17 Hot Channel Coolant Conditions 3-51 3-18 DNB Ratios in the Fuel Assembly Channels 3-6h

                            ~

3-19 Clad Circumferential Stresses 3-78 5 3-20 LRD Fuel Swelling Irradiation Program 3-84 3-88

  • 3-21 Control Rod Drive Design Data

, 3-22 Control Rod Assembly Design Data 3-93 l7 ' _3-23 Axial Pot;r Shaping Rod Assembly Data 3-9h

O 00000i43

                                                 '? iii  (Revised 7-15-69)

_ _ _ _ . _ _ _ _. . LIST OF FIGURES (At rear of Section)

Figure No. Title 3-1 Baron concentration versus core Life 3-2 Axial Peak to Average Power versus Xenon Override Rod Insertion

                .3-3                      Axial Power Profile, Xenon Override Rods 55 Per Cent Inserted 3-4                      Moderat.or Temperature ccefficient versus Boron concentration 3-5                      Modeator Temperature coefficient versus mderator Tsap=mture and Various Boron Levels 3-6                      Per Cent Initial Power versus Time Following Trip 3-7                      Effect of Fuel Temperature (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 IRB Ratio ' 3-11 Power Shape Reflecting Increased Axial Power Peak for 144-Inch Core 3-12 Distribution of Fuel Rod Pading 4 3-13 Possible Fuel Rod DNB's for M uimum Design Conditions - 36,816-Rod core 3-14 Possible Fuel Rod IRB's for Most Probable Conditions - 36,816-Rod core 3-15 Distribution of Population Protected, P, and 1-P versus Number Rods for m st Probable Conditions 3-16 DNB Ratios (BAW-168) versus Reactor Power 00000I4;4 3-17 Maximum Hot channel Exit Quality versus Reactor Overpower

                                                                                                                                      ;        ,r 3-18                     Thermal conductivity of vo 2

3-19 Fuel Center Temperature at the Hot Spot versus Linear Pcnter ( .. Q 3-20 Number of Data Points versus 4E/kC 13-21 Hot Channel Factor versus Per Cent Population Protected 3-iv

    . . . . . . _ . _ . _ . _ . _ . _ . _ _ .              _ _ . . . _ . . . . . . _ _ _ . . _ _ _ _ . . _ . . _ . . _ _ _ _ . _ _           _ . _ . - _ .
                                                                                    '
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FIGURES '(Cont'd) Figure No. Title 3-22 Burnout Factor versus Population for Various Confidence Levels 3-23 Fods 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

3-30 Ratio of Experimental to Calculated Bumout Heat Flux 3-31 'atio of Experimental to Calculated Burnout Heat Flux O
     \

3-32 natie of 8xPer1 e=ta1 to C 1ce1 ted Ber===* Heat r1== l 3-33 Ratio of Experimental to Calculated Bumout Heat Flux 3-34 Ratio of Experimental to Calculated Burnout Heat Flux 3-35 Ratio of Experimental to Calculated Burnout Heat Flux 3-36 Ratio of Experimental to Calculated Burnout Heat Flux l 3-37 Ratio of Experimental to Calculated Burnout Heat Flux i ' 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 DesiFn and Nominal Channel Exit Quality versus Reactor Power (without Engineering Hot Channel Factors) 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 Channe 00145 3-45 Flow Regime Map for Wall Channel 3-v

 -     . . _ . --.        - , _ - . _ -               .._.- ... . - .....             . - . . .-.- .    . . - - -
                                                                                                                  . - .

FIGURES (Cont'd)

                                                                                         '

k Figure No. Title 3 h6 Hot Channel DNB Ratio Comparison 3-h7- Reactor Coolant Flow versus Power 3 h8 Thermal Conductivity of 95 Per Cent Dense Sintered UO Pellets 2 3-h9 Fuel Center Temperature for Beginning-of-Life Conditions 3-50 Fuel Center Temperature for End-of-Life Conditions ' 3-51 Per Cent Fission Gas Released as a Function of the Average Temperature of the UO Fuel 2 I~ 3-52 Axial Local to Average Burnup and Instantaneous Power Comparisons 3-53 Fission Gas Release for 150 and 1.70 Max / Avg Axial Power Shapes 3-5h 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 Patios

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3-56 Maximum Fuel Rod Power Peaks and Cell Exit Enthalpy Fise Ratios

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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 Licit Local 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 1 3-62 . Fuel Assembly 3-63 Orifice Rod Assembly 3-6h Control Rod Drive - General Arrangement 3-65 Control Rod Drive - Vertical Section 3-66 Control Rod Drive System and Trip Block Diagram

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,  g'      3-f 3-68.

f"*****)

(Deleted) 00000146 7
          .3-69    ' control Rod Assenbly 3-70      Axial Power Shaping Rod Assembly .                                7 3-vi (Revised 7-15-69)

D ' d 3 REACTOR

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31 DESIGN BASES The reactor is designed to meet the performance objectives specified in 31.1 without exceeding the limits of design and operation specified in 3 1.2. 3 1.1 PERFORMANCE OBJECTIVES The reactor is designed to operate initially at 2,452 MWt with sufficient design margins to accommodate transient operation and instrument error without damage to the core and without exceeding the pressure at the safe-ty valve settings in the reactor coolant system. The ultimate operating power level of the reactor is expected to be 2,5kk MWt, but additional operating information vill be required to justify operation at this higher power level. Thus, this section of the report describes only reactor op-eration at the initial power level. The fuel rod cladding is designed to maintain its integrity for the antic-ipated core life. The effects of gas release, fuel dimensional changes, and corrosion- or irradiation-induced changes in the mechanical properties of cladding are considered in the design of fuel assemblies. Reactivity is controlled by control rod assemblies (CRA's) and chemical

   ~N   poison dissolved in the coolant. Sufficient CRA vorth is available to (k     shut the reactor down (kerr 5 0 99) in the hot condition at any time dur-ing the life cycle with the most reactiv0 CRA stuck in the fully with-drawn position. Redundant equipment is provided to add soluble poison to the reactor coolant to insure a similar shutdown capability when the re-actor 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 I cause a transient capable of damaging the reactor coolant system or caus-ing significant fuel failure. 3 1.2 LIMITS 3 1.2.1 Nuclear Limits The core has been designed to the fol ; wing nuclear limits: I '

a. Fuel has been designed for sa average burnup of 28,200 MWD /MTU and for a maximum burnup of 35,000 MWD /MTU.
b. The power coefficient is negative, and the control system is capable of compensating for reactivity changes resulting from l nuclear coefficients, either positive or negative.

Control systems will be available to handle core xencn insta- ' c. { l) b' ' bilities should they occur during operation, without jeopAr-dizing the safety conditions of the system. 000147

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l 3-1

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d. The core vill have sufficient excess reactivity to produce the de-sign power level and lifetime without exceeding the control capacity or shutdown margin. g
e. I Controlled reactivity irertion rates have been limited to 1.1 x lg"k Ak/k/see for a single regulating CRA group withdrawal, and 7 x 10-Ak/k/see for soluble boron removal.
f. Reactor control and maneuvering procedures vill not produce peak-to-average power distributions greater than those listed in Table 3-1.

The lov vorth of CBA groups inserted during power operation limits power peaks to acceptable values. 3.1.2.2 Reactivity Control Limits The control system and the operational procedures vill provide adequate control of the core reactivity and power distribution. The following control limits vill be met:

a. Sufficient control vill be available to produce a shutdown margin of at least IT Ak/k.
b. The shutdown nargin vill be naintained with the CRA of highest worth stuck out of the core.
c. CRA withdrawal limits the reactivity insertion rate to 1.1 x 10 Ak/k/see on a single regulating group. Boron dil 7 limited to a reactivity insertion rate of 7 x 10 gtion is also Ak/k/sec.

3.1.2.3 Thermal and Hydraulie Limits The reactor core is designed to meet the following limiting thermal and hydrau-lie conditions:

a. No central melting at the design overpower (llh per cent).
b. A 99 per cent confidence that at least 99.5 per cent of the fuel rods in the core are in no jeorardy of experiencing a departure from nu-cleate boiling (DNB) during continuous operation at the design over-power.
c. Essentially 100 per cent er.*idence that at least 99.96 per cent of the fuel rods in the core in no jeopardy of experiencing a D?!B during continuous operation at rated power.
d. The generation of net steam in the hottest core channels is permis-mible, but steam voids vill be lov enough to prevent flow instabilities.

The design overpower is the highest credible reactor operating power pernitted by the safety system. Normal overpower to trip is significantiv less than the d: sign overpower. Core rated power is 2,h52 trit. 00000148

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O 3-2 (Revised 7-15-69)

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4 3'l.2.k

                                                           .         Mechanical Limits-3 1.2.4.1        Reactor Internals e          0                                                                                                                                                     '   '

Tha reactor internal components are designed to withstand tihe stresses resulting

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from startup; steady state operation with one, two, three, or four reactor cool-ant pumps running; and shutdown conditions. No damage to the reactor internals

                                                   .will_ occur as a result of loss of pumping power.

Reactor internals 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 nomal reactor operation and transients. Structural integrity of all core support assembly circumferential~ velds will be assured by compliance with

ASME Code Sections-III and IX, radiographic inspection acceptance standards, and welding qualifications. . The core support structure will be designed as a Class I structure, as defined in Appendix 5A of this report, to resist the effects of seismic disturbances. The basic design guide for the seismic analysis will be AEC publication TID-7024,1" Nuclear Reactorc and Earthquakes". j Lateral 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. l The structural internals will be designed to maintain their functional integrity in the event of a major loss-of-coolant accident as described in 3 2.k.1. The dynamic loading resulting from the pressure oscillations because'of a loss-of-

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coolant accident will not prevent CRA insertion. 4 1 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 per=it 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. The cold-worked Zircaloy-4 cladding is designed to be free-standing. Fuel rods are held in place by mechanical spacer grids that are de-signed to maintain dimensional control of the fuel rod spacing throughout the design life without impairing cladding integrity. Contact loads are limited to prevent fretting. 5 The spacer grids are also designed to permit differential thermal expansion of the fuel rods without restraint tnat would cause distortion of the rods. The fuel assembly upper end fitting and the ~ control rod guide tube in the internals structure are both indexed to the grid plate above the fuel assemblies, thus insuring continuous alignment of the guide channels for the CRA's. The control

           .

a

             ;              y rodLtravel is designed so that the rods are always engaged in the fuel assembly

- 3-3 (Revised 1-15-68) 00000149 i' g. +ug .;.. .

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guide tubes, thus insuring that CRA's can always be inserted. The assembly structure is also designed to withstand handling loads, shipping loads, and earthquake loads. ,g ; Stress and strain for all anticipated normal and abnomal 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 pemitting 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 en11 defomations of the material, and the single occurrence of which will not make a significant contribu-tion to the possibility of a failure, will be pemitted to exceed the yield strength of the material. Where such stresses exceed the mate-rial yield strength, strain limits vill 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 exsmple of this "*pe of stress is the themal stress resulting from the thermal grad 4 ' across the clad thickness.
c. Combinations of these two types of stresses, in addition to the in-dividual treatment outlined above, vill be evaluated on the low-cycle fatigue basis of Item b. Also, clad plastic strain due to diameter increases resulting from themal ratcheting and/or creep, including g

the effects of internal gas pressure and fuel swelling, vill 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, on short time collapse, at end void. (2) End void must not collapse (cust 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 te=perature through the clad ! vall. (4) Clad must be freestanding at design pressure on a short time basis at =725 F hot spot average temperature through the clad vall. i ! 00000150 0 3h

                                             . _

es 3 1.2.h.3 control Rod Assembly (CRA)

 - 'C               ,
                              ,
                                                   ;        .
           - The control rod clad is designed to the some criteria as the fuel clad, as ap-plicable. Adequate clearance vill be provided between the control rods and the guide tubes, which position them vithin 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, including earthquake.

Overstressing of the CRA c.omponents during a trip will be prevented by minimizing the shock loads by snubbing and by providing adequate strengt1. 3.1.2.h.h control Bod Drive Deleted sentence. 7 The control rod drives provide control rod assembly (CRA) insertion and with-drawal rates. consistent with the required reactivity changes for reactor op-erational load changes. This rate is. based on the vorths of the various rod groups, which have been established to limit power-peaking flux patterns to design values. The maximum reactivity addition rate is specified to limit the magnitude of a possible nuclear excursion resulting from a control system or operator malfunction. The normal insertion and withdrawal velocity has been established as 30 in./ min. 7 The control rod drives provide a " trip" of the CRA's which results in a rapid

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shutdown of the reactor for conditions that cannot be handled by the reacter v 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 command of a CRA has been estab- 1 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. All pressure-containing components are designed to meet the requirements of the ASME Code, Section III, Nuclear Vessels , for Class A vessels. 7 j l Materials selected for the control rod drive are capable of operating within l the specified reactor environment for the life of the mechanism without any l deleterious effects. Adequate clearance vill be provided between the stationary l and moving parts of the control rod drives so that the CRA trip time to full in-l- sertion vill not be adversely affected, by mechanical interference under all op-l erating conditions and seismic distrubances. Structural integrity and adherence to a' lovable 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.

e . .. 00000151 n

  - gg -

3-5 (Revised 7-15-69)

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32 REACTOR DESIGN 3 2.1 GENERAL

SUMMARY

The important core design, thermal, and hydraulic characteristics are tabulated in Table 3-1. Table 3-1 Core Design, 'Ihermal, and Hydraulie Data

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Reactor 1ype Pressuu zed Water Rated Heat Output, MWt 2,452 Vessel Coolant Inlet Temperature, F 555 Vessel Coolant Outlet Temperature, F 602.8 Core Outlet Temperature, F 604 3 Operating Pressure, psig 2,185 Core and Fuel Assemblies Total Nu=ber of Fuel Assemblies in Core 177 Number of Fuel Rods per Fuel Assembly 208 Number of Control Rods per Control Rod Assembly 16 Nu=ber of Incore Instrumentation Pcsitions per Pael Assembly 1 Fuel Rod Outside Diameter, in. 0.420 Clad Thickness, in. 0.026 Pael Rod Pitch, in. 0 558 Fuel Assembly Pitch Spacing, in. 8.587 Unit Cell Metal / Water Ratio 0.80 Clui Material Zircaloy-4 (cold-worked) Fuel Material UOg Form Dished-End, Cylindrical Pellets , Diameter, in. O.362 l Active Length, in. 144 Density, % of theoretical 95 i l Heat Transfer and Fluid Flov at Rated Power l Total Heat Transfer Surface in Core, ft2 48,578 Average Heat Flux, Btu /hr-ft2 167,620 Maximum Heat Flux, v.u/hr-ft2 543,000 Average Power Density in Core, kv/l 79.60 Average Themal Output, kv/ft of fuel rod 5.4 l Maximum Thermal Output, kv/ft of fuel rod 17 49 Maximum Clad Surface Temperature, F 654 Aversge Core Fuel Temperature, F 1,385 Maxi =um Fuel Central Temperature at Hot Spot, F 4,160 00000152 3-6 _ a gw

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Table 3-1 (Cent'd) fm Total Reactor Coolant Flov, lb/hr 131 32 x 10 6 d Core Flow Area (effective for heat transfer), ft2 Core Coolant Average Velocity, fps

                                                                                                   '47 75
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13 7 Coolant Outlet Te.perature at Hot Channel, F 644.4 Power Distribution Maximum /AveragePowerRatio,radialxlocal (Fa nuclear) 1.85 Maximum Average Power Ratio, axial (F: nuclear) 1 70 Overall,PowerRatio(Fq nuclear) 3 15 Power Generated in Fuel and Cladding, % 97 3 Hot Channel Factors Power Peaking Factor (Fq) 1.008 Flow Area Reduction Factor (F ) 0 992 Iccal Heat Flux Factor (Fqa) A 1.013 Hot Spot Maximum / Average Heat Flux Ratio (Fq nue. 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

     /%

U 3 2.2 NUCLEAR DESIGN AND EVALUATION The basi : 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 pemit safe reac-o r opera *, ion and shutdown at all times during core lifetime.

3 2.2.1 Nuclear Characteristics of the DesiEu, 3 2.2.1.1 Excess Reactivity The nuclear design characteristics are given in Table 3-2. The excess reactivities associated with various core conditions are tabulated in Table 3-3 The core vill operate for 410 full power days for the first

   ,    e,       cycle and. vill have a 310 full power day equilibrium cycle. Design
 ,

limits vill be held with respect to reactivity control and power distri-

      .          bution. Incore instrumentation will be used to insure proper power peaking levels. Single fuel asse=bly reactivity information is also in-cluded in Table 3-3 N

' V 00000153 37

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           - - ,                                         ,    , -    ,    , , ~ - . . ,    ,   , -     --   ,   . - . , -

T.blo 3-2 Nuclear Design Data Fuel Assembly Volume Fractions Fuel 0.285

                                                                               'O l Moderator                                                      0 590 Zircaloy                                                       0.099 Stainless Steel                                                0.011 Void                                                           0.015 1.000 Total UO2, metric tons                                               91.61 Core Dimensions, in.

.

     . Equivalent Diameter                                            128 9 Active Height                                                   14k.0 Unit Cell H W 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, MWD /MrU First Cycle Average                                            12,460 Succeeding Cycle Average                                        9,410 Feed Enrichments, v/o U-235 First Cycle                                   2.29/2.64/290(byzone)

Equilibrium Cycle 2 94(a) t Control Data Control Rod Material Cd-In-Ag Number of Control Rod Assemblies 69 Total Rod Worth (A k/k), % 10.0 l3 Control Rod Cladding Material Type 304 SS (a) Average feed enrichment. , 1 00000154 [ 3-8 (Revised 3-1-68) _ _ e

                            .

Table 3-3

    ,-                                Excess Reactivity Conditions
 '

Effective Multiplication - BOL(8) Cold, Zero Power 1.302 1 Hot, Zero Power 1.2L7 - Hot, Rated Power' 1.229 Hot, Equilibrium Xe, Rated Power 1.192 Hot, Equilibrium Xe and S=, Rated Power (b) 1,153 Single Fuel Assembly (C) 0 77 Hot (d) Cold o,37 (a)BOL - Beginnin6-of-Life. IO) Includes burnup until equilibrium sa=arium is reached. ( } Based on highest probable enrich =ent of 3 5 veight per cent. (d)A center-to-center assembly pitch of 21 in. is required for this keff in cold, nonborated water with no xenon or sa=arium.

   -A t

u The minimu= critical = ass, with and without xenon and sa=arium poisoning, may be specified in a variety of for=s, i.e., single assembly, multiple i assemblies in various geometric arrays, damaged or crushed assemblies, ! etc. The unit fuel assembly has been investigated for comparative pur-poses. A single cold, clean asse=bly containing a caxi=um probable en-rich =ent of 3 5 vt % is suberitical. Two assemblies side-by-side are supercritical except when both equilibrium xenon cnd sc=nrium are present , Thrae assemblies side-by-side are supercritical with both equilibriu= l xenon and samarium present. ! l 3 2.2.1.2 Reactivity Control Dietribution Control of excess reactivity is shown in Table 3-4. l Table 3-h First Cycle Reactivity Control Distribution

                                                                                   % Ak/k
           -1. Controlled by Soluble Boren
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         -
a. Moderator Temperature Deficit (70 to 4040155 3.u b., _Equilibritm Xenon and Sa-arium 2.5 l 1 3-9 (Revised 1-15-68)
                         .              _    . _ .

Table 3-4 (Cont'd) O 5 Ak/k

c. Fuel Burnup and Fission Product Buildup 16.0 1
2. Controlled by Inserted Control Rod Assemblies
a. 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.

d. Minimum Available CRA Worth Minimum Movable CRA Worth Available 70 5.6 g Explanation of Items Above ( l. Soluble Boron Boron in solution is used to control the following relatively slov-moving reactivity changes;

a. The moderater deficit in going from ambient to operating tempera-l tures. The value shown is for the maximum change which would i occur toward the end of the cycle,
b. Equilibrium sa=arium and a part (approximately 1.h% Ak/k) of the

, equilibrium xenon. l l c. The excess reactivity required for fuel burnup and fission prod-uct buildup throughout cycle life. 00000156 g 3-10 (Revised 1-15-68)

          )-       _. _

Figure 3-1 shows the typical variation in baron concentration with life for Cycle 1 and the equilibrium cycle. Cq Control rod assemblies (CRA's) vill be used to control the reactivity changes associated with the fellowing:

2. Inserted Control 1

. (DELETED) . O l l l_ Sufficient rod worth remains inserted in the core during normal opera-tion to overcome the peak xenon transient following a power reduction. This override capability facilitates the return to nomal operating conditions without extended delays. The presence of these rods in the core during operation does not produce pcVer peaks above the design value, and the shutdown margin of the core is not adversely affected. Axial power peak variation, resulting from partial or full insertion of xenon override rods, is described fully in Figures 3-2 and 3-3 The loss of movable reactivity control due to the insertion of this , group producea no shutdown difficulties and is reflected in Table 3-5 1 3 _ Movable Control I  ; e. . Powe$*.,levelchanges(Doppler)andregulation.

b. The portion;of the equilibrium xenon not controlled by soluble boron, I

approximately'1% A k/k, is held by movable CRA's. O ' d00'00157

3-11 (Revised 1-15-68) l
   . _ _ ~ , . . .   -        __  . _ _  , , . . -
m. _ _ . . . . , _ , . _ _ , . , _ - . _ , . _ ,.. _. - - _
c. Betvxn z:ro and 15 per cent of full power, reactivity compensation by CRA's may be required as a result of the linear increase of reactor coolant temperature from 520 F to the normal operating value.
d. Additional reactivity is held by a group of partially inserted CRA's O (25 per cent insertion maximum) to allev periodic rather than contin-uous soluble boron dilution. The CRA's are inserted to the 25 per cent limit as the boron is diluted. Autcmatic withdrawal of these CRA's during operation is allowed to the 5 per cent insertien 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 required as part of the reactivity controlled by CRA's.
4. Rod Worth A total of 4.0% Ak/k(a) is required in movable control. Analysis of the 1 69 CRA's under the reference fuel arrangement predicts a total CRA vorth of at least 10.07. Ak/k. 'Ihe stuch-out CRA vorth was also evaluated at a value no larger than 3.0% A k/k(b). ' Ibis evaluation included selection of the highest worth CRA under th9 irst CRA-out condition. The minimum available CRA vorth of 5 6% Ak/k(a(/ is sufficient to meet movable control requirements.

3 2.2.1 3 Reactivity Shutdown Analysis The ability to shut down the core under both hot and cold conditions is illus-trated in Table 3-5 In this tabulation ' oth the first and equilibrium cycles are evaluated at the beginning-of-life (BOL) and the end-of-life (EOL) for shut-down capability. h .. -. (a)Does not include transient control. See Table 3-h. First cycle. See Table 3-u. 00000158 .

                                                                                   'O 3-12 (Revised 1-15-68)

_

Table 3-5 Shutdown Reactivity Analysis O First Cycle Equilibrium

                                                                                                                                       -

Reactivity Effects, % Ak/k BOL EOL BOL ' EOL 1

1. Maximum Shutdown CRA Requirement Doppler (100 to 0% Power) 1.2 15 1.2 1.5 .

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 33 2.2 33 , 2. Maximum Available CRA Worth (a) _10,o 10,o _to,o 10,o Transient Xe Insertion Worth 1.4 1.4 1.4 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.h -9.S OneCRAStuck-Out(b) -5.h -5.h -5.h -6.8

4. 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 r 6. h + 3. 0 t 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 t 3.6 4 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 (*) Total worth of 69 CRA's. (b)CRA of highest reactivity value. .j } g ( ) Includes changes in CRA worth, moderator deficit, and

                                                                                                     ~
         '

equilibrium Xe held by soluble boron.

           .(d)No boron addition.
                 '
               '

3-13 (Revised 1-15-68)

   . - .               -
                             - . . . - ,      - - - - . - . . . - - . . , . . - . . . - .. .                      -.- - .         ..   . - . .
    .                                 .       ._        ..      _ _ . - . .  ..

Exa=in; tion of T;bl9 3-5 for Minimum Hot Shutdown Margin (Item k) shows that, with the highest worth CRA stuck out, the core can be maintained in a suberitical condition. Normal conditions indicate a minimum hot shut-down margin of 51% ak/k at end-of-life. l1 g 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 a least 1% A k/k. The reactivity changes that take place between the hot zero power to cold conditions are tabulated, and the corresponding increases in soluble boron are listed. Beginning-of-life boron levels for several core conditions are listed in Table 3-6 along with boron vorth values. Additional soluble boron could be added for situations involving more than a single stuck CRA. The con-ditions 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, kerr = 0 99 No CRA's In 1,820 1 All CRA's In 1,290 One Stuck CRA 1,h50
2. Hot, Zero Power, ke rr = 0 99 h No CRA's In 2,080 All CRA's In 1,080 One Stuck CRA 1,380 3 Hot, Rated Power Nc CRA's In 1,860
4. Hot, Equilibrium Xe and Sm, Rated Power l No CRA's In 1,360

! Core Condition Boron Worth. (% N-/k)/ ppm l l Hot 1/100 Cold 1/75 P0000!60 00000160 3-14 (Revised 1-15-68) 0 __ __ _ __

O 3 2.2.1.4 Reactivity Coefficients Reactivity coefficients form the basis for analog studies. involving normal 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 func-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 10-5 to -1 7 x 10-5 A k/k/F.
b. Moderator Void Coefficient B e moderator void coefficient relates the change in neutron multi-plication to the presence of voids in the moderator. Cores controlled by appreciable amounts of soluble boron may exhibit a small positive coefficient for very = mall void levels (several per cent void), while The highervoidlevelsproduceincreasinglynegativecoeffgeients.

expected range for the void coefficient is +1.0 x 10- to -3 0 x 10-3 A k/k/% void. l

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. Bis 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 would 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 fewer negative moderator temperature coefficients than do cores controlled solely by movable or fixed CRA's. The major temperature effect on the cool-ant is a change in. density. An increasing coolant temperature pro-duces a decrease in water density and an equal percentage reduction in boron concentration. The concentration change results in a posi-tive reactivity component by reducing the absorption in the coolant.

l The magnitude of this component is proportional to the total reac- ' tivity held by soluble boron.

 ,           The moderator temperature coefficient has been parameterized for the 4           reference core in terms of boron concentration and reactor, coolant temperature. The results of the study are shown in Figures'3-4 and
                                                .u.;

i 00000161 h"

                                  ,,      3-15

_ 3-5 Figure 3-4 shows the coefficient variation for ambient and op-erating temperatares as a function of soluble boron conepntration.

                                                                                 $

De operating value ranges from approxim +1.0 x 10-4 at the be-ginningofthefirstcycleto-30x10gtely/k/Fattheendofthe Ak equilibrium cycle. Figure 3-5 shows the moderator temperature coef-ficient as a function of temperature for various poison concentra-tions for the first cycle. The coefficients of the equilibrium cycle vin be more negative than those of the first cycle since the boron concentration levels are considerably lover. S e positive temperature coefficient during the initial portion of each cycle viu not constitute an operational problem. Se Doppler deficit represents a much larger reactivity effect in the negative direction and, together with the CRA system response, vill provide adequate control. -

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 itas such as clad materials, fuel assembly crud deposition, system average temperature, and prior system vater chemistry.

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 approximately 5 5 to a hot calculated value of 7 8 with 1,400 , ppm boron and 3 ppu KOH in solution at the beginning of life. l Lifetime bleed dilution to 20 ppm boron vill reduce pH by ap-proximately 0.8 pH units to a hot calculated pH value of 7 0. ! (2) Considering the maximum system makeup rate of 70 gym, the cor-responding changes in pH are 0.071 pH units per hour for boron dilution and 0.231 pH units per hour for KOH dilution. Apply-l ing pH vorth values of 0.16% Ak/k per,pH unit, as observed at ! Saxton, insertion rates are 316 x 10-o% Ak/k see and 1.03 x 10-5% Ak/k sec, respectively. Rese insertion rates correspond , to1.03percentpower/hourand34percentpover/ hour,re-l spectively, which are easily compensated by the operator or the l automatic control system. , .

                                                                              -

O

                                              -

00000162 3-16 _ .__

g-4,- (/ 3.2.2.1 5 Reactivity . Insertion Pates - Figure 7-7 displays the integrated rod worth of three overlapping rod banks as a function of distance withdrawn. The indicated groups are those used in the core during power operation. Using approximately 1.21 Ak/k CRA groups and a 30 in./ min drive speed in conjunction with the reactivity l7 response given in Figure 7-7 yields a naximum reactivity insertion rate of 1.1 x 10~ (Ak/k)/sec. he maximun reactivity insertion rate for soluble l7 boron removal is 7 x 10-6 (Ak/k)/second. 3.2.2.1.6 Neutron Flux Distribution and Spectrum Figure 3-6 displays the beginning-of-life power decay curve,s for the two least effective CFA verths as outlined in Table 3-5, Item No. 3. The power decay is initiated by the trip release of the CPA's with_ e 300 msec 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 vall are given in Table 3-7 Table 3-7 Exterior Neutron Levels and Srectra Neutron Flux Levels n/cm2 /nec(" Interior Vall of Flux Core Edge Pressure Vessel Groun (x 1013) (x inlo) 1 0.821 Mev to 10 "ev 6.0 3.h 2 1.230 Kev to 0.821 Mev 90 75 3 0.hlh ev to 1.230 Kev 6.2 57 h Less than 0.hlh ev 71 2.1 (a)These values include the maxinum axial peak-to-average power ratio of 1.7

         - The calculations were performed using The Babcock & t!ilcox Connany's LIFE code (BAW-29        Section 3.6.3) to generat input data for the transport code, TOPIC. b A h-group edit is obtained from the LTFE outnut which in-cludes diffusion coefficients, absorption, removal, and fission cross sec-tions, and the zeroth and first moments of the scattering cross section.

TOPIC is an Sn code designed to solve the 1-dinensional transnort equation , in cylindrical coordinates for up to six groups of neutrons. For_the ra-dial and azimuthal variables, a linear approximation to the transport

 . im U                 .
                      ,

00000163 3-17 (Revised 7-15-69)

                         '_____ _ . __1

equation is used; for the polar angle, Gauss quadrature is used. Scatter-ing functions are represented by a Legendre series. The azimuthal angle can be partitioned into 4 to 10 intervals on the half-space between 0 and

r. The number of mesh points in the radial direction is restricted by the number of these intervals. For the core exterior flux calculations, four intervals on the azimuthal were.used. This allows the maximum num-ber of mesh points (240) in the "r" direction to describe the shield com-plex. An option is available to use either equal intervals on the azi-r.:uthal angle or equal intervals on the cosine of the angle- Equal inter-vals on the cosine were chosen since this provides more detail in the forward direction of the flux (toward the vessel). Five Gauss quadrature points were used on the cosine of the polar angle in the half-space be '

tween 0 and r. Results from the above method of calculation have been co= pared with ther-mal flux measyrgments through an array of iron and water slabs in the LIDO pool reactor.(21 Although this is not a direct comparison with fast neu-tron measurements, it does provide a degree of confidence in the method since the magnitude of the ther=al flux in shield regions is governed by fast neutron penetration. Results of the comparison showed that fluxes predicted by the LIFE-TOPIC calculation vere lover, in general, by about a factor of 2. Results of the fast flux calculations are, consequently, increased by a facter 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 O density of 41 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.'T vas used. This is about 30 per cent greater than the 1 3 expected over life.

Uncertainties in the calculations include the following:

1. The use of only four nertron groups to describe the neutron energy spectrum.
2. Use of the LIFE code to generate the 4-group cross sections.

In the LIFE program, the 4-group data in all regions are com-puted from a fission spectrum rather than a leakage spectrum. 3 Having only four intervals, i.e., n = 4 in the Sn calculation, to describe 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 underesti-mates the effectivenese of the ther=al 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 i ' 000001M 3-18

 -   -

O cross section of iron increases rapidly with energy. Therefore, the assump-tion of a fission spectrum to compute cross sections in the thermal shield, and the use of a fev-group model to cover the neutron energy spectrum, would underestimate the neutron energy loss in the thermal shield and the subsequent attenuation by the ya}er between the vessel and thermal shield. The results from 34-group P3) Cit 3/ calculations show that reduction of the flux above 1 Mev by the themal shield is about a factor of 4 geater than that computed from the 4-group calculations. 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 84 and an S12 calculation was made in penetration through hydrogen. The results for a variety of energies over a penetration range of 140 cm showed the S4 calcu-lation to be lower than the S 21 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 themal flux ita indicates a possible underestimate. Uhtil a better comparison with data can be made, we have assumed that the underestimate is correct and accordingly have increased the flux calculations by a factor of 2 to predict the nyt 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. 3O The effect of this steel-water configuration on (a) the neutron irradiation, and (b) the themal stresses in the reactor vessel vall, were evaluated as follows:

a. Neutron Irradiation Calculations were performed in connection with the reactor vessel design to detemine the relative effects of varying the baffle and thermal shield thicknesses on tne neutron flux (> 1 Mev) at the ves-sel vall. TheJe the P3101 code (3)using calculations were perfomed 34 fast neutron groups. with The the P1 option results showed of that tne neutron flux level at the vessel vall is dependent, for the most part, on the total metal and water thickness between the care and the vessel. However, there was some variation in fluxes devend-ing upon the particular configuration of steel-water lamination .

Also, the gain in neutron attenuation by replacing water with steel diminishes somewhat vi-h increasing steel thickness. In general however, the results showed that for total steel thick-nesses 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 vould be reduced, 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 p b attenuation in this reactor vessel with those in San Onofre, Turkey Point 3 and 4, Indian Point 2, and Ginna. The total distance be-tween 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 00000165 3-19

                                                                          .
                                                                                                                                                                     -
  .- -- -        ,.---, - ,,,_,..,,....,                  - . -        -, . . . - .   -...--.----.---e    , - - . - - . . - . .       ,  _ -     - - - - . , - - . ,
                                                                               .

vessel than in the other reactors. For neutrons above 1 Mev it was found that this additional distance would provide additional attenu-ation ranging from a factor of 1.1 to 5 times greater than that in the other WR's considered,

b. Themal Stresses The gamma heating in the reactor vessel is produced by primary gam-mas from the core and by secondary gammas originating in the core liner, barrel, themal shield, and the vessel itself. In this reac-tor design the major portion of the heat is generated by Samma rays from the core and by secondary gamma rays from the core liner and barrel.

Since the gammas from each of these sources must penetrate the ther-mal shield to reach the vessel, the vessel heating rate is dependent on the themal shield thickness. For designs which employ thicker themal shields, or in which inter-nals are to be exposed to higher neutron fluxes, gam =a rays originat-in6 in the themal chield or in the vessel itself may govern the ves-sel heating rates. Since gamma rays from these sources would have to penetrate only portions or none of the themal shield to reach the vessel, the vessel heating in such cases vould 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 vesse' heating was dependent on the gam =a ray attenuation provided by the tnemal shield. This approach would be conservative since, as noted above for some designs, gamma sources other than those attenuated by the thermal shield may con-tribute appreciably to the vessel heatin6 The results of the com-parison showed that the difference in gamma attenuation between this reactor and other PWR's ranged fmm negligible difference to a factor of 5 3 less for this reactor 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-l ' atively low value, and no problems are anticipated from themal stresses in the reactor vessel vall. 3 2.2.2 Nuclear Evaluation ' Analytical models and the application of these models are discussed in this l s2ction. Core instabilities associated with xenon oscillation are also mem-i tiened, with threshold data evaluated under reference conditions. 3 2.2.2.1 Analytical Models R; actor' design calculations are made with a large number of co=puter codes. The choice of which code set or sets to use depends on which phase of the de-sign is being analyzed. A list'of codes used in core analysis with a brief discussion follows in 3 2.2.2.2. $ 3-2o 00000l66 _

                                                     .      -   .    .        .
,   4 O              a. 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 on the type of calculations to be made. In a clean type of a calculation where there are no. strong localized absorbers of a type differing from the rest of the lattice, 1-dimensional analysis is satisfactory. This type of problem is handled quite well by the BisW 1-dimensional depletion package code LIFE. LIFE is a composite of MUFT (Ref. 4), KATE (Ref. 5), RIP, WANDA (Ref. 6), and a depletion routint. Nomally the MUFT portion is used with 34 energy groups, an exact treatment of hydrogen, the Greuling-Goertzel approximation for elements of mass less than 10, and Femi age for all heavier ele-ments. The KATE portion nomally uses a Wigner-Wilkins spectrum. 4 In WANDA, 4 ene:gy groups are utilized. Disadvantage factors for in-put to the themal 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. Recen' check calculations on critical experiments have a standard deviatic.a 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-yO 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 re-activity coefficients, fuel cycle enrichments, lifetimes, and solu-ble poison concentrations can be found to improve the initial condi-tions specified 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 i cosfiguration which can be shown explicitely and are often stud $ d "'.th 'tuarter core symmetry. . Symmetry is desirable in the design. Wrf ci '.os in generality occurs. I The geometric description includes e*A 4 assembly and as much de- , tail as is possible, i.e., usually ech Mc in the fuel assembly. ! Analysis of this type pemits detailed power digt;. 'bution studies as well as reactivity analysis. The power distributice in a large PWR , core which has zone loading cannot be predicted reliatly with 1-dimensional calculations. This is particularly true when local power peaking as a function of power history is of interest. It is.neces-sary to study this type of problem with at least a 2-dimensWal code, and in some cases 3-dimensional calculations are necessary.

          .
                         . Use of the 2-dimensional programs requires the generation of
                                                   ~                                                  gi  7 constants as a function of material composition, power histo ,              vI  I
      - p(/. .         s geometry. For regions where diffusion theory is valid, MUFT and

_ , KATE with THERMOS disadvantage factors are used to generate epither-

mal and thermal coefficients.- This would apply at a distance of a 00000167 M 0010 3_,1
            . - _ -          - _ - - - _ - . - . -                              . _ -             - - .

_. _

                                                                                                                                         .

few mean paths from boundaries or discontinuities in the fuel rod lattice. Discontinuities refer to fuel assembly can, water channels, instrumentation 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 struc-tural material, water ch mnels, and adjacent fuel rod rows can be represented well in slab geometry. This problem is ana4 zed by P3)G (Ref. 3) which is effective in slab geometry. The coefficients ( so generated are utilized in the epithermal energy range. Coeffi-i cients for the thermal energy range are generated by a slab THERMOS calculation. The regions adjacent to an interface of material of different enrichment are also well represented with the P3no code. The arrange.snt 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 be-tween the codes. The flux shape is calculated by DTF-IV and cross sections by the others. The outer boundary of the core where there is a transition from fuel to reflector and baffle is also represented by the DTF-IV code. The 3-dimensional analysis is accomplished by extending the techniques of 2-dimensional representation.

b. Control Rod Analysis g B&W has developed a procedure for analyzing the reactivity worth of small Ag-In-Cd rods in fuel lattices. Verification of this pzocedure was made by the comparative analysis of 14 critical experiments vit*.

varying rod and rod assembly configurations.(13,l'+) 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 vLich bracket the ref-erence design. Water holes, simulating withdrawn rods, were includ-ed as part of the lattice study. The resulting comparison of the analytical and experimental worths 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 .ssembly - 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 14-E l' 20 0 1 54' l.52 5-E 2 20 292 3 05 3 01

                                                         .

3-22 00000168

 .--          - - ~ .    . - - , - - , - .      -y ~ u     . - - . -   --.-r~--    . . - - . - , - - - --           - - - . . ~ - , -- -   --

O The mean error in calculating these-configurations is shown to be less than 1 per cent. Comparison of the power shape associated with the 16-rod reference assemblies showed good similarity. Point-to-average power had a maximum variation of less than 2 per cent with experimental data. The analytical method used for this analysis is based on straight dif-fusion theory. Thermal coefficients for a control rod are obtained  ; from THERMOS by flux-weighting. Epithermal coefficients for the upper energy groups are generated by the B&W LIFE program. The re-sulting coefficients are used in the 2-dimensional code PDQ to obtain the required eigenvalues. GAKER and LIBPM are used to prepare data for THERMOS. GAKER gener-ates scattering cross sections for hydrogen by the Nelkin technique. LIBPM uses the Brown and St. John free gas model for generating the remaining scattering crocs 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-p ond step is to use THERMOS in a calculation where the Ag-In-Cd rod is V surrounded by fuel. This is used to generate the flux-weighted con-trol rod cell coefficients as a function of boron concentration. As a check on the validity of the THERMDS approach, extrapolation dis-tances were compared to those given by the Spinks method.(15) The agreement was within 2.2 per cent for a set of cases wherein the num-ber densities of A6-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 layout where each fuel rod cell is shown separately.

c. Determination of Reactivity Coefficients This type of calculation is different from the reactivity analysis only in application, i.e., a series of reactivity calculations being required. Coefficients are determined for moderator temperature, voiding, and pressure, and for fuel temperature. These are calcu-lated from small perturbations in the required parameter over the range of possible values of the parameter.

The moderator temperature coefficient is detennined 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 ing This section contains a brief description of codes mentioned in 't@he gr

          ~

sections.

                          '

0@ g.

                .

3-23

  .                                         .   .-                  .

THERMOS (Ref. 7) This code solves the integral form of the Boltzmann Transpor Equation for the neutron spectrum as a function of po-sition. A diaBonalized connection to the isotropic transfer matrix has been incorporated allowing a degree of anisotropic scattering. MTT (Ref. 4) - This program solves the P1 or By multigroup equation for the first two Legendre coefficients of the directional neutron flux, and for the isotropic and anisotropic components of the slowing down densities due to a cosine-shaped neutron source. Coefficients are generated with MUFT for the epithermal energy I range. KATE (Ref. 5) - The code solves the Wigner-Wilkins differential equation for a homogeneous medium moderated by chemically unbound hydro-gen atoms in thermal equilibrium. 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 MTT, Pilc , and P3M?r. 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 combination mechanizes the procedures for using the codes separately. g GAM (Ref.11) - This code is a multigroup coefficient generation program that solves the F1 equations and includes anisotropic scattering. Inelastic scattering and reconance parameters are also treated by GAM. P310 (Ref. 3) - The code solves the multienergy transport equation in various geometries. The code is primarily used for epithermal coefficient generations. DIF (Ref.10) - This code solves the multigroup,1-dimensional Boltmann transport equation by the method of discrete onlinates. DIF al-lows multigroup anisotropic scattering as well as up and down scattering. i PDQ (Ref. 8) - This program solves the 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. g 00000170 3-24 _ _ _ _ - - - .-.

               .       _       __ _                   . _ . .

i i <,,. l 1

 -

O

      .
             '

TNT (Ref. 9) - This code is similar in application to TURBO, but is a 3-

                     .      dimensional code extended from DRACO.

l 3 2.2.2 3 xenon stability Analysis l , Initial studies of the reference core, where realistic fuel temperatures are generated by thermal-nuclear iteration, indicate no instability at any time during the life cycle. These results are encoura6 1 ng, but until more detailed analyses are completed, it will be assum M that axial xenon oscillations are possible. _ Azimuthal oscillations are un11ely, and radial oscillations will not occur. Since the size, flux level, and power coefficient of current PWR designs are conducive to xenon oscillations, an extensive investigation must be completed before the stability of a core can be ascertained. An adequate solution can be found by first using analytical techniques 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. O -

a. Summary of Results (1) Threshold Analysis In the threshold analysis axial, azimuthal, and radial oscilla-tions slightly were investigated dished for beginnin6 7)of-life, power distributions.(1 flattened, The results are and 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 spatial oscillation is increased as the dimension of the core increases. (c) The large size of current PWR 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 worse. (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.

t 0000017I

                                    '
                          .            ~g.              3-25 u__
                                        . .-       _      . . _ .      _-
 .
                                                          .

g

                                                                                    . . .

(f) A large, negative power coefficient tends to dampen oscil-lations. If this coefficient is sufficiently large, os-cillations cannot occur regartiless 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 power distributions. However, there exists a finite prob-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 hours) is long enough to pemit 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, azimuthal oscillations are unlikely, and radial oscillations vill not occur. (2) Depletion Analysis O Diffusion-depletion calculations coupled with heat transfer equations were employed to investigate further the axial sta-bility of the core since the ulalytical 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-l ' tial cycle. An average fuel temperature of 1,400 F was maintained during the cycle.

(b) The threshold for axial instability near the end of the initial cycle was found to coincide with a core average fuel temperature of 900 F. Diffusion theory was also used to examine the problem of con-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 sec-tions which are moved up and down about the midplane of the core to offset oscillatory power shifts. O 00000172 3-26

                            -
                               -.                .     ..         -     --.  . -. . _.-
                                                              .__       _.
                                                                                                                               !

, O (b) - Detailed power profiles will be available to the reactor , operator as output from the instrumentation. The large l period of the oscillation will allow partial rod movement  ; such that axial power peaks are held well within allowabic l 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: (a) Core size. (b) Flux level. (c) Degree of flatness in the power distribution. 1 (d) Power coefficient. (e) Reactivity held by saturation xenon. In addition, slightly dished power distributions were investi-gated to show that any dishing resulting 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 examined 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 lower value was computed with a zero power coeffi-cient and was not conservative without modal coupling. The higher value wac computed 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 dimensions are also shown for comparison pur-poses in.the following discussion. Table 3-10 shows the threshold dimensions for first mode in-stability as a function of flux flattening. The percenta6e of flattening is defined mJ 100 per cent times the. ratio of the flattened power distribu11on to the total physical dimension under consideration. Tta parameters of Table 3-9 at two full power days were used sir.ce they are virtually the same as those at 150 days but are more conservative. Axial depletion studies 3-n 00000173 _ - _ _ _ - - - - _ - - . - - - _ - - - - . - . - ..

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

show that power distributions are flattened by 0, 63,'and 73

                                       -per cent at 2, 150, and 354 full power days, respectively. A m

W 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 oscinations are unlikely, and radial oscillations viu not oc-Cur. Threshold dimensions for second mode oscinations 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. Table 3-9 Reference Core Parameters Two Full (Rated) 150 Full (Rated) ~ Power Days Power Days M2 , cm2 57 0 57 0 2 Tth,n/cm-sec 3 9 x 1013 3 E x 1013 ax (reactivity held by saturation xenon),ak/k O.034 0.033 DopplerCoefficient,ak/k/F -1.1 x iG'3 -1.1 x 10-5 Moderator Temperature Coefficient Positive but Small Negative i a T(powercoeff.),Ak/k/ unit flux = -2.2 x 10-1 = -2 3 x 10-1 , , Equivalent Dimensions, ft ! l Height 12.00 l Diameter 10 74 Radius 5 37 Table 3-10 First Mode Threshold Dimensions and Flatness Flatness, % Threshold Dimensions, ft 0 50 80 ! ' Thresholdheight(axialoscillations) 18 5 14.1 11.8 g

                                                                                                                          ,

Threshold diameter (azimuthal oscillation) 20.4 16 5 14.0 Threshold radius (radial oscillation) 16.8 16 7 14 5

                                 .

3-28 00000174

t

                                                                                       --                                   .
 - - . - - _ _ _ _ _                            .           .. .    ..         . . .              .          _.
                                                                                                                             .-.

O Table 3-11showsthevaluesofH/Dversuspowerflatnessfor-equal likelihood of axial, azimuthal, and radial first narmonic oscillations, i.e., if the core is just at the axial threshold for axial oscillations, it can-also be expected that there will beazimuthalandradialoscillationsprovidedthevalueofH/D in Table 3-11'is satisfied. H/Dforthisreferencereactoris 1.12.

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

Ratio O 20 50 80 100 H/D(axialversusazimuthal) 0 91 .o.87 o.86 o.86 0.85 H/D(axialversusradial) 0 55 0.49 _O.42 0.41 0.41 The modal methods used to examine the xenon oscillation problem made use of core-avera6ed quantities such as flux, power coeffi-l cient, and reactivity held by saturation xenon. In addition, L - flux distributions were limited to

l (a) Geometric distributions. (b) Partially or totally flat. (c) Slightly dished. , The power distribution during early life is such that no xenon l instabilities wil2 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 l 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 B8M LIFE depletion program was modified to include axial heat transfer. The equations and it m tion scheme are outlined below:

 .fN                              (a) The average fluid temperature for each axial core region k                                               is computed from a previously known power density.d(stribu-i tion as follows:                                           }/J
                                                                                                                           .

3 29

     - . - . . . - - .    - . . - . . . - . . . - - . . . . - . - -                        - . - _ .      -    - . - - - .       - - -
   -
 -       _._     _ . _                          ..     --         . - - - - - -
                              '
                                                                "

ATi = (Tout - Tin)1=C[Z PD(Z)dZ (A) in where ATg = temperature change in region "1" PD (Z) = power density in Z direction Zin, Zout = region "i" boundaries and

                                                     #*

C= (B) J PD (Z) dZ where H = active fuel height. Equation (A) is solved to Tout of region "1". Since Tin 18 known from core inlet conditions, the average fluid temper-ature is defined as follows:

                                   -

T

  • Tout + Tin fluidi 2 g (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. Local or bulk boiling is not permitted.

l (c) The average fuel temperature for each axial core region is , then computed from the average fluid temperatures and power densities: fueli

  • K i+ fluid i N where FDi= coversge power density of region "i" and K is defined by T
                                       -fuel - T-fluid         core
                                     ,

PDcore ! l l l (d) After the new fluid temperatures, moderator densities, and i 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. 6

     -

00000176 3-30 _ _ __ . . - . .- - __ . _ _ _ -.

This analysis ussd an sxact solution in that the cpectrum was recalculated for each zone (11 axial zones described the reac-tor) for each iteration at every time step. This included the A effects of the moderator coefficient. V This LIFE package was used to determine the 6 ?fects of the un-certainty in the power Doppler on the stability of the core. The uncertainty in the Doppler was more than compensated with a

                 - reduction in fuel temperature of 500 degrees. The reference core was analyzed with core average fuel temperatures of 1,h00 F and 900 F. Figure 3-7 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               i studies were made at beginning-of-life boron levels of approxi-mately 1,900 ppm. This level is approximately 200 ppm above the predicted beginning-of-life level, and consequently reflects a more positive moderator temperature 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 calcula-tion. Figure 3-8 shows the effect of fuel temperature toward the end of life. It is easily verified that the 900 F fuel tem-perature case approached the threshold condition for axial oscil-lation in this core. On the basis of the information presented, it can be said thtt for a realistic fuel temperature this core does not exhibit ax.al 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 l power Doppler. Normally, this would produce a divergent oscil-l p lation as shown in Figure 3-9 A study was completed wherein ! V a 1% ok/k rod bank with a 3-ft-long section of regular control rod material was cuccesPfully maneuvered to control the core after'a perturbation of .epowershapeatapointabout3/hof the way through Cycle 1. The controlled results are also shown ' in Figure 3-9 The miniman rod motion was one foot, and the . time step employed was 4.8 hours. More precise rod movement over shorter time periods would produce a much smoother power ratic 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 cotion 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-themal iterations. A detailed quantitative analysis of core stability and control procedures, employing either partial or nomal control rods, is to be undertaken with the new program. A U

    .
      .,                                                           000001/7 i-                                        3-31 (Revised 1-15-68)
                                                                 .   . _ . . . -, - - . . - . .-
                                                                               .

323 THERMAL AND HYDRAULIC DESIGN AND EVALIATION 3231 Thermal and Hydraulic Characteristics g 3 2 3 1.1 Fuel Assembly Heat Transfer Design

e. . Design Criteria The criterion for heat transfer design is to be safely below Depar-ture from Nucleate Boiling (DNB) at the design overpower (114 per cent of rated power). A detailed description of the analysis is given in 3 2 3 2.2, statistical Core Design Technique.

The input infomation 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. (4) 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, g

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

q" = (1.83 - o.000415 P) x 90,000 9 o.3987 e o.001036 A T ,c - 1.027 x 10-6 (A Tm )2

                   .2%

where q" = critical heat flux as predicted by the bestfitform, Btu /hr-ft2 P = core operating pressure, psia G = channel mass velocity, lb/hr-ft 2 l S = channel equivalent diameter, ft L = length up the channel to the point of interest, i't AT,,e = inlet subcooling (Tsat - Tinlet) b Tsat = coolant saturation temperature corre-sponding to P, F

                     "

3 32 00000178

                                                     .   .                _.

_

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

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 ',,o detemine a IIIB. ratio, i.e., the ratio of the heat flux at which a IRB is predicted to occur to the heat flux in the channel being investigated. . This DEIB 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 IRB versus P shown is for a 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 138 DtlB ratio corresponds to a 99 per cent confidence that - at least 94 5 per cent of the population of all suca hot channels are in no jeopardy of experiencing a IItB. This statement is a cor-ollary to the total core statistical statement given in 3 1.2 3, Themal'and Hydraulic Limits. The criterion for evaluating the themal design margin for individual channels or the total core is the confidence-population relationship. The DtiB 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 FA h = 1 79 Fz = 170 Fq = 3 04 (2) The design nuclear peaking factors for the worst time in core life are Fa h = 1.85 Fz = 1 70 Fq = 3 15-00000179

                                        .

8

i. 3-33 l

___. . _ . _~ _ . . . _ . , - - - , . _ _ , . . _ , _ . , , - . . . . _ . . _ . . . _ . . . _ , _ . . . . _ , _ . . . . . _ _ _ - , _ _ _ . - . . - . . . - -

                                                                          - _ - _ . _ .

_ l l FA h = max /av; 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 179 to 1.85 to obtain a more conservative design. The axial nuclear factor, Fz, is illustrated in Figure 3-11. The distribution of power expressed as P/P is shown for two conditions of reactor operation. The first condition is an inlet peak with a max /avgvalueof170resultingfrompartialinsertionofaCRA group for transient control following a power level change. This condition results in the maximum local heat flux and maximum linear heat rate. The second power shape is a symmetrical cosine which is indicative of the power distribution with xenon override rods with-drawn. The flux peak max / avg value is 150 in the center of the ac-tive core. Both of these flux shapes have been evaluated for ther-mal DNB limitations. The limiting condition is the 1 5 cosine power distribution. The inlet 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. Thic e feet has been demonstrated in DNB tests of nonunifom flux shapes. 9) The 15 cosine axial shape has been used to determine individual h channel DNB limits and make the associated statistical analysis. The nuclear factor for total radial x local rod power, Fa h, is cal-culated for each rod in the core. A distribution curve of the frac-tion of the core fuel rods operating above various peaking factors is shown in Figure 3-12. Line B shows the distribution of the maxi-mum calculated 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 1.85/179 or 1.033 to pmvide l conservative results. Determination of the peaking distribution for l the design condition in this mr ier has the effect of increasing re-actor power by about 3 per cen . This assumption is conservative since the distribution with a maximum peak F A h of 1.85 vill follow a line similar to Line C where the average power of all rods in the core is represented by an F Ah of 1.0. The actual shape oi the dis-tribution curve is dependent upon statistical peaking relationships, CRA positions, moderator conditions, and operating history. The shape of the distribution curve vill be more accurately described during the detailed core design.

    '

00000180 ' O-l ~ , l 3-34

                                                                                                                                        -.
  .
d. Engineering Hof, 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 dimensions, and flow channel geometry from perfect physical quantities and dimensions.

The application of hot channel factors is described in detail in

                                   .3 2 3 2.2, Statistical Core Design Technique. The factors are de-termined statistically from fuel assembly as-built or specified data where F is a heat input factor, F an is a local heat flux factor at a hot s t, and F is a flow area reduction factor describing the variationincoolbtchannelflowarea. Several subfactors are com-

, bined statistically to obtain the final values for Fq, Fqa, 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. Table 3-12 Coefficients of Variation CV No. Description e i CV 1 Flow Area 0.00075 0.17625 0.00426 O's 2 Local Rod Diameter 0.000485 0.420 0.00116 3 Average Rod Diameter 0.000485 0.420 0.00116 (Die-drawn, local and average ssme) 4 Local Fuel Loading - 0.00687 Subdensity 0.00647 0 95 0.00681 Subfuel area 0.000092 0.1029 0.00089 (Diameter effect) 5 Average Fuel Loading 0.00370 Subdensity 0.00324 0 95 0.00341

Sublength 0.16181 144 0.00112 l Subfuel area 0.000092 0.1029 0.00089
                                                                                                                                    '

l (Diameter effect) - 6 Local Enrichment 0.00323 2.24 0.00144

7. Average Enrichment 0.00323 2.24 0.00144 CV Coefficient of Variation e/E
                                '

e Standard Deviation of Variable i Mean Value of Variable

    .
        .

(Enrichment values. are for worst case normal assay batch; 000001M j maximum variation occurs for minimum enrichment. )

  • I,

( t i l !, 3-35 i. i _ . . _ . _ , - _ _ _ . . _ , . _ _ _ _ _ _ _ _ _ . _ ..__. _.

                  . _ _ _ .           -__                                    _

l 1 4

e. Core Flow Distribution Hot Channel Factors g

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 intemals is being per-formed to demonstrate the adequacy of the internal arrangements. The final variations in flow will be determined when the tests are com-pleted. Interim fsetors for flow distribution e.ffects have been cal-l culated from test data on reactor vessel models for previous pres-surized water reactor designs. i A flow distribution factor is determined for each fuel assembly loca-l tion in the core. The factor is expressed as the ratio of fuel as-l sembly flow to average fuel assembly flow. The finite values of the l ratio may be greater or less than 1.0 depending upon the position of ! the assembly being evaluated. The flow in the central fuel assemblies l 13 in general larger than the flow in the outermost assemblies due to l the inherent flow characteristics of the reactor vessel. l l The flow distribution factor is related to a particular fuel assembly i 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 applying model test isothermal flow distribution data and heat input effects g

                                                                                    .

, l at power as outlined in 3 2 3 2.ki. Two assumptions for flow distri-l bution 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 channe!, 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 will result in improved statistical statements for the maximum design condition.
f. Maximum Reactor Design 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 protertive 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, will 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' analyze the reactor at the maximus design conditions described previously. The total number of fuel rods in the core that have a 00000182 3-36

                                                                                                      -
                    .
                                                                                       ~

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 Dr 114 per cent power with the design FA h nuclear of 1.85 Line 2 was calculated using the maximum calculated value for F Ah nuclear of 179 to show the margin between maximum calculated and design con-ditions. It is anticipated that detailed core nuclear analyses will permit a lowering of the maximum design value for F Ah. The number o'f fuel rods that may possibly reach a DNB at the maximum design condition with an FAh of 1.85 and at 114 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 12 the core are in no jeopardy of experiencing 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 sumarized as follows in Table 3-13 Table 3-13 DNB Results - Maximum Design Condition (99 per cent Confidence Level) Power, Poesible Population Point $ of 2,452 Wt Fah DNB's Protected, % A 114 1.85 184 99 50 B 114 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 Design Condition 1.nalysis Sumary 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 hao the mechanical and heat transfer character-istics 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 characteristics. A more realistic indication of the number of fuel rods in jeopardy may be obtained by the application of the l

statistical heat transfer data to average rod power and mechanical , conditions.

    'O               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 I 3 37 00000183

 ..    - . .        .                 .      _   . -     - -    .. _ _ . . . -             .   .-  --   -

__ 1 l analysis are shown in Figure 3-14. The analysis may be summarized as 1 follows in Table 3-14. ,

                                                                                            ,

Table 3-14 DNB Results - Mast Probable Condition Power, Possible Population Point  % of 2,452 Mit Fah DNB's Protected,% F 100 1 79 2 99 994 G 114 1 79 32 99 913 H 118 1 79 70 99 815 The analysis was made from Point F at 100 per cent power to Point H at 118 per cent power to show the sensitivity of the analysis with power. The worst condition expected is indicated by Point G at 114 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 (DF. . This result forms the basis for the following statistical sta ement 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.

1. Distribution of the Fraction of Fuel Rods Protected The distribution of the fraction (P) of fuel rods that have been sh'own statistically to be in no jeopardy of a DNB has been calculated for the maximum design and most probable design conditions. The com-puter programs used provide an output of (N) number of rods and (P)

, fraction of rods that will not experience a DNB grouped for ranges of l (P). The results for the most probable design condition are shown l 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 condi.tions as shown in Figures 3-13 and 3-14. 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 3-14. 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 3-14. The upper curve shows P and (1-P) at the 114 per cent' power condition represen-l ted by Point G of Figure 3-14. The integral of N and (1-P) of the i upper curve forms the basis for the statistical statement at the most ,

 '

probable design condition described in para 6raph h above. 00000184 3-38

                                          .
                                            . . - , , -                        . - -

_._

                                   .
                                                                                  -
    ?           j. Hot Channel Performance Snmmary The hottest unit cell with all surfaces heated has been axamined for hot channel factors, DNB ratios, and quality for a range of reactor powers. The cell has been examined for the maximum value of Fah nuclear of 1.85    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 07 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 reflec-ted as an increase in AT of the coolant.

Error bands of 65 psi operating pressure and i 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. 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 138 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 ' F ., = 1.013 q F ==0 992 A The hot channel exit quality for various powers is shown in Figure 3-17 The combined results may be summarized as follows: Reactor Power, % DNB Ratio (BAW-168) Exit Quality, % 100 1.60 o 107 5 (trip setting) 1.47 2.6 114 (maximum power) 1 38 54 149 1.00 23 0 i 3 2 3 1.2 Fuel and Cladding Thermal Conditions ,

a. Fuel 000001h,-j A digital computer code is used to calculate the fuel temperature The program uses uniform volumetric heat generation across the fuel l - disneter, 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 o.f the temperature variation. Values for fuel conductivity were used as

 -
                                     '.
                                     .;

l 3-39 L t

    . . .         _             ..               .- _. _. ~ _...__ -          . . - _ _      , _ , . _ . _ .     - , . - _ . . _ . .
                                                            -.
                                                                       - .

shown in Figure 3-18, a plot of fuel conductivity versus te=perature. The heat transfer from the fuel to the clad is calculated with a fuel g 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 kv/ft is shown in Figure 3-19 for beginning-of-life conditions. The linear heat rate at the maximum overpower of 11h per cent is 19.9 kv/ft. The corresponding center fuel te=perature shown in Table 1-2 is h,h00 F. The center and average temperatures at 100 per cent power are k,160 and 1,385 F as shown in Table 3-1. The peaking factors used in the calculation are FAh = 1.85 Fg = 1.70 F ,, = 1.03 q F .(nue, and mech.) = 3.2h q . A conservative value of 1.03 was assu=ed for the heat flux peaking factor, F ,,. g The assigned value corresponds to a 99 per cent con-fidence and 99.99 per cent population-protected relationship as de-scribed in the statistical technique.

b. Clad The assumptions in the preceding pa agraph were applied in the calcu-l 1ation of the clad surface temperature at the maximum overpower.

l ' Boiling cond tions pervail 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 claculated clad temperature is 65h F at a system operating pressure of 2,185 psig. 1 00000186 4 3-h0 _

_ O.

 '

3232 Thermal and Hydraulic Evaluation 3232.1 Introduction Summsu y results for the characteristics of the reactor design are presented in 3231. The Statistical Core Design Technique employed in the design repre-sents a refinement in the methods for evaluating pressurized water reactors. Corresponding single hot channel DNB data were presented to relate the new method with previous criteria. A comprehensive description of the new tech-l 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 DNB relationships included in this section. A comparison of the BAW-168 correlation with other correlations in use is also included. 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. I 3 2 3 2.2 Statistical Core Design Technique The core thermal design is based on t. 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-I ematical treatment of the calculation. The method reflects the performance of the entire core in the resultant power rating and provides insight into the i reliability of the calculation. This section discussee 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 thrs yrformance of the entire core. The result then will provide a good measure of t'.e core safety and reliability since the method provides a statistical statement for the total core. This statement also reflects the conservatism or design aargin in the calculation. l A reactor safe operating power has always been determined by the ability of the l coolant to remove heat from the fuel material. The criterion that best measures

  • this ability is the DNB, which involves the individua.1 parameters of heat flux, coolant temperature rise, and flow area, and their intereffects. The DNB cri-r terion is commonly applied through the use of the departure from nucleate boil-

! ing ratio (DNBR). This is the minirun ratio of the DNB heat flux (as computed I by the DNB correlation) to the surface heat flux. The ratio is a measure of l the. margin'betiieen the operating power and the power at which a DNB might be l expected to occur in that channel. The DNBR varies over the channel length, and it is the min 4=nm value of the ratio in the channel of interest that is V used. .

                                                            -

00000187 3-41

     - -                  ._ -  _.         .

The calculation of DNB heat flux involves the coolant enthalpy rise and coolant 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 statistical hot channel factors required are a heat input factor, Fq, and a flow area fac-tor, FA. In addition, a statistical heat flux factor, Fq , ais required; the heat flux factor statistically describes the variation in surface heat flux. The DNBR is most limiting when the burnout heat flux is based on mini == flow area (small F A) and maximum heat input (large Fq), and when the surface heat flux is large (large F Q"). 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 YM 1 DNBR = B.F.

  • A' Q
  • Q"surface xFnQ involves as parameters statistical hot channel and DNB factors. The DNB factor (B.F. ) above is usually assigned a value of unity when reporting DNB ratios so that the margin at a given condition is shown directly by a DNER greater than 1.0, i.e., 1 38 in the hot channel. -

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 ( &C ). A typical histogram is shown in Figure 3-20. g 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 ( casily obtained for the histogram. These same mathematical parameters are the i basis for the statistical burnout factor (B.F. ). 1 ( 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 l tolerances for the proposed core. Regardless of the source of data, the sub-i factors can be shown graphically (Figures 3-21 and 3-22). All the plots have the same characteristic shape whether they are for subfactors, hot channel factors, or turnout factor. The factor increases with either in-creasing population or confidence. The value used for the statistical hot chan-nel and burnout factor is a function of the percentage of confidence 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 frequently used assumption in statistical analyses is that the data available represent an infinite sample of that data. The implications of this assumption should be - noted. For instance, if limited data are available, such an assumption leads g

                                          ,_,

00000188

_

 !D U'    to a somewhat opt'imistic result.' The assumption also 'impl'ies 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 PE4 calculationa'.

procedure does not make this assu=ption, but rather uses the specified sample size to yield a result that is much more meanin6ful and statisti-cally 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 F,=1tKa p ,, q where F qi, is the statistical hot channel factor for heat flux K is a statistical multiplying factor ep ,, is the standard deviation of the heat flux fac-tor, including the effects of all the subfactors If 7 pg ., = 0.05 for 300 data points, then a K factor of 2.608 is re-i quired to protect 99 per cent of the population. 'The value of the hot channel factor then is . Fq,, = 1t (2.608 x 0.050) = 1.1304 and vill provide 99 per cent ec.nfidence for the calculation. If, in-stead of using the-300 data points, it is assumed that the data repre-sent an infinite sa=ple, then the K factor for 99 per cent of the popu-lation is 2 326. The value of the hot channel factor in this case is l F,=1t q (2 326 x 0.050) = 1.1163 ! which implies 100 per cent confidence in the cale at en. The values l of the K factor used above are taken from SCR-607, 21 The same basic techniques can be used to handle any situation involving variable confi-dence, population, and number of points. Having established statistical hot channel factors and statistical DIO 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 peakin6 factor. Since this fraction is known, the maxi-mum fraction in jeopardy is also known. It should be recognized that every rod in the core has an associative DIG ratio that is substantially greater than 1.0, even at the design overpower, and that theoretically no rod can have a statistical population factor of 100 per cent, no cat-ter how large its DIG ratio. f._ 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 3-u3 00000189

                                        -
                                         . - -                  . -    -.

can be obtained by simple summation. Se 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 peak-g ing factor summed over all peaking factors can be made in a statistically rig-orous manner only if the confidence for all populations is identical. If an infinite sample is not assumed, the confidence varies with population. To form this 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-protected 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: 2 ere is at 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 continuous operation at the design overpower. Se 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 F Ah by a factor of 1.033 to provide additional design margin.
b. The maximum value for Fz (nuclear max / avg axial fuel rod heat input) is determined for the limiting transient or steady state condition.
c. Every coolant channel in the core is assumed to 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 maldistribution.
e. Every fuel rod in the core is assumed to have a heat input greater than the maximum calculated value. This value is represented by engineering hot channel heat input factors, Fq and Fqn, which are greater than 1.0.
f. Every channel and associated fuel rod has a heat transfer margin above the experimental best-fit limits reflected in DNB ratios greater than 1.0 at maximum overpower conditions.

The statistical core design technique may also be used in a similar manner to evaluate the entire core at the most probable mechanical and nuclear conditions to give an indication of the most probable degree of fuel element jeopardy. The result of the technique based on the most probable design conditions leads to a statistical statement which is a corollary to the maximum design statement: 00000l90 , O 50 h bl 3-% r u -

                                                         =*-w p                   -- W

_

                                                                                   ~

There is at least a 99 per cent confidence that at least 99 9'per ~

 '

cent of the rods in the core arc in no jeopardy of experiencira a DNB, even with continuous operation at the design overpower. The most probable design conditions are assumed to be the same as the maximum design conditions with these exceptions:

a. Every coolant channel is assumed to have the nominal flow area (FA = 1.0).
b. Every fuel rod is assumed to have (1) the maximum. calculated value of heat input, and (2) Fq and Fqa are assigned values of 1.0.
c. The flow in each coolant channel is based on core flow and power distributions.
d. Every fuel red is assumed to have a nominal value for FA h nuclear.

The full meaning of the maximum and most probable design statements re-quires additional co= cent. 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 fever than O V 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 nu=ber could occur for the conditions assumed. In summary, the calculational procedure outlined here represents a sub-stantially improved design technique in two ways:

a. It reflects the performance and safety of the entire core in the resultant power rating by considering the effect of each rod on the power rating,
b. It provides infor=ation on the reliability of the calculation and, therefore, the core through the statistical statement.

32323 Correlation of Heat Transfer Data The BAW-168 report (Ref. 18) serves as a reference for the "best-fit" fom of the design relationship used by B&W. This heat transfer corre-lation has been found to be the most satisfactory in the representation of both uniform and nonuniform heat flux test data. The BAW-168 correla-tion 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 mint::um ist selected as the DNB point, and that value of the ratio at that" point is the DNB ratio (DNBR) for that chan-U nel.

                              .

g, 00000191

                         -. _      -                  .  .    ..      .-.   .-

_ _ _

                                                 .

This particular discussion deals with the comparison of DNB data to g three particular correlations. The Tc correlation in the case of BAW-168,(e1 trelations selected

                                            ) a cor gtion  withwere whichthethe B&W industry is familiar in the case of WAPD-188, eci and a correlation re-cently proposed for the case of W-3.s23)use in the design of pressurized water reactors in The data considered for the purpose of these comparisons were taken from the following sources:
a. WAPD-188 (Ref. 22).
b. AEEW-R213 (Ref. 24).
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 Cc=pany Euratom Data (Ref. 30).

The co=parison of data to the BAW-168 correlation is presented as histo-grams of the ratio of the experimental DNB heat flux ($ E ) to the calcu-lated heat flux (4C). The data from each source were Brouped by pres-sure and analyzed as a 6roup; batches were then prepared including com-con pressure groups from all sources. Altogether there are 41 different & data Broups and batches considered. Histograms for only the BAW-168 W correlation are presented to minimize the graphical material. 'Ihe in-fomation required for the generation of histogra=s of the other two correlations was also prepared. The co=parison of the various correlations to each other is facilitated throu6h the use of tabulations of pertinent statistical parameters. The standard deviation and mean value were obtained from the computed values l ' of ($ E/? C) for each group or batch. A comparison of standard deviations is somewhat indicative of the ability of the correlation to represent the data. ! However, differences in mean values from group to group and correlation to correlation tend to complicate this type co=parison. A relatively simple method may be used to compare the correlations for various data; ! l this method uses the coefficient.cf variation (Ref. 31) which is the ratio of the standard deviation (e) to the mean Y. The coefficient of variation may be thought of as the standard deviation given in per cent; it essentially nomalizes the various standard deviations to a ec= mon mean value of 1.0. ' Table 3-15 is a tabulation of the data source, heat flux type, and cor-responding histogram nu=bers. The histo 6ra=s are shown on Figures 3-24 through 3-39 Y 00000192 3-k6

                                                    -    -
            .            __                 -                                   .

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  • Q' Table 3-15 Heat Transfer Test Data' Histogram Figure Source Heat Flux Type Number Number WAPD-188 Unifom 1-9 3-24 3-25 3-25 AEEW-R-213 Unifom 10-14 3-26 3-27 3-28 Columbia Unifom 15-19 3-28 3-29
                        '

3-30 ANL Unifom 20 3-30 B&W Unifo m 21 3-31 B&W-Euratom Unifom 22-24 3-31 k[ 3-32 Combined Data (500-720 psia) Unifom 25 > 3-32

         ,

Combined Data (1,000 psia) Unifom 26 3-33

       ,

Combined Data (1,500 psia) Unifo m 27 3-34 Combined Data (2,000 psia) Unifom 28 3-35 Combined Data (1,750-2,750 psia) Unifor= 29 3-36 B&W-Euratom Chopped Cosine Nonuniform 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) Nonunifo m 39 3-39 Combined Nonunifom (1,500 psia) Nonunifom 40 3-59 Combined Nonunifom (2,000 psia) Nonunifo m 41 3-39 . D ' g 00000193

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1 The histograms graphically demonstrate the distribution of ($ each data group. The Gaussian type distribution of ($E/@C) a ou the

                                                                             ) for     g mean for the group is apparent in the large data grour,. Some data groups are too amall to provide meaningful histograms, but they are pre-sented in order to complete this survey.

The data were used as presented in the source for the calculation of ($E/@C); n p ints were discarded for any reason. A good correlation should be capable of representing DNB data for a full range of all per-tinent parameters. The result of the comparison on this basis is dem-onstrated in Table 3-15 The data source, pressure, histogram figure nu=ber, heat flux type, and number of data points in the group are tab-ulated. For each of the three correlations the following data are indi-cated: e/i The coefficient of variation based on all available data in the group. nR The number of data points rejected using Chauvenet's crite-rion (Ref. 31). This criterion is statistical in nature and is applied to the values of ($ /@C E . ap n s @at fall outside certain limits with respect to the main body of data are rejected. (e/5)' The coefficient of variation based on the original data sam-ple less those points rejected by Chauvenet's criterion, i.e., based on n-ng values of ($ /@C)* E h, It is unfortunate that Chauvenet's criterion must be applied to the values of ($E/@C) rather than to the original data, since application to ($ 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 so=e bad data. The logical choice then f.s to present data both ways, i.e., with art. without Chauvenet's criterion ap-plied. Of the 41 groups and batches analyzed the following is observed from Table 3-15: l Groups and Batches of Data Groups and Batches of Data With Smallest er/i Without With Smallest e/5 With , Correlation Chauvenet's Criterion Chauvenet's criterion l l BAW-168 38 36 WAPD-188 2 3 00b00194 0 3-48 ! _ _ _

_ Chauvenet's criterion rejected the following number of points for each correlation: Uniform Nonunifom Total BAW-168 (Groups only) 32 1 33 BAW-168 (Batches Oniv) 39 0 39 WAPD-188 (Groups only) 34 2 36 WAPD-188 (Batches only) 33 o 33

 ,              W-3 (Groups only)               59        12          71 W-3 (Batches only)              50          9         59 Several no able peculiarities exist in the tabulation of Table 3-16
                   .

The Columbia 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 (e/T )' are shown as not available (N.A.). Neither the BAW-168 nor the WAPD-188 predicted any negative DNB heat fluxes; the W-3 predicted 93 negative values for unifom data. The fact that only 59 were re-jected for this correlation indicates that the remaining 34 uniform points which were negative (93-59'= 34) were close enough to the body of the data to be considered statistically significant. Table 3-16 may be consolidated somewhat as below by 'abulating the number of groups C)

     '"'  " '*"'" ' " ' " "" " "***" ' '' " ""'" ' ""*"

interval for each correlation. (e/I) Interval BAW-168 BAW-168:(a) WAPD-188 WAPD-188'(a) w.3 _ w 3e(a) Negative o o o o 2 o o-o.1 6 8 o o o 1 0.1-0.2 24 24 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-o.5 1 0 5 7 5 6 0 5-o.6 o o 6 5 3 4 o.6-o.7 o o 3 2 1 1 o.7-o.8 o o 2 1 7 8 o.8-o.9 1 o 0. 0 1 5 0 9-1.o o o o o 1 o Greater than 1.0 o o 2 0 16 7 Total 41 40 41 40 41 40 ('}Chauvenet's criterion applied. (

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_ As is seen from the tabulation the column for.BAW-168 with Chauvenet's criterior spplied 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 74018. For W-3 the spread is still greater, and a maximum value of 1 7483 is noted. The negative values of DNB heat flux predicted by the W-3 correlation are in part respon-sible for the large spread in (e/ T). The ability of the BAW-168 correlation to fit both unifom and nonuni-form heat flux data over a vide range of pertinent variables leads us to believe that it is the best DNB correlation available. 3 2 3 2.4 Evaluation of the Thermal and }{ydraulic Design

a. Hot Channel Coolant Quality and Void Fraction An evaluation of the hot channel coolant conditions provides additional confidence in the themal 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 40 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 FA h of 1.85 Additional calculations for a 10 per cent increase in FA h to 2.035 were l p made at 114 per cent power. The significant results of both j Q calcula ions are summarized in Table 3-17 The effects of using at Fah of 179 are shown in Fig'tre 3-40 I Table 3-17 Hot Channel.Cociant Conditions Exit Exit Void Operating Pcuer, % Fah gality,% Fraction, % Pressure, psig 100 1.85 (-)2.4(D) 0 5(a) 2,185 l 114 1.85 2.8 13 5 2,185 130 1.85 9.4 36 9 2,185 114 2.035 8.7 2,185 100 1.85 0 35 380(a) 2,120 114 1.85 5.4 25 2 2,120 130 1.85 1r.1 45 2 2,120 114 2.035 11.' 43 4 2,120 l (")Subcooled volds. ( (b) Negative indication of quality denotes subcooling l of 10.2 Btu /lb. l ! O 00000197 3-51'

   .         _
                   . _ _ _ -           -_         _
                                                             , _ .      _.
                                                     - _

The conditions of Table 3-17 vere determined with all of the hot channel factors applied. Additional calculations were made for unit cell channels without engineering hot channel factors to show h the 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 assembly flow distribution hot channel factors at 2,185 psig as shown on Figure 3-41. These results show that the exit qual-ities from the hottest cells should in general be considerably lower than the maximum design conditions.

b. Core Void Fraction The core void fractions were calculated at 100 per cent 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 as-sembly flow distribution was checked by determining the total voids for both 100 and 95 per cent total core flow for the two pressure conditions.

The results are as follows: Flov, % Pressure, psig Core Void Fraction, % 100 2,185 0.007 100 2,120 0.033 95 2,185 0.041 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 program uses a combination of Bowring's (33) model with Zuber's (3k) correlation between void fraction and quality. The Bovring model considers three different regions of forced convec-tion boiling. They are: (1) Highly Subcooled Boiling In this region the bubbles adhere to the vall 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 Icttes equation. 'lhe highly subcooled region ends when

                          -T bulk = n6V                           (A) sat 00000198      9
             .

3-52 sf^

_ O ' r-

                               &.1calheatf1ux, Btu /hrf12 y = 1.863 x 105 (14 + o.co68p)

V=velocityof" coolant,ft/sec p = pressure, psia 2e 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 frcm the void fraction as follows: 1 x{= , pf 1 (B)

 '

lt 1 h ad where , xd = nonequilibrium quality at end of Region 1 ad : void fraction at Tsat - Tbdk* l Pf : liquidcomponentdensity,lb/ft3 L A l V Pg c vaporcomponentdensity,1b/ft3 (2) Slightly Subcooled Boiling . In this region the bubbles depart from the vall and are trans-ported along the channel (condensation of the bubbles is ne-glected). Bis 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 regica is conipUted from the'

                                                                              ~
                                                                                                              ~

l l following formula: i

  • P x*=xj+m.  !

Z (& - &gp)dz (C) hfg(1 + e ) d , x* = nonequilibrium quality in Region 2 l ! hfg=latentheatofvaporization, Btu /lb l 1 = fraction of the heat flux above the single 1+e phase heat flux that actually goes to pro- ' ducing voids

       .
   .[% -                          &gp c    singlephaseheatflux, Btu /hr-ft2
            , .. . c. '  -

n0000199 3-53 __. - _ _ - .-- -. . -- - - , . . . . . - - .

_ b = mass flow rate, lb/hr Ph = heated perimeter, ft O' z: channel distance, ft 2e void fraction in this region is computed from

  • a= X
                                            -                                    (D)
                    ',                          38 3 Af Pg 'agge(P f - Pg )" 1/4 C    x o        + Pg /Pf (1 - x*)  +     ,

m

                    .                       .              .

f . Where g: accelerationduetogravity,ft/sce2 g  : " e constant in Newton's Second Law : 32.17 2 Co : Zuber's distribution parameter A : flow area, ft2 f e: surface tension Equation (D) results from rearranging equations found in Re-ference (34) and assuming bubbly turbulent flov in determining the relative velocity between the vapor and the fluid. Zuber has shown that Equation (D) results in a better prediction of 31 W 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 usin placing x*g Equation (D) with the themodynamic quality, x, re-

                       .

!

c. Coolant Channel Hydraulic Stability l

l A flow regime map was constructed to evaluate channel h;/draulic sta-l bility. Se transition from bubbly to ar3nular flow at high mass ve-locities was determined using Bakers's(3b> correlation, and the tran- ! sition from bubbly to slug was determined with Rose's(flgv which occurs at low mass velocities 37/ correlation. The tr9n flowto(angularflowwasdeterminedbyHaberstroh'sug4tionfromslug> correlation. Bergles 39) found that these correlations, which were developed from l l adiabatic data, are adequate for locating flow regime transitions I with heat addition, and that they adequately predict the effects of pressure. Figure 3-42 shows the flow regime map on which has been i 00000200 3-54 __

plotted a point representing operating conditions in the hot channel at 114 per cent overpower. To aid in assessing the conservatism of the design, an additional point is plotted at 130 per cent overpower. - Inspection shows that both pointo lie well within the bubbly flow regime. Since the bubbly flow regime is hydraulically stable, no flow instabilities should occur. This flow regime map was prepared for the hot unit cell at the me.ximum design condition characteristics outlined in 3 2 3 1.1. 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 t'e sensitivity of the analysis with respect to mass flow rate, channei dimensions and mixing intensity in unit, corner, and wall-type cells. The resulto are shown in Figures 3-43, 3-44, and 3-45 The mass velocity and quality in each type of channel for the two cases are plotted on *,t.e. figures. The conditions assumed for the nominal and postulated erst case are given in 3 2 3 2.4 J. Data from the burnout tests performed by B&W on a 9-rod bundle simu-lating the core geometry 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 bundle 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 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 determined for the BAW-168 and W-3 correlations. The results are shown it. Figure 3-46. DNB ratios for both correlations are shown for the 1 50 axial max /avs symmetrical cosine flux shape from 100 to 150 per cent power. me 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 calculated 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 quality condition of 20 per cent. The W-3 calculation is accurate to about 130 per cent power, but because of quality limitations it cannot be used to examine the channel at the 150 per cent power condition.

The sensitivity of DNB ratio with F Ah and Fz nuclear was examined from 100 to 114 per cent power. De 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 00000201 3-55

 .
                  . .                  ,           -

l I i 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-DIB ratio plot in 3 2 3 2.2, statistical core De-g sign Technique. The influence of a change in F Ah 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 design value of 1.85 'lhe resulting BAW-168 DNB ratio is 1.22 and 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 Flov 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 detemined to insure that the required system flow is obtained in the as-built plant. The experi-mental programs previously outlined in Section 1 vill confirm the pressure drop and related pump head requirements. It is anticipated that the as-built reactor flow will exceed tne design value and vill 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 138 (BAW-168), and (b) the statistical core design criteria. are shown in Figure 3-47 The results of the analysis The power shown is the 100 per cent rating, h and the limiting condition is 114 per cent of the rated power. An examination of the slope of the curve indicates stable characteristics, and a 1 per cent change in flow changes the power capability by only about 1/2 per cent.

f. Reactor Inlet Temperature Effects The influence of reactor inlet temperature on power capability at a given flow was evaluated in a similar manner. A variation of 1 F in reactor inlet temperature vill 2 esult 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 columnnr 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 equisxed grain growth - 2,700 F. (2) Outer limit of columnnr grain growth - 3,200 F. (3) Outer limit of molten fuel (UO2) -5,000F.g 3-56

                         ._

_

 $ .                      Data from References 40 through 43 were used to compare calculated ana 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 the fuel over the temperature range considered This model combined with the model for calculating the heat tranaf cefficient was ccm-pared with the model developed by Notley et al of A3CL. The dif-ference in fuel growth for the two calculation models was less than the experimental scattei -f data. Se 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. Se 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. D e conductivity of the gas in the gap is detemined as a function of burnup and subsequent release of fission product gases. In the event of fuel clad contact, contact coefficients are determined on the basis of methods suggested by Ross and Stoute(48). Se contact coefficient is determined as a function of the mean conductivity of the interface materials, the contact p pressure, the mean surface roughness, the material hardness, and the f 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 sccording to the new ge-ometry. De program 2tes a polynomial fit; relationship for fuel taermal con-l ductivity. " ree relationships were used to evaluCe the effects of ! conductivity A comparison of the reference design 5) is shown C1fg' -)142@thesein Figure 3-48.conductivity De values relatio suggested in GEAP-4624\ e.nd CVNA-246 @7) are very similar up to l 3,000 F and the former values are more conservative above 3,000 F. McGrath 471 concludes that the CVNA-246 values are lover 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 evalus,e the margin to central melting at the maximum overpower and to show the sensitivity of the calculation with respect to themal conductivity. Since the power peaks vill be burned off with irradiation, the peaking factors used are conservative at end-of-life. I aesu1ts 00000203 2e results of the analysis with the methods described above are shown n V in Figures 3-49 and 3-30 for beginning and end-of-life conditions. Re beginning and end-of-life gas conductivity values are 0.1 and 0.01 Btu /hr-ft-Frespectively. 2 The calculated end-of-life center fuel

                                                       .3-57

! ._ _- _ _ _ . _ . ., _ .__ _ . _ - __ . . _ . _ _ ~_ _ -- _ . _ _

temperatures are higher then the beginning-of-life values because of the reduction in the conductivity of the gas in the gap. The effect is apparent even though a contact condition prevails. The calcula- g tion does not include the effects of fuel swelling due to irradiation. S e calculated contact pressures are conservatively lower than those expacted at end-of-life conditions in the hottest fuel rods, and the fuel temperatures shown in the above figures are conservatively higher. Se BW model gives very good results when compared 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 am11 differences between the BW curve and the other curves indicate the relative insensitivity of the results to the shape of the conductivity at the elevated te=peratures. The most conservative assumptions, using GEAP-4624 data 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 max hum design value of 19 9 kv/ft at 114 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-4624 fuel thermal conductivity curve to reflect a conservative value for the max hum average te=perature and stored energy in the fuel. Use of this curve results in a his;her tempera-ture and therefore a lover Doppler coefficient, since it decreases g with temperature. Bus the resultant Doppler effect is also con-servative.

h. Fission Gas Release

, The fission gas release is based on results rted in GEAP-4596. b9) ! Additional data from GEAP-431k(50), AECL-603 , and CF-60-12-14(52) have been compar9d v1th the suggested release rate curve. Se re-lease rate curvel4 9) is representative of the upper limit of release data in the temperature region of most importance. A design release ! rate of 43 per cent and an internal gas pressure of 3,300 psi are used to determine the fuel clad internal design conditions reported in l 3 2.4.2:. Fuel Assemblies. , The design values for fission gas release from the fuel and for the I 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 l provided without utilizing tne initial porosity voids present in the UO2 fuel. A detailed analysis of the design assumptions for fission gas release, and the relationship of burnup, fuel growth, and initial dia=etral 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. O 3-58 00000704

                                                       .     -
            ..

H . (1) Design Assumptions - (a) Fission Gas Release Rates h e fission gas release rate is calculated as a function of fuel temperature at the design overpower of 114 per cent. Se procedures for calculating fuel temperatures are dis-

                      . cussed in 3 2 3 2.4 g. S e 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. Se fuel temperatures were calculated using the GEAP-4624 fuel thermal conductivity curve to ob-tain conservatively high values for fuel temperatures.

(b) Axial Power and Burnup Assumptions Se temperature conditions in the fuel are detemined for the most severe axial power peaking expected to occur. Two axial power shapes have been evaluated to determine the max-imum release rates. Researe150and170 max /avgshapes

                      'as shown_in Figure 3-11 and repeated as part of Figure 3-52 of this analysis. Se 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.

Be quantity of fission gas produced in a given axial loca-tion is obtained from reactor core axial region burnup studies. B ree 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. S e 930-day, or end-of-life condition, is e : condition with the maximum fission gas inventory. Le average burn-up at the end of life in the hot fuel rod is 38,150 MWD /M which has been determined as follows: CalculatedHotBundleAverageBurnup, MWD /m 33,000 Hot Fuel Rod Burnup Factor 1.05 Margin for Calculation Accuracy 1.10 . HotRodMaximumAverageBurnup, MWD /m 38,150 l Re local burnup along the length of the fuel rod is te product of the hot rod maximum average value above the local to average ratio shown in Figure 3-52. De ting hot rod local maximum burnup for the 930-day, end- fe condition is about 42,00016fD/M. Bis is the maxis _ ,

   ,O   ~

calculated value. However, local values to 55,000 MWD /m l have been evaluated to insure adequate local fuel cladding l 00000205 L 3-59

    - .

__ 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 a siptions De naximum hot rod total power occuring 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. Bis results in a max-imum Jinear heat rate of 19 9 kv/ft which corresponds to 114 per cent of the maximum linear heat (17.49) shown in l Table 3-1. This is a conservative assumption when coupled l vith the end-of-life fission gas inventory since bundle ara individual fuel rod poaer 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. me peak bundle ratio of 1.69 (1.85 + hot rod ratio) vill only occur during the first two fuel cycles when the fission gas in-ventory is less than the maximum value. (d) Fuel Growth Assumptions The fuel growth was calculated as a function of burnap as indicated in 3 2.4.2.1. Fuel pellet dimensions in the ther-mal temperature and ges 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 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 when the fuel expands enough to contact the clad. 2 based ca 43 per cent A gas conductivity of 0.01 Btu /hr-ft -F release of fission gas at the end-of-life condition was used in the analysis. Diametral clearances of 0.0025 to 0.0075 in. reflecting minimum and maximum clearances after fuel growth were analyzed. S e contact heat transfer coefficients were calculated as suggested in Reference 48. (2) S - m of Results The fission gas release rates were determined in the first eval-uation. Rates were found for various cold diametral clearances and axial power peaking and burnup shapes. The results are shown in Figure 3-53 The levest curve is the expected condition for a 1 70 axial power shape with a 930-day axial burnup distribution l as shown in F16 ure 3-52. Se increase in release rate with , 1 3-6 00000206

              . -                          .
        ..

i diametral clearance results from the fact that the fuel' tempera-ture 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 produces the max - imum clad stress due to fuel growth with irradiation. Se assem-bly of maximum size pellets with minimum internal diameter clad-L ding will produce this condition after fuel growth. In the event a few hot pellets have the maximum diameter and the remainder have the minimum diameter, then the average cold gap would be 0.0035 in. producing a slightly larger release rate. The re-lease rate of 33 per cent for the maximum diametral clearance wil1 not occur with the maximum stress condition due to fuel growth, since the fuel nan grow into the cle vance. 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 2e 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 witha170 max /avgratioassnownonFigure3-52. similar re-sults are shown for the 1 50 max / avg ratio. Rese curves show the release rates expected are not strongly influenced by the various power and burnup shapes. The second evaluation shows the resulting internal pressures due , to the release of fission product gases. Plots of pressures for the expected 930-day axial burnup distribution and a 1.70 max / avg axial power shape are shown in Figure 3-5h. 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). Se upper data band is for a closed pore condition with all released gas contained outside the fuel pellets in spaces between the ex-panded dished ends of the pellets, the radial gaps (if any), and ! the void spaces at the ends of the fuel rods. Se band of data shown reflects the effect of fuel densieication and grain growth described in 3 2 3 2.4. The upper limit is for an ideal thermal model without grain growth or densification; the lover limits are for the design model. The calculation of the maximum pres-I sure is also relatively insensitive to the axial burnup distri-i bution as shown by the dashed line in Figure 3-54 for a 150 maximum to average axial power and burnup shape. (21s corre-sponds to a local burnup peak of 37,000 MWD /MTU.) The allowable design inte nal 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 condition. A modest increase in average fuel burnup can be tolerated within the prescribed internal pressure design limits. It has been indicated in Reference 44 and in AECL-1598 that the _ UO2 fuel is plastic enough to flow under low stresses when the temperature is above 1,800 F. That fraction of the fuel below

                   .. ,
           -
                   -

3 61 0%WU

   .-        .-. .         _.   .
                                                       -- .             - -

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 vill 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 fraction 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 40 0.005 20 0.0075 5 The retention of fuel porosity in the lov temperatvre and lov burnup regions will result in modest reductions in internal gas pressure.

1. Hot Channel Factors Evaluation (1) Rod Pitch and Boving A flow area reduction factor is determined for the as-built fuel assembly by taking channel flow area measurements and statisti-cally determining an equivalent hot channel flow area reduction facter. A fuel asse=bly has been measured with the results h 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 shovn in Figure 3-21 for 99 per cent confidence.

With a 99 per cent confidence and 94.5 per cent population re-lationship described in 3 2 3 1.1 for the hot channel, the area reduction factor is 0 992. D e approximate limit of this factor j is obtained by examining the value in Figure 3-21 as the popula-l tion prote ted approaches 100 per cent. FA at 99 99 per cent of l the population protected is 0 983 2 e hot channel value is ! shown in Table 3-1. l l ' Special attention is given to the influence of water gap varia-tion between fuel assemblies when detemining 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 vater 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, vall cell, I and corner cell flow conditions. Y 3-62 00000208

                 ._                                                _                                          _ _ . _           ..

_ l (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 combined statistically as
-

described 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. Se local heat flux factor, Fqn, for 99 per cent confidence and 94.5 per cent population is 1.013 These hot channel values are shown in Table 3-1. Se corresponding values of Fq and Fqn with 99 99_per cent popula-tion protected are 1.017 and'1.03 respectively. A conservative value of Fqn of 1.03 for 99 per cent confidence and 99 99 per

 ;                                         cent population is used for finding the maximum fuel linear heat rates as shown in 3 2 3 1.2.

R ese factors are used in the direct solution for channel en-thalpies and are not expressed as factors on enthalpy rise as is often done. The coefficients of variation will be under con-tinuous review during the final design and develognent of the fuel assembly. (3) Flow Distribution Effects

'

Inlet Plenum Effects , F The final inlet plenum effects wi.L1 be determined from the 1/6 i scah model flow test now in progress. "he initial runs indi-cate satisfactory flow' distribution. Although the final nuclear i 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 isothennal 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 Dissimilar Coolant Conditions l ! The hot fuel assembly flow is less than the f_sw 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. B is effect is allowed for by making a direct calculation for the hot assembly flow. Se 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. Se worst combination of effects has been assumed in the initial design, and the hot as-f -. sembly flow has been calculated to be about 85 per cent of the S average assembly flow at 114 per cent overpower. Actual hot as-sembly flows are calculated rather than applying an equivalent ihot channel enthalpy rise factor. ~

               *
                       - - :/
                           .; :<; : ,                                           3-63                                               000002M
   ._. ..           __         _ . .         __ _ __ _ . _ _ _ . _                  . . _ _ _ _ . _ . . _ . . _ _ . _ _ . _ . .            . . . . . _

_ Physical Mixing of Coolant Between Channels The flow distribution within the hot assembly is calculated with a mixing code that alicvs an interchange of heat between channels. Mixing coefficients have been determined from multirod mixing tests. The fuel assembly, consisting of a 15 x 15 array of fuel rods, is divided into unit, vall, and corner cella as shown by the heavv lines in Figure 3-55 The mixed enthalpy for every cell n,t.etermined simultaneously so that the ratio of cell to average assembly enthalpy rise (Enthalpy Rise Factor) cad the corresponding local enthalpy are obtained for each cell. Typical enthalpy rise factors are shown in Figures 3-55 and 3-56 for cells j surrounding the hottest fuel rod located in the corner of the as-l sembly. The assumptions used to described the channels for the peaking and enthalpy rise factors shown are given in Wall and Corner Channels Evaluation, 3 2 3 2.4 j, which follows. J. Evaluation of the DNB Ratios in the Unit, Wall, and Corner Cells i DNB Results The DNB ratios in the hot unit cell at the maximum design condition ! described in 3 2 3 1 are shown in Figure 3 46. The relationships i shown are based on the a ' data in the BAW-168(10)and pplicat4on of) single W-3(23,66 channel An correlations. heatadditional transfer 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 channel data. , l The sensitivity of the assembly design with respect to variations of mass flow rate (G), channel spacing, mixing intensity, and local peak-i ing en the DNB ratios in the fuel assembly channels has been evaluated I by analyzing the nominal conditions and a postulated vorst case (sn-dition. The su= mary results are shown below in Table 3-18. Table 3-18 DNB Ratios in the Fuel Assembly Channels Nominal Case Cell Type G,1b/hr-ft2 x 10-6 DNBR Corner 1 59 2.20 Wall 1 90 2.11 Unit 2 52 2.01 Postulated Worst Case Cell Type G,1b/hr-ft2 x 10-6 DNBR Corner 1 32 1 70 Wall 1.6h 1.65 Unit 2.29 1 73 h

                                                                     '

3-64

                                           .               .             .      _ . .

The DNBR's sbove are ratios of the' limiting heat flux to the local flux along the length of the channels. The limiting heat flu::es have been determined from the 9-rod assembly DNB test data. S e 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 unit cell channel. The postulated worst case conditions are more severe than the required maximum design conditions. Tht rssults of the assembly tests and this evaluation show that the perfeinance of the vall and corner cells is more sensitive to local enthalp 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 appropriate correction for a nonuniform axial power shape. Typical results are shown in Figures 3-57 and 3-58 for the nominal and vorst case condi-tions in the corner cell. The line defined by a best fit of the data

   ~

is shown on cach figure as a solid line. A design limit line, sbovn as dotted, hau been determined by lowering the best-fit line to account (\.>} for the effects of nonuniform flux shapes. Se magnitude of the re-duction was determined by ecm:parison with the results of the. stom nonuniform test data Reference 19 and the results of more rec .t non-uniform tests conducted by B&W. The limiting best-fit lines were derived from a ;-rod fue.1 assembl*; test section 72 in. long with rod diameter, pitch spacing, and sacer grids of the type to be used in the reference design. A total of 513 data points between 1,000 psi and 2,450 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 ranges. The ranges of test variables for the 162 data points used were: t Pressure - 1,800 to 2,450 psi i Mass Flow Rate - 1.0 to 3 5 x 106 lb/hr-ft2 l 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 ' 1* 2 e nominal case loc'l-to-average a rod powers and the local-to-average exit enthalpy rise ratios are shown in Figure 3-55 for the het corner, 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 ' 3-65

          -.    ,     _-    ._          , , - .- - - . - -.- .. _ - - - . -                                    . . - - - -   .
 -
                                                                                              .

fuel assemblies with nominal rod-to-vall spacing, with nominal flow to g th9 hot fuel assembly, and with a nominal intensity of turbulence, w a (*), equal to 0.03 Additional tests are being run to determine the maximum values of in-tensity of turbulence associated with the fuel assembly. The expected 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 loccl-to-average rod powers and exit en-l thalpy rise ratios in the hot fuel assembly are shown in Figure 3-56. The factors were determined for this case with twice the nominal water gaps between the hot fuel assembly and adjacent fuel assemblies with minimum rod-to-vall spacing, with minimum flow to the hot fuel assem- , bly, and with a minimum assumed intensity of turbulence, a, equal to ! 0.01. ( l In neither the nominal nor the postulated vorst case analysis haa any credit been taken for the coolant which is flowing in the water gaps between the fuel assemblies and which serves to reduce enthalpi.es in the peripheral cells of the hot fuel assembly by mixing with the cool-ant in those cells through the can panel perforations. In both cases, however, the effective roughness of the can panel perforations and its effect on reducing the flow in the peripheral cells of the fuel assem-bly has b^en accounted for. B e magnitude of the effective roughness was obtained from the results of a series of flow tests performed on a mockup of the outer two rews of fuel rods and the can panels of two M W 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 reducing the flow in the peripheral cells. The nominal distance from the center of the outside rods to the can panel is 0 324 in. 2 e corresponding postulated vorst case dimension was assumed to be 0 310 ir. l Fuel Assembly power and Flow Conditions ! The nominal and postulated vorst cases were run at 114 per cent re-actor power with the nominal and worst F Ah factora shown in 3 2 3 1.1 c.

                he 150 modified cosine axial power shape of Figure 3-11 was used to describe the worst axial condition.
   .....

(a)The intensity of turbulence,e , is defined as V{/V whereV{isthetransverseco=ponentofthefluctuatingturbulentvelocity, l and V is the coolant velocity in the axial direction. This method of com-I puting mixing is described by Sandberg, R. O., and Bishop, A. A., CVTR g Thermal-Hydraulic Design for 65 K4 Gross Fission Power, CVNA-227 W 3-66 f2

                                      .    .     .                                     .
     ._

The hot assembly flow under nominal conditions without a flow mal- ' distribution effect is 93 per cent of the aversge 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 combined effects of heat input and flow maldistribution. Summary Analysis of all' B&W bundle data to date indicates tnat the B&W method vill correlate data with less deviation than previous methods. In-dications are that this is also true when considering nonuniform axial power distributions. Additional bundle testa vill be conducted with nonuniform axial power distribution to confinn that the use of a power shay correction factor based on single channel and annular specimens is conservative. Completion of the test programs outlined in this report and evaluation of the experimental data vill provide final design correlations and flow relationships that will give complete confidence in the conser-vatism of the design and the B&W analytical procedures. It should be noted that the postulated verst 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 perfonnance of the various channels. This I indicates that the outside cell geometries have been compensated cor-l rectly to account for vall effects. ! i ! I t . 1 I

l l

                                                                   * .
                                      .

3-67

___. .- -- 3.2.h MECHfdlICAL DESIGN LAYCGT 3.2.4.1 Internal Lavout Reactor internal co=ponents include the upper plenu= assembly, the core support O assembly (consisting of the core support shield, vent valves, core barrel, lover grid and flow baffle, thermal shield, and surveillance specimen holder tubes), l1 and the incore instrument guide extensions. Figure 3-59 shows the reactor ves-sel, reactor vessel internals arrangement, and the reactor coolant flow path. Figure 3-60 shows a cross section through the reactor vessel, and Figure 3-61 shows the core flooding arrangement. Reactor internal co=ponents do not include fuel assemblies, control rod assem-blies (CRA's), surveillance specimen asse=blies, or incore instrumentation. Fuel assemblies are described in 3.2.L.2, control rod assemblies and drives in 3.2.4.3, surveillance specimen assemblies in h.k.3, and incore instrumentation in 7.3.3. The reactor internals are designed to support the core, maintain fuel asse=bly align =ent, limit fuel assembly movement, and maintain CRA guide tube align =ent between fuel assemblies and control rod drives. They also direct tie flow of reactor coolant, provide ga==a and neutron shielding, provide guides for incore instrumentation between the reactor vessel lover head and the fuel asse=blies, support the surveillance specimen assemblies in the annulus between the thermal shield and the reactor vessel vall, and support the internals vent valves. 1 These vent valves are provided to relieve pressure generated by steaming in the core following an inlet pipe rupture so that the core vill re=ain sufficiently covered with coolant. All reactor internal components can be removed frc= the reactor vessel to allow inspection of the reactor internals and the reactor ves-sel internal surface. A shop fitup and checkout of all internal components in an as-built reactor ves-sel mockup will insure proper alignment of =ating parts before shipment. Du==v fuel assemblies and control rod. assemblies vill be used to check fuel asse=bly clearances and CRA free movement. In anticipation of lateral deflection of the lower end of the core support as-se=bly as a result of hori:cntil seismic loadings, integral veld-attached, de-flection-limiting spacer blocks have been placed on the reactor vessel inside vall. In addition, these blocks limit the rotation of the lover end of the core support assembly whien could conceivably result from flow-induced torsional loadings. The blocks allow free vertical movement of the lower end of the in-ternals for thermal expansion throughout all ranges of reactor operating con-ditions, but in the unlikely event of a flange, circumferential veld, or bolted joint failure the blocks vill limit the possible core drop to 1/2 in or less. The final elevation plane of these blscks vill 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 stcp. blocks for a 1.'k in. core drop would be approxi-mately 5 g's total. Llock location and geometry will be evaluated and deter-mined to transfer this loading through the. vessel support skirt to the reactor building concrete. A significant reduction in impact loading can be achieved through proper stop block design and detailed analysis. A 1/2 in, core drop will not allow the lower end of the CRA poison rods O

  '

3-68 (Revised 1-15-68) o.ononp74

              .-
                                       -          -          ,                            .            .           .

_ _ -. -_ .. 1 l l j

        -

to disengage from their respective fuel assembly guide tubes if the CRA's

 ,         y          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.

vill not result in a significant reactivity change. The core cannot re- , tate and bind the. drive lines because rotation of the core support assen- , bly is prevented by the stop blocks. The failure of the core support shield and core barrel upper flanges, or related flanges and other circumferential joints, is not considered cred-ible on the basis of the conservative design criteria and large safety factors employed in the internals design. The final internals design vill be espable of withstanding various combinations of forces and load-ings resulting from the staf,1c veight of internals (179,000 lb total), core with control rod drive line (303,000 lb total), dynamic load from trip (10 g's gives 207,000 lb), seismic (0.10 g vertical gives 48,000 lb), coolant flow hydraulic loading (230,000 lb), and other related loadings. The algebraic sum of this simplified loading case is 507,000 lb. This I results in a tensile stress of about 700 psi in the core support shield '

shell, which is approximately 4 per cent of the material yield strength.

Final internals component weights, seismic analysis, dynamic loadings from flow-induced vibration, detailed stress analysis with consideration for thennal stress during all transients, and resolution of fabrication details such as shell rolling tolerances and weld joint preparation de-tails will increase the stress levels listed above. As a final design criterion, the core support components will meet the stress requirements of the ASME Code, Section III, during normal operation and transients.

      .              The structural integrity of all core support circumferential weld joints

( in the internals shalls vill be insured by compliance with the radio-graphic inspection requirements in the code above. The seismic analysis will include detailed calculations to' determine the maximum structural response of the reactor vessel and internals. T'-te analysis will be per-i formed as described in 3 1.2.h l. I In the event of a major loss-of-coolant accident, such as a 36 in. diam-eter reactor coolant pipe break near the reactor vessel outlet, the fuel assembly and vessel internals vould be subjected to dynamic loadings re-sulting from an oscillating (approximately sinusoidal) differential pres-sure across the core. A preliminary analysis of this postulated accident indicates that the fuel assemblies vould move upward less than 3/8 in. Some deflection of the internals structures would occur, but internals component failure vill not occur. The occurrence of a loss-of-coolant accident and resulting loadings will be evaluated during the detailed de-sign period for the fuel assemblies and related internals structural com-ponents. The deflections and movements described above vould not prevent CRA in-sertion because the control rods are guided by split tubes throughout their travel, and the guide tube to fuel assembly alignment cannot change regaviless of related component deflections. CRA trip could conceivably be delayed momentarily as a result of the oscillating pressure differen-tial. However, the CRA travel time to full insertion vould remain rela-tively unaffected as transient pressure oscillations are dampened out in approximately 0 5 sec. On this basis, the CRA travel time to 2/3 inser- 1 tion on a trip command will be approximately 1.55 see instead of the spec- ! itied 1.40 sec. Also, 3-69 (Revised 1-15-68) 00000215

   - . . . - - . - -    .   - - - , - - - . - - -          . - - - , - _ . _ , . - _ - , - . - . - , , - - , , , ,

this possible initial minor delay in trip initiation vould not contribute to the severity of the loss-of-coolant accident because at the initiation of CRA g trip, the core would be suberitical from voids. W Material for the reactor internals bolting vill be subjected to rigid quality control requirements to insure structural integrity. The bolts vill be dye-penetrant inspected for surface flav 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 vill be lock-velded to insure assembly integrity. 3 2.4.1.1 Upper Plenum Assembly The upper plenum a?nenbly is located directly above the reactor core and is re-moved as a single component before refueling. It consists of upper and center grid assemblies, CRA guide tubes, and a flanged cylinder with openings for re-actor coolant outlet flow. The upper grid is a series of parallel flat bars intersecting to form square lattices and is velded to the plenum cylinder top flange. A machined upper end on each CRA guide tube is located and welded to the plenum cover which is attached to the upper grid bars. CRA guide tubes provide CRA guidance and protect the CRA from the effects of coolant cross-flow. Each CRA guide tube consists of an outer tube housing and sixteen slotted tubes which are properly oriented and brazed to a series of castings. As the tubes tre slotted for their full length, the brazement provides continuous guidance for the CRA full stroke travel. Design clearances in the guide tube vill ac- s commodate some degree of misalignment between the CRA guide tubes and the fuel assemblies. Final design clearances vill be established by tolerance studies and by the results of the Control Rod Drive Line Facility (CRDL) prototype tests. Preliminary test results are described in 3 2.4 3 5 The center 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 lover flange. The center grid assembly locates the lower end of the individual CRA guide tube relative to the upper end of the correspond-ing fuel assembly. Locating slots in the upper plenum assembly top flange engage the reactor ves-cel top flange locating devices to align the upper plenum assembly with the re-cetor vessel, reactor closure head control rod drive penetrations, and the core support shield. The bottom of the upper plenum assembly is guided and aligned by locating blocks attached to the inside 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, lover grid and flow baffle, thermal shield, and surveillance specimen holder tubes. Static loads from the assembled components and fuel assemblies, and dynamic loads from CRA trip, hydraulic flov, thermal expansion, seismic disturbances, and loss-of-coolant accident considerations, are all carried by the core sup- , port assembly. g 34 00000216 _

n (") The core support asse=bly co=ponents are described as follows:

a. Core Support Shield The core support shield is a large flarged cylinder which =ates with the reactor vessel opening. The top flange rests on a circu=ferential ledge in the reactor vessel top closure flange.

The core support shield lower flange is bolted to the core bar-rel. The cylinder vall has two nozzle openings for reactor coolant outlet flow. Locating blocks on the inside of the cyl-inder vall near the botto= Suide and align the upper plenu= chamber relative to the core support shield. The reactor vessel outlet nozzles are sealed to the =atire com-ponents of the core support shield by the differential ther=al expansion between the carbon steel reactor vessel and the stain-less steel core support shield. The nozzle seal surfaces are finished and fitted to a predetemined cold gap providire clearance during core support asse=bly installation and re= oval. At reactor operating te=perature the mating =etal surfaces are in contact to =ake a seal without exceeding allowable stresses in either the reactor vessel or internals.

b. Core Barrel The core barrel supports the fuel asse=blies and lover grid and flow baffle, and directs the reactor coolant flow through the vessel. The 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 horizon-t< spacers to fo= an inner vall enclosing the fuel asse=blies.

Construction of the core barrel vill be si=ilar to that of the l reactor internals co=ponent developed by B&W for the Indian l Point Station Unit No. 1. Coolant flow is downward along the outside of the core barrel cylinder and upward through the fuel asse=blies contained in the core barrel. A s=all portion of the coolant flows upward throu6h the space between the core barrel outer cylinder and the inner plate vall. ! Coolant pressure in this space is maintained slightly lover than l the core coolant pressure to avoid tension loads on the bolts l ' attachirq the plates to the horizontal spacers. The vertical plate inner vall vill be carefully fitted together to reduce reactor coolant leakage to an acceptable rate. The upper flange of the core barrel outer cylinder is bolted to the =ating lever flange of the core support shield, and the lower flange is bolted to the =ating flange of the lover grid , p) - and flow baffle. All bolts will be inspected and installed as ' V i

                   ~                           "

00000?I7

      ._.
                                                  .

d: scribed in 3 2 3.1, and vill be lock-velded aft:r final assembly.

c. Iover Grid and Flow Baffle he lover grid provides alignment and support for the fuel assemblies and aligns the incore instrument guide extensions with the fuel as-sembly incore instrument tubes. We lower grid consists of two flat plate and bar lattice structures separated by short tubular columns surrounded by a flanged cylinder. The top flange is bolted to the lower flange of the core barrel. Se lover grid top flange also positions and supports the themal shield.

Se flow baffle is a dished plate with an external flange which is bolted to the bottom flange of the lower grid. The flow baffle is perforated to distribute the reactor coolant entering the bottom of. the core.

d. Thermal Shield A cylindrical stainless steel thermal shield is installed in the annulus between the core barrel outer cylinder and the reactor ves-sel inner vall. The thermal shield reduces the neutron and gaiana internal heat generation in the reactor vessel vall and thereby re-duces the resulting thermal stresses.

The themal shield is supported on, positioned by, and attached to the lover grid top flange. Also, the thermal shield upper end is positioned by spacers between the thermal shield and the core barrel outer cylinder to minimize the possibility of thermal shield vibra- g tion. The thermal shield attachment is designed to prevent fasteners W from being loaded in shear. Fasteners are lock-velded after final assembly.

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

, The tubes extend from the top flange of the core support shield to l the lower end of the thermal shield. The tubes will be rigidly at-tached to prevent flow-induced vibration. Slip joints at the lover end of the core support shield will allow the shield to be removed from the core support assembly without destructively removing the surveillance specimen holder tubes.

f. Intern'als Vent Valves 1 Internals vent valves are installed in the core support shield to t

prevent a pressure unbalance which might interfere with core cooling l following a loss-of-coolant accident. In its natural state and under all normal operating conditions, the vent valve vill be. closed. In

the event of a loss-of-coolant accident in the cold leg er the reactor 3-72 (Revised 1-15-68)
              .

on m 8

                                                                                  -

m -w- ,.,-._w--,- -

                                                                           ,-, w.

_

            .

(~j loop, the valve vill open to permit steam generated in the core to 1 y_/

   -

flow directly to the leak and will prevent the core from bacoming more than 1/2-uncovered after emergency core coolant has been sup-plied to the reactor vessel. The preliminary design of the laternals vent valve is shown in Figure 3-61a. Each valve assembly consists 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 velded in the core support shield vall. The mounting ring contains the neces-sary features to retain and seal the perimeter of the valve _ assembly. Also, the mounting ring includes an alignment device to maintain the correct orientaticn of the valve assembly for hinged-disc operation. Each valve assembly vill be remotely handled as a unit for removsl 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 vill include an integral arm hook, eye, or other device for remote inspection of disc function. The preliminary arrangement consists of lk-in. diam check valve assem-blies installed in the cylindrical vall of the internals core support shield. The valve centers are coplanar and are h2 in, 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 i vall. h',",N)

                                                  .

The hinge design vill consist of a shaft, two valve body journal re-

                                                                     .
  '

ceptacles, two valve disc journal receptacles , and four flanged shaft journals (bushings). Loose clearances vill be used between the shaft and journal inside diameters, and between the journal outside diameters and the,ir 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 unha=pered disc-free motion. In the worst case, at least four . clearances must bind or seize solid to adversely affect valve disc-l free motion. l * ! In additien, the valve disc vill contain a self-alignment feature so that the external differential pressure vill adjust the disc seal face to the valve body seal face. This feature minimizes the possi-l

bility of increased leakage and pressure-induced deflection loadings l on the hinge parts in service.

! l The external side of the disc vill be contoured to absorb the impact I l load of the edqsc on the reactor vessel inside vall without transmit-

                                              .

ting exe,es;si've impact loads to the hinge parts as a result of a loss '

                             .

of-coolant accident. L 0.0000219 L-3-72a (Revised 1-15-68)

          . ...            _
                                .    . _- ,      .-.      -  -.     .. .- - ..                .-

3 2.4.1 3 Incore Instrument Guide Extensions The incore instrument guide extensions guide the incore instrument assemblies between the instrument penetrations in the reactor vessel bottom head and the h instrument tubes in the fuel assemblies. Sufficient clearance in the instru-ment guide extensions provides for minor misalignment between the reactor ves-sel inistrument penetrations and the instrument guide extension tubes. A per-forated shroud tube, concentric with the instrument guide tube, adds rigidity

 .

O . OnMc. ao l l l 1 3-72b (Revised 1-15-68) ! l _ l - . - - - --. . _ . . _ _ . _ . _ _ ._ ___._ _ _

( ( / to the assembly and reduces the effect of coolant flow forces. Fifty-two in-core instrument guide extensions are provided. The incore instru=ent guide extensions are designed so that they will not be affected by the core drop described in 3 2.4.1. 3 2.4.2 Fuel Assemblies 3 2.4.2.1 Description

a. General Description The fuel for the reactor is sintered pellets of lov 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 bun-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 177 fuel assemblies. All assem-blies are identical in mechanical construction, i.e., all are de-signed to accept the control rod assemblies (CRA). However, only 69 have CRA's to control the reactivity of the core under operating con-ditions. In the 108 fuel assemblies containing no CRA during a given core cycle, the guide tubes are partially filled at the top by an
                " Orifice Rod Assembly" (Figure 3-63) in onier to minimize bypass cool-(q g             ant flow. These orifice rod assemblies also tend to equalize coolant flow between fuel asse=blies with CRA's and those with orifice rod assemblies.

Fuel assembly components, materials, and dimensions are listed below. Item Material Dimensions, in. l Fuel UO2 Sintered 0 362 diam. l Pellets Fuel Clad Zircaloy-4 0.420 OD x 0 368 ID x 152-7/8 l long l Fuel Rod Pitch 0 558 l Fuel Asse=bly Pitch 8 587 l Active Fuel Length 1.44 Overall Length =165 Control Rod Guide Zircaloy-4 0 530 OD x 0.015 vall p Tube

 ;b
                        ,                                              00000221 3-73 f

1

                                                                                    !

Item Materini Dimensions, in. Incore Instrument Guide Extension Zircaloy-4 0 530 OD x 0.075 van hl ' Spacer Grid Stainless Steel, Spacedat21-7/16in. Tp-304 Can Panel Stainless Steel, 0.031 thick TP-304 End Fitting Stainless Steel, Tp-304

b. Fuel The fuel is in the form of sintered and ground penets of ura-nium dioxide. The pellets are dished on each end face to min-imize the difference in axial themal expansion between the fuel and cladding. The density of the fuel is 95 Per cent of theoretical.

ofthefuelis28,200 MWD /MTU. Peak Averagedesignburnup/MTU. burnup is 55,000 MWD At the peak burnup, the fuel growth iscalculatedtobe9-1/2volumepercentbythemethodgiven in Reference 53 This growth is accommodated by pellet poros-ity, by the radial clearance providea cetween the pellets and the cladding, and by a small amount of plastic strain in the h cladding. Each fuel column is located, at the bottom, by a thin-vall stainless steel pedestal and is held in place during handling by a spring at the top. The spring allows axial differential themal expansion between fuel and cladding, and axial fuel growth. The bottom pedestal is also co napsible, thus provid-ing a secondary buffer to prevent excess cladding axial strain. l Fissien 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 i fuel rod. ,

c. Fuel Assembly Structure 1

(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 velded together at l the corners for the entire length. The spacer grids are velded to the panels, and the lover and upper end fittings

                                                                                  ~

! are velded to the panels to complete the structure. The upper end fitting is not attached until the fuel rods, 3-w 00000222

                                                          .     -
                                            -  w
                                                                                       .

l O suide tude , =

                                                                             ~

1= tr# e=t *1o= *=de h ve dee= 1= t 11 a. At each spacer grid assembly 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 suf-ficient to restrain it so that flov-induced vibrational amplitudes are minimal. However, to avoid undesirable bowing of the fuel rods, the spring loads 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 made of square tubing.

The ferrule has a portion of each side formed into spring sections which have hydrodynamically shaped " dimples" that l contact the fuel rods. The ferrules are joined together i ' by brazing to Iorm the spacer grids. The grids, which provide the desired pitch spacing between fuel rods, are spot-velded at intervals to the per' ~ated stainless steel can panels. ! l (3) Lover End Fitting The lover end fitting is constructed from Type 304 stain-p les: steel members which when joined together form a box k' structure. Four deep cross members serve as the position-l ing curfaces for the fuel assembly when it is inserted in-to the lover core support structure. The assembly includes a grid structure which provides a support base for fuel rods while maintaining a maximum inlet flow area for the coolant. (4) Upper End htting The upper end fitting is similar to the lover end fitting. It positions the upper end of the fuel assembly and pro-l vides coupling between the fuel assembly and the hmM14ng equipment. A hollow post, velded in the center of the as-sembly, is designed to provide a means of uncoupling the CRA-to-drive connection and to retain the orifice rod as-sembly. In order to identify a fuel assembly under water, a serial number is milled into a flat, chrome-plated sur-face which is velded to the box frame. (5) Control Rod Guide Tubes The Zirealoy guide tubes serve to guide the control rods within the fuel assembly during operation. The tubes are restrained axially by the upper and lover end fittings in l , the fuel assembly and radially by the spacer grids in the same manner as the fuel rods.

 . -
                                                                                '

Onn00723 3-75

       . . -
                               - - . . - . - - - - . . .

3 2.4.2.2 Evaluation

a. Fuel Rod Assembly g

(1) General The basis for the design of the fuel rod is discussed in 3 1.E 4. Materials testing and actual operation in reactor service with Zircaloy cladding has demonstrated that Zircaloy-4 material has smple corrosion resistance and sufficient mechanical prcperties to maintain the integrity an,i 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 severest. 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 vill result during operation. The actual clad stresses are considerably below the yield strength. $ Circu=ferential stresses due to external pressure, calculated using those combinations of clad dimensions, ovality, and eccen- h 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

is 32,000 psi at design pressure, 27,000 psi at operating pres-1 sure. At this stress, the material may creep sufficiently to allow an increase in ovality until further creep is restrained by support from the fuel. Contactloadsontheorderof20lb/

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 I 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 1/4 of the yield strength and therefore is not a potential source l of short time burst. The possibility of stress-rupture burst ! has been investigated using finite-difference methods to esti-l mate the long time effects of the increasing pressure on the clad. The predicted pressure-time relationship produces stresses i l that are less than 1/3 of the stress levels that would produce h 3-76 Nb h')24 _

                                                             +-h -
                                                                               ~ - ,, ,

1 N

          '

stress rupture at the end.of-life. Outpile stress-rupture data were used, but the greater than 3:1 margin on stress is :nore 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, (e) initial water vapor and atmospheric gases adsorbed on the fuel, and (3) fission product gases. 'Ihe water vapor present initially is expected to disso-
ciate over the life of the fuel and enter into hydriding and l

' 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 fluxeg 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. O O 00000225 _...

                                         .
               #

3-Tl

 .,.__ _ _ _     .-  . _ _._,. _ . -        _ . . , . .
                                                                                                  ._

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

33,000 46,000

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

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

  -3    EOL - Sh_utdown 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 vill be approximately 75,000 psi yield strength (0.2 per cent offset) and 85,000 psi ultimate tensile strength. ( ) Cladding stresses due to fuel swelling are discussed further on another page of 3 2.4.2.2.

.

g 3-78

                                                   .__

-w - o- -- - *rr----+

  • w -*

I

                                           ,

Table 3-19 (Cont'd) Ultimate Calc. Yield Tensile Stress, Stress, Stress, operating Condition psi psi psi 3 Hours Later (50F/hrPressurizerCooldownRate) Fuel Rod Internal Pressure - 1,05o psig System Pressure - 680 psig Average Clad Temperature - Approxi-mately 425 F 3,300 52,000 55,000 The total production of fission gas in the hottest fuel rod assembly is based on the hot rod average bumup of 38,000 WD/MrU. The corresponding maximum burnup at the hot fuel rod midpoint is 55,000 WD/MrU. O The fission gas. release is based on temperature versus re-lease fraction experimental data.(49) Fuel temperatures are calculated for small radial and axial increments. The total fission gas release is calculated by integrating the incremental releases. The maximum release and gas pressure buildups are deter-mined by evaluating the following factors for the most con-servative conditions: (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 re-lease 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.

             ' ~

3 79 00000227

The fuel temperatures used to determine fission gas release N and internal gas pressure have been calculated at the re-actor overpower condition. Fuel temperatures, total free gas volume, fission gas release, and internal gas pressure have been evaluated for a range of initial diametral clear-ances. This evaluation shows that the highest internal pressure results when the maximum diametral gap in assumed because of the resulting high average fuel temperature. The release rate increases rapidly with an increase in fuel temperature, and unrestrained axial growth reduces the rel-atively cold gas end plenum volumes. A conservative ideal thermal expansion model is used to calculate fuel tempera-tures 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 maxi-mum fission gas release rate is 43 per cent. (3) Collapse Margins Short time collapse tests have demonstrated a clad collaps-ing pressure in excess of 4,000 psi at expansion void maxi-zum temperature. Collapse pressure margia is approximately 17 Extrapolation to hot spot average clad temperature ( = 725 F) indicates a collapse pressure of 3,500 psi and a margin of 1.4, which also greatly exceeds requirement. s Outpile creep collapse tests have demonstrated that the clad meets the long time (creep collapse) requirement. (4) Fuel Svelling Fuel rod a"erage and hot spot operating conditions and de-sign parameters at 100 per cent power, pertinent to fuel swelling considerations, are listed below. Average Maximum HeatFlux, Btu /ft-hr 2 167,620 543,000 LinearHeatRate,kv/ft 5.4 17 5 l Fuel Temperature, F 1,385 4,160 Burnup (IGD /MI'U) at Equilibrium 28,200 55,000 Nominal Values Pellet OD, in. 0 362 l Pellet Density, % of Theoretical 95 Pellet-Clad Diametral Gap at Assy., in. 0.004 - 0.008 Clad Material Cold-Worked Zr-4 Clad Thickness, in. 0.026 9 3-80 00000228 _ w- + y-w.- -- ,- p -a-1s ay- ,-----e.y - - , -- - 7 - ,e

                . - . .                                                                                                                 -.

1 l The capability of Zircaloy-clad 002 fuel in solid rod form to perfom satisfactorily in PWR service has been amply . demonstrated through operation of the CVTR and Shippingport cores, and through results of their supplementary develop-ment programs, up to approximately 40,000 !WD/)EU.

 >                                             As outlined below, existing experimental information sup-ports the various individual design parameters and oper-ating conditions up to and perhaps beyond the maximum burnup of 55,000 MWL/IEU, but not in a single experiment.

However, the LRD irradiation test program, currently in . progress, does combine the items of concern in a single experiment, and the results are expected to be available to contribute to final design' confirmation. (5) Application of Experimental Dtta to Design Adequacy of the Clad-Fuel Initial Gap to Accommodate Clad-Fuel Differential Thermal Expansion Experimental Work Six rabbit capsules, each containing three Zr-2 clad rods of 5 in. fuel len Test Reactor (45)atpowerlevelsupto24kv/ft. gth, vere irradiated in the Westinghouse The 94 per cent theoretical density (T.D.) UO2 pellets (0.430 OD) had initial clad-fuel'diametral gaps of 6, 12, and 25 mils. t'O No dimensional changes were observed. Central melting oc-curred at 24 kv/ft only in the rods that had the 25 mil initial gap. Two additional capsules were tested.(53) The specimens were similar to those described above except for length and initial gap. Initial gaps of 2, 6, and 12 mils were used in each capsule. In the A-2 capsule, three 38-in.- longrodswereirradiatedto3,450 MWD /)CUat19kv/ft maximum. In the A-4 capsule, four 6-in.-long rods were irradiated to 6,250 MWD /}EU at 22.2 kv/ft maximum _. No central melting occurred in any rod, but diameter in-creases up to 3 mils in the A-2 capsule and up to 1 5 mils in the A-4 capsule were found in the rods with the 2 mil initial' gap. Application In addition to demonstrating the adequacy of Zircaloy-clad UO2 pellet rods to operate successfully at the power' levels of interest (and without central melting), these experiments demonstrate that the design initial clad fuel. gap of 4 to 8 mils is adequate to prevent unacceptable clad diameter in-crease due to differential themal expansion between the clad and the fuel. A maximum local diametral increase of less than 0.001 in. is indicated for fuel rods having the O. minimum initial gap, operating at the maximum overpower

  • condition. \
                                        '

00000229 3-81 ! ! . - . . _ , , , _ ._ ~..- _ , . . . . . _ _ _ . . . . . - . . _ _ _ , . . _ _ , . . . _ . _ _ , _ , . . . . . . . . . _ _ _ _ . , . _ , . . . _ , _ . _ ,

(6) Adequacy of the Available Voids to Accommodate Differential - ' Expansion of Clad and Fuel, Including the Effects of Fuel Svelling Experimental Work Zircaloy-clad, UO2 pellet-type rods have performed success-fully in the Shippingport reactor up to approximately 40,000 wD/Mru. Bettis Atomic Power Laboratory (53) has irradiated plate-type UO2 fuel (96-98 per cent T.D.) up to 127,000 MWD /MrU and at fuel center 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/ccafterinternalvoidsarefilled. This point of

       " breakaway" appears to be independent of temperature over the range studied and dependent on clad restraint and the void volume available for collection of fission products.

The additional clad restraint and greater fuel plasticity (from higher fuel temperatures) of rod-type elements tend to reduce these swelling effects by providing greater re-sistance to radial swelling and lover resistance to longitu-dinal swelling than was present in the plate-type test speci-mens. This is confirmed in part by the work of Frost, Bradbury, g and Griffiths of Harwell(56) in which 1/4 in. diameter UO2 W pellets clad in 0.020 in. stainless steel with a 2 mil diametral gap were irradiated to 53,300 WD/Mru at a fuel center temperature of 3,180 F vithout significant dimensional change. In other testing (57) 0. '.50 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 . ail diametral gaps have been irra-diated to 77,000 MWD /MrU at fuel temperatures high enough to approach central melting without apparent detrimental , results. Comparable results were obtained on rods svaged ( to 75 per cent T.D. and irradiated to 100,000 WD/MrU. Application l Based on the BAPL experimental data, swelling of the fuel I rods is estimated as outlined below. Fuel is assumed to swell uniformly in all directions. Clad- , pellet differential themal 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 themal expansion. If the initial gap exceeds the mini-mum, the additional gap volume is assumed available to j accommodate swelling. This additional void volume may g initially tend to be filled by pellet themal expansion W 33, 00000230 _

i because of the low contact pressure and resultant low con- , tact coefficient, but as the fuel swells, the contact pres- )

 '

sure must increase if the clad is to be stretched. Where I

            -fuel cracking tends to fill the radial gap, it is assumed                         l 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/ccuntilthe5percentinitialvoidin the 95 per cent T.D. pellets is filled at about 9 x 1020 f/cc. From that time on, swelling is assumed to take place at 0 7% AV/1020 f/cc until the maximum burnup of 13 6 x 1020 r/ce (55,000 WD/Mru) is reached. Total fuel volume increaseis4-1/2 percent,whichresultsina1-1/2per cent diameter increase in a rod with the 0.004 in. mini-mum initial gap. Clad stress is estimated at 22,000 psi, so that the elastic strain is about 0.2 per cent. Net plastic strain is 1 3 per cent. similar calculations indi-cate that fuel rods with maximum burnup and the nominal clad-fuel gap (0.006 in, at assembly) will 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 ISBR(30) suggests that' swell-ing rates for this design may exceed those indicated by the BAPL data because of the higher fuel temperatures. However, O ts the A.E.R.E. tests (56) and the General Electric tests (57) do not support more than a small increase in post " break-away" swelling rates at temperatures of interest. Fuel Swelling Studies - LRD Irradiation Program (59) . Dimensional stability of UO2 under inpile conditions simu-l lating large reactor environments is under investigation. This study is currently being carried out under USAEC Con-tract AT(30-1)-3269, "Large Closed-Cycle Water Reactor Research and Development Program". Parameters contributing to swelling are burnup, heat rating, fuel density and grain size, and clad restraint. These are systematically being studied by irradiating a series of capsules containing fuel rods. These experiments were as-signedbytheAECtoETR/MIR. Test variables are shown in Table 3-20. O '

          -
         ..

P 00000231 l l- 3-83 l '

      -.        .   . _ . . - .   . . .  . . . - . . . . - _ _ _ . .       . . . . . _ . _ . .
                                                 .-

O i l Table 3-20 l LRD Fuel Swelling Irradiation Program Initial coal Heat Rating (b) capsule (a) Enrichment, kv/ vatts/ Fuel Density, Burnup, WAPD-49  % ft em  % T.D. MWD /MrU A 18 7 12 394 94 and 97 38,000 B 18 7 12 394 94 and 97 38,000 c 18 7 12 394 90, 94, and 97 38,000 D 16.0 18 591 90 and 97 47,000 g E 13 5 18 591 94 and 97 47,000 F 13 5 18 591 90, 94, and 97 47,000 o 16.0 18 591 90 and 97 47,000 H 17.o 24 788 94 and 97 56,000 l I 18 7 24 788 94 and 97 56,000 l ' J 20.0 24 788 94 and 97 56,000 K 20.0 24 788 90 and 94 56,000 L 20.0 24 788 94 and 97 56,000 l ("Fourrods/ capsule. ( } Fuel center temperatures vary from 1,570 to 4,110 F. 00000232 Y , 3-84 , o m 1 - . . .-. .- . . .. . - - . .-_ , .-

1

                                                                                                            )

r' .

 . (_)/                       Effect of Zircaloy Creep The effect of Zircaloy creep on the amount of fuel rod growth due to fuel swelling has been investigated. Clad creep has the effect of producing a nearly constant total pressure on the clad ID by permitting the clad diameter to increas9 ag the fuel di-ameter increases. Based on out-of-pile data, tN1 1 per cent creep will 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 = 720 F average temperature through the cicd at the hot spot. At the start of this high swelling period (roughlythelast1/3ofthecorelife),thereactorcoolant system pressure would more or less be balanced by the rod in-ternal pressure, so the total pressure to produce the clad stress of 22,000 psi vould have to come from the fuel.      Contact pres-sure would be 2,400 psi. At the end-of-life, the rod internal pressure exceeds the system pressure by about 1,100 psi, so the clad-fuel contact pressure would drop to 1,300 psi. Assuming that irradiation produces a 3:1 increase in creep rates, the clad stress for 1 per cent strain in 10,000 hr would drop to about 15,000 psi. Contact pressures would be 1,800 psi at the beginning of the high.svelling 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 b--}                        to irradiation) when the pellet is slightly hotter than calcu-lated. The effect of this vould be a slight increase in pellet thermal expansion and therefore in clad strain. Considering the improbability that irradiation vill 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. dia=etral clearance between the control rod guide tube and the control rod is provided to cool the control rod and to insure adequate freedom to insert the control rod. As indi-cated below, studies have shown that fuel rods vill not bow suf-ficiently to touch the guide tube. Thus, the guide tube vill not undergo deformation caused by fuel rod bowing effects. Initial lack of straightness of fuel rod and guide tube, plus other ad-verse tolerance conditions, conceivably could reduce the 0.083 in. nominal gap between fuel rod and guide tube to a minirum of about 0.045 in., including amplification of bowing due to axial friction loads from the spacer grids. The maximum expected flux gradient of 1.176 across a fuel rod vill produce a temperature difference of 12 F, which will result in a thermal bow of less than 0.002 *in. Under these conditions, O 00000233
                                                                                                            .

3-85

                                                                                                            !

_ _ . , _ _ . . m-._. . - _ - . . . , - - _ _ _ -- _

for the fuel rod to touch the guide tube, the thermal gradient across the fuel rod diameter would have to be on the order of 300 F. The effect of a DNB occurring on the side of a fuel rod adja-O cent to a guide tube vould result in a large temperature differ-ence. In this case, however, investigation has shown that the clad temperature would be so nigh that insufficient strength would be available to generate a force of sufficient magnitude to cause a significant deflection of the guide tube. In addi-tion, the guide tube would experience an opposing gradient that would resist fuel rod bowing, and its internal cooling would raintain temperatures much lower than those in the fuel rod cladding, thus retaining the guide tube strength. (2) Vibration The semiempirical expression developed by Burgreen was used to calculate the flow-induced vibratory anplitudes 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 a=plitude for the fuel assembly, a direct measure-ment will be obtained for a full-size prototype fuel assembly during testing of the assenbry in the Control Rod Drive Line Facility (CRDL) at the B&U Research Center, Alliance, Ohio. (3) Demonstration lh 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 experinental program in the CRDL. 3.2.h.3 Control Rod Drive System 3.2.h.3.1 Description 7 The control rod drive system includes drive mechanisms which actuate control rod assemblies and axial power shaping rod assemblies , drive controls, p;ver supplies , position indication, operating panels and indicators , safety devices, enclosures, housings, and mountings. Criteria applicable to drive mechanisms for both control rod assemblies and axial power shaping rod assemblies are given in 3.2.h.3.1.1. Additional requirements for the mechanisms which actuate only control rod assemblies are given in 3.2.h.3.1.2. 3 2.h.3.1.1 General Design Criteria

a. Sincle Failure 00 flo' nv 2 3 A'
                                                       '

No single failure shall inhibit the trotective action of the con-trol rod drive system. The effect of a single failure shall he

         . limited to one control rod drive.

O 3-86 (Revised 7-15-69) __.

l

    -
  ;i y i      f O                b. l Uncontrolled Withdrawal
                         -Ho single; failure or chain of failures shall cause uncontrolled         7
                       ~ withdrawal .of any control ' rod assembly (CRA).
                   .c. Ecuinment Removal-The disconnection of plug-in. connectors , modules , and subassemblies -
                         - from the ' protective circuits shall be annunciated or shall cause a reactor trip.
d. Position Indication Continuous position indication, as well as an upper and lower posi-tion limit indication, shall be provided for esch control rod drive.

The accuracy. of the position indicators shall be consistent with the tolerance set by reactor safety analysis,

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

These include rod position deviation and power supply voltage,

f. Drive Sneed N' The control rod drive control system shall provide for. single uni-form speed of the mechanism. The drive controls , or. mechanism and motor combination, shall have an inherent speed limiting. fea-ture. The speed of.the mechanism shall be 30 in./ min for both insertion and withdrawal. The withdraval speed shall be limited so as not to exceed 25 per cent overspeed ir. the event of speed control fault.
g. Mechanical Stons i

Each control rod drive shall have mositive mechanical stops at' both ends of the stroke or travel. The stons shall be capable of receiving the full operating force of the mechanisms without fail-ure. 3.2.4.3 1.2 . Additional Design Criteria The following criteria are applicable only to the mechanisms which actuate control rod assemblies.

a. CRA Positionine The control rod drives shall provide for controlled withdrawal or insertion of the control rod assemblies (CRA) out of, or into, the reactor core to establich and hold the cover level required.

j 'l

   '
        ;                The drivi s are' also capable of rapid insertion or trip for emergency
      -

reactor c:nditions. Y

                  $ !I:l,
          '

00000235

                                                      '3-87   (Revised 7-15-69)'
b. CRA Trin g
                                                                                           ,

7 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 delav, and shall not require external power. Circui interrupting devices shall not prevent reacter trip. Fuses, where used, shall be urovided with blown indicators. Circuit breaker position information shall also be indicated,

c. Groun Withdrawal The control rod drive system allows only two out of three regulat-ing CRA groups to withdraw at any time subject to the conditions  !

described in 7.2.2.1.2. 3 2.h.3.2 control Rod Drive Mechanisrm The control rod drive mechanisms provide for controlled withdrawal or ins-~ tion of the control red assemb12-s out of or into the core and are capable of rapid insertion or trip. The drive mechanisms are hermetically sealed, reluctance motor-driven screw units. The CRDM data is listed in Table 3-21. Table 3-21 Control Pod Drive Mechanism Desien Data Axial Power Mechanism Punction Shim Safety Shapine h Type Roller Nut Drive Poller Nut Drive Quantity 61 8 Loc ation Top-mounted Top-nounted Direction of Trip Down Does not trip Velocity of Normal Withdrawal 30 30 and Insertion, in./ min. Maximum Travel Time for Trip 2/3 Insertion, s 1.h0 Drive has no trip function L;ngth of Stroke, in. 139 139 D; sign Pressure, psig 2,500 2,500 D: sign Temperature, F 650 650 , Weight of Mechar. ism ( App. ) 9ho lb 9ho Ib _ O 3-88 (Revised 7-15-69) 000002*36 _ _

a

                                                                                 +

I 3 2.h.3.2.1 Shim Safety Drive vechanisn [v~) The drive mechanism consists of a motor tube which houses a lead screw and its c rotor assembly and a buffer. The end of the motor tube is closed by a cap and vent assembly. A motor stator is placed down over the motor tube pressure ves-sel, and position indication svitches are arranged outside. the motor tube ex-tension. The control rod drive output elenent is a translating screw shaft which is cou-pled to the contrcl rod. The screw is driven by an anti-friction nut element which is rotated magnetically by a motor stator located outside the pressure boundary. Current impressed on the stator causes the senarqble nut halves to engage; a mechanical spring causes ~them to disengage the -screw in the absence of a current. For rapid insertion, the nut separat3s to release the screw shaft which then falls into the core by gravity. A hydraulic buffer within the upper housing decelerates the falling assembly to a lov speed a short distance above its ftll-in position. The final deceleration- is accommodated by the down-stop buffer spring. This mechanism incorporates proven principles and material combinations and is based on extensive analytical, developmental, design, test, and manufacturing experience obtained over the years for the Shippingport and ae Naval Nuclear Program. The control rod drive is shown in Figures 3-6h and 3-65 Subassemblies of the control rod drive are described as follows: q a. Motor Tube

   '#              The motor tube is a three-piece velded' assembly designed and manu-factured in accordance with the requirements of the ASME Code, Sec-tion III, for Class A nuclear pressure vessel. Materials conform to ASTM or ASMS, Section II, Material Specifications. All velding shall be performed by personnel qualified under ASME Code, Section IX, Welding Qualifications. The motor tube vall between the rotor as-sembly and the stator is constructed of magnetic material to present a small air gap to the motor. This region of the motor tube is of lov alloy steel clad on the inside diameter with stainless steel or with Inconel. The upper end of the motor tube functions only as a pressurized enclosure for the withdrawn lead screw and is made of stainless steel transition-velded to the upper end of the low alloy steel notor section. The lover end of the lov allov steel tube sec-tion is velded to a stainless steel machined forring which is flanged l                   at the face which contacts the vessel control rod nozzle. Double

! gaskets , which are separated by a ported test annulus , seal the flanged connection between the motor tube and the reactor vessel.

b. Motor The motor is a synchronous relu'tance unit with a slip-on stator. The rotor assembly is described in F tragraph (f). The stator is a h8-slot four-pole arrangement with water 'ooling coils wound on the outside of its casing. The stator is encansulated after vinding to establish a hermetically sealed unit. It is six phase star-connected for operation in a pulse-stepping mode and advances 15 mechanical degrees per step.
                                                           ~

The stator assembly is mounted over the motor tube h6t[Mng as shown

.('O.

l_ in Figure 3-65 000.0.0237 3 89 (Revised 7-15-69) L

                                                                                .
                                                               .
c. Can and Vent Valve 17 s The upper end of the motor tube is closed by a cap containing a vapor bleed port and vent valve. The bleed port and vent balve and the cap-to-motor tube closures have double seals. The car is retained by a bolting ring threaded to the outside of the notor tube. The retaining bolts are made long so as to be elastic enough te provide positive seal preload at any assembly temperature fron 20 to 650 F. The j

minimum preload is equal to the 3750 psig proof pressure force. , I

d. Actuator l

The actuator consists of the translating lead screw, its rotating nut assembly, and the toraue taker assembly on the screw. The actuator lead screw travel is 130 inches.

e. Lead Screw The lead screw has a lead of 0.750 in. The thread is double lead with a single pitch spacing of 0.375 in. Thread lead error is held to 0.0005 in, maximum in any 6 in. for uniform loading with the roll-er nut assemblies. The thread form is a modified ASMZ vith a flant angle that allows the roller nut to disengage without lifting the screw.
f. Fotor Assembly The rotor assembly consists of a bal] bearing supported rotor tube carrying and limiting the travel of a pair of scissors arms. Each of the two arms carry a pair of ball bearing supported roller (nut) assemblies which are skewed at the lead screv helix angle for engare-ment with the lead c;t av. The current in the motor stator (two of a six vinding stator) causes the arms that are rivoted in the rotor tube to move radially toward the motor tube vall to the limit urovided thereby engaging the four raller nuts with the centrally located lead screv. Also, four separating springs mounted in the scissor arms keep the rollers disengaged when the power is renoved from the stator coils. A second radial bearing mounted to the urner end of the rotor tube has its outer race pinned to both scissor arms thereby synchro-nizing their motion during engagement and disencarement. 'Jhen a three phase rotating magnetic field is ap;11ed to the motor stator, the resulting force produces rotor assembly rotation.
g. Torcue Extension Tube and Tornue Taker .

The torque extension tube is a separate tubular assembly containing a keyvay that extends the full length of the lead screw travel. The tube assembly is supported against rotation and in elevation by the upper end of the motor tube extension. The lower end of the tube assembly supports the buffer and is the down stop. A set of index-ing serrations mate to prevent rotation and orient the toraue exten-sion tube with the motor tube below the can and vent valve assembly. An integral shoulder at the top of the tube rests against a step in the motor tube inside diameter to provide a vertiell support. 3-90 (Revised 7-15-69) 00000238 __

                                                                       -

_

 .g N,)
  • The toraue taker assembly consists _of the position indicator perma- 7 nent magnet, the buffer piston, and a positioning key. The torque

' taker key fixed at :the top of the lead screw is mated with' the torque extension t'.b'e keyvay to provide both radial and. tangential position-ing of'the lead screv.

h. Buffer The buffer assenbly is canable of decelerating the ' translating mass
              .fron the unpressurized terminal velocity to zero velocity without anplying greater than ten times- the gravitational force on the con-trol rod. -The vater buffer consists of a piston fixed to the top end of the screv shaft and a cylinder which is . fixed to the lover.

end of the toroue extension tube. Twelve inches above the bottom stop, the piston at the top of the screw enters the cylinder. Guid-ing is accomplished because the piston and torque key are in a single part, and the cylinder and keyway are in a single mating part. As the piston travels .into the cylinder, water is driven'into the center of the lead screw through holes in the upper section which produce the damping pressure drop. The number of holes presented to the buf-fer chamber is reduced as the rod moves into the core, so. that the danping coefficient increases as the velocity reduces, thereby pro-viding an approximately uniform deceleration. A large helical buffer spring is used to take the kinetic energy of the drive line at the end rs of the water buffer stroke. The buffer spring accepts a five-foot (_) per .-cond impact velocity of the drive line and control rod with an inscEataneous overtravel of one inch past the normal down stop. The inclusion of this buffer spring permits practical clearances in the water buffer.

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

In order to obtain trip travel times of acceptably small values, it is necessary to provide an auxiliary flow path around the guide bush-ing. The larger area path is necessary to reduce the pressure dif-ferential reauired to drive water into the nechanism to equal the screw displacement. The auxiliary flow paths are closed for small pressure differentials (several inches of water) by ball check valves which prevent the convection flow but open fully during trip. y3

 " J                                                           00000239
t 3-90a L(Revised T-15-69)
                                                                                          .
j. Position Indications 7

Two methods of position indication are provided; one, an absolute position indicator and the other, a relative position indicator. lh The absolute position transducer consists of a series of magnetically opertted reed switches mounted in a tube parallel to the motor tube extension. Each switch is hermetically sealed. Switch contacts close when a nermanent magnet mounted on the upper end of the lead screw extension comes in close proximity. As the lead screw (and the control rod assembly) moves, switches operate sequentially producing an analogue vo]tage proportional to position. .The ac-curacy of the analogue signal is + 1.h 7er cent and produces a readout of approximately + 2 5 per ce. accuracy. Additional reed switches are included in the same tube with the absolute position transducer to provide full withdrawal and insertion signals. The relative position indicator consists of a small tulse-stepping motor driving a potentiometer that generates a signal accuracy of

           + 0.7% producing a position readout of + 1.7% accuracy.                          -
k. Motor Tube Desien Criteria The motor tube 6esign complies with Section III of the ASME Boiler and Pressure Vessel Code for a Class A vessel. The operating tran-ient cycles, which are considered for the stress analysis of the re-actor pressure vessel, are also considered in the motor tube desirn.

Quality standards relative to material selection, fabricatien, and inspection are specified to $nsure safety function of the housings essential to accident prevention. Materials conform to ASTM or ASPE, Section II, Material Specifications. All velding shall be perfonned by personnel cualified under ASME Code, Section IX, Welding Qualifi- ll cations. These design and fabrication procedures establish quality assurance of the assemblies to contain the reactor coolant safely at operating tenterature and pressure. In the highly unlikely event that a pressure barrier conronent or the control rod drive assenhly does fail catastrophically, i.e., ruptured completely, the following results vould ensue:

1. Control Rod Drive Nozsle The assembly would be ejected upward as a missile until it was stopped bv the missile shield over the reactor. This upward motion would have no adverse effect on adjacent assemblies.
2. Motor Tube The failure of this component anywhere above the lover flange veuld result in a missile-like ejection into the missile shield-ing over the reactor. This upward motion would have no adverse effect on adjacent mechanisms.

3.2.h.3.2.2 Axial Power Sharing Rod Drive For actuating the partial length control rods which maintain their set position during a reactor-trip of the shin safety drive, the CRDM is modified so that the roller nut assembly vill not disengage from the lead screw on a loss of power to the stator. Except for this modification, the shin drives and the axial power shaping rod drives are identice.l. nn 3-90b (Revised 7-Ir.69) I !! O -

b. O ? 4 0

_

__ -- . . - . _ _ - _ _ - O

                                                                                                 <

(DELETED) 1

;

O t i ! 00000241  :

                 .

. O:

                                .

3-91 (Revised T-15_69)

  ----- _ . _ __ _ _._ _ _____ _ _ __    . _ _ _ _ _ ,

F 7 Paragraph deleted. 3 2.4 3 3 Control Rod Drive System Evaluation

a. Desien Criteria The system vill be designed, tested, and analyzed for com-pliance with the design criteria. A preliminary safety analysis of the control rod drive motor control subsystem was conducted to determine failures of logic functions. It was concluded that no single failure in any CRA control vould prevent CFA insertion, nor cause inadvertent CRA withdrawal of another CFA or CRA eroup.
b. Materials Selection Materials are selected to be compatible with, and operate in, the reactor coolant. Certified mill test reports con-taining chemical analysis 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 Desien Termerature All parts of. the control rod drive exposed to reactor cool-ant are designed to operate at 650 F, although it is ex-pected that all parts vill operate considerably cooler.

Some tests have been completed, and additional tests are planned, to closely determine the operating temperature gradients throughout the drive mechanism during all phases of operation. These tests will also provide an indication of the amount of convection that takes place within the water space of the mechanism.

d. Desien Life The design life of the control rod drive control system is as follows:

00000242 0 3-92 (Revised 7-15-69)

T p (1) Structural portions , such as flanges and pressure ( ) housings - h0 years. (2) 14oving parts, such ns lead screw and roller nuts - 20 years. (3) Electronic control circuitry - 20 years. 3.2.h.3.h. Control Rod Assembly (CFA) Each control rod assembly is made up of 16 control rods which are coupled to a single Type 30h stainless steel spider (Figure 3-69). Each control rod con-sists of an absorber section of silver-indium-cadmium poison clad with cold-worked, Type 304 stainless steel tubing and Type 30h 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 poison. 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 assembled, the CRA cannot be withdrawn far enough to cause disen-gagement of the control rods from the incore guide tubes. Pertinent design data are shown in Table 3-22. Table 3-22 p Control Rod Assembly Design Data

 %J Item                                              Data         T Number of CRA                                    61 Number of Control Rods per Assembly              16 Outside Diameter of Control Rod, in.             0.hh0 Cladding Thict< ness , in.                       0.020 Cladding Material                                Type 30h SS, Cold-Worked End Plug Material                                Type 30h SS, Annealed Spider Material                                  SS Grade CF3M Poison Paterial                                  805 Ag, 155 In, 55 Cd Female Coupling !!aterial                        Type 30h SS, Annealed Length of Poison Section in.                     134 Stroke of Control Pod, in.                       139 This type of CRA has been developed nder the USAEC Large Peactor Development

- Program and offers the following sign)ficant advantages:

a. More uniform distribution e e.bsorber throughout the core volume.
b. Shorter reactor vessel and shorter internals owing to elimination of control rod followers.

N

 
c. Lower reactor building requirements owing to reduction of reactor
                '. coolant inventory.
          . d. Better core power distribution for a given CRA vorth.

3-93 (Revised 7-15-69) 000 002

                                                        -  .        .  -

- -n - ~_ .- ~ ._ 3.2.h.3 5 Axial Power Shaping lod Assembly ( APSRA) 7 Each axial power shaping rod assembly (Figure 3-70) has 16 axial power shaping ' rods, a stainless steel spider, and a female coupling. The 16 rods are attached to the spider by means of a nut threaded to the upper shank of each rod. After assembly all nuts are lock velded. The axial power sharing rod drive is coupled to the APSRA by a bayonet connection. The female couplings of the APSRA and CRA have slight dinensional differences to ensure that each type of rod can only be coupled to the correct type of drive mechanism. When the APSRA is inserted into the fuel assenbly it is guided by the guide tubes of the fuel assembly. Full length guidance of the APGRA is provided by the control rod guide tube of the upper plenum assembly. At the full out posi-tion of the control rod drive stroke, the lower end of the APSRA remains within the fuel assembly guide tube to maintain the continuity of guidance throughout the rod travel length. The A"SRA's are designed to termit maximum conformity with the fuel assembly guide tube throughout travel. Each axial power shaping rod has a section of neutron absorber material. This absorber material is an alloy of silver-indium-cadmium and is clad in cold-vorked, stainless steel tubing with stainless steel upper and lower end pieces. The end pieces are velded to the clad to form a water and pressure-tight container for the absorber material. The tubing provides the structural strength of the axial power shaping rods and urevents corrosion of the absorber material. Above the section containing the absorber material is a tubular follower made of cold-worked Zircaloy-h tubing, with Zircaloy h upper and lower cnd pieces. The end pieces are velded to the tubing and are vented to permit the coolant-moderator to fill the follower. The follower and absorber sections are fitted together, pinned, and lock welded to form a complete g axial power shaping rod. Pertinent data on the APSRA is shown in Table 3-23. T Table 3-23 Axial Power Shanine Rod Assemb1v Data Item Data Number of Axial Power Shaping Rod Assemblies 8 Humber of Axial Power Shaping Rods per Assembly 16 Outside Diameter of Axial Power Shaping Rod, in. 0.hho Cladding Thickness, in. 0.021 Cladding Material Tyne 30h SS, cold-worked Clad End Plug Material Type 30h SS, annealed Follover Tube Material Zircalov-h, cold-worked Follower End Plug Material Zircaloy-h, annealed Absorber Section to Follower Pin Material Type 30h SS , annealed Poison Material 80% Ag, 15% In, 5% Cd Spider Material RS, Grade CF3M Fcrale Coupling M terial Type 30h SS, Annealed Length of Poison Section in. 36 g Stroke of Control Rod, in. 139 00000244 3-9h (Revised 7-15-69) _ __

                                                                                                 ,
                                                         .

D's ./ - 7

        '

These axial power shaping rods are designed to withstand all operating loads _ including those resulting from hydraulic forces and thermal gradients. i The ability of the axial power shaning rod clad to resist collapse due to the systen nressure has been demonstrated by an extensive collapse test - progran on Zircaloy h tubing. Internal pressure is not generated within the clad since the Ag-In-Cd alloy does not yield gaseous products under irradiation. Svelling of the absorber naterial is negligible, and will not cause unacceptable clad strain.

              . Mechanical interference between axial power shaping rods and the ~ fuel assembly gu!de ubes can be tolerated, since the mechanical interference between axial pver shaping rods and the fuel assembly guide tubes must be. expected. The parts involved are flexible and result in very small friction drag loads. Thernal distortions of the rods are small because of the low heat generation and adequate coolinF. Consequently, the APSRA's vill not encounter significant frictional resistance to their. notion in the guide tubes.

l D , ! V i l

   ~

! !

 '
     'ny' 00000245
              *                               .

3-pha (Revised 7-15-69)

33 TESTS AND INSPECTIONS 331 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 vorth of the CRA under various conditions similar to those f - the reference core. These parameters include control rod arrangement in .L CRA, fuel enrichments, fuel element geometry, CRA materials, and soluble boron concentration in the , moderator. Gross and local power peaking were also studied, and three-dimensional.pover-peaking data were taken as a function of CRA insertion. Detailed peaking data vere also taken between fuel assemblies and around the water holes left by withdrawn CRA's. The experimental data are being analyzed and will become part of the experimental bench mark for the analytical models used in the de-sign. 3 3 1.2 Zero Power, Approach to Power, and Power Testing Boron worth and CRA vorth (including stuck-CRA vorth) vill be determined by physics tests at the beginning of each core cycle. Recalibration of boron worth and CRA vorth is expected to be performed at least once during each core

    -cycle. Calculated values of boron worth and CRA vorth vill be adjusted to the

, l 0 test values as necessary. The boron vorth and CRA vorth at a given ti=e in core life vill be based on CRA position indication and calculated data as ad-i justed by experimental data. ! ' The reactor coolant will be analyzed in the laboratory periodically to deter-mine the boron concentration, and th- reactivity held in boron vill then be calculated from the concentration and the reactivity worth of boron. The method of =aintainin6 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 vill be monitored by reactivity calculations, and inter-l locks vill be provided to prevent CRA patterns that produce excessive power peakingand/orreductionofshutdownmargin. Operation under all power conditions will be monitored by incore instrumenta-tion, and the resulting data vill be analyzed and compared with multidimen-sional calculations in a continuing effort to provide sufficient support for further power escalations.

      .q 332           THERMAL AND HYDRAULIC TESTS AND INSPECTION 332.1           Reactor Vessel Flow Distribution and Pressure Drop Test
   - A 1/6-scale model of the reactor vessel and internals will be tested to measure l                                                                           00000246 3-95
              , _                                        _    _   - _ _ _ ~ , - -,.., _ _ .

l l

a. Tha flow distribution to c ch fu;l ass;mbly of the rr.ctor cora and to l 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.
                                                                                        /g
c. The overall pressure drop between the vessel inlet and outlet nozzles, and the pressure drop between various points in the reactor vessel flow circuit.

The reactor vessel, thermal shield, flow baffle, core barrel, and upper plenum assembly are made of clear plastic to allow use of visual flow study techniques. All parts of the model except the core are geometrically similar to those in the prototype reactor. However, the simulated core was designed to maintain dynamic similarity between the model and prototype. Each of the 177 simulated 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 vill be measured in the inlet and outlet no sles of the reactor model and at the inlet 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 instrumentation will provide the data necessary to accomplish the objectives set forth for the tests. 3.3.2.2 Fuel Assembly Heat Transfer and Fluid Flow Tests B&W is conducting a continuous research and development program for fuel assembly heat transfer and fluid flow applicable to the design of the reference reactor. Single-channel tubular and annular test sections and multiple rod assemblies have been tested at the B&W Research Center. The reactor thermal design is based upon burnout heat transfer experiments with 1 (a) multiple rod, heated assemblies with uniform heat flux, and (b) single rod, annular heaters with nonuniform axial heat flux, at design conditions of pressure h and mass velocity. These experiments are being extended to test nonuniform mul-tiple rod heater assemblies as described in 1 5.h. The results of these tests will be applied to the final thermal design of the reactor and the specification of op-erating 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 l obtained. References to uniform flux data are given in BAW-168 and 3.2.3.2.3 of l this report. The effect on the critical heat flux caused by nonuniform axial power generation in a tubular test section at 2,000 psi pressure was investigated as early as 1961.(29) This program was extended to include pressures of 1,000, 1,500, and 2,000 psi and mass velocities up to 2.5 x 10 6 lb/hr-ft2.(63) The effect on the critical heat flux caused by differences in the radial and axial power distribution in an annular test section van recently investigated at reac-tor design conditions.(64) Data were obtained at pressures of 1,000, 1,500, 2,000, l and 2,200 psi and at mass velocities up to 2.5 x 10 6 lb/hr-ft2, The tubular tests included the following axial heat flux shapes where P/P is local to average pover:

a. Uniform Heat Flux (P/P) = 1.000 constant
b. Sine Heat Flux (T/P) max = 1.396 @ 50% L 00000247 O
            ,

3-96 (Revised 1-15-68)

c. Inlet Peak Heat Flux (P/P) = 1 930 @ 25% L O d. Outlet Peak Heat Flux (P/E) = 1 930 @ 75% L Tests of two additional, nonunifom, 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/5) = 1.65 6 28% L
b. Outlet Peak Heat Flux (P/E)mu = 1.65 @ 72% L These tests, still in progress, vill cover approximately the same range of pressure, mass flow, 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: 0 5 T gs 250 1,000 5 P 5 2,400 0.2 x 106 5 G 5 3 5 x 10 6 where AT g= M et s % coo % , F P = pressure, psia G = mass velocity, lb/hr-ft 2 The 6eometry of this section consisted of nine rods of 0.420 in. diameter on a 0 558 in. square pitch. Analysis of the last data of this set is in process. 3 3 2.2 3 Fuel Assembly Flow Distribution and Pressure Drop Tests Flow visualization and pressure drop data have been obtained from a 10-times-full-scale model of a single rod in a square flow channel. These data have been used to refine the spacer ferrule designs with respect to mixing turbu-lence and pressure drop. Flow distribution in a square 4-rod test assembly has been measured. A salt solution injection technique was used to detemine the average flow rates in the simulated reactor assembly corner cells, vall cells, and unit cells. In-terchannel mixing was obtained for the same assembly. These data have been used to confirm the flow distribution and mixing relationships employed in the

 ,

core thermal and hydraulic design. Additional mixing, flow distribution, and 04 .- 00000248 3-97 _ _ _ _ _ _ , .__ _

                                                       -                                     -  .

pressure drop data vill be obtained to improve the core power capability. The following fuel assembly geometries vill be tested to provide additional data:

a. A 3 x 3 array identical to that for which critical heat flux data 9l '

I have been obtained to provide additional interchannel mixing data. l

b. A 4 x 6 array divided in half by a perforated plate simulating adja- '

cent fuel assemblies to provide data on mixing between assemblies.

c. A full scale 15 x 15 rod fuel assembly to provide additional flow distribution, mixing, and pressure drop information applicable to a complete assembly.

1 (DELETED) 333 FUEL ASSliNBLY, CONTROL ROD ASSEMBLY, AND CONTROL ROD DRIVE MiiUHANICAL 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 performed, are in progress, or are in the final stages of prepara-tion. 3331 Prototype Testing @ A full scale prototype fuel assembly, CRA, and control rod drive is presently being tested in the Control Rod Drive Line (CRDL) Facility located at the B&W Research Center, Alliance, Ohio.. This full-size loop is capable of simulating reactor environmental conditions of pressure, temperature, and coolant flow. To verify the mechanical design, operating compatibility, and characteristics 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 perfomed with maximum misalignment conditions. Equipment is available to record and verify data such as fuel assembly pressure drop, I vibration characteristics, hydraulic forces, etc., and to demonstrate control l rod drive operation and verify scram times. Al'. prototype ccmponents will be Gxa=inedperiodicallyforsignsofmaterialfretting, wear,andvibration/ fatigue to insure that the mechanical design of the equipment meets reactor operating requirements. Preliminary test results are given in 3 2.4 3 5 After the prototype fuel assembly has been tested under simulated reactor operating conditions, it will be installed in the full-size, low pressure loop to verify specific fuel assembly design data. These data include pressure drop, coolant interchannel mixing, and coolant velocity profiles. 3332 Model Testing Many functional improvements have been incorporated in the design of the proto-type fuel assembly as a result of model tests run to date. For example, the g l 3-98 (Revised 1-15-68) 00000249 ! I _ -. _ . , . - . . - . - _ , - _ . _ _ _ - - _

p V spacer Brid to fuel rod contact area was fabricated to 10 ti=es reactor size and tested in a loop simulating coolant flov Reynolds numbers of interest. Thus, visuany, the shape of the fuel rod support areas vas optimized with re-spect to mini =1 zing the severity of flow vortices. Also, a 9-rod (3 x 3) ac-tual size model was fabricated (using production fuel assembly materials) and tested at 6h0 F, 2,200 psi, and 13 fps coolant flov. Principal objectives of thistestweretoevaluatefuelrodcladdin6tospacergridcontactwear,and/ or fretting corrosion resulting from flow-induced vibration._ A vide range of contact loads (including small clearances) was present in this specimen. No significant wear or other flow-induced da= age was observed after 210 days of loop operation. 3333 co=ponent and/or Material Testine; 33331 Fuel Rod Claddin6 Extensive short time collapse testing was performed on Zircaloy-4 tube speci-mens as part of the B&W overall creep-collapse testing program. Initial test specimens were 0.k36 in. OD vith van thicknesses of 0.020 in., 0.024 in., and 0.028 in. Ten 8-in.-long specimens of each thickness vere individually tested at 680 F at slowly increasing pressure until collapse occurred. Collapse pres-sures for the 0.020 in. van thickness specimens ranged from 1,800 to 2,200 ps18, the 0.024 in. specimens ranged from 2,800 to 3,200 psig, and the 0.028 in, specimens ranged from h,500 to 4,900 psig. The material yield strength of these specimens ranged from 65,000 to 72,000 psi at room temperature, and was 35,800 psi at 680 F. 1"I Additional Zircaloy-4 short t1=e collapse specimens were prepared with a ma-terial yield stress of 78,000 psi at room temperature and 48,500 psi at 615 F. Fif teen specimens having an OD of 0.k10 in, and an ID of 0 365 in. (0.0225 in, nominal vall thickness) were tested at 615 F at increasing pressure until col-lapse occurred. Collapse pressures ranged from h,470 to h,960 psig. Creep-collapse testing was performed on the 0.h36 in. OD specimens. Twelve specimens of 0.024 in. vall thickness and 30 specimens of 0.028 in vall thick-ness were tested in a single autoclave at 680 F and 2,050 psig. During this test, two 0.02h in. van thickness specimens collapsed during the first 30 days and two collapsed between 30 and 60 days. None of the 0.028 in. vall thickness specimens had conapsed after 60 days. Creep-collapse testing was then per-formed on thirty 0.410 in. OD by 0 365 in. ID (0.0225 in. nominal vall) speci-mens 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.h10 in. OD specimens showed no indication of incipient collapse. The 60-day period for creep-collapse testing is used since it exceeds the point of primary creep of the material, yet is sufficiently long to enter the ' ye 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 ps j testing is scheduled .for 1967 using specimens fabricated to the reference de-

 '  sign fuel clad dimensidns, material specifications, and operating
                    ~

s.

                      .

250 3-99

                                                                                       -.
                                                                                   -_

3.3.3.3.2 Fuel Assembly Structural Components The mechanical design of the prototype can panel assemblv is the result of an extensive can panel design and structural evaluation procram. The full-size, simulated loop, functional testing noted in 3.3.31 is expected to verify can panel design criteria. Prototype static and dynamic load testing is underway to verify can panal structural adequaev for vibration, handling, operation, and seismic loads. In the mechanical design of the spacer grids , narticular attention is given to the ferrule-to-fuel-rod contact points. Sufficient load must be applied to position the fuel rods and to minimize fuel rod vibration, yet allow axial thermal differential expansion, and not produce fretting vear in che fuel rod cladding. Static load and functional testing of the prototvre grids vill demonstrate their adequacy to perform within the design requirements. 3.3 3.h cr itrol n o d Drive Tests and Inspection 3.3.3.b.1 Control Pod Drive Developmental Tests The prototype roller nut drive is under test at the B&W Research Center, 7 Alliance, Ohio. Near characteristics of critical components have indicated that naterial com-natibility and structural desiFn of these components vill be adequate for the li fe of the mechanism. The development program has been ecmpleted and the complete prototype con-trol rod drive is being subjected to environmental testing under simulated reactor conditions (except radiation) in the Control Rod Drive Line (CRDL) Facility at Alliance. Environmental tests include: Onerational Tests nperating speeds. Temperature profiles. Trip times for full and partially withdrawn control rod assemblies (CRA) for var $ous flow-induced pressure drops across the CRA. Life Tests (Uith internals assembled to maximum misaliennent permitted by drawing dimensions and tolerances. ) No. of Partial Strohe Span of Control Rod Stroke Stroke Cveles Lencth, in. From " Pull-Tn" Desition, in. 1,550 83 From 56 to 130 5, hon 50 71 121 8.900 ns 11h 130 8,50n 13 126 139 No. of Trip

         ' "
  • From 0 to 139 0000025 O
          $00                     130 3-100 (Revised 7-15-69)

1 l l p () t Misalignment Tests , 100 full strokes and 100 full stroke tries with internals tolerances altered to l'.5 times maximum allowable misalignment. Coupling Tests Complete check of coupling operations after testing. The cycles above meet the: total test reauirements of 5,000 full strokes and 500 trips. The assembly will be comoletely disassembled and inspected at various B&W facilities after completion of environmental' testa. 3 3.3.h.2 Control Rod Drive Control System Developmental

    ,                         Tests A control rod drive power supply unit has been built in group nrototype form.      Following the combined test of the power supply and mechanisn, thermal, life, and simulated failure tests vill be conducted. The sinulated failure test will be designed to verify the safety analvnis.

The control rod drive control system vill be tested in conluction with the con-trol rod drive notor control to insure proper eneratien. cirulated failure test-ing vill also be perforced on the ecmbined systen to insure that protective requirements are being met. () The position indicator and limit switch subsystem has been built in prototype fons and life-tested mechanically under expected environnental conditiens. Further testing, both mechanical and electrical, vill be done under expected environmental conditions at the B&W Pesearch Center. Characteristics to be determined will include accuracy, repeatability, linearitv, short term stability, and long term stability. 3 3.3.h.3 Production Tests Production tests discussed in this section vill be nerforced either on the drives installed, or. on drives nanufactured to the same snecifications. The finished control rod drive vill be proof-tested as a connlete systen, i .e. , mechanism, motor control, and system control working as a system. Thin proof-testing vill be above and beyond any developnental testing performed in the product development stages. Mechanism production tests vill include

a. Ambient Tests Coupling tests.

Operating speeds.

                 .

'

 'l                                                                        00000252
  • 3-101 (Revised 7-15-69) 4

_- Position indication. Trip tests. .O \

b. Operational Tests i

Operating speeds. l

                                                                                                                       )

Position indication. Partial and full stroke cycles. Partial and full stroke trip cycles. l l Control system production tests will be performed as described in the follow-j ing paragraphs. l 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 will include such things as automatic opera-tion, manual group operation, trim or single CM operation, position indication of all CM's, travel limit on all CM's, trip circuit operations, IN command, 01/r 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 vill be tested. Power supply tests will be performed to detemine the upper and lower operating voltage and to prove inrounity to switching transients. Fault conditions vill be simulated to prove that no unsafe action results fmm defective components, circuits, or viring. Ability to detect unsafe fault con- , ditions at the operating console vill be determined. Typical of faults to be l simulated are i a. Defective limit switch or circuit. l

b. Impmper UM group 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'CM control circuit or switch.

J. Defective povar supply.

k. Defective motor translator.

1. m. Defective motor cable.

            .Defectivc g,sition transmitter.

00000253 g 3-102 l l '

                                           .           . . - - . - . .- - - . . . . - . - , , - . . , , . - - .-     ,
       -_

() D The finished hardware vill be visually inspected for quality of workman-ship. This inspection vill include an examination of the enclosure, cable entrances, dust-tightness, maintenance features, drawers and cable retractors, fasteners, stiffeners, module counts, wire harnesses, and other similar details. 334 INTERNALS TESTS AND INSPECTIONS The internals upper and love" plenum hydraulic design vill be evaluated andguidedbytheresultsfromthe1/6-scalemodelflowtestwhichisde-scribed in detail in 3 3 2.1. These test results vill indicate areas of gross flow maldistribution and allow verification of vessel flow-pressure drop computations. In addition, the test results will provide measured pressure pulses at specific locations to aid in assessing the vibration response characteristics of the internals components. The effects of internals misalign=ent will be evaluated on the basis of the test results from the CRDL tests described in 3 3 3 4. 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 en-viron=ent. These tests will not include the effects of neutron flux ex-posure. After completion of shop fabrication, all internals components will be N shop-fitted and asse= bled to final design requirements. The assembled , internals components will be installed in a mockup of the as-built reac-

'

ter vessel for final shop fitting and alignment of the internals for the mating fit with the reactor vessel. Dummy fuel and CRA's will be used to check out and insure that a=ple clearances exist between the fuel and in-ternals structures guide tubes to allow free movement of the CRA through-out its full stroke length in various core locations. Fuel assembly mat-ing fit vill be checked at all core locations. The dummy fuel and CRA's vill be identical to the production components except that they will be manufactured to the most adverse tolerance space envelope; even though the assembly weights will be representative of the production units, the du==y components will not contain fissionable or poison materials. Internals shop fabrication quality centrol tests, inspection, procedures, and methods vill be similar to the pressure vessel tests described in de-tail in h.1.4. With regard to the internals surveillance spe:imen holder tubes, the ma-terial irradiation surveillance program is described in 4.4 3 All internal emironents can be re=oved from the reactor vessel to allow inspection of all vessel interior surfaces (see 4. +.J.). Internals com-ponents surfacer ean be inspected when the internals are removed to the canal stors6e loc'ation. 0000 W \ 3-103

                                          .

The internals vent valves will be designed to relieve the pressure generated 1 by steaming in the core following the LOCA so that the core vill remain suffi-ciently covered. The valves will be designed to withstand the forces resulting from rupture of either a reactor coolant inlet or outlet pipe. Testing of the valves vill consist of the following:

a. A full-size valve assembly (seat, locking mechanism, and socket) will be tested at steady-state conditions at the maximum pressure expected to result during the blowdown.
b. Sufficient tests will be conducted at zero pressure to determine the frictional loads and clearances in the hinge assembly, the inertia of the valve cover, and the deflections resulting from impact of the cover so that the valve response to cyclic blowdown forces may be determined analytically.
c. The valve assembly will be pressurized to determine what pressure differential is required to cause the valve to begin to open. A de-termination of the pressure differential required to open the valve to its maximum open position will be simulated by mechanical means.
d. A valve assembly will be installed and removed remotely in a test stand to judge the adequacy of handling equipment.

Since the temperature differential existing across the valve assembly during normal operation in the reactor is only approximately 55 F, and since the same material is used for the valve seat, socket, and cover, there is no need to conduct tests at elevated temperatures. The valves are located in a region of relatively low velocity and turbulence, O and preliminary analysis indicates that there is insufficient energy in the coolant to cause vibrational proble:rs. Therefore, no testing to prove the vi-bration adequacy of the valve is planned. During refueling outages after the reactor vessel head and the internals ple-num assembly have been removed, the vent valves vill be accessible for visual and mechanical inspection. A remote inspection tool vill be provided to en-gage.with the previously mentioned valve dise hook or eye. With the aid of this tool, the valve disc can be manually exercised to evaluate the dise free-dom. The hinge design will incorporate special features, as described in

3. 2. h .1( f) , to minimize the possibility of valve disc motion impairment dur-ing its service life.

Remote installation and removal of the vent valve assemblies will be performed with the aid of another tool which will include unlocking and operating fea-tures for the wedge ring. This handling tool design will be functionally de-veloped and tested on a fu]l-size mockup of the vent valv'e installation config-uration prior to valve manufacture. With the aid of the above described inspection tool, a visual inspection of the valve body and disc sealing faces can be performed for evaluation of observed surface irregularities. O 3-103a (Revised 1-15-68) 00000255

the valve disc, hinge shaft, shaft journals (bushings), disc journal recepta-

                 -

1 clas, and valve body journal receptacles will be designed to withstand without failure the internal and external differential pressure loadings resulting 7

   -

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 press'ure vessels. The hinge materials will be selected on the basis of their corrosion resistance, surface hardness, antiga111ng 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 impaired. In the unlikely event that a hinge part shauld fail during normal operation, the most significant indication of such a failure would be a change in the free-disc mo-tion as a result'of altered rotational clearances. 4 O

                          '
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00000256 1 3-103b (Revised 1-15-68)

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                                                        -

_ 34 REFERDiCIS (1) Putnam, G. E., TOPIC - A Fortran Program for Calculating Wansport of Particles in Cylinders, IDO-16968, April 1964. g (2) Ave:y, A. F., The Prediction of Neutron Attenuation in Iron-Water Shields, AEEW-R125, April 1962. (3) Bohl, H., Jr., et _al_., P3)C1, A One-Dimensional Multigroup P-3 Program for the Philco-2000 Computer, WAPD-TM-272. (4) Bohl, H., Jr. and Hemphill, A. P., MUFT-5, A Fast Neutron Spectrum Pro-gram for the Philco-2000, WAPD-TM-218. (5) Amster, H. J. and Callaghan, J. C., KATE-1, A Program for Calculating Wigner-Wilkins and Maxwellian-Averaged Themal Constants on the Philco-2000, WAPD-TM-232. (6) Marlove, O. J. and Suggs, M. C., WANDA-5, A One-Dimensional Neutron Dif-fusion Equation Program for the Philco-2000 Computer, WAPD-TM-241. (7) Honeck, H. C., THERMOS, A Thermalization Transport Theory Code for Reae-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-Dep3etion Problem, WAPD-TM-477 (9) Marlove, O. J., Nuclear Reactor Depletion Programs for the Philco-2000 Computer, WAPD-TM-221. (10) Lathrop, K. P., DIF-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-454. (13) Clark, R. H. and Pitts, T. G., Physics Verification Experiments, Core I, BAW-TM-455 (14) Clark, R. H. and Pitts, T. G., Physics Verification Experiments, Cores II and III, BAW-TM-458. l (15) Spinks, N., "The Extrapolation Distance at the Surface of a Grey Cylin-drical Control Rod",13 lear Science and Engineering 22, pp 87-93, 1965 (16) Clark, R. H., Batch, M. L., and Pitts, T. G., Lumped Burnable Poison Pzugram - Final Report, BAW-3492-1. (17) Neuhold, R. J., Xenon Oscillation, BAW-305, 1966. O 3-104 00000257

O (18) Wilson, R. H. and Ferrell, J. K., Correlation of Critical Heat Flux for Boiling Water in Forced Circulation at Elevated Pressures, The Babcock

         & Wilcox Company, BAW-168, November 1961.

(19) U.S.-Euratom Joint R&D Program, Burnout Flow Inside Round Tubes With Non-unifom Heat Fluxes, The Babcock & Wilcox Company, BAW-3238-9, May 1966. (20) Jens, W. H. and Iottes, P. A., Analysis of Heat Transfer Burnout, Pres-sure Drop, and Density Data for High Pressure Water, ANL-4627, May 1951. (21) Owen, D. B., Factors for One-Sided Tolerance Limits and for Variable Sa=pling Plans, SCR-607, March 1963 (22) DeBortoli, R. A., et_ al., Forced Convection Heat Transfer Burnout Studies for Water in Rectangular Channels and Round Tubes at Pressures Above 500 psia, WAPD-188, Bettis Plant, Pittsburgh, Pennsylvania, 1958. (23) USAEC Docket 50-244, Exhibit D-3, entitled " Rochester Gas and Electric Corporation, Brookwood Nuclear Station Unit No.1", (Third Supplement to: Preliminary Facility Description and Safety Analysis Report, February 28,1966). (24) Lee, D. H. and Obertelli, J. D., An Experimental Investigation of Forced Convection Burnout in High Pressure Water. Part 1, Round Tubes With Uniform Flux Distribution, AEEE-R-213, August 1963 O (25) Matzner, B. and Griffel, J., Bimonthly Progress Report (MPR-XIII-11 r2 12-63), Task XIII of Contract AT(30-3)-187, Basic Experimental StudieE~ of Boiling Fluid Flov and Heat Transfer at Elevated Pressures, for 'M . vember and December 1963, January 27, 1964. (26) Matzner, B. and Griffel, J., Monthly Progress Report (MPR-XIII-6-63), Task XIII of Contract AT(30-3)-187, Basic Experimental Studies of Boil-l ing Fluid Flow and Heat Transfer at Elevated Pressures, for June 1963, l June 28, 1963 (27) Matzner, B., Monthly Progress Report (MPR-XIII-5-63), Task XIII of Con-tractAT(30-3)-187, Basic Experimental Studies of Boiling Fluid Flow and Heat Transfer.at Elevated Pressures, for May 1963, May 31,1963 (28) Internal Me=o, Weatherhead, R. J. to Inttes, P. A., Critical Heat Flux i (Burnout) in Small Diameter Tubes at 2000 psia, December 29, 1956. l (29) Svenson, H. W., Carver, J. R., and Kakarala, C. R., The Influence of Axial Heat Flux Distribution on the Departure Fmm Nucleate Boiling in a Water Cooled Tube, ASME Paper 62-WA-297 (30) Nonunifom Heat Generation Experimental Pmgram, Quarterly Progress Re-Port No. 7, January - March 1965, BAW-3238-7, Joint U.S.-Euratom B&D Program, A E Contract No. AT(30-1)-3236. . h (31) Hald, A.,: Statistical Theory With Engineering Applications, John Wiley & Sons, Inc., New York, 1955 9 $0000258 3-105

                                                                                ._ _

(32) Worthing, A. G. and Geffner, Wiley & Sons, Inc., New York, 1943 J., Treatment of Experimental Data, John g (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 Reaktor Project, December 1962. (34) zuber, N. and Findlay, J. A., Average Volumetric Concentrations in Two Phase Flov Systems, Presented at the ASME Winter Meeting, 1964. To be published in tne ASME Transactions. , (35) Maurer, G. W., A Method of Predicting Steady-State Boiling Vapor Frac-l tions in Reactor Coolant Channels, Bettis Technica: Review, WAPD-BT-19 1 (36) Baker, 0., Simultaneous Flov of 011 and Gas, Oil and Gas Journal, Vol. 11, pp 185-195, 1954. (37) Rose, S. C., Jr., and Griffith, P., Flow Properties of Eubbly 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 1964. (39) Bergles, A. E. and Suo, M., Investigation of Boiling Water Flow Regimes at High Pressure, NYO-3304-8, February 1, 1966. (h0) Notley, N. J. F., The Themal Conductivity of Columnar Grains in Irradi- O ated U02 Fuel Elements, AECL-1822, July 1963 (41) Lyons, M. F., et al., UO2 Fuel Rod Operation With Gross Central Melting, , GEAP-426k, October 1963 (42) Notley, M. J. F., g _al_., zircaloy-Sheathed U0 2 Fuel Elements Irradiated at Values of Integral kd6 Detween 30 and 83 w/cm. AECL-1676, December 1962. (43) Bain, A. S., Melting of UO2 During Irradiations of Short Daration, AECL-2289, August 1965 (44) Notley, M. J. F., et al,., The Iongitudinal and Diametral Expansions of UO2 Fuel Elements, AECL-2143, November 1964. (h5) Duncan, R. N., Rabbit Capsule Irradiation of UO ,2CVNA-142, June 1962. I (46) Lyons, M. F., et al. UO Pellet Thermal Conductivity From Irradiations With Central M Et Eg, G -4624, July 1964.

                                         ,

l (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-246. O 00000259 3-106

 -P w+                            cy g     r p -    o     -
                                                                                          ,-c--   -e.

_ _ _ _ _-. O (48) Ross, A. M. and Stoute, R. L., Heat Transfer Coefficients Between UO2 and Zircaloy-2, _AECL-1552, June 1962. (49) Hoffman, J. P. and Coplin, D. H., The Release of Fission Gases From Uranium Dioxide Pellet Fuel Operated at High Temperatures, GEAP-4596, September 1964. (,') Spolaris, C. N. and Megerth, F. H., Residual and Fission Gas Release From Uranium Dioxide, GEAP-4314, July 1963 (51) Robertson, J. A. L. Fuel, ABCL-603,195b.et al., Behavior of Uranium Dioxide as a Reactor

                                                      -

(52) Parker, G. W., e_t,al., t Fission Product Release From U02 by High Tempera- , ture Diffusion and Melting in Helium and Air, CF-60-12-14, ORNL, Febru-j ary 1961. ! (53) Daniel, R. C., e_t al_., Effects of High Burnup on Zircaloy-Clad, Bulk U028 Plate Fuel Element Samples, WAPD-263, September 1962. (54) Biomeke, J. O. and Todd, Mary F., Uranium Fission Product Production as a Function of Thermal Neutron Flux, Irradiation Time, and Decay Time, ORNL-2127, Part 1, Vol. 1 and 2. (55) Duncan, R. N., CVTR Fuel Capsule Irradiations, CVNA-153, August 1962. (56) Frost, Bradbury, and Griffiths (AERE Harwell), Irradiation Effects in Fissile Oxides rmd Carbides at Iow and High Burnup Levels, Proceedings of IAEA Symposium on Radiation Damage in Solids and Reactor Materials, Venice, Italy, May 1962. (57) Gerhart, J. M., The Post-Irradiation Examination of a Pu0 -UO2 2 Fast Re-actor Fuel, GEAP-3833 (58) Atomic Energy Clearing House, Vol. 12, No. 3, P 11. (59) Large closed-Cycle Water Reactor Research and Development Program Prog-ress Report for the Period, January 1 to March 31, 1964, Westinghouse Electric Corporation, Pittsburgh, Pa., 1964, WCAP-3269-2. Also WCAP-3269-3 for period from April 1 to June 30, 1964. (60) Physical and Mechanical Properties of Zircaloy-2 and -4, WCAP-3269-41, Figure 18. , (61) Burgreen, D., Byrnes, J. J., and Benforado, D. M., " Vibration of Rods Induced by Water.in t Parallel Flow", Trans. ASME 80_, p 991, 1958. (62) Large Closed-Cycle Water Reactor IED Program, Progress Report for the Period January 1 to March 31, 1965, WCAP-3269-12. p (63) Burnout for Flow. Inside Round Tubes With Nonuniform Heat Fluxes, BAW-d 3238-9, May 1966.

                                               '

00000260 3-107 _. , .. - - - - . .. - .. . _ - _ . - - . _ . , . - . - - - _ - - . . . . - _.

(64) Nonunifom Heat Generation Experimental Program, BAW-3233-13, July 1966. g (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. (67) Clark, R. H., Physics Verification Experiment, Axial Power Mapping on Core IV, BAW-TM-255, December 1966. (68) Larsen, P. S., et al., DNB Measurements for Upwanis Flow of Water in an l ' Unheated Square Channel with a Single Unifomly Heated Rod at 1600-2300 psia, Proceedings of the Third International Heat Transfer Conference, August 1966. I l 00000261

      -

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                                     '

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add 'uo!)eagueouog uoJos metsig guetoog Jo}sesy 00000262 BORON CONCENTRATION VER$US CORE LIFE g

 \J                                                                                                     CRYSTAL RIVER UNITS 3 & 4
                                                                                                                               -

FIGURE 3 1 AMEND.1 (1 15-68)

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l i O l 1 Axial Power Profile for 557. insertion is shown on Figure 3-3.

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AXlAL PEAK TO AVERAGE POWER YER5US XENON OVERRIDE ROD INSERTION CRYSTAL RIVER UNITS 3 & 4 000263 8 IA FIGURE 3-2

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00000264 ( O , AXIAL POWER PROFILE, XENON OVERRIDE ROD 5 55 PERCENT INSERTED CRYSTAL RIVER UNITS 3 & 4 E- FIGURE 3-3

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Moderator Boroi. Concentration, ppm boron x 10 1 l i f 00000265 l MODERATOR TEMPERATURE COEFFICIENTS VERSUS BORON CONCENTRATION CRYSTAL RIVER UNITS 3 & 4

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a l l l PERCENT INITIAL POWER VER$US TIME FOLLOWING TRIP CRYSTAL RIVER UNITS 3 & 4 000007.67

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l l m FIGURE 3-6 AMEND.1 (115 68)

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2. Oscillation insteated at T 2 days.

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00000268 (3

 'V EFFECT OF . FUEL TEMPERATURE (DOPPLER)

ON XENON OSCILLATIONS BEGlHNING OF LIFE CRYSTAL RIVER UNITS 3 & 4 aois

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

Case I - Temperature iteration with 7, . 1,400 F. Case 2 - Temperature lieration with 7, . 900 F.

2. Oscillation initiated at T . 300 days.

00000269 EFFECT OF FUEL TEMPERATURE (DOPPLER) OH XENON OSCILLATIONS NEAR END OF LIFE CRYSTAL RIVER UNITS 3 & 4 i -

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2. Oscillation initiated at T . 200 days.

l l 00000270 Q couraot or ixiit oscittirios wirn einriit noos CRYSTAL RIVER UNITS 3 & 4

                                                                         -         FIGURE 3 9
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60 - g' 50 - i i i i 1.0 1.2 1.4 1.6 1.3 2.0 DNB Ratio POPULATION INCLUDED IN THE STATISTICAL 00000271 STATEMENT VERSUS DNB RATIO CRYSTAL RIVER UNITS 3 & 4

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0.. l \ l 1 L Fuel MidPlane 0.2 Core Core T l 0.1 ' T 144" {  ; 0.0 0 20 40 60 80 100 120 140 10 30 50 70 90 110 130 Distance from Bottom of Active Fuel, in. 00000272 O POWER SHAPE REFLECTlHG INCREASED V AXIAL POWER PEAK FOR 144 INCH CORE CRYSTAL RIVER UNITS 3 & 4 E

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FIGURE 3 11

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EE I.85 g 0 90 - (1) Line A x Line B oe 1.79 s

         "_          o,go .    (2) Line C is for illustration only                  's
         .I                    (3) Line 8 is based on detailed data                    N 0.70 -         from a rod by rod printout of a                        g PDQ (two dimensional power and CAO--                                                                   g flux calculation for worst time                                              -

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0 10 PO 30 ko 50 60 70 80 90 100 Percentage of Fuel Rods with Higher Peaking Factors Than Point Values. 7 00000273 DISTRIBUTION OF FUEL ROD PEAKING CRYSTAL RIVER UNITS 3 & 4 _

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         -                FIGURE 3- 12 l

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Line 2 100 FAh Nuclear = 1.79s

I

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j 80

                                    #

p 7 ,,1,, 60 C  ! u go - A r e f

  • 20
                      /    -

l A l

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ib I 100 102 104 106 108 HO H2 H4 H6 H8 120 Rated Power (2,452 MWt),7 00000274 , POSSIBLE FUEL ROD DNB's FOR MAXIMUM DESIGN CONDITIONS - 36,816 ROD CORE CRYSTAL RIVER UNITS 3 & 4 E- FIGURE 3 13 _ _

1M g I e '

  • H l

t , 7 ' a \

      =
      -

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o I
      '                                                                                                                l
     '*        50 l      l
E 40 ' /

3 Fah Nuclear = 179 -- l 2 [ 30 g 8 vi--

     =                                                                                                                       overpower 20                                                                                                j      '
  • r / l
     }e       10                                                      "    /                                          l                                                    h E
                                                     #
                                                           /                                                          l 0                                                                                                     l 100 102                       104   106   108                        110              112 114    116             n8 120 Rated Power (2,452 MWt), %

POSSIBLE FUEL ROD DNB's FOR MOST PROBABLE CONDITIONS - 36.816 ROD CORE CRYSTAL RIVER UNITS 3 & 4 nggo0275 0' I"d"" - FIGURE 3- 14

                  .- -... .- - -._ , . - . - . . . .                  , . - . - . . , . - . . . . _ _ _ , _ , _                 , . . . - - .    -. . - . . . . . . - - .-

(1 P) (P) 0.1 09 0.01 , 0 99

                     \
                      \
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                 '

00000276 l

             '

DISTRIBUTION OF POPULATION PROTECTED P, & 1 P VERSUS NUMBER RODS

 ,tp/                                                                     FOR MOST PROBABLE CONDITIONS CRYSTAL RIVER UNITS 3 & 4 i                                                                                  =

l == FIGURE 3- 15

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    -    2 a'
    -  ' k; ,    .
    -6 100   no     120         130          1ho        150 Rated Power (2.452 MWt). %

00000278 p %) MAXIMUM HOT CHANNEL EXIT QUALITY VERSUS REACTOR OVERPOWER CRYSTAL RIVER UNITS 3 & 4 E.no FIGURE 3- 17 _-

                                                                                                                            .

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Data Based on CVNA - 142 (Jtme 1962) I 1.m i , O 1000 2000 3000 4000 5000 Temperature, F 00000279 THERMAL CONDUCTIVITY OF UO CRYSTAL RIVER UNITS 3 & 4 l m)s_ eiOuRe 3.ie !

                            - _ . , . _ _ _ _ _ _ _ _ . - _ _ . - _ _ . . . - - - -               _ _ . . _ _ _ . _ _ _ _ _            _ - _ _ ..   .

O 6000 l UO Melting l Temperature l 5000 _ _ _ -. _ _ _ ___-. - - p-- 1 1

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l I O l l 0 5 10 15 20 25 LinearHeatRate,kw/ft 000002gg FUEL CENTER TEMPERATURE ST THE HOT SPOT Ox VERSUS LINEAR POWER CRYSTAL RIVER UNITS 3 & 4 b= FIGURE 3- 19

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                                                                                                                                                                                                                                        ;

i l l

                                                                                                                                                                                                                                        !

l HUMBER OF DATA P0lHTS VERSUS E$ /DC CRYSTAL RIVER UNITS 3 & 4 n0000281 s E. FIGURE 3-20 i t _... _ , . . . , _ , . , . _ . . - _ _ _ _ _ _ . . _ , . , _ . . ,___ ,_ _.,_, _ __ _ , _ . _ . , _ _ _ , , _ _ _ , . , _ , _ . _ _ _ , , , _ . _ , _ _ _ . . _ , _ _ , , _ , _ _ , _ _ _ ,

                                                                                                                                                                           . .                                           __.

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l. 025 -

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l T A l .985 -

 .O                                                                                                                                                                                             Population Protected, f.

l

   ..

t 00000282 HOT CHANNEL FACTOR VERSUS A() PERCENT POPULATION PROTECTED CRYSTAL RIVER UNITS 3 & 4

                                                                                                                                                                                                                               -
                                                                                                                                                                                                                               "*"
                                                                                                                                                                                                                               --       FIGURE 3-21
        . _ . . _ ..._ _ _._.._.-, _ . . _ _ . . _ . _ _ - _ . _ _ _ _ . . _ _ . . . _ . _ _ _ _ _ _ _ _ _ _ . _ . _ _ , . . - _ _ . . _ _ . _ _ _ . - _ _ _ _ _ . _ _ _ _ _ _ . . . _ . _ _ _ . . _ _ -
                                                                                                     ~ _ _ _ - .  ..                 ..

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1.0 l.2 l'. 4 d6 $8 2.0 Burnout Factor, DNB Ratio

!

BURNOUT FACTOR VER$US POPULATION FOR VARIOUS CONFIDENCE LEVELS 4 CRYSTAL RIVER UNITS 3 & 4 "al= = FIGURE 3 22

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         --___.___---_._.-_-_-.--.-.....-_.-,-....,vm.,--..--w,---,---.=
                                    --

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O l.00 1.05 1.10 1.15 1.20 Fraction of Rated Power (2,452 MWt) 00000284 RODS IN JEOPARDY VERSUS POWER CRYSTAL RIVER UNITS 3 & 4

                                                                     'C"" o      FIGURE 3-23
   . . _ _ - _           _          - - - - - - - . -
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          -

0lll"- FIGURE 3-24

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                                          "

10 - 0 M i i i i i i i i i i e i i i iiiiiii (9) 0 .2 .4 .6 .8 1.0 1.2 1.4 1.6 1.8 2.0> 2 g

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lll" FlGURE 3-27

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                                                                                    #/C  E Columbia 720 psia Data and BAW-168 l

l RATIO OF EXPERIMENTAL TO CALCULATED BURNOUT HEAT FLUX CRYSTAL RIVER UNITS 3 & 4 E~ FIGURE 3-28 fiOI)OO2h !

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                   '
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                                                                                                          ==           F!GURE 3 31
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1

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         .,
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Y: h h FIGURE 3-34 _ a---y v - y w-- gewv-- -w =---gv ,--ev ---c.y --*-o.-W- -w--y- -%yw g ,9g y w-,e, , - - --9m 9y.w-999- , -.-wmn - . , ,w w

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                                                                                   ==        FIGURE 3 35

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{  !

        =

l0" 0 i i i i i . iiiiie i i e i iie i i 0 .2 .4 .6 .8 1.0 1.2 1.4 1.6 1.8 2.0 > 2 (32) 4 # E C Euratom Chocoed-Cosine 2000 psia Data _ and BAW-168 0 i i i i ii,i hh M. . . . . . , ,iii, 0 .2 .4 .6 .8 1.0 1.2 1.4 1.6 1.8 2.0 > 2 (33) i, / . Euratom and B&W inlet Peak 1000 osia Data and BAW-168 L] 00000298 RATIO OF EXPERIMENTAL TO cA'cu'Ars0 suawouT NEAT etux CRYSTAL RIVER UNITS 3 & 4 E- FIGURE 3 37

     ..       -                    -               .-                - - - .           --                            - -
                                                                                                                                    ,

10 m  ! O  ! O r, i 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 2.0 :e (N) ., s., Euratom and B&W Inlet Peak 1500 psia Data and BAW-168 1D -

                                                          .-

0 & i i i i i e i i e 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 "c
            -
                                                                                /
                                                                            "E ' 'c

[ (35) Euratom and B&W Inlet Peak 2000 osia Data 10 - and BAW-168

           $

x 0 [% n r-rfl

                                      '      '

O' .'2 .'k .6 ' .h ' 1.'O ' 1 2 ' 1. 4 ' 1.6 ' 1. 8 ' 2. 0 >2

                                                                                                                       ' '

(36) sg/"'c Euratom and B&W Outlet Peak 1000 psia Data and BAW-168 10 -

                                                      -

O i i i i N iiie i i i iie i i i 3 i i i 0 .2 .4 .6 .8 1.0 1.2 1.4 1.6 1.8 2.0 >2 (37)

                                                                            . . I.c

! Euratom and B&W Outlet Peak 1500 psia Data and BAW-168

                                                                                                                       ,

RATIO OF EXPERIMENTAL TO CALCULATED BURNOUT HEAT FLUX  ; CRYSTAL RIVER UNITS 3 & 4

         -

h FIGURE 3-38 00000299 3Q A\ G

                           , . - . .         .- .             . -   . _ _ .                                - . , _ . . ..      -.

_ _ _ - -

                                                             /
                                                                 /
 ,q          10 -                              __

Gi _ _ 0 r i i i i i i m d i i .. i .,ii. 0 .2 .4 .6 .8 1.0 1.2 1.4 1.6 1.8 2.0 M (38) *E 'c Euratom and B&W Outlet Peak 2000 psia Data and BAW-168 10 - _ _

                                                                             ~

W Cn

                                                             -

0i nT i e i i a i i i i i i i i i e i i i . i i

        ,
       ;            O     .2           4     .6                 .8        1.0                     1.2                   1.k     1.6 1.8      2.0 M 2 (39)                                                                  'E        c 0
        ,                        All 1000 osia Non-Uniform Data and BAW-168 E
  • B 10 -
       =                                                      _
                                                                        -

_ 0 _ f- _~ q i i i i i i. . . i i i . . . i i i i i . . Qw g -0 .2 .4 .6 .8 1.0 1.2 1.k 1.6 1.8 2.0 M (40) 'E "C All 1500 psia Non-Uni orm f Data and BAW-168 20_ l _

                                                  -

10- - - _- l

                                          -

__ 0 i i i ,m i a iiiiie h i i i e i i i e i 0 .2 .4 .6 .8 1.0 1.2 1.4 1.6 1.8 2.0>e (41) e.gl .'c All 2000 psia Non-Uniform Data and BAW-168 RATIO OF EXPERIMENTAL TO CALCULATED BURNOUT HEAT FLUX [J

 '

00000400 CRYSTAL RIVER UNITS 3 & 4 3* 3 FIGURE 3 39 AMEND.1 (1-15-68)

    .. ,.     - .                                , - -            .       . , . . . - . - , - - - - - . . - - , . - - .                          . - . .        .           - - -
                                                                                                                                                                                                                                              ..-        .     . _ .         ,

1 O

                                                                             +18                                                                                                                                                          88
                                                                             *
                                                                                                                                                                                                                             /     f Design                                                                                                 N Overpower (ll4%)                                                                                                   g
                                                                            +14                                                                                                             e
                                                                                                                                                                                                  /                                       95 f                     .

i ' 412 -

                                                                                                                                                                           /                     7                         ;
                                                                                                                                                                                                                               /        ' 96 N
                                                                                                                                                                                                                                                     -

x k

                                                                            +10 g                        g                                        100        $
                                                                            +8                                                                                                                                                            104    e
                                                                                                                                                                                                                                                     -
  • a 3E
                                                               -
                                                                            +6                                                i
                                                                                                                                .j'/ff              /                                                                                             25 g              ,
                                                                                                                                                 ,

109 g .5 A.,

                                                                                                                          '
                                                                                                                    /           /[                                                                                                       u6     j$
                                                                            +2                               r'                                                                                                                          126
                                                                                                      /            fs     f OE

{ O 7

                                                                                                                   ,

0"'I 144 $ Subcooled = l 5

                                                                                                                       ---            2,120 psig                                   (Fah= 1.85)                                                      $

2.185 psig (Fan . l.85) f

                                                                            -k                                                                                                                                                                      a
                                                                                                                     -           -    2,185 psig                                   (Fah . l.79)                                                      *
                                                                            -6 100                      110            120                              130                                            140                  150 Rated Power (2,452 MWt). %

MA7l MUM HOT CHANNEL EXIT QUALITY VERSUS REACTOR POWER 00000T01 - CRYSTAL RIVER UNITS 3 & 4 g l dha_ rioURe 3 40 AMEND.1 (1 15 68) I i

                                                                                                                                                                                                           -.
 . _ _ _ _ _ _ _ _ _ _ . . _ _ _ _ _ _ _ _ . _ _ . . . . - . . . . . _ . . . _ _ _ . . . . . . - . . . . . . _ .
                                                                                --
                                                                                                                                          - . . , , , . , _ , _ , - . . . . . - . . . . - . . _ - - . _ , - . - . . . . . . , .                            ..        . . . -

_ _ j l

                                                                                                  '

O

+14 l I Il i I i l

, 5% below Average Assembly Flow.

                  +12               --- Average Assembly Flow.                               ,

Il i I

                  +10                                Fah = 1.85 a

p

                  ,g                                 Fah = 1 7
                                                                            '
                  +6                                              /         -
                                                                         /
                - .4                                     .         .   -

5

                , .2                             .
                                                   / // -.
                                                                         '
                                                                            '
                                                      /          /

0 A ' I 4"*1$t:' l i O -2

                         /,///.                  ,
                                                   '

r K ah F = 1 79 Subcoo:.ed

                                     /      /                       l i
                  ,,    / f                                   Fan - 1.85
                       /'/
                  -6   #

I[ "'** "" IllO)

                  -8                           l 100                 110             120             130              140 s

Rated Power (2,452 MWt), % HOTTEST DESIGN & NOMINAL CHANNEL EXIT QUALITY VERSUS REACTOR POWER O.% (WITHOUT ENGINEERING HOT CHANNEL FACTORS) dJ CRYSTAL RIVER UNITS 3 & 4

 . _ _ _ . _ -.          - -___._ _
                                                                 $s_              ,,euRe 2 4,

__

3.0 g

                                                                  \

l14% Power \ 2.5

  • l 130% Power
                                                                    \,
                                                                     "

2~0

       )                                !              f                    \

l-2  : i 3 1.5 I

                                        .
  • i
                                                                                          '
                                        '                                     3
        $            Bubble to                            $

[

        'S           Slug 3                                                  }

(Griffith and J Rose) k 3  % Bubble to E I.0 t Annular g (Baker)

                                                                \

Slug to Annular

                         /           (Haberstroh)                 \

i f \

             $                         #

i g Bubble to Slug N . (Baker) N

  • i 3 l

l 0 10 20 30 40 Quality (Ib vapor / total Ib), % t FLOW REG!ME MAP FOR UNIT CELL CHANNEL 0 CRYSTAL RIVER UNITS 3 & 4 E FIGURE 3-42 03 l l . - . - . _ . . . . . - .

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

                 + Hot Unit Cell Nominal Conditions 3.0
                                                          \        -

6 4 s.A a \

                                                              \
                                                               \
 .
2. 5
                             .
                             '                                   \
  ' o_
   .

m S A A O da A I a 6 3- l

2. 0 -

4 5 8 A d E

E 6 i.5 > Bubble to Slug (Griffith and Ro se) 1.0 Bubble to Annular x (Baker)

                        /                                                                -
        .5 Bubble to Slug (Baker)                          ,

_ 0 10 20 30 40 Quality (Ib vapor / total Ib), 7 FLOW REGIME MAP FOR UNIT CELL CHANNEL CRYSTAL RIVER UNITS 3 & 4 5- 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

                                                                              *
                                                                                                     \
                                                                                                       \
                                                                                                         \
2. 5 g

! O O i O

       .o
  • S 2.0 W
      %                                                   @     @                           E 7

a 2-l 4 h 1.5 N j Z o y e ! O O ! l0 r  ;

                                                                                                                        ,

Bubble to Slug Bubble to Annular * (Griffith (Baker)

                                                                                                                      '

5 . andRose) j Bubble to Slug (Baker) i I

I
                                                                                                                    -
                                                                                          <_         :

0 10 20 30  % Quality (Ib vapor / total Ib), % FLOW REGIME MAP FOR CORNER CHANNEL CRYSTAL RIVER UNITS 3 & 4 9n.)00s05 e _ h FIGURE 3-44 l l - _ - _ _ , _ _ . _ . _ , . _ _ _ _ . _ _ _ _- .-- --_ - _-- - - - - - - -- - ~ . - - - - - - - - -- - - - - - - -

__ g Bundle Burnout Test Conditions where stable operations were observed. 5 Hot Cell Worst Conditions

                + Hot Cell Channel Nominal Conditions 3.0
                                                              \

e 8

                                                                \
                                                                  \
                                                                    \
2. 5 g

g s e e e

  'o M

u, 2.0 i

   $

e

                     +h                     $
     -

g 1.5

  • i g g . Bubble
   .;                                   to Slug (Griffithand Rose) 1.0 d6 Bubble to                  i Annular                    I

( Baker) f

                   '
         .5 Bubble to Slug (Baker)

0 10 20 30  % Quality (Ib vapor / total Ib), 7

              '

k FLOW REGIME MAP FOR WALL CHANNEL CRYSTAL RIVER UNITS 3 & 4

                                                                        =
                                                                        -        FIGURE 3-45
                            .                      - - - - --             ._           .     -
                                                                                                                                    - -                                                   - . . -

9

                                       %                                                        Design Overpower (ll%)

g 2.0 - g\

                                       \ \                                                     1.65 Cosine (W-3) 1.8      -
                                              \

g

                                                           \          N i

s 1.80 Cosine (W-3) l 1.6 \ '

                                                                              ..        g
                          $                                              3'                   \                                 MW-168 Design E    1.4     -                                              ,,,,_ __

_,___,,D M (1 58) E & - W-3 Design

                          --  1.2     -                                                                                             DM (130) 1.65 Cosine (MW-168)\

d 1.0 - \

                         ~                                     1.80 Cosine (MW-168)                                                        \

E \ 0.8 - N 1 50 Cosine (MW-168) 0.6

                                                                                                                                                                 \
                                      -

g N 150Cosins(W-5) O.4 - 0.2 - 0 l l l l 100 110 120 130 140 150 RatedPower(2,t452MWt),7 HOT CHANNEL DNB

  • RATIO COMPARISON CRYSTAL RIVER UNITS 3 & 4 0%00507 O E. FIGURE 3-46 w

w_m-- '---_.-----*---e =e .---w --w--m--e.-w.- e. w- *--93 w -

                                                                                           -e    e ---we---v-    w.wg-ew---wew-       -eW-   r--  3p-w---q-me +-sy--+  -wt  w,.-g, ww w ww grw-w-- wy--w- g
    .

O ( 150 l g Design Power '

                            *

(2,452MWt) I Kl

                              '

140 m i D

                                   . 130
                                                                    }

b '

                                                              ,
u. -

l' 8 l 1* 120 [ l 3

                                #                                      1

'

 .O i                                                                     j 110 s                          i         .

2300 2400 2500 2600 Reactor Core Power, MWt ! l i i l (;) 00000.$08 REACTOR COOLANT FLOW VERSUS POWER CRYSTAL RIVER UNITS 3 & 4 E. FIGURE 3 47

      . . _ _ - . _ - - _ _ _ _ _ _.-.
                                                                                                                                          ,

4.00 h

                                                                                                                                          '

I 002 He ts s

         ~.              T\
            ,

i

                           \

l 2 \ 3.* 6 y i I

                              ,

5 2 i

                         \
                          \
                                \\                                                                                l
            $                     '

Ba# Design Value (CVNA - 142) [ [ CVNA - 2M

                                                                                                          #

E g hn

'

3

2*

7 i \

            !                             \                                                     /                               e 5                                4S
                                                -
                                                                    -
                                                                                  /
                                                                                     /                           J
                                                                                                                  '
                                                                                                                  -
                                                                                                                  '

TI- GEAP h624 l 1.00 l . O 1000 2000 3000 4000 5000 Temperature, F l THERMAL CONDUCTIVITY OF 95r. DENSE SINTERED UO2 PELLETS CRYSTAL RIVER UNITS 3 & 4 _ lllll%. FIGURE 3-48 00000209

                                       ..
                                                    ,,,,,-.,,,-.,.m.-                 ,,..,.,eg   ..wep+    g99Wemw*muv e *-=e+Wv-t='-9-*

1 O

   /

6000 5500 - Design Overpower (1147.)

                                                                                                  /

f ,e

                                                                                            #

1005 Power U0 eltingTemoerature's.\ / / gj, 2 50 . _ - _ . _ _ _ _ _ r_ _____ l  ! / w

                                                              \
                                                                                ,/ j   /

i

         '5"                                                               /     //

i < y

       ;                                                             /,l e                                                            //

e

                                                               ,
       $ 4000                                               #

t o I 3500 Y l l 3000

                              '
                                 /
                                //

BW Design Value (CVNA-142)

                                         -- _- Ref. 42 (GEAP-4624) jr            ---- Ref. 43 (CVN A-246) 1 2500             h,                   e      e           1         t        l 6      8      10         12    14     16       18           20       22     24     26       28  30 Linear Heat Rate, kw/ft 00000410

' O FUEL CENTER TEMPERATURE FC;i h , BEGINNING-OF LIFE CONDITIONS CRYSTAL RIVER UNITS 1 & 4 b FIGURE 3-49

     -                -           -. .-                                         .-
                                                                            ~

l i 6000

                                                                                                                     '
                                                       '
             $$oo
                   ._ Design Overpower (ll4%)                                              ,                      j' l                  /

l 100% Power / , UO,, Melting Temperature 3l\ / / / Smo

                          '

s- r  ; j

                                                                                '

__ / /__ _ _ __

                                                          '
                                                                          /

i / / / 2 u,. /// e

u /

                                                   /l/
                                                 ,//

Y 4000 ///

     -
             ,,.            > v'                                                                                             e

[ - -- B&W Design Value (CVNA-142) Ref. 42 (GEAP-4624)

                                ----- Ref. 43 (CVN A-246) 3000 g

i 2500 6 8 10 12 14 16 18 20 22 24 26 28 3o Linear Heat Rate. kw/f t n FUEL CENTER TEMPERATURE FOR END-OF LIFE CONDITIONS CRYSTAL RIVER UNITS 3 & 4 _

     "'""                                                                         00000311
     -~             FIGURE 3-50 me
       ,w ,.                             --                 --%w--   - -             e        , - -   .,-.w            e p -
   ,-..

( )

   ,..-

100.00

. . .  :. :2: --- y
                                                                                                                     ,

y -. .L, . _1..

                                                                                                                                              -
                     ~                                                                                   . _ _]' .         '       '

_

                                                                                                                                                  ~

50.00 --.. -- __

                                                                                      "                                      G
                                                                   -  e     -A                                 CH
7. *
                          .
                                                            '.                      ,,
                                                                                                             -       ;

I ,

               ,0,00
                          *
                                        .
                                                  ,
                                                    ,/      8                   *                          .
                                                                                                                ..

I I

                                                                                                                                     -
                                                                                                                                      .

A _

          *                                     [                                                                          ,
            *                                 '
          ,     5.00                                '

5  : 'Y t .A

                                            /                                                            i ~                       i        i
                                                                                                           '
          =
          .

3C) { e + J

  • 8
          ;     1.00                 ,

_ i .

                                   #      '                             -
                                                                                                                                  '

0.50

                                  /                                                                     i                 i       ,

e

  • 4 k i i t i Y A I l l l ,

f'\ t /

  \~/                                                                                                O GEAP - 4596
 '

D +

  • GEAP 4314 O.10
                                                                                                     + AECL - 603

_I e i a CF-60-12-14 (0RitL) 0.05 v i . i i i i i i i i f l I I400 1600 1800 2000 2200 2400 2600 2800 3000 3200 3s00 3600 Volumetric Average Temperature, F l l l l 00000512

 ,r '                                                              PERCENT FISSION GAS kELEASED AS A FUNCTION k.y)                                                          OF THE AVERAGE TEMPERATURE OF THE UO2 FUEL CRYSTAL RIVER UNITS 3 & 4 no m
                                                                                                  ~
                                                                                                  -                    FIGURE 3-51
          .

l.8 I e, I

                ~
                                            '  ~N                   P/P - 1.70 (Partial Rod 1.6                                                          '             Insertient
                                       /                                   I I*                     /
                                   /
                                                              \   /                'N            P/ P - 1. 50 f            l           \         (Modified Cosine)

[ \ 100 Oay

               
                                                                                             '
                             /             ff y                                                       '

30i0.,s 1.1 i

                              '
                                        ;-i                   %                                 ,

n

                                                                                                          /; ,,0 0.,s

_ l x

               *         !Nf                /                           l        \              N\

i VD / l \ VN ) 2'

                        /#               /                                !              \                \\ \\
                

N/ / i

                                                                                           '

x \\ \\

                      /h            /                                     l                      \            \\ \\

e

                  '
                      'g ,/                                              l                               N % \\

V

                                                                                                         '
0. 4 ,

i

                                                                         !                                    x    Y#

0.3 f I \ \\) i F Fuel Midplane 0.2 ' Core I Core Bottom l Top 0.1 \

                         ,        ,
                                                                        ,
                                                                                                                 ,            3
c IW" = .
                         '        '       '     '           '                      '   '      '                  '         '

0.0  ! I I , 20 40 60 80 100 120 140 10 30 50 70 90 110 130 j Distance from Bottom of Active Fuel, in. l l l ! AXlAL LOCAL TO AVERAGE BURNUP AND tO l ! INSTANTANEOUS POWER COMP ARISO'45 QOh # CRYSTAL RIVER UNITS 3 & 4 m)=-giouRE352

                           ~

l

                                                                                                                                       .--

_ _ _ . ,_. __ _ . .

                                                                                     .

O

 /

50 Design Limit 40 m

             '8
               .
                                                      /        /
                                                                   /
                  *
             -

0 1.70 BU and / // / a: Axial Shape. #

             *
              $       l.50 BU and Axial Shape j f

[

             .g  20   930 Day BU and 1.70 Axi l-C        Shape.

O 10 0 0 1 2 3 4 5 6 7 8 Cold Diametral Clearance, in. x 10 ' 000003i$L

 ,A   '
       -

FIS$10N GAS RELEASE FOR 1.50 AND k *t i i 1.70 MAX / AVG AXlAL POWER SHAPES

           '

CRYSTAL RIVER UNITS 3 & 4

                                                               ~

h FIGURE 3-53

    -
                                                                                               .                                                 _              .  . .

3500 Design Limit O 3000 l losed Pores

                             .

2500 l \ '

                                                                                                                                                     '
                                                                                                                                                       /Y

__

                             ,
                             .-
                            $           2000              1.5 Axial Power 5                             and Burnup Shape.
                             ._
-

2 1500 -- -

                               !                                            /

2 1 7 Axial Power npS pe. Open Pores a

                                                                                         '

l l 500 l l l l

                                                                                                                                                              '

! O l 2 3 4 5 6 7 8 Cold Diametral Clearance, in. x 10

                      .

GAS PRESSURE INSIDE THE FUEL CLAD FOR VARIOUS AXIAL BURNUP AND POWER SHAPES CR'YSTAL RIVER IJNITS. 3 & 4 5\b h=- - eGuRE3 54 @

                         . --
 ---w-e-   , , . ,-wn-w,    --w-o g- an  , . - , , -         -,,    -c,noe   , , - - , - - ,     ,.--,,m  ,--,,,.,,,,,,,maaw------,-,,--.--,-
                                                                                                                             -
                                                                                                                                                                       ,m.s,-
                                -

O ( l l

     ---
r: 1. 3.4 .9-9 . . ;e . ): . .e .
                                                                                                 --
                                                                                                                 .:    ---
               .. :e     -.003    , ,. . . : .,   C.)s-               :i-                    .Li'
             -d3                     =       d3-e                          =       s
                            )           ..x.                                  ...:2                             :
                                                  ' ' ~ ~                                   *I I                5
                                                                   ?

6 ebe

 '

14s

              - ?~E
                                         =      43                                  .. :8           1. ".12 - 1.015
                           -

9 9 2.9~: 9~ . 01-

                                                                                    ..:12           . . f.6      . 016
              . ,.
              ..                                 .. ...         : . 3,e-       .m         ..
                                                                                          .....              ...e n; ear restir.g ra:::r I                                  Er.thalpy Fise Fs:*.:r 00000ig
                   '
           '

A NOMINAL FUEL ROD POWER AKS AND kb CELL EXIT ENTHALPY RISE RATIOS l

                     '

CRYSTAL RIVER UNITS 3 & 4 E FIGURE 3 55

   .

_ . - -

__ _ _ _ _ _ . _ _ . _ . . . _ . . _ _ . _ . . _ _ _ . _ . _ . . _ . _ _ _ _ _ _ _ _ . _ _ _ _ . ~ _ _ . _ _ . _ . . . . . _ . . _ _ _ _ _ _ _ _ . _ . . - _ _ 99-C 3HOOld -

                                                                                                                                                                                                   -

O s

                                               \
                                                         .>                                                                                                                  F T C SilNO H3AlH 1Y1SAH3 Sol 1VM ESl3 Ad lVH1H3 ilX31133 QNY SXY3dl3kod 00H 190d WnWlXYW l

Jou ng asTd Ai'9 7^.*J3 gg : - g J.. .-- -

                                                                                        '
                                                                                         /                                               = , u ., w .d                    . c.                      ,I        l
-6C ' ' "" 2 " 2?C " ~C"  ?:  ??:'I Gio'- ':*-

ge-* ;6C'~  !?:*-

                                                 ~*C'T                E66*'                     9 t'                        ?E 6
  • 66*: CT0 " 096*" *6*'
                                                                                                                                                                                                -
                                             "           =~~                  "='~

O  ??:" Eid*C C0d

                                                                                                 ?E6':                      .0-*-                     E20..

Ed '66*: dif': L*C' 56* @ :: " 1.20** 900 " 696*: 6*6*

.e: - .ee c te: - te -  ;:0 :  ::0 1 ;96 C 6-6 :
                                                 ?? ** ,              ?!C "                                                90C*!                      t0C'*                   I      ?96*C 9:C*T               966*:                      :66*                      696 0                        '96*C              9s6*      is6*C      96*:
                       --                         2'C'                   6*:                   6M*: a                     6*6*:                         6-6 *(           096*:         s6':   J.2C     __

O , I

                                                                                                                                       ,

l l b a /

                *
                             \                                 l              l      l 1.3                                          G            l.59 x 106 l b/ hr-f t2 N    \s 1.2              \
                                  \                       Best Fit l.I                    3
                                         \

1.0 o Design Limit x 0.9 g i

            ;  0.8                          ,
                                                          \

x

                                                                 \

7*' Minimum DNBR 2.20 N

            **

0.7 \

                                                                               \

d \ C ' 0.6

            %                                                                      \

S 0.5 \ i j g

                                                                                                      \

3 0.4 g

                      /                       \                                                           x.g Calculated 0.3 0.2 0.1 580 600   620    640    660        680   700              720  740                  760   780     800 Local Enthalpy, Btu /lb a o o 5000

_

                    ,

CALCULATED AND DESIGN LIMIT LOCAL HEAT l',, FLUX VS ENTHALPY IN THE HOT CORNER /~ 33'i (",T/

                       ~
                         ,

CELL AT THE HOMINAL CONDITION CRYSTAL RIVER UNITS 3 & 4

                                                                                                -
                                                                                                "
                                                                                                -               FIGURE 3 57 AMEND.1 (1*15 68)
     . _ .                                      ..__ -      .. . _ . . - _ .           . . _ , . _ . _                 . . _ . _ .   .
                                                                                                                        '

l.4 " 'G O i i i i i

                                            '
1. 3 G 1.32 x 106 l b/ h r-f t2 _.

N l.2 1.1

                                                    \              8est Fit
                ,       1.0                           \       T x    0.9            Design Limit           g N                                               \

U \s \ L 0.8 e s

                  $
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