ML19322A782

Chapter 3 of Oconee 1,2 & 3 PSAR, Reactor. Includes Revisions 1-6

ML19322A782
Person / Time
Site:	Oconee
Issue date:	12/01/1966
From:	DUKE POWER CO.
To:
References
NUDOCS 7911250001
Download: ML19322A782 (150)

Text

. . . -

O N

\,

O =

a O

=

J O

7911 250 OO2 -

mm em,nn UU uusuw

/4Lx __

d

TABLE OF COITfENTS Section P_ age 3 REACTOR 3-1 31 DESIGN BASES 3-1 3 1.1 PERFOINANCE OBJETIVES 3-1 3 1.2 LIMITS 3-1 31.2.1 Nuclear Limits 3-1 3 1.2.2 Reactivity Control Limits 3-2 31.23 Thermal and Hydraulic Limits 3-2 31.2.4 Mechanical Limits 3-3 32 REACTOR DESIGN 3-6 32.1 GENERAL SUM 4ARY 3-6 O 3 2.2 NUCLEAR DESIGN AND EVAWATION O 3-6 32.2.1 Nuclear Characteristics of the Design 3-6 3 2.2.2 Nuclear Evaluation 3-18 323 THEPMAL AND HYDRAULIC DESIGN AND EVAWATION 3-25 3231 Thermal and Hydraulic Characteristics 3-25 3232 Themal and Hydraulic Evaluation 3-33 32.4 14ECHANICAL DESIGN IAYOUT 3 49 3 2.4.1 Internal Layout 3 49 32.4.2 Fuel Asse- .ies 3-54 32.4.3 Centrol System 3-65 33 TESTS AND INSPECTIONS 3-$ 7 5~

331 NUCIEAR TESTS AND INSPECTION 3-W 75' 3 3 1.1 Critical Experiments 3 'M 75 3 3 1.2 O\ Zero Power, Approach to Power, and Power Testins 3-78 2*

/

mn 1 vu nntni vviv, 3-1 14 5 J

CONTENTS (Cont'd)

Section Page 3.3.2 THERMAL AND HYDRAULIC TESTS AND INSPECTION 3-76 3.3.2.1 Reactor Vessel Flow Distribution and Pressure Drop Test 3-76 3.3.2.2 Fuel Assembly Heat Transfer and Fluid Flow Tests 3-76 3.3.2.3 Preoperational Testing and Postoperational Testing 3-77 3.3.3 FUEL ASSEMBLY, CONTROL CLUSTER ASSEMBLY, AND CONTROL ROD DRIVE ASSEMBLY MECHANICAL TESTS AND INSPECTION 3-77 3.3.3.1 Prototype Testing 3-78 3.3.3.2 Model Testing 3-78 3.3.3.3 Component and/or Material Testing 3-78 3.3.3.4 Control Rod Drive Assembly Tests and Inspection 3-94L 76 3.3.4 INTERNALS TESTS AND INSPECTIONS 3-83

3.4 REFERENCES

3-84 4

O 00 00102

=

3-11 (4-29-67)

)

LIST OF TABLES

[^

G '))

Table No. TJty Page 3-1 Core Design, The'. ::al, and Hydraulic Data 3-6 3-2 Nuclear Design Data 3-8 3-3 Excess Reactivity Conditions 3-9 3-4 First Cycle Reactivity Control Distribution 3-10 3-5 Initial Cycle Reactivity Shutdown Analysis 3-12 3-6 Soluble Boron Levels and Worth 3-13 3-7 Exterior Neutron Levels and Spectra 3-16 3-8 Reference Core Parameters 3-23 3-9 First Mode Threshold Dimension and Flatness 3-23

'N 3-10 Threshold Ratio c. ' Power Flatness 3-24 (O 3-11 Coefficients of variation '

3-28 3-12 DUB Results - Maximum Design Condition 3-30 3-13 DUB Results - Most Probable Condition 3-31 3-14 Heat Transfer Test Data 3-39 3-15 Comparison of hs it Transfer Test Data 3-43 3-16 Hot Channel Coolant Conditions 3-4 3-17 clad Circumferential Stresses 3-58 3-18 LRD Fuel Swelling Irradiation Program 3-63 1 3-19 Control hod Drive Design Data 3-G 6 7 3-20 Control Cluster Assembly Design Data 3-W 7 3 l

l LJ a

An n A 1 O 7-3-111 t  ;

n FIGURES (Cont'd)

Figure No. Title 3-44 Fuel Center Temperature for Beginning-of-Life Conditions 3-45 Fuel Center Temperature for End-of-Life Conditions 3-46 Reactor Vessel and Internals General Arrangement 3-47 Reactor Vessel and Internals General Arr.:ngement - Cross Section 3-48 Core Flooding Arrangement 3-49 Fuel Assembly 3-50 orifice Cluster Assembly 3-51 Control Rod Drive - General Arrangement 3-52 Control Rod Drive - Vertical Section 3-53 (Deleted by 4-29-67 Amendment) 3-54 (Deleted by 4-29-67 Amendment) 3-55 (Deleted by 4-29-67 Amendment) 3-56 (Deleted by 4-29-67 Amendment) 3-57 Drive Mechanism Control Block Diagram 3-58 Limit Signal and Position Indication System 3-59 Reactor Trip Circuit 3-60 Control Cluster Assembly i

O' s c 3-vi (Revised 4-29-67) 0000l46 L

[] FIGURES (Cont'd)

'O Figure No. Title 3-20 Burnout Factor Versus Population for Various Confidence Levels 3-21 Rods in Jeopardy Versus Power 3-22 Ratio of Experimental to Calculated Burnout Heat Flux 3-23 Ratio of Experimental to Calculated Burnout Heat Flux 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 V 3-30 Ratio of Experimental to Calculated Burnout Heat Flux 3-31 Ratio of Experimental to Calculated Burnout Heat Flux 3-32 Ratio of Experimental to Calculated Burnout Heat Flux 3 33 Ratio of Experimental to Calculated Burnout Heat Flux 3-34 Ratio of Experi= ental 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 3-37 Ratio of Experimental to Calculated Burnout Heat Flux l

3-36 Maximum Hot Channel Exit Quality Versus Reactor Power 3-39 Hottest Design and Nominal Channel Exit Quality Versus Reactor Power (Without Engineering Hot Channel Factors) 3-40 Flow Regime Map at 2185 psis 3-41 Hot Channel DUB Ratio Comparison f) 3-42 Reactor Coolant Flov Versus Power ,

v I 3-43 Ther=al Conductivity of 95fo Dense sintered vo2 h 3-v 0000ly(

FIGURES (Cont'd)

Figure No. Title 3-44 Fuel Center Temperature for Beginning-of-Life Conditions 3-45 Fuel Center Temperature for End-of-Life Conditions 3-46 Reactor Vessel and Internals General Arrangement 3-47 Reactor Vessel and Internals General Arrangement - Cross Section 3-48 Core Flooding Arrangement 3-49 Fuel Assembly 3-50 Orifice Cluster Assembly 3-51 Control Rod Drive 3-52 Position Indicator Transmitter Assembly 3-53 Nutating Disc Actuator O 3-54 Spool Piece and Iower Guide Tube Assembly 3-55 Snubber 3-56 Cap and Drive Line Vent Assembly 3-57 Drive Mechanism Control Block Diagram 3-58 Limit Signal and Position Indication System j 3-59 Reactor Trip Circuit 3-60 Control Cluster Assembly

~

00 00l@{)

3-v3

pg 3 REACTOR v 31 DESIGN BASES The reactor is designed to meet the performance objectives specified in 3 1.1 without exceeding the limits of design and operation specified in 31.2.

3 1.1 PERFORMANCE OBJECTIVES l

The reactor is designed to operate at 2,1&52 mwt with sufficient design margins to accomodate transient operation and instrument error without damage to the core and without exceeding the pressure at the safety valve settings in the i reactor coolant system.

1 The fuel rod c3 adding is designed to maintain its integrity for the anticipated core life. The effects of gas release, fuel dimensional changes, and corrosion-  !

or irradiation-induced changes in the mechanical properties of cladding are j considered in the design of fuel assemblies, i Reactivity is controlled by control cluster assemblies (CCA) and chemical poi-son dissolved in the coolant. Sufficient CCA worth is available to shut the reactor down (keff 5 0 99) in the hot condition at any time during the life  !

cycle with the most reactive CCA stuck in the fully withdrawn position. Re-dundant equipment is provided to add soluble poison to the reactor coolant to insure a similar shutdown capability when the reactor coolant is cooled to am-bient temperatures.

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

I 3 1.2 IDETS l

31.2.1 nuclear Limits The core has been designed to the following nuclear limits:

a. Fuel has been designed for an average burnup of 28,200 WD/ICU and for a maximum burnup of 55,000 WDhEU.

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

c. Control systems will be available to handle core xenon instabilities should they occur during operation without jeopardizing the safety conditions of the system.

d. The core will have sufficient excess reactivity to produce the de- i sign power level and lifetime, without exceeding the control capacity l

/' or shutdown margin.

J j 00 0015 i 3-1

e. Controlled reactivity insertion rates have been limited to: 5.8 x l 5 a see for a single regulating CCA group withdrawal, and 7 x 10 j' a see for soluble boron removal.

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

The low worth of CCA 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 will provide adequate control of the core reactivity and power distribution. The following control limits will be met:

a. Sufficient control will be available to produce a shutdown margin of at least 1% a k/k.

b. The shutdown cargin will bc maintained with the CCA of highest worth stuck out of the core,

c. CCA withdrawal limits the reactivity insertion to 5 8 x 10-56k/k/see on a single regulating group.

reactivityinsertionof7x10gorondilutionisalsolimitedtoa Ak/k/sec.

31.23 Thermal and Hydraulic Limits The reactor core is designed to meet the following limiting themal and hydrau-lic conditions:

a. No central melting at the design overpower (114 per cent).

b. A 99 per cent confidence that at least 99 5 per cent of the fuel rods in the core are in no jeopardy of experiencing a departure from nu-cleate boiling (DNB) during continuous operation at the design over-power.

c. Essentially 100 per cent confidence that at least 99 96 per cent of the fuel rods in the core are in no jeopardy of experiencing a DN3 during continuous operation at rated power.

d. The generation of net steam in the hottest ? ,N channels is permis-sible, but steam voids will be low enough to , 'avent flow instabili-ties.

The design overpower is the highest credible reactor operating power permitted by the safety system. Normal overpower to trip is significantly less than the j design overpower. Rated power (2,452 mwt) is the licensed operating power.

O 6

. ~(

O j U 3-2

3.1.2.4 Mechanical Limits 3.1.2.4.1 Reactor Internals The reactor internal components are designed to withstand the stresses result-ing from startup; steady state operation with one, two, three, or four 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 l be designed within the allowable stress levels permitted by the ASME Code, j Section III, Table N-421. Structural integrity and allowable stress levels of all core support assembly welds will be assured by compliance with ASME Code Sections III and IX, radiographic inspection acceptance standards, and welding qualifications.

The reactor internals will be designed as a Class I structure as defined in Appendix SA of this report. In addition, the internal structural components j will be designed to resist the effects of seismic disturbances. The basic de-l sign guide for the seismic, analysis will be AEC publication TID-7024, " Nuclear i i Reactors and Earthquakes."

Lateral deflection and torsional rotation of the lower end of the core barrel will be limited to prevent excessive movements resulting from seismic distur-bances and thus prevent interference with control cluster assemblies (CCA's).

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

I The structural internals will be designed to maintain their functional integ-rity in the event of a major loss-of-coolanti accident as described in 3.2.4.1.

j The dynamic loading resulting from the pressure oscillations because of a loss-j of-coolant accident will not prevent CCA insertion.

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

I 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 thennal strain. The cold-worked Zircaloy-4 cladding is designed to be free-standing.

Fuel rod assemblies are held in place by mechanical spacers that are designed

to maintain dimensional control o'f the fuel rod spacing throughout the design life, without impairing cladding integrity. Contact loads are limited to pre-

. vent fretting.

The spacers are also designed to permit differential thermal expansion of the fuel rods without restraint that would cause distortion of the rods. The fuel assembly upper end fitting.and the guide tube in the internals structure are l O~ both indexed to the grid plate above the fuel assemblies, thus insuring con-  ;

tinuous alignment of the guide channels for the CCA's. The_ control pin I

e i

3-3 00 00153 h

1 l

l l

travel is designed so that the pins are always enga6ed in the fuel assembly ,

guide tubes, thus insuring that CCA's can always be inserted. The assembly I structure is also designed to withstand handling loads, shipping loads, and earthquake loads. I Stress and strain for all anticipated nor=al and abnormal operatin6 conditions will be limited as follows:

a. Stresses which are not relieved by small deformations of the material vill be prevented from leading to failure by not permitting these stresses to exceed the yield strength of the materia? nor to exceed levels which would use in excess of 75 per cent of t,ne stress rup-ture life of the material. An example of this type of stress is the circumferential me=brane stress in the clad due to internal or ex-ternal pressure.

b. Stresses which are relieved by small deformations of the material, and the single occurrence of which will not make a significant con-tribution to the possibility of a failure, will be pemitted to ex-ceed the yield strength of the material. Where such stresses ex-ceed the material yield utren6t h, strain limits will be set, based on low cycle ftti6ue techniques, using no more than 90 per cent of the material fatigue life. Evaluations of cyclic loadings will be based on conservative estimates of the number of cycles to be expe-rienced. An example of this type of stress is the thermal stress resulting from the thermal gradient across the clad thickness.

c. Combinations of the above two types of stresses, in addition to the individual treatment outlined above, will be evaluated on the low-cycle fatigue basis of Ites b. Also, clad plastic strain due to diameter increases resultin6 fromther=alratchetingand/orcreep, including the effects of internal gas pressure and fuel swelling, will be limited to about 1 per cent.

d. Minhm clad collapse pressure margins will be required, as listed below:

(1) 10 per cent mar 81 n over system desi6n pressure, on short time collapse, at end void.

(2) End void must not collapse (must be either freestanding or have adequate support) on a lon6 time basis.

i l

(3) lo per cent margin over system operating pressure, on short time collapse, at hot spot average temperature through the clad wall.

(h) Clad must be freestanding at design pressure on a short time basis at ~ 725 F hot spot average temperature through the clad wall.

O C

e 34 jhb\ .

3.1.2.4.3 Control Cluster Assembly (CCA)

The control pin clad is designed to the sams criteria as the fuel clad, as applicable. Adequate clearance will be provided between the control pins and the guide tubes, which position them within the fuel assembly, so that control pin overheating will be avoided, and so that unacceptable mechanical interfer-ence between the control pin and the guide tube will not occur under all op-erating conditions, including earthquake.

Overstressing of the CCA components during a trip will be prevented by mini-mizing the shock loads by snubbing and by providing adequate strength.

3.1.2.4.4 Control Rod Drive Assembly Each control rod drive assembly is provided with a pressure breakdown seal to prevent leakage of reactor coolant water external to the reactor coolant sys-tem. All pressure-containing components are designed to meet the requirements of the ASME Code,Section III, Nuclear Vessels for Class A vessels.

The control rod drive assemblies provide control cluster assembly (CCA) inser-tion and withdrawal rates consistent with the required reactivity changes fc:

reactor operational load changes. This rate is based on the worths of the var-ious rod groups, which have been established to limit power-peaking flux pat-terns to design values. The maximum reactivity addition rate is specified to limit the magnitude of a possible nuclear excursion resulting from a control O system or operator malfunction. The normal insertion and withdrawal velocity has been established as 25 in./ min.

The control rod drive assemblies provide a " trip" of the CCA's which results in a rapid shutdown of the reactor for conditions that cannot be handled by the reactor control system. The trip is based on the results of various re-actor emergency analyses, including instrument and control delay times and the amount of reactivity that must be inserted before deceleration of the CCA occurs. The maximum trip time for a 2/3 insertion of a CCA has been estab-lished as 1.4 sec.

The control rod drive assemblies can be coupled and uncoupled to their respec-tive CCA's without any withdrawal movement of the CCA's.

Materials selected for the control rod drive assembly are capable of operating within the specified reactor environment fo; the life of the mechanism without any deleterious effects. Adequate clearance will be provided between the sta-tionary and moving parts of the control rod drive assemblies so that the CCA trip time to full insertion will not be adversely affected by mechanical in-terference under all operating conditions and seismic disturbances.

Structural integrity and adherence to allowable strtss limits of the control rod drive line and related parts during a trip will be achieved by establish-ing a limit on impact loads through enubbing.

O ~

An Ani nO _

uv vvTvi 3-5 (Revised 4-29-67)

32 REAC'IOR DESIGN O

3 2.1 GEIERAL

SUMMARY

The important core design, thermal, and hydraulic characteristics are tabulated in Table 3-1. The nuclear design characteristics are presented in Table 3-2.

3 2.2 NUCLEAR DESIGN AND EVALUATION The basic design of the core satisfies the following requirements:

a. Sufficient excess reactivity is provided to achieve the design power level over the specified fuel cycle.

b. Sufficient reactivity control is provided to permit safe reactor operation and shutdown at all times during core lifetime.

3 2.2.1 Nuclear Characteristics of the Design 3.2.2.1.1 Excess Reactivity The excess reactivities associated with various core conditions are tabulated in Table 3-3 Data are shown for two-Unit fuel sharing. Unit 1 of the two-Unit site will operate for one cycle of 272 full power days. Fuel from Unit 1 will then be loaded into Unit 2, and both Units will be operated on an equilib-rium cycle of 310 full power days. The reactivity data shown in Table 3-3 as- '

suces that 50 per cent of the fuel from Unit 1 is shared with Unit 2. The safe operation of the two Units will not be jeopardized by fuel sharing. De-sign limits will be held with respect to reactivity control and power distri-bution. Incore instrumentation will be used to back up the analytical selec-tion of the particular fuel assemblies to be shared and thereby insure proper power peaking levels.

Table 3-1 Core Design, Thermal, and Hydraulic Data Reactor Type Pressurized Water Rated Heat Output, twt 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 ,

TotalNumbeihfjFue[AssembliesinCore 177

3-6 lN

i D

l e

Table 3-1 (Cont'd)

Number of Fuel Rods per Fuel Assembly 208 Number of Control Pins per Control Rod Cluster Assembly 16 Number of Incore Instrumentation Positions per Fuel Assembly 1 Fuel Rod Outside Diameter, in. 0.420 Clad Thickness, in. 0.026 Fuel Rod Pitch, in. 0 558 Fuel Assembly Pitch Spacing, in. 8 587 Unit Cell Metal / Water Ratio 0.80 Clad Material Zircaloy-4 (cold-worked)

~

Fuel Material UO2 Form Dished-end, cylindrical pellets Diameter, in. 0 362 Active Iength, in. 144 Density, % of theoretical 95

/3 Q Heat Transfer and Fluid Flow at Rated Power Total Heat Transfer Surface in Core, ft 48,578 Average Heat Flux, Btu /hr-ft 167,620 MaximumHeatFlux, Btu /hr-ft 543,000 AveragePowerDensityinCore,kw/1 79 60

/.verage Thermal Output, kw/ft of fuel rod 5.4 MaximumThermalOutput,kw/ftoffuelrod 17.49 Maximum Clad Surface Temperature, F 654 Average Core Fuel Temperature, F 1,385 Maximum Fuel Central Temperature at Hot Spot, F 4,160 TotalReactorCoolantFlow,lb/hr 131 32 x 10 6

Core Flow Area (effective for heat transfer), ft2 47,75 Core Coolant Average Velocity, fps 15 7 Coolant Outlet Temperature at Hot Channel, F 644.4 Power Distribution Maximum / Average Power Ratio, radial x local (F nuclear) 1.85 g

Maximum /Ar-arage Power Ratio, axial (F nuclear) 1 70 Overall Power Ratio (F nuclear) g 3 15 UV vvisv

Table 3-1 (Cont'd)

Power Generated in Fuel and Cladding, 5 97 3 Hot Channel Factors Power Peaking Factor (Fq ) 1.017 Flow Area Reduction Factor (FA ) 0 983 Local Heat Flux Factor (Ff' ) 1.030 Hot Spot Maximum / Average Heat Flux Ratio (F nue. and mch. ) 3 24 DIG Data Design Overpower Ratio 1.14 DNB Ratio at Design Overpower (BAW-168) 1 38 DNB Ratio at Rated Power (BAW-168) 1.60 Table 3-2 Nuclear Design Data Fuel Elemnt Volum Fractions Fuel 0.285 Moderator 0 590 Zircaloy 0.099 Stainless Steel O.011 Void 0.015 1.000 Total UOa, metric tons 91.61 Core Dimnsions, in.

Equivalent Diamter 128 9 Active Height 144.0 Unit Cell H 2O to U Atomic Ratio Cold 2 97 Hot 2.13 Full Power Lifetim, days Unit 1 Unit 2 First Cycle 272 310 Each Succeeding Cycle 310 310 0 -

, 3.r'

'~

3-8 'g

p Table 3-2 (Cont'd)

V Fuel Irradiation, MID/MIU Unit 1 Unit 2 First Cycle Average 8,260 9,410 Succeeding Oycle Average 9,410 9,410 FeedEnrichments,w/oU-235 First Cycle 2.24/2.47/277 2.47(*)

(by initial zone)

Control Data Control Pin Material Cd-In-Ag Number of Control Cluster Assemblies 69 Total Rod Worth (Ak/k), % 10 Control Pin Cladding Material Type 304 SS

• Average feed enrichment.

Single fuel assembly reactivity information is also included in Table 3-3 Table 3-3 Excess Reactivity Conditions Effective Multiplication - BOL

• Unit 1 Unit 2 Cold, Zero Power 1 312 1.255 Hot, Zero Power 1.258 1.201 Hot, Rated Power 1.242 1.184 Hot, Equilibrium Xe, Rated Power 1.205 1.154 Hot, Equilibrium Xe and Sm, Rated Power ( 1.167 1.119 SingleFuelAssembly(b)

Hot 0 77 Cold (c) o,g7

(" BOL - Beginning-of-Life.

Based on highest probable enrichment of 3 5 weight /per cent.

( A center-to-center assembly pitch of 21 in. is required for this O k*f*, in cold nonborated water with no xenon or samarium.

(d) Includes burnup until equilibrium samarium is reached.

an vv nn***

vvi1I 3-9 l 5 c'

The =inimum critical mass weight, with and without xenon and sa:::arium poison-ing, may be specified in a variety of for=s, i.e., single assembly, multiple assemblics in various geometric arrays, da= aged or crushed assetiolies, etc.

The unit fuel asse=bly has been investigated for comparative purposes. A single cold, clean assembly containing a maximum probable enrichment of 3 5 veight/per cent is suberitical. Two assemblies side-by-side are supercritical except when both equilibrium Xe and Sm are present. Three assemblies side-by-side are supercritical with both equilibrium Xe and Sm present.

3 2.2.1.2 Reactivity Control Distribution Control of excess reactivity for both Unit 1 and Unit 2 is shown in Table 3 4.

Table 3-4 First Cycle Reactivity Control Distribution Controlled by Solus e Enron - % Ak/k Unit 1 Unit 2

1. Moderator Temperature Deficit (70 to 520 F) 33 36

2. Equilibrium Xenon and Samarium 25 2.4 3 Fuel Burnup and Fission Froduct Buildup 16.0 12 3 Total Soluble Boron Worth Required 21.8 18 3 Controlled by Control Cluster Assemblies - % ak/k

1. Doppler Deficit (0 to 100% power) 1.0 1.2

2. Transient Xenon 1.4 1.4 3 Equilibrium Xenon 1.0 0.8

4. Moderator Temperature Deficit (0 to 15% power at end-of-life) 0.6 0.6 5 Dilution Control 0.2 0.2

6. Shutdown Margin 1.0 1.0 Total CCA Worth Required 52 52 Available Cluster Control Assembly Worth - 4 ok/k

1. Total Movable CCA Worth 10 10

2. Stuck CCA Worth (CCA of hi6 hest reactivity value) (-)3 (-)3 Minimum CCA Worth Available 7 7 Boron in solution is used to control the following relatively slow moving re-activity changes:

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

The value shown is for the maximum change which would occur toward '

the end of the cycle.

,a 3-10 h

b. Equilibrium samarium and a part (approximately 15% ak/k) of the equilibrium xenon.

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

Figure 3-1 shows the typical variation in boron concentration with cycle life for Unit 1 and Unit 2.

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

a. Power-level changes (Doppler) and the resultant short-term xenon transients are considered here. Sufficient rod worth remains in-serted in the core during normal operation to overcome the peak xenon

- transient following a power reduction. This override capability fa-cilitates the return to normal operating conditions without extended delays. The presence of these rods in the core during operation does not produce power peaks above the design value, and the shut-down margin of the core is not adversely affected. A typical con-tour drawing of relative power distribution with the xenon override rods inserted is shown in Figurt 3-2. All isopower lines are taken relative to the highest power peak. Axial power peak variation, re-sulting from partial or full insertion of xenon override rods, is described fully in Figures 3-3 and 3-4.

b. The portion of the equilibrium xenon not controlled by soluble boron, approximately1%ak/k,isheldbymovableCCA's.

c. Between zero and 15 per cent of full power, reactivity compensation by CCA's may be required as a result of the linear increase of reac-tor coolant temperature from 520 F to the normal operating value.

d. Additional reactivity is held by a group of partially inserted CCA's (25 per cent insertion maximum) to allow periodic rather than con-tinuous soluble boron dilution. The CCA's are inserted to the 25 per cent limit as the boron is diluted. Automatic withdrawal of these CCA's during operation is allowed to the 5 per cent insertion limit where the dilution procedure is again initiated and this group of 1 CCA's is reinserted.

e. A shutdown margin of 1% ok/k to the hot critical condition is also required as part of the reactivity controlled by CCA's.

A total of 5 2% Ak/k is required in movable control. Analysis of the 69 CCA's under the reference fuel arrangement predicts a total CCA worth of at least 10%ak/k. The stuck-out CCA worth was also evaluated at a value no larger than 3 0% Ak/k. This evaluation included selection of the highest worth CCA under j the first CCA-cut condition.

p The minimum available CCA vorth of 7 0% ak/k is sufficient to meet movable con-( trol requirements.

QA Ant in 3-11 L* v VUi IC

3 2.2.1 3 Reactivity Shutdown Analysis s The ability to shut down the core under both hot and cold conditions is illus-trated in Table 3-5 In this tabulation Unit 1 and Unit 2 are evaluated at the beginning-of-life (BOL) and the end-of-life (EOL) for shutdown capability.

Table 3-5 Initial Cycle Reactivity Shutdown Analysis Unit 1 Unit 2 Reactivity Effects - % ok/k BOL EOL BOL EOL

1. Maximum Shutdown CCA Requirements Doppler (100to0% power) 1.0 15 1.2 15 Equilibrium Xenon (") 25 2.4 2.2 2.1 Moderator Deficit (15 to 0% power) 0.0 0.6 0.0 0.6 Total 35 4.5 3.4 4.2

2. Maximum Available CCA Worth ( -10.0 -10.0 -10.0 -10.0 Transient Xenon CCA Worth 1.4 1.4 1.4 1.4 3 Minimum Available CCA Worth All CCA's In -8.6 -8.6 -8.6 -8.6 One Stuck CCA(c) -5 6 -5.6 -5.6 -5 6

4. Minimum Hot Shutdown Margin All CCA's In -5 1 -4.1 -5 2 -4.4 One Stuck CCA -2.1 -1.1 -2.2 -1.4 5 Hot-to-Cold Reactivity Changes (d)

All CCA's In +0 5 +3 5 +1 5 +6.4 One Stuck CCA +0.1 +2.8 +0.8 +5 8

6. Cold Reactivity Condition (*)

All CCA's In -4.6 -0.6 -3 7 +2.0 One Stuck CCA -2.0 +1 7 -1.4 +4.4 Notes:

Includes 15%Ak/knormallyheldbysolubleboron.

Total worth in 69 CCA's.

CCA of highest reactivity value.

Includes rod worth and all temperature changes.

No boron addition.

\hb 3-12

Table 3-5 (Cont'd)

Unit 1 Unit 2 Reactivity Effects - % ok/k BOL EOL BOL EOL 7 Soluble Boron concentration (ppm)

Addition Required for k ff = 0 99 (cold)

All CCA's In 0 25 0 190 One Stuck CCA 0 175 0 335 Examination of Table 3-5 for Minimum Hot Shutdown Margin (Item 4) shows that, with the highest worth CCA stuck out, the core can be maintained in a suberiti-cal condition. This is true even if the soluble boron concentration is not in-creased to control the reactivity released by equilibrium xenon decay (see Note a ). Normal conditions indicate a minints hot shutdown margin of 4.1% Ak/k at end-of-life (Unit 1 only).

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

Table 3-6 Soluble Boron Ievels and Worth BOL Boron Ievels, ppm Core Conditions Unit 1 Unit 2

1. Cold, keff = 0 99 No CCA's In 1,860 1,600 All CCA's In 1,290 1,030 One Stuck CCA 1,MO 1,200

2. Hot, Zero Power, ke rr = 0 99 No CCA's In 2,150 1,770 All CCA's In 1,150 770 One Stuck CCA 1,450 1,070 3 Hot, Full Power p No CCA's In 1,950 1,550 l V h. Hot, Equilibrium Xe and Sm, Full Power 1,430 1,060 e

3-13 k

Table 3-6 (Cont'd) -

Boron Worths - % Ak/k/ ppm Unit 1 Unit 2 Hot 1/100 1/100 Cold 1/75 1/75 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. This information vill apply to both Units except as noted,

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 r9n;c for the Doppler coefficient under operating conditions is expected to be -1.1 x 10-5 to -1 7 x 10-5 Ak/k/F.

b. Moderator Void Coefficient The moderator void coefficient relates the change in neutron multipli-cation to the presence of voids in the moderator. Cores controlled by appreciable amounts of soluble boron may exhibit a small positive coefficient for very small void levels (ceveral per cent void),

whilehighervoidlevelsproduceincreasinglynegativecogfficients.

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

c. Moderator Pressure Coefficient The moderator pressure coefficient relates the change in moderator density, resulting from a reactor coolant pressure change, to the corresponding effect on neutron production. This coefficient is op-posite in sign and considerably smaller when compared to the modera-tor temperature coefficient. A typi over a life cycle would be -1 x 10 $ cal range to +2 of pressure x 10-6 Ak/k/ psi.coefficients

d. Moderator Temperature Coefficient The moderator temperature coefficient relates a change in neutron multiplication to the change in reactor coolant temperature. Reac-l tors employing soluble boron as a reactivity control have less nega-tive moderator temperature coefficients than cores controlled solely l by movable or fixed CCA's. The major temperature effect on the coolant is a change in density. An increasing coolant temperature l

, ,. 4 3-14

mi produces a decrease in water density and an equal percentage reduc-tion in boron concentration. The concentration change results in a positive reactivity component by reducing the absorption in the cool-ant. The magnitude of this component is proportional to the total reactivity held by soluble boron.

The moderator temperature coefficient has been parameterized for the reference core in terms of boron concentration and reactor coolant temperature. The results of the study are shown in Figures 3-5 and 3-6. Figure 3-5 shows the coefficient variation for ambient and operating temperatures as a function of soluble boron cogeentration.

The operating value ranges from approximately +0.8 x 10- to -1 7 x 10 b ak/k/F. Figure 3-6 shows the moderator temperatum coefficient as a function of temperature, for various poison concent.ations. The coefficients of Unit 2 will be more negative than Unit 1, since the boron concentration levels are considerably lover.

The positive temperature coefficient during the initial portion of Cycle 1 for each reactor will not constitute an operational problem.

The Doppler defi:it represents a much larger reactivity effect in the negative direction, and together with the CCA system response will provide adequa ,e control.

e. pH Coefficierf p)

\

U At the present time 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, Saxton, and Con Edison nave experienced reactivity changes at the time of pH changes, but there is no clear-cut evidence that pH is the direct in-fluencing variable without considering other items such as clad ma-terials, fuel assembly crud deposition, system average temperature, and prior system water chemistry. 1 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 results of applying Saxton ob-servations 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 ppm KOH in solution at the beginning-of-life.

Lifetime bleed dilution to 20 ppm boron will 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 gpm, the cor-responding chan6es in pH are 0.(T(1 pH units per hour for boron dilution and 0.251 pH units per hour for KOH dilution.

tonof0.16%ak/kper ApplyingpHworthvaluesobservedatSg%ak/kseeand103x pH unit, insertion rates are 3 16 x 10 .

} 10-5%ak/ksec,respectively. These insertion rates correspond sy to1.03percentpower/hourand3.4percentpower/hourrespec-tively, which are easily compensated by the operator or the au- <

tomatic control system.

GG 00ii4 3-15 g

3 2.2.1 5 Reactivity Insertion Rates Figure 7-7 displays the integrated rod worth of four overlapping rod banks as a function of distance withdrawn. The indicated groups are those used in the core during power operation.--Using approximately 1.2% ok/k CCA groups and a 25 in./ min drive speed in conjunction with the reactivity response given in Figure 7-7 yields a =aximum reactivity insertion rate of 5 8 x lo S ak/k/s Themaxi=umreactivityinsertionrateforsolubleboronremovalis7x10gc.

ok/k second.

3 2.2.1.6 Power Decay Curves Figure 3-7 displays the beginning-of-life power decay curves for the two least effective CCA worths as outlined in Table 3-5, Item No. 3 The power decay is initiated by the trip release of the CCA's with a 300 msec delay from initiation to start of CCA motion. The time required for 2/3 rod insertion is 1.4 sec.

3 2.2.1 7 Neutron Flux Distribution and Spectrum The neutron flux levels at the core edge and the pressure vessel wall are given in Table 3-7 Table 3-7 Exterior Neutron Ievels and Spectra 2

NeutronFluxIevels,n/cm/sec("}

Flux Interior Wall of Group Core Edge Pressure Vessel 1 0.821 cev to 10 cey 6.0 x 101 ) 3.4 x 10 1.230 key to 0.821 mey 13 10 2 9 0 x 10 7 5 x 10 6.2 x 1013 10 3 0.414 ev to 1.230 kev 5 7 x 10 4 Iess than 0.414 ev 7 1 x 10 5 2.1 x 10 l

l (a)These values include the maximum axial peak-to-average power ratio l of 17 l

The calculations were perforced using The Babcock & Wilcox Company LIFE code (BAW-293, Section 3.6 3) to generate input data for the transport code, TOPIC.

A 4-group edit is obtained from the LIFE output which includes diffusion co-efficients; absorption, removal, and fission cross sections; and the zeroth and first coments of the scattering cross section. TOPIC is an S code designed to solve the 1-di=ensional transport equation in cylindrical coorEinates for up to six groups of neutrons. For the radial and azi=uthal variables a linear c

Q b/

approximation to the transport equation is used; for the polar angle Gauss quadrature is used. Scattering functions are represented by a Iegendre series.

The azimuthal angle can be partitioned into 4 to 10 intervals on the half-space between 0 and r. The number of =esh points in the radial direction is restrict-ed by the number of these intervals. For the core exterior flux calculations, four intergls on the azimuthal were used. This allows the maximum number of mesh points (240) in the "r" direction to describe the shield complex. An op-tion is available to use either equal intervals on the azimuthal angle or equal intervals on the cosine of the angle. Equal intervals 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 between 0 and r.

Results from the above method of calculation have been compared with thermal flux mea ements through an array of iron and water slabs in the LIDO pool reactor. Although this is not a direct comparison with fast neutron mea-surements, it does provide a degree of confidence in the method, since the mag-nitude of the thermal flux in shield regions is governed by fast neutron pene-tration.

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

O The following conservatisms were also incorporated in the calculations:

V

a. Neutron fluxes outside the core are based on a maximum power density of 41 vatts/cc at the outer edge of the core rather than an estimated average of 28 watts /cc over life, resulting in a safety margin of about 45 per cent.

b. A maximum uial power peaking factor of 17 was used. This is about 30 per cent greater than the 13 expected over life.

Uncertainties in the calculations include the followin6:

1. The use of only four neutron 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 h-group data in all regions is computed from a fis-sion 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 will ec..servatively pre-diet a high fast neutron flux at the vessel wall because it underestimates the effectiveness of the thermal shield in reducing the fast flux. In penetration through water, the average energy of the neutrons in the group above 1 mev in-

[ creases above that of a fission spectrum, i.e., the spectrum in this group bj e

3-17 g]

hardens. For neutrons above 1 cov, the non-elastic cross section of iron in- s creases rapidly with energy. Therefo're, the assumption of a fission spectrum to co=pute cross sections in the ther=al shield, and the use of a few-group model to cover the neutron energy spectrum, would underesticate the neutron energy loss in the ther=al shield and the subsequent attenuation by the water between the vessel and ther=al shield. The results from 34-group P3m(3) cal-culations show that reduction of the flux above 1 mov by the thermal shield is about a factor of 4 greater than that computed from the 4-group calculations.

The effect of 3 above is expected to underestimate the flux at the vessel wall.

In calculations at ORNL using the S ntechnique, a comparison between an Sg 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 calcula-tion to be lower than the S12 by about a factor of 2 at maximum. Good agree-ment was obtained between the S12 and coments method calculations.

The above uncertainties indicste that the calculation technique should overesti-mate the fast flux at the reactor vessel wall. However, the comparison with thermal flux data indicates a possible underestimate. Until a better compari-son 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 pre-dict the nyt in the reactor vessel.

3 2.2.2 Nuclear Evaluation Analytical models and the application of these models are discussed in this sec-tion. Core instabilities associated with xenon oscillation are also mentioned, with threshold data evaluated under reference conditions.

3 2.2.2.1 Analytical Models Reactor desjgn calculations are made with a large number of computer 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.

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 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 B&W 1-dimensional depletion package code LIFE. LIFE is a composite of MUFT (Ref. 4), KATE (Ref. 5), RIP, WANDA (Ref. 6), and a depletion routine. Normally the MJFT portion is used with 34 energy groups, an exact treatment of hydrogen, the Greuling-Goertzel approximation for elements of mass less than 10, and Fermi age for all heavier elements.

The KATE portion nor= ally uses a Wigner-Wilkins spectrum. In WANDA, 4 energy groups are utilized. Disadvantage factors for input to the thermal group are calculated with the THERMOS (Ref. 7) code. This

! code set has been shown to give reliable results for a reactivity cal-culation of this type. Recent check calculations on critical experi- c mentshaveastandarddeviationoflessthan05%ok/k.

),yg gh t _

A A 1-dimensional analysis of a geometric arrangement, where there are localized strong absorbers such as CCA's, requires a preliminary 2-dimensional analysis. The required properties of the 1-dimensional system are then matched to the 2-dimensional analysis. In this man-ner, it is possible to analyze the simpler 1-df ceasional 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 reac-tivity coefficients, fuel cycle enrichments, lifetimes, and soluble poison concentrations can be found to improve the initial conditions specified for 2-dimensional analysis.

Two-dimensional reactivity calculations are done with either the PDQ (Ref. 8) or TURBO (Ref. 9) diffusion and/or depletion codes. These codes have mesh limitations on the size of a configuration which can be shown explicitly and are often studied with quarter core symmetry.

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

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

Analysis of this type permits detailed power distribution studies as well as reactivity analysis. The power distribution in a large PWR core which has zone loading cannot be predicted reliably with 1-di-mensional 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-dimensional code, and in some cases 3-dimensional celculatiens cre necessary.

Use of the 2-dimensional pro 6 rams requires the generation of group constants as a function of material composition, power history, and geometry. For regions where diffusion theory is valid, MJFT and KATE with THERMOS disadvantage factors are used to generate epithermal and l ther=al coefficients. This would apply at a distance of a few cean ,

paths from boundaries or discontinuities in the fuel rod lattice. l Discontinuities refer to fuel assembly can, water channels, instru-mentation ports, and CCA guide assemblies. The interfaces between regions of different enrichment are considered to be boundaries as well as the outer limit of the core.

To generate coefficients for regions where diffusion theory is inap-propriate several methods are utilized. The arrangement of structural material, water channels, and adjacent fuel rod rows can be repre- ,

sented well in slab geometry. This problem is analyzed by P3MG (Ref. l

3) which is effective in slab geometry. The coefficients so generated l are utilized in the epithermal energy range. Coefficients for the thermal energy range are generated by a slab TEFMis ::alculation.

The regions adjacent to an interface of materP 1 c f different enrich-ment are also well represented with the P3F a sdt The arrangement of instrumentation ports pd cou31 pin guide assem-blies lends itself to cylindrical geometry. DTF-IV (Ref.10) is quite i g effective in the analysis of this arrangement. Input to PTF-IV is t

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 *

.Dn nne <<

3-19

"" "U' 'U~

l

[ t> C ,

I 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. Th 3-dimansional analysis is accomplished by extending the techniquez of 2-dimensional representation.

b. Control Pin Analysis The analysis of control pins and their guide assemblies requires transport methods similar to techniques mentioned in the previous section. The difference arises in application. Instead of calculat-ing group constants, logarithmic boundary conditions are used. These conditions are derived as a function of material, geometry, and mate-rial condition. The material condition part refers to temperature and power history.

The logarithmic boundary conditions are calculated with DTF-IV with input from GAM and THERMOS or KATE. These conditions are then trans-formed to a PDQ or TURBO representation. This step is necessary be-cause of the differences arising between the transport calculation in cylindrical geometry and the square mesh representation in the 2-dimensional diffusion codes.

This method is at present being verified by comparison with critical experiments containing Ag-In-Cd and boron poison rods. The critical data is reported in References 12, 13, 14, and 15

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 calculated from small perturbations in the required parameter over the ran6e of possible values of the parameter.

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

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

3 2.2.2.2 Codes for Reactor Calculations This action contains a brief description of codes mntioned in the preceding section.,

THERMOS (Ref. 7) - This code solves the integral form of the Boltzmann Transport Equation for the neutron spectrum as a function of posi-tion. A diagonalized connection to the isotropic transfer matrix has been incorporated allowing a degree of anisotropic scattering.

MUFT (Ref. h) - This program solves the F1 or B1 multigroup equation for the first two Ingendre coefficients of the directional neutron flux, and for the isotropic and anisotropic components of the c

'I

- ij: fiU V 3-20

Q b

slowing down densities due to a cosine-shaped neutron source, Coefficients are generated with WFT for the epithermal energy range.

KATE (Ref. 5) - The noce solves the Wigner-Wilkins differential equation for a homogeneous medium moderated by chemically unbound hydro-gen atoms in ther=al equilibrium. Coefficients for the thermnl energy range are generated by KATE.

RIP - This program averages cross sections over an artitrary group struc-ture, calculates resonance integrals for a set of resolved peaks, and computes L-factors for input to MUFT, P1m, and P3MG.

WANDA (Ref. 6) - This code provides numerical solutions of the 1-dimen-  !

sional fev-group neutron diffusion equations. )

LIFE - This is a 1-dimensional depletion package code which is a com-bination of MUFT, KATE, RIP, and WANDA. 'Ihe combinatfon mecha-nizes the procedures for using the codes separately. j GAM (Ref.11) - This code is a multigroup coefficient generation program that solves the P1 equations and includes anisotropic scattering.

Inelastic scattering and resonance paramters are also treated by GAM.

t P3m(Ref. 3) - The code solves the multienergy transport equation in l various geomtries. The code is primarily used for epitherinal v l coefficient generations.

]

DTF (Ref. 10) - This code solves the multigroup, 1-dim nsional Boltzmnnn transport equation by the mthod of discrete ordinates. DTF al-lows multigroup anisotropic scattering as well as up and down scattering.

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

I 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.

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

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

3 2.2.2 3 Stability Analysis The core has been examDed for xenon stability. A modal analysis of axial, azi-

' muthal, and radial spatial xenon oscillations for large pressurized water l H

l l

nn nnies l 3-21 vu uviIr l 17k .

moderated reactors has been completed.(16) The method used in this analysis is '

an extension of the 1-group treat =ent including power coefficient introduced by Randall and St. John. One- and 2-group treata nts have been compared, and the conclusion drawn that a 1-group model is satisfactory for large cores. For all three geometries, data vere generated as a function of:

a. Core size.

b. Flux level,

c. Degree of flatnes; in the power distribution.

d. Power coefficient.

e. Reactivity held by saturation xenon.

In addition, slightly dished power distributions were investigated to show that any dishing resulting from high depletion is not sufficient to require corree-tions to data based on replacing the dished segment with a flat power distribu-tion.

The effect of modal coupling has been examined and shown to be of no conse-quence for cores similar to the reference reactor design. Values of the criti-cal dimnsion varied no more than 1 to 2.8 per cent for the same core with and without modal coupling. The lower value was computed with a zero power coeffi-cient and was not conservative without modal coupling. The higher value was computed with the reference power coefficient and was conservative without modal couplicg.

Table 3-8 sn-m izes those parameters for the reference core which affect the xenon atability threshold. The parameters were calculated at two substantially different times in core life. Reference physical dimensions are also shown for comparison purposes in the following discussion.

Table 3-9 shows the threshold dimnsions for first mode instability as a func-tion of flux flattening. The percentage of flattening is defined as 100 per cent times the ratio of the flattened power distribution to the total physical dimension under consideration. The parameters of Table 3-8 at two days were used since they are virtually the sam as those at 150 days but more conserva-tive. Axial depletion studies show that power distributions are flattened by 0, 63, and 73 per cent at 2, 150, and 354 full power days. A maximum flatness of approximately 80 per :ent may be expected for long core life.

An examination cf the data in Table 3-9 shows that with the maximum flatness axial oscillations are possible, acimuthal oscillations are unlikely, and radial oscillations will not occur.

l Threshold dimnsions for second mode oscillations were 50 per cent larger in magnitude than those shown in Table 3-9 for the first mode. Oscillations in the second mode will not occur in the reference core.

O C

O Q Table 3-8 Reference Core . irameters Two Full 150 Full Power Days Power Days M , em 57 0 57 0 13

(,n/cm-sec 3 9 x 10 3 8 x 1013 a x(reactivityheldbysaturation xenon), Ak/k O.034 0.033 DopplerCoefficient,ak/k/F -1.1 x 10-3 -1.1 x 10~3 Moderator Temperature Coefficient Positive but Small Negative

~

a (power coeff. ), ok/k/ unit flux = -2.2 x 10 =-2 3 x 10 Equivalent Dimensions, ft Height 12.00 Diameter 10 74 Radius 3 37

\_ Table 3-9 First hkxie Threshold Dimensions and Flatness Flatness, %

Threshold Dimensions, ft 0 50 80 Threshold height (axial oscillations) 18 5 14.1 11.8 Threshold disceter (azimuthal oscillation) 20.4 16 5 14.0 Threshold radius (radial oscillation) 16.8 16.7 14.5 Table 3-10 shows the values of H/D versus power flatness for equal likelihood of axial, azimuthal, and radial first harmonic oscillations, i.e., if the core is just at the axial threshold for axial oscillations, it can also be expected thattherewillbeazimuthalandradialoscillationsprovidedthevalueofH/D in Table 3-10 is satisfied. H/D for this reactor is 1.12.

O U

nn m.

nntin

.. iv 3-23

Table 3-10 s Threshold Ratio and Power Flatness Flatness,% Ratio 0 20 50 80 100 H/D(axialversusazimuthal) 0 91 0.87 0.86 0.86 0.85 H/D(axialversusradial) 0 55 0.49 0.42 0.41 0.41 The modal methods used to examine the xenon oscillation problem made use of core-averaged quantities such as flux, power coefficient, and reactivity held by saturation xenon. In addition, flux distributions were limited to:

a. Geometric distributions.

b. Partially or totally flat.

c. Slightly dished.

In view of these limitations, a study is underway to investigate xenon oscilla-tion using nuclear depletion programs coupled with heat transfer equations. The axial heat transfer equations are included in the BE LIFE-5 diffusion de-pletion program and are being used for extensive axial analysis, including con-trol requirements for the suppression of axial oscillations should they occur. Additional analysis will include the study of the first mode azimuthal instabil-ity. The power distribution of Unit 1 during early life is such that no xenon in-stabilities will occur. The power flattening effect of fue.' burnup with time renders the core more susceptible to xenon oscillations. Assuming that oscillation does occur, adequate notification vill be received from incore instrumentation. The long period of the oscillation (approximately 30 hr) provides ample time for system evaluation and the initiation of correc-tive measures. These corrections vould range from CCA adjustments for dampirs out the oscillation to a temporary reduction in core power level. \ O s 4;< 1 3-24 l p(-

    /9  323        THEMAL AND hTDRAULIC DESIGN AND EVALUATION Q'

3231 Thermal and Hydraulic Characteristics 3 2 3 1.1 Fuel Assembly Heat Transfer Design

a. Design Criteria The criterion for heat transfer design is to be safely below Depar-ture from Nucleate Boiling (DND) 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 infor=ation for the statistical core design technique and for the evaluation of individual hot channels consists of the following: (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) ::aximum reactor overpower J These inputs have been derived from test data, physical measurements, and calculations as outlined below.

b. Heat Transfer Equation and Data Correlation The heat transfer relationship used to predict limiting heat transfer conditions is presented in EX4-168.(17) The equation is as follows:

q" = (1.83 - 0.000h15 P) 2 x 90,000 G 0 3987 + 0 001036 aT e3c - 1.027 x 10 6(a Tese) where: q" = critical heat flux as predicted by the best fit form, Btu /hr-ft2  ; 1 P = core operating pressure, psia G = channel mass velocity, lb/hr-ft2 l l S = channel equivalent diameter, ft i L = length up the channel to the point of interest, ft N.J c 3-25 L. L 5 l

AT esc = inlet subcooling (Tsat - Tinlet),F , Tsat = coolant saturation temperature corresponding to P, F This equation was derived from experimental heat transfer data. An analysis of heat transfer data for this and other relationships is described in detail in 3 2 3 2 3, Correlation of Heat Transfer Data. Individual channels are analyzed to determine a DNB ratio, i.e., the ratio of the heat flux at which a D!B is predicted to occur to the heat flux in the channel being investigated. This DIE ratio is re-lated to the data correlation as in Figure 3-8. A confidence and population value is associated with every DNB ratio as described in the Statistical Core Design Technique. The plot of DN3 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 DUB ratio corresponds to a 99 per cent confidence that at least 94.5 per cent of the population of all such hot channels are in no jeopardy of experiencing a DIE. This statement is a corollary to the total core statistical statement given in 31.2 3, Thermal and Hydraulic Limits. The criterion for evaluatin6 the thermal design margin for individual channels or the total core is the confidence-population relationship. The DIE 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: F ah = 1 79 Fz = 1 70 Fq = 3 04 (2) The design nuclear peaking factors for the worst time in core life are: FAh = 1.85 Fz = 1 70 Fq = 3 15 0 - t . *

                                 -                                    j

i Fah= max /avgtotalpowerratio(radialxlocal nuclear) Fz = max / avg axial power ratio (nuclear) Fq = F Ah 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, FA h, from 1 79 to 1.85 to obtain a more conservative design. The axial nuclear factor, Fz, is illustrated in Figure 3-9 The dis-tribution of power expressed as P/P is shown for two conditions of reactor operation. The first condition is an idlet peak with a max / , avg value of 170 resulting from partial insertion of a CCA 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 i the power distribution with xenon override rods withdrawn. The flux peakmax/avgvalueis150inthecenteroftheactivecore. Both of these flux shapes have been evaluated for t'.ermal 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 CJsine peak farther up the channel results in a less fa-O vorable flus to enthalpy relationship. This effegte has been demon-strated in INB tests of nonuniform flux shapes.(1W The 1 5 cosine axial shape has been used to determine individual channel DIB limits and make the associated statistical analysis. The nuclear factor for total radial x local rod power, Fah, is cal-culated for each rod in the core. A distribution curve of the frac-tion of the core fuel rods ojerating above various peaking factors is shown in Figure 3-10. L12 e B shows the distribution of the maxi-mum calculated values of Fa h for ncminal conditions with a maximum value of 1 79 The distribution of peaking factors for the de' sign condition is obtained by inct easing the maximum calculated value for all rods in the core by the ratio of 1.85/179or1.033toprovide conservative results. Determination of the peaking distribution for the design condition in this manner has the effect of increasing re-actor power by about 3 per cent. This assumption is conservative since the distribution with a maximum peak F Ah of 1.85 will follow a line similar to Line C where the average power of all rods in the core is represented by an Fah of 1.0. The actual shape of the dis-tribution curve is dependent upon statistical peaking relationships, CCA positions, moderator conditions, and operating history. The shape of the distribution curve vill be more accurately described i during the detailed' core design. l 1

d. Engineering Hot Channel Factors-Power peaking factors obtained from the nuclear analysis are based on mechanically-perfect fuel assemblies. Engineering hot channel factors are used to describe variations in fuel loading, fuel and +

3-27 'Oh bbi2U

clad dimensions, and flow channel geometry from perfect physical s quantities and di=ensions. The application of hot channel factors is described in detail in 3 2 3 2.2, Statistical Core Design Technique. The factors are de-termined statistically from fuel assembly as-built or specified data where Fq is a heat input factor, Fqvi is a local heat flux factor at s hot spot, and FA is a flow area reduction factor describing the variation in coolant channel flow area. Several subfactors are com-bined statistically to obtain the final values for Fq, F q,,, and FA-These subfactors are shown in Table 3-11. A description of the fac-tor, the coefficient of variation, the standard deviation, and the mean value are tabulated. Table 3-11 Coefficients of Variation CV No. Description e 3E CV 1 Flow Area 0.00075 0.17625 0.00426 2 Incal Rod Diameter 0.000485 0.42 0 0.00116 3 Average Rod Diameter 0.000h85 0.h20 0.00116 (Die-drawn, local and average same) h Local Fuel Loading 0.00687 Subdensity 0.00647 0 95 0.00681 Subfuel area 0.000052 0.1029 0.00089 (Diametereffect) 5 Avera6e Fuel Ioading 0.00370 Subdensity 0.00324 0 95 0.00341 Sublength 0.16181 1h 0.00112 Subfuel area 0.000092 0.1029 0.00089 (Diametereffect) 6 Incal Enrichment 0.00323 2.24 0.001W 7 Average Enrichment 0.00323 2.24 0.001% C'l Coefficientofiariation,e/Tc

               & Standard Deviation of Variable x Mean Value of Variable (Enrichment values are for worst case normal assay baten; maximum variation occurs for minimum enrich =ent.)

c

             , 0 0 0 ":

3-28 g

O e. Core Flow Distribution Hot Channel Factors U The physical arrangement of the reactor vessel internals and nozzles results in a nonuniform distribution of coolant flow to the various fuel assemblies. Reactor internal structures above and below the active core are designed to minimize unfavorable flow distribution. A 1/6 scale model test of the reactor and internals is being per-formed to demonstrate the adequacy of the internal arrangements. The final variations in flow will be determined when the tests are completed. Interim factors for flow distribution effects have teen calculated from reactor vessel model test data for previous pres-surized water reactor designs. A flow factor is determined for each fuel assembly location in the core. The factor is expressed as the ratio of fuel assembly flow to average fuel assembly flow. The finite values of the ratio may be greater or less than 1.0 depending upon the position of the assen-bly being evaluated. The flow in the central fuel assemblies is in general Jarger than the flow in the outermost assemblies due to the inherent flow characteristics of the reactor vessel. The flow distribution factor is related to a particular fuel assem-bly 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 expected in the hottest fuel assembly is five per cent less than average flow. Two assumptions for flow distribu- { s tion have been made in the thermal analysis of the core as fo11cus: (1) For the maximum design condition and for the analysis of the hottest channel, it has been assumed that all fuel assemblies receive five per cent less than average flow, regardless of assembly power or location. l (2) For the most probabic design conditions predicted flow facters have been assigned for each fuel assembly consistent with loca-tion and power. The flow factor assumed for the maxinm design condition is conservative. Application of vessel flow test data and individual assemb1 ficw factors in the detailed core design will result in improve. statistical statements for the maximum design condition.

f. Maximum Reactor Design Overpower Core performance is assessed at the maxinnn design overpower. The selection of the design overpower is based on an analysis of the re-actor protective system as described in Section 7 The reactor trip point is 107 5 per cent power; and the maximum overpower, which is 111+ per cent, will not be exceeded under any conditions.

g. Maximtc Design Conditions Analysis Summary The Statistical Core Design Technique described in 3 2 3 2.2 was used L to analyze the reactor at the maximum design conditions described previously. The total number of fuel rods in the core that have a 9 l

3-29 nn vv nn191 vvi. r (Y

possibility of reaching DIE is shown in Figure 3-n for 100 to n8 per cent overpower. Point A on Line 1 is the maxhmm design point for 114 per cent power with the design F Ah nuclear of 1.85 Line 2 was calculated using the maximum calculated value for Fah nuclear of 179 to show the mrgin between maxi =um calculated and design con-ditions. It is anticipated that detailed core nuclear analyses will permit a lowering of the maximum design value for FA h. The number of fuel rods that may possibly reach a DIS at the maximum design condition with an Fah of 1.85 and at 11h per cent overpower, represented by point A on Figure 3-11, foms the basis for this sta-tistical statement: There is a 99 per cent confidence that at least 99 5 per cent of the fuel rods in the core are in no jeopardy of experiencing a departure from nucleate boiling (DIB) during continuous op-eration at the desi n6 overpower of 114 per cent. Statistical results for the maximum desi n 6 condition calculation shown by Figu e 3-11 may be summarized as follows in Table 3-12. Table 3-12 DNB Results - Maximum Design Condition (99% Confidence Level) Power, Possible Population Point  % of 2,452 mwt FAh DNB's Protected,% A 11h 1.85 184 99.So B 114 1 79 loo 99 73 c loo 1.85 17 99 95 D 100 1 79 lo 99 98 E 118 1 79 184 99 50

h. Most Probable Design Condition Analysis Su=cary The previous maximum design calculation indicates the total nuur of rods that are in jeopardy when it is conservatively assumed taai, every rod in the core has the mechrc.ical and heat transfer character-istics of a hot channel as described in 3 2 3 2.2. For exa=ple, all channels are analyzed with FA (flow area 14ctor) less than 1.0, Fq (heat input factor) greater than 1.0, and with mini =um fuel assembly flow. It is physicany impossible for au channels to have hot chan-nel characteristics. A more realistic indication of the number of fuel rods in jeopardy may be obtained by the application of the sta-tistical heat transfer data to average rod power and mechanical con-ditions.

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 e

    'b' lh 3-30 t

analysis are shown in Figure 3-32. The analysis may be summarized v as follows in Table 3-13 Tabl'e 3-13 DNB Results - Most Probable Condition Power, Possible Population Point  % of 2,452 mwt FAh DIB'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 (DIB) . This result forms the basis for the following statistical statement for the most probable design conditions: There is at least a 99 per cent confidence that at least 99 9 per cent of the rods in the core are in no jeopardy of experi-encing a DIB, 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 shown statistically to be in no jeopardy of a DIS has been calculated for the maximum design and most probable design conditions. The  !

computer programs used provide an output of (N) number of rods and l (P) fraction of rods that will not experience a DIE grouped for ranges of (P). The results for the most probable design condition are shown in Figure 3-13 The population protected, (P), and the population in jeopardy, (1-P), are both plotted. The integral of (1-P) and the number of fuel rods gives the number of rods that are in jeopardy for given conditions j as shown in Figures 3-11 and 3-12. The number of rods is obtained I from the product of the percentage times the total number of rods being considered (36,816). The two distributions shown in Figure 3-13 are for the most probable condition analysis of Points F and G on Figure 3-12. The lower line of Figure 3-13 shows P and (1-P) at the 100 per cent power condition represented by Point F of Figure i 3-12. The upper curve shows P and (1-P) at the 114 per cent power O V' condition represented by Point G of Figure 3-12. The integral of n and (1-P) of the upper curve forms the basis for the statistical g 00191.5&\\Olvv . -- 3-31 1

statement at the most probable design condition described in pars-graph h above.

j. Hot Channel Performance St m=ary The hottest unit ec11 with all surfaces heated has been examined for hot channel factors, DIB ratics, and quality for a range of reactor powers. The cell has been examined for the maximum value of FA h nuclear of 1.85 The hot channel was assumed to be located in a fuel assembly with 95 per cent of the average fuel assembly flow. The heat generated in the fuel is 97 3 per cent of the total nuclear heat. The re=aining 2 7 per cent is assumed to be generated in the coolant as it proceeds up the channel within the core and is reflected as an increase in AT of the coolant.

Error bands of 65 psi operating pressure and +2 F are reflected in the total core and hot channel thermal margin calculations in the direction producing the lowest DIB ratios or highest qualities. The DIO ratio versus power is shown in Fi 6ure 3-14. The DNB ratio in the hot channel at the maximum overpower of 114 per cent is 1 38 which corresponds to a 99 per cent confidence that at least 94 5 per cent of the fuel channels of this type are in no jeopardy of experi-encing a DIO. 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 q., = 1.013 FA= 0 992 The hot channel exit quality for various powers is shown in Figure 3-15 The combined results may be su=marized as follows: Reactor Power, #g DIS Batio (BAW-168) Exit Quality, de 100 1.60 0 107 5 (trip setting) 1.47 2.6 114 (m vi m m power) 1 38 5.4 149 1.00 23 0 3 2 3 1.P. Fuel and Cladding Thermal Conditions

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

The program uses uniform volumetric heat generation across the fuel diameter, and external coolant conditions and heat transfer coef-

ficients determined for thermal-hydraulic channel solutions. The fuel ther=al conductivity is varied in a radial direction as a e

          -      ,3:
               . U ti 4 $h 3-32

f function of the temperature variation. Values for fuel conductivity Q] were used as shown in Figure 3-16, a plot of fuel conductivity versus temperature. The heat transfer from the fuel to the clad,is calcu-lated with a fuel and clad expansion model proportional to tempera-tures. The temperature drop is calculated using gas conductivity at the beginning The Btu-ft/hr-F-ft. 2of-life gasconditions conduction when the gas model conductivity is used is 0.1 in the calculation until.the fuel thermal expansion relative to the clad closes the gap to a dimension equivalent to a contact coefficient. The contact coef-ficient is pressure- and gas conductivity-dependent. A plot of fuel center temperature versus linear heat rate in kw/ft is shown in Figure 3-17 The linear heat rate at the maximum overpower of114percentis199kw/ft. The corresponding center fuel temper-ature is 4,400 F. The following peaking factors were used in the calcuhtion: Fah = 1.85 F3= 1 70 F qi, = 1.03 A conservative value of 1.03 was assumed for the heat flux peaking factor, Fq ,. The assigned value corresponds to a 99 per cent confi-O t dence and 99 99 per cent popuntion-protected relationship as de-scribed in the statistical technique.

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

Boiling conditions prevail at the hot spot, and the Jens and Inttes relationship (19) for the coolant-to-clad A T for boiling was used to determine the clad temperature. The resulting maxi =um calculated clad temperature is 654 F at a system operating pressure of 2,185 psig. 3232 Thermal and Hydraulic Evaluation 3232.1 Introduction Summary results.for the characteristics of the reactor design were presented in 323 1. The statistical Core Design Technique employed in the design rep-resents a refinement in the methods for evaluating pressurized water reactors. Correspondin6 single hot channel DNB data were presented to relate the new method with previous criteria. A comprehensive description of the new tech-nique is included in this section to permit a rapid evaluation of the methods used, p 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 determine the confidence, population, and DNB relationships c 3-33 0000193

included in this section. A comparison of BAW-168 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 DIB ratio, quality, and fuel temperatures. BAW-168 DIB ratios have been compared with W-3 dis ratios to fccilitate a, comparison of the desi6n with PWR reactor core designs previously reviewed. 3232.2 statistical core Design Technique The core themal design is based on a Statistical Core Desi6n Technique devel-oped by B&U. The technique offers many substantial improvements over older methods, particularly in desi6 n approach, reliability of the result, and math-ematical treatment of the calculation. The method reflects the perfomance of the entire core in the resultant power rating and provides insight into the reliability of the calculation. This section discusses the technique in order to provide an understandin6 of its engineering = erit. The statistical core desi 6n technique considers all parameters that affect the safe and reliable operation of the reactor core. By considering each fuel rod the method rates the reactor on the basis of the performance of the entire core. The result then vill provide a good measure of the core safety and reliability since the method provides a statistical statement for the total core. This statement also reflects the conservatism or design cargin in the calculation. A reactor safe operating power has always been detemined by the ability of the - coolant to remove heat from the fuel =aterial. The criterion that best measures this ability is the DIB, which involves the individual parameters of heat flux, coolant temperature rise, and flow area, and their intereffects. The DIB cri-terion is co==only applied throu6h the use of the departure from nucleate boiling ratio (DI3R). This is the mini =um ratio of the DIB heat flux (as computed by the DIB correlation) to the surface heat flux. The ratio is a measure of the =argin between the operating power and the power at which a DIB might be ex-pected to occur in that channel. The DIBR varies over the channel length, and it is the mini =:n value of the ratio in the channel of interest that is used. The calculation of DIB 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 possib]e to separate these two effects; the statistical hot channel factors required are a heat input factor, Fo, and a flow area fac-tor, FA. In addition, a statistical heat flux factor, F q n, is required; the heat flux factor statistically describes the variation in surface heat flux. The DIBR is most limiting when the burnout heat flux is based on minimum flow area (s=all Q and maximum heat input (lar6e Fq), and when the surface heat flux is larGe (large Fqn). The DIB correlation is provided in a best-fit fom, i.e., a fom that best fits all of the data on which the correlation is based. To affoni 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, O. -

            ,  s c       p

lt, .,- 3-34 4 i

V b"IB 1 DIER = 3,7, x f(FA 'EQ )

• qn xy" 3 surface q involves as parameters statistical hot channel and DIB factors. The DIS factor (B.F.) above is usually assigned a value of unity when reporting DIE ratios so that the margin at a given condition is shown directly by a DIER creater than 1.0, i.e., 1 38 in the hot channel.

To find the DIB correlation, selected correlations are compared with DIE 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 DIB heat flux ($ E ) to the calculated value of the burnout heat flux ( $C )" A typical histogram is shown in Figure 3-18. A histogram is obtained for each D'iB 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 . 3-18), mathematical parameters that quantitatively rate the correlation can be I easily obtained for the histogram. These same mathematical parameters are the basis for the statistical burnout factor (B.F.). In analyzing a reactor core, the statistical information required to describe s 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 toler-ances for the proposed core. Regardless of the source of data, the subfactors can be shown graphically (Figures 3-19 and 3-20). All the plots have the same characteristic shape whether they are for subfactors, hot channel factors, or burnout 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 i 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 in-
;    stance, if limited data are available, such an. assumption leads to a somewhat optimistic result. The assumption also implies that more information exists for a given sample than is indicated by, the data; it implies 100 per cent con-fidence in the end result. The B&W calculational procedure does not make this assumption, but rather uses the specified sample size to yield a result that is much more ccaningful and statistically rigorous. The influence of' che aw unt

, of data for instance can be illustrated easily as follows: Consider the heat flux factor which has the form Fn q = 1 + Kcr y O v c 00 00lS6 3-35

l l where-F ,, is the statistical hot channel factor for heat flux 'l Q K is a statistical =ultiplying factor a is the standard deviation of the heat flux factor, 4" including the effects of all the subfactors If 7 p,,, = 0.05 for 300 data points, then a K factor of 2.608 is required to protect 99 per cent of the population. The value of the hot channel factor then is Fq,, = 1 + (2.608 x 0.050) = 1.1304 and will provide 99 per cent confidence for the calculation. If, instead of using the 300 data points, it is assumed that the data represent an infinite sample, then the K factor for 99 per cent of the population is 2 326. The value of the hot channel factor in this case is Fq,, = 1 + (2 326 x 0.050) = 1.1163 which implies 100 per cent confidence in th9 eqiculation. The values of the K factor used above are taken from SCR-607.g201 The same basic techniques can be used to handle any situation involving variable confidence, population, and number of points. Having established statistical hot channel factors and statistical DIB factors, we can proceed with the calculation in the classical =anner. The statistical factors are used to deterrdne the minimum fraction of rods protected, or that are in no jeopardy of experiencing a DIB at ,each nuclear power peaking factor. Since this fraction is known, the maximum fraction in jeopardy is also known. It should be recognized that every rod in t,he core has an associative DIB 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 =atter how large its DIS ratio. Since both the fraction of rods in jeopardy at any particular nuclear power peaking factor and the number of rods operating at that peaking factor are known, the total number of rods in jeopardy in the whole core can be obtained by simple su=mation. The calculation is =ade as a function of power, and the plot of rods in jeopardy versus reactor overpower is obtained (Figure 3-21). The su==ation of the fraction of rods in jeopardy at each peaking factor su=med over all peaking factors can be =ade in a statistically rigorous manner only if the confidence for all populations is identical. If an infinite sample is not assumed, the confidence varies with population. To form this su==ation then, a conservative assumption is requin.i. P&W's total core model assumes that the confidence for all rods is equal to that for the least-protected rod, i.e., the minimun possible confidence factor is associated with the entire calcula-tion. The result of the foregoing technique, based on the -mimm design conditions (114 per cent power), is this statistical statement: C

f. - ' Djq[}

l f' 3-36 {[g

O There 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, evcn with continuous operation at the desi6n overpower. The maximum design conditions are represented by these assumptions:

a. The maximum design values of Fah (nuclear max /av6 total fuel rod heat input) are obtained by increasing the maxi =um calculated-value of FAh by a factor of 1.033 to provide additional design margin.

b. The maximum value for2F (nuclear max / avg axial fuel rod heat input) is determined for a transient condition followins a power level change with partial insertion of a rod group for control.

c. Every coolant channel in the core is assumed to have less than the nominal flow area represented by engineerin6 hot channel area factors, FA,.less than 1.O.

d. Every channel is assumed to receive the minim = flow associated with core flow maldistribution.

e. Every fuel rod in the core is assumed to have a heat input greater than the maximt.m calculated value. This value is represented by engineering hot channel heat input factors, Fq and Fqu, which are greater than 1.0.

V f. Every channel and associated fuel rod has a heat transfer margin above the experimental best-fit limits reflected in DIE ratios greater than 1.0 at maximum overpower conditions. ) The statistical core design technique may also be used in a simihr manner to evaluate the entire core at the most probcble mechanical and nuclear conditions i 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: There is at least a 99 per cent confidence that at least 99 9 per cent of the rods in the core are in no jeopardy of experiencing a DIB, even with continuous operation at the design overpower. The most probable design conditions are assumed to be the same as the =aximum 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 =aximum calculated value of heat input, and (2) Fqand F qn are assigned value; of 1.0.

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

d. Every fuel rod is assumed to have a nominal value for FA h nuclear.

3-37 00001957

The full meaning of the maxMm and most probable design statements requires additional ecm=ent. 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 will be in jeopardy, and that in gen-eral the number in jeopardy will be fewer than 0 5 per cent or 0.1 per cent. The statements do not mean to specify a given number of DUB's, but only ac-knowledge the possibility that a given number could occur for the conditions assumed. In su==ary, the calculational procedure outlined here represents a substantially improved desi6 n technique in two ways:

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

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

32323 Correlation of Heat Transfer Data The BAU-168 report (Ref.17) serves as a reference for the "best-fit" form of the design relationship used by bel. This heat transfer correlation has been found to be the most satisfactory in the representation of both uniform and nonuniform heat flux test data. The BAU-168 correlation is used by comparing the integrated average heat flux along a fuel red to a DUB heat flux limit predicted by the correlation. For uniform heat flux the integrated avera6e 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 DIS heat flux to the integrated average heat flux is a minimum is selected as the DIE point, and that value of the ratio at that point is the DIB ratio (DIBR) for that channel. This particular discussion deals with the comparison of DIB data to three par-pccorrelationsselectedwere: The BET correlation in ticularcorrelations(17/acorrelationwithwhichtheindustryisfamiliarin thecase the case of BAW-168, (21) and a correlation recently propsed for use in the of WAPD-188, desi6n of pressurized water reactors in the case of W-3.s22) The data considered for the purpose of these comparisons were taken from the following sources:

a. WAPD-188 (Ref. 21) .

b. AEEU-2213 (Ref. 23) .

c. Columbia University Data (Ref. 24, 25, and 26) .

l l 1 y 3-38

d. Argonne National Iaboratory Data, AIIL (Ref. 27).

'd c. The Babcock & Uilcox Cc=pany Data, B&;l (Ref. 28).

f. The Babcock & Wilcon Company Zuratom Data (Ref. 29).

The comparison of data to the EAW-168 correlation is presented as histograms of the ratio of the experimental DIG heat flux ( &,) to the calculated heat flux ( &C ). The data from cach source were Grouped by pressure and analyzed as a Group; batches were then prepared includin6 co=on pressure groups from all sources. Altogether there are 41 different data groups and batches con-sidered. Histograms for only the 2Ati-168 correlation are presented to minimize the graphical naterial. The information required for the generation of histo-Brams of the other two correlations was also prepared. The comparison of the various correlations to each other is facilitated through the use of tabulations of portinent statistical parameters. The standard devia-tion and mean value were obtained from the computed values of ( & 3 /$ 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 comparison. A relatively simple

    =ethod may be used to compare the correlations for various data; this method uses the coefficient of variation (Ref. 30) which is the ratio of the standard g  i deviation ( 7) to the mean X. The coefficient of variation may be thought of

( as the standard deviation given in per cent; it essentially norma 11zes the var-ious standard deviations to a comon mean value of 1.0. Table 3-14 is a tabulation of the data source, heat flux type, and corresponding histogra= numbers. The histograms are shown on Figures 3-22 through 3-37 Table 3-14 Heat Transfer Test Data Histogram F16ure Source Heat Flux Type Number Number WAPD-188 Ur.iform 1-9 3-22 3-23 3-24 AEN-R-213 Uniform 10-14 3-24 3-25 3-26 Columbia Uniform 15-19 3-26 3-27 A 3-28 U ANL Uniform 20 3-28 m 3 39 0000)F

Table 3-14 (Cont'd) s Source Heat Flux Type Number Number B&W Uniform 21 3-29 B&J-Euratom Unifom 22-24 3-29 3-30 Combined Data (500-720 psia) Unifom 25 3-30 Combined Data (1,000 psia) Uniform 26 3-31 Cembined Data (1,500 psia) Unifom 27 3-32 Combined Data (2,000 psia) Uniform 28 3-33 Combined Data (1,750-2,750 psia) Unifom 29 3-34 B&W-Euratom Chopped Cosine Nonunifom 30-32 3-35 B&W-Euratom and B&W Inlet Peak Nonunifom 33-35 3-35 3-36 Euratom and B&W Outlet Peak Nonuniform 36-38 3-36 3-37 Combined Nonunifo m (1,000 psia) Nonunifom 39 3-37 Combined Nonunifom (1,500 psia) Nonunifom 40 3-37 Combined Nonuniform (2,000 psia) Nonunifom 41 3-37 The histo 6 rams graphically demonstrate the distribution of ( & g/$ C ) for each data group. The Gaussian type distribution of ( &p/&C) about the mean for the group is apparent in the lar6e data groups. Some data Eroups are too small to provide meanin6ful histo 6 rams, but they are presented in order to complete this survey. Thedatawereusedaspresentedinthesourceforthecalculationof(&g/&)> C no points were discarded for any reason. A good correlation should be capable of representin6 DIO data for a full range of all pertinent parameters. The result of the comparison on this basis is demonstrated in Table 3-15 The data source, pressure, histo 6 ram figure number, heat flux type, and number of data points in the gmup are tabulsted. For each of the three correlations the fol-lowing data is indicated: r/E The coefficient of variation based on all available data in the group. n a The number of data points rejected usin6 Chauvenet's criterion (Ref. 31). This criterion is statistical in nature and is applied to the -

o. ' ' b,1,i! ,'

3 ho ,

                                                                         . 6

(9 (,,) values of ( & -g&C ). Dati points which fall outside certain limits with respect to the main body of data are rejected. ( e / s)' The coefficient of variation based on the original data sample less those points rejected by Chauvenet's criterion, i.e., based on n-nR values of (& g/$C )* It is unfartunate that Chauvenet's criterion must be applied to the values of (& C) rather than to the original data, since application to (&-jC & ) leads to t e rejection of points for either of two reasons:

a. Bad data points,

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

It is not desirable to reject points for the second reason, and yet one might expect to encounter some bad data. The logical choice then is to present data both ways, i.e., with and without Chauvenet's criterion applied. Of the 41 groups and batches analyzed the following is observed from Table 3-15: Groups and Batches of Data Groups and Batches of Data With Smallest a/7 Without With Smallest e/f With Correlation Chauvenet's Criterion Chauvenet's Criterion BAW-168 p/ U WAPD-188 38 2 36 3 W-3 1 2 Chauvenet's criterion rejected the following numbers of points for each corre-lation: Uniform Nonuniform Total BAW-168 Groups Only) 32 1 33 BAW-168 Batches Only) 39 0 39 laPD-188 (Groups Only) 34 2 36 WAPD-188 (Batches Only) 33 0 33 W-3 { Groups Only) 59 12 71 IJG sBatchesOnly) 50 9 59 Several notable peculiarities exist in the tabulation of Table 3-15 The Columbia data 500 psia group contained only five data points of which four were rejected by Chauvenet's criterion leaving one point. A standard devia-tion cannot be computed for one point; therefore all three values of (e/y)' are shown as not available (N.A.). Neither the BAW-168 nor the WAPD-188 pre-dicted any negative DNB heat fluxe : the W-3 predicted 93 negative values for uniform data. The fact that only s were rejected for this correlation indi-cates that the remaining 34 unifor points which were negative (93-59 = 34) V were close enough to the body of the data to be considered statistically sig-nificant. Table 3-15 may be consolidated somewhat as below by tabulating the e g1 000010

nuder of groups and batches of data having coefficients of variation within a specified interval for cach correlation. - (e/I) Interval EAW-168 BAW-168'(3) WAPD-188 iaPD-188'( ) W-3 W-3'(*) Negative 0 0 0 0 2 0 0-0.1 6 8 0 0 0 1 0.1-0.2 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 o.4-o.5 1 o 5 7 5 6 0 5-o.6 o o 6 5 3 4 0.6-o.7 o o 3 2 1 1 0 7-o.8 o o 2 1 7 8 0.8-o.9 1 0 0 o 1 5 0 9-1.0 0 0 0 0 1 0 Greater than 1.0 0 0 2 0 16 7 Total 41 40 41 40 41 40 ("}Chauvenet's criterion applied. As is seen from the above tabulation the colu=n for EAW-168 with Chauvenet's criterion applied indicates a grouping of 0.1 to 0.2, and a =aximum value of 0.28780 is noted from Table 3-15 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 DIB heat flux predicted by the W-3 correlation are in part responsible for the large spread in (e/I). The ability of the BAW-168 correlation to fit both uniform and nonuniform heat flux data over a wide range of pertinent variables leads us to believe that it is the best DNB correlation available. O h . I

  ,,*,.;  U:                      3 42 fr

O} x_/ c a l-V e e 88 s 31 5 Es me t: a a sa n

• a.s 4' -o o

o --- 71 o .

                                                     "         32 o o -

1%"S5

                                                                                  -oooo

a 1

o o g l o. o

            ] glwmocono-o              -mmog aooom g n ooo nemgg oon oom ooa oo.
                    -SQg$ add 9Ig              dEQR%:~                   82. om-og &#~                  ggg 0               *$m
                                                          $3 m b Ed.4gjdR3
                = S S *: . 1':':

g"~ - 4 8 42A

                                                                         *:-t

• AOig ~

                                                                                      'i's . *:@.g    $ *~:a 1 a

• i .4 Ia

                                                                                                                          *:oo.9
                    -oooowooo          = = 7oo. ':oooo-a: .1-1  a .1g  1 oo-        .
                                                                                 ~a-mm ooy ooa oo-
               -      =        =          2           . y : A R*               Musgt                           &

d R 4 3 GgO M j A 4. oo 4 o 1 o

                                                * =

o 1 9 1 3 o o o o 9.11" ooooo o 2 f gjomooooooo omoo- ao.on . - ooo a.m.a ooo ooo aoo ooo 2 2200Ra * ~24  : aR g nRE 223 gds #M GOR e= hx.exss,: 3.3:  :': n : s..-s~ es nn.ns .g g- A

e a esei':s .ma

4. : *:s n

e gs a$ e0 s': : -s :-: n .1 -vin d * * * ** *A* * * ** * ** **

e b e .

• Amm
• E a

0 S . 2 2 =GO O A R 3=

                      --m                                e        ~              ~ ~~.                     R                       '

8 1022 B  % 4 8 Edd42 2 ' y 3 edid & & & & 6 & & l ddddd & c 3

         %  7 in m     glo--maaooo monou aomon ~ o ooo m.aag ooo                                                o-,     ooo ooo

[ ' Y 4 v-asg8 g=~ tE **g g e a - aegs : g gja mg-+a~ g a a ag m e gm:23 -~ .gs.dg.

                                                                         -a      neg           *ag s      na*=e s~             =g=2 o  4
               ; 2400o08E8           r. m       a R$$~a 2             % dds     MO 22          0-Ota &&ddd ddd did ddd did      20     R$          I H  e         ddddddddd dadds &&d&& & a                                                                                      l D  @

d M 8  % O C e O %s a og U e j' mgd3R,c***

                        -               32R2f e -      *R2 -   2 R~ % 22d aS yg.E.T ARO $24 22A 002 m  a3 0

0 1 1 838 jij

                                                                                                 "

• gu 111 EEE 3

c, 2 122 222 EEEEEEEEE gEEEE EEEEE E E EEE EEEEE *-

            ,k     222222222 ~2222 2222e 2 2 222 22222 "gg 11111        111 11111               1111a 1a 1 ill *tt*         ,,, usu AA2 94 J**       tama === mamma                    uoan         ===111     11111 amama                66 Ett~~-      d" d' d" === 8 I4 -~-a-----

a an==, =2=== a = ==, =2=== axx =,n xxx as: E o a i assae:zas

                        ---~~~~        asiaa---     asgaa     - - i~ a
                                                                     -- gag gggv van                 g gag         aaa men s
                                                                                                       *e 14, % %

mm%. l l ......

              . uusn==g..au muuuu      2000:                                                                                       .

\', 22-2 ... 111 s1 ... aae aan ,a,:1 e e 4444j1111 --* **- 22

            -4     :deaaaeda siiltsi          aims agggg i 111:                  aal-sss-33333      Ex:

233 lll ll! 333 9 g3 00 00l($ = 3-5

O 'V) 3232.4 Evaluation of the Ther=al and Hydraulic Desi6n

a. Hot Channel Coolant Quality and Void Fraction An evaluation of the hot channel coolant conditions provides additional confidence in the thermal design. Sufficient coolant flow has been provided to insure low quality and void fractions. The quality in the hot channel versus reactor power is shown in Figure 3-38. The sensi-tivity of channel outlet quality with pressure and power level is shown by the 2,185 and 2,120 psig system pressure conditions eynmined.

These calculations were made for an Fa h of 1.85 Additional calcula-tions for a lo per cent increase in Fah to 2.035 were made at 114 per cent power. The significant results of both calculations are summmn-rized in Table 3-16. The effects of using an FAh of 179 are shown in Figure 3-38. Table 3-16 Hot Channel Coolant Conditions Exit Exit Void operating Power,% Fah Quality,% Fraction,% Pressure, psig O loo 1.85 (-)2.4(b) o,5(a) 2,185 114 1.85 2.8 13 5 2,185 130 1.85 94 36 9 2,185 n4 2.035 87 35.o 2,185 loo 1.85 o 38(*) 2,12 0 uh 1.85 5.4 25 2 2,120 130 1.85 12.1 45 2 2,120 uk 2.035 n.3 43 4 2,12 0 ("}Subcooled voids. (b) Negative irdication of quality denotes subcooling of 10.2 Btu /lb. The conditions of Table 3-16 vere determined with all of the hot chan-nel factors applied. Additional calculations were made for unit eell l' channels without engineering hot channel factors to show the .soa. ant (N conditions more likely to occur in the reactor core. Valuer for F ah h of 179 and 1.85 vere examined with and without fuel assembly flow distribution hot channel factors at 2,185 psig as shown on Figure 3-39 These results show that the exit qualities from the hottest a

I l cells should in general be considerably lower than the maximum de-sign conditions.

b. Core Void Fraction The core void fractions were calculated at 100 per cent pas ; for the normal operating pressure of 2,185 psig and for the minimum operating pressure of 2,20 psis. The influence of core fuel assembly 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: Flow,$ Pressure, psig Core Void Fraction, %

             '100            2,185                0.007 loo            2,120                0.033 95            2,185                0.041 95            2,120                0.227 i      The most conservative condition of 95 per cent flow at 2,120 psig l       results in no more than 0.13 per cent void volume in the core. Con-servative maximum design values for F Ah nuclear described by Line A of Figure 3-10 were used to make the calculation.

c. Coolant Channel Hydraulic Stability A flow regime map was constructed to evaluate channel hydraulic sta-bility. The transition from bubbly t9 agnular flow at high mass ve-locities was determined using Baker's(32; correlation, and the tran-sition from bubbly to slug was determined with Rose's(flgw which occurs

33) correlation. at low The mass velocities trsnsition from slug flow to annular flow was determined by Haberstroh's(34) correlation.

Bergies(35) found that these correlations, which were developed from adiabatic data, are adequate for locating flow regime transitions with best addition, and that they adequately predict the effects of pres-sace . Figure 3 ho shows the flow regime map on which have been plotted points 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 points lie vs11 within the bubbly flow regime. Since the bubbly flow regime is hydraulically stable, no flow instabilities should occur. l l d. Hot Channel DIE Comparisons DIB ratios for the hottest channel have been determined for the EAW-168 and W-3 correlations. The results are shown in Figure 3 41. DIB l ratios for both correlations are shown for the 150 axial max / avg l symmetrical cosine flux shape from 100 to 150 per cent power. The BAW-168 DIG ratio at the maximum design power of 114 per cent is 138; the corresponding W-3 value is 172. This compares with the suggested - il iii 3-45  !$ 10

( W-3 design value of 13 It is interesting to note that the calcu-lated DIG ratio reaches a value of 1.0 at about 150 per cent power with the BAW-168 equation which adequately describes DIS 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 con-dition. The sensitivity of DIE ratio with Fah and Fz nuclear was examined from 100 to 114 per cent power. The detailed results are labeled in Figure 3 41. A cosine flux shape with an F: of 1.80 and an FAh of 1.85 results in a W-3 Dra ratio of 1.45 and a BAW-168 ratio of 133 The W-3 value is well above suggested design values, and the BAW-168 value of 133 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-DIO ratio plot in 3 2 3 2.2, statistical core De-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 179 and 10 per cent above the mximum design value of 1.85 The resulting BAJ-168 DIS 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. l

   ) e. Reactor Flow Effects v                                                                                                   )

1 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 re- i quired system flow is obtained in the as-built plant. The experimental ' programs previously outlined in Section 1 will confirm the pressure drop and re3ated pump head requirements. It is anticipated that the as-built reactor flow will exceed the design value and will lead to increased power capability. An evaluation of reactor core flow and power capability was made by determining the m ximum steady state power rating versus flow. The analysis was mde by evaluating the hot channel at the overpower con- l ditions while mintaining (1) a DIE ratio of 138 (BAW-168), and (2) the statistical core design criteria. The results of the analysis are shown in Figure 3 ha. The power shown is the 100 per cent rating, and the limiting condition is 114 per cent of the rated power. An evaminntion 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 l The influence of reactor inlet temperature on power capability at a

p. given flow was evaluated in a similar manner. A variation of 1 F i ) in reactor inlet te=perature vill result in a power capability change of slightly less than } percent. m 3 46 00001% .

1

g. Fuel Te=perature and Fission uas Release Evaluation Temperature Determination s A fuel te=perature and gas pressure computer code was developed to calculate fuel temperatures, expansion, densification, equiaxed and colu=nar grain growth, center piping of fuel pellets, fission gas release, and fission 6as pressure. Program and data comparisons were made on the basis of the fraction of the fuel diameter within these structural regions:

1) Outer limit of equiaxed grain growth - 2,700 F

2) Outer limit of colu=nar grain growth - 3,200 F

3) Outer limit of molten fuel (UO2 ) - 5,000 F Data from Rt.ferences 36 throush 39 vere used to compare calculated and experimental fractions of the rod in grain growth and central melting.

The radial expansion of the fuel pellet is computed from the mean fuel temperature and the average coefficient of linear expansion for the fuel over the te=perature range considered. This model combined with the model for calculating the heat transfer coefficient was com-pared with the model developed by Not%y et al (Ref. 40) of AECL. The difference in fuel growth for the. >- calculation models was less than the experimental scatter of data. The program uses a polynomial fit relationship for fuel thermal con-ductivity. Three relationsnips were used to evaluate the effects of conductivity. A comparison of these conductivity relationships with the reference design C The values suggested in GFAP k624 IA-142(hl) is sh wn)in

                              ) and CVNA-246       Figure 3 are  very3similar 43     up to 3,000 F and the fomer values are more conservative above 3,000 F.

McGrath(43) concludes that the CVNA-246 values are lower limits for the hi6h tempra ture conditions. Fuel center temperatures for all three of the cc,aductivity relationships at the peaking factors given in 3 2 31.2 have been calculated to evaluate the margin to central i melting at the maximum overpower, and to show the sensitivity of the calculation with respect to thermal conductivity. Since the power peaks will be burned off with irradiation, the peaking factors used are conservative at end-of-life. The temperature drop from the fuel to clad is calculated with a sin-ple 6as gap conduction model until thermal expansion of the fuel in-traduces a contact effect. A contact coefficient is calculated con-sidering contact un by Ross and Stoute. ) pressure and gas gap conduction as suggested The results of the analysis with the methods described above are shown in Fi6 ures 3 4k and 3-45 for beginning and end-of-life condi-tions. The beginning and end-of-life gas conductivity values are 0.1 and 0.01 Btu /hr-ft2 -F respectively. The calculated end-of-life center fuel temperatures are higher than the beginning-of-life values because of the reduction in the conductivity of the gas in the 6ap. The effect is apparent even thou6h a contact condition prevails. The 3 47 '

                                                             .{

i

model does not include the effects of fuel swelling due to irradiation; consequently the ca. Llated contact pressures are conservatively lower than those expected at end-of-life conditions in the hottest fuel rods. The B&4 model gives very Good results when compared to the results of others in the field as is shown in F1 6ure 3 45 In the linear heat range of most interest, i.e., approximately 20 kv/ft, there is only about 300 F difference between the mv4== and min 4== values calcu-lated. Also the small difference . between the BEJ curve and the other curves indicates the relative insensitivity of the results to the shape of the conductivity at the elevated temperatures. The most conservative ass.mtptiens, using GEAP-4624 data with relatively little increase in thermal conductivity above 3,000 F, result in cen-tralfuelmeltingatabout22kw/ft,whichis2kw/fthigherthanthe maximumdesignvalueof199kw/ftat114percentpower. Further evaluation of the two fi6ures shows that central fuel melting is pre-dicted to occur between 22 and 26 kW/ft depending on the time-in-life and conductivity asstunptions.

h. Fission Gas Release

                                                                            @5)

The fissiondata Additional gasfrom release is based on results GEAP-4314,(46) re orted AECL-603,( in GEAP-4596(40)

7) and CF-60-12-14 have been compared with the su6gested release rate curve. The release rate curve (45) is representative of the upper limit of release data in the temperature region of most importance. A rele.ase rate of 51 per cent is assumed above 3,500 F.

The fission 6as release rates may be summarized as follows for end-of-life conditions with maximum burnup. Conductivity Model Release Rate, % CVUA-142 43 CVHA-246 h2 GEAP-4624 40 A release rate of h3 per cent is used to detemine the fuel clad in-ternal design cor. iittons reported in 3 2.4.2, Fuel Assemblies. I i 1 O v l h, 3 us 0000J$

3 2.4 lECHANICAL DESIGN IAYOUT - 3 2.4.1 Internal Layout Reactor internal components include the upper plenum assembly, the core sup-port assembly (consisting of the core support shield, core barrel, lower grid and flow baffle, thermal shield, and surveillance specian holder tubes), and the incore instrum nt guide extensions. Figure 3-46 shows the reactor vessel, reactor vessel internals arrangement, and the reactor coolant flow path. Fig-ure 3 47 shows a cross section through the reactor vessel, and Figure 3-48 shows the core flooding arreagement. Reactor internal compoconts do not include fuel assemblies, control cluster assemblies (CCA's), surveillance specimen assemblies, or incore instrumentation. Fuel assemblies are described in 3 2.4.2, control cluster and drive assemblies in 3 2.4.3, surveillan::e specima assemblies in 4.4 3, and incore instrumnta-tion in 7 3 3 The reactor internals are designed to support the core, maintain fuel align-ment, limit fuel assembly movement, and maintain CCA guide tube alignment be-tween fuel assemblies and control rod drive assemblies. They also direct the flow of reactor coolant, provide gam =a and neutron shielding, provide guides for incore instrum ntation between the reactor vessel lower head and the fuel assemblies, and support the surveillance specimen assemblies in the annulus be-tween the ther=al shield and the reactor vessel vall. All reactor internal components can be removed from the reactor vessel to allow-inspection of the reactor internals and the reactor vessel internal surface. A shop fitup and checkout of all internal components in an as-built reactor vessel cockup will insure proper alignment of =ating parts prior to shipment. Dummy fuel assemblies and control cluster assemblies will be used to check fuel assembly clearances and CCA free movement. In anticipation of lateral deflection of the lower end of the core support as-sembly as a rese.; of horizontal 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 lower end of the core support assembly which could conceivably result from flow-induced tor-sional loadings. The blocks allow free vertical movemnt of the lower end of the internals for thermal expansion throughout all ranges of reactor operating conditions, but in the unlikely event of a flange, circumferential veld or boltedjointfailure,theblockswilllimitthepossiblecoredropto1/,2in. or less. The final elevation plane of these blocks will be established on the same elevation as,the vessel support skirt attachment to minimize dynamic load-ing effects on the vessel shell or bottom head. Preliminary calculations indicatetheimpactloadingonthestop'blocksfora1/4in.coredropwould be approximately 5 g's total. Block location and geometry vill be evaluated and detemined to transfer this leading through the vessel support skirt to the reactor building concrete. A significant reduction of the above impact loading can be achieved through proper stop block design and detailed analysis. O 5 1 -

               'It                 3-49
                                                                          !          I

hU A 1/2 in. core drop will not allow the lower end of the CCA poison pins to dis-- engage from their respective fuel assembly guide tubes if the CCA's are in the full-out position, as approximately 6-1/2 in. of pin length would remain in the fuel assembly guide tubes. A core drop of 1/2 in. Will not result in a significant reactivity change. The core cannot rotate and bind the drive lines because rotation of the core support assembly is prevented by the stop blocks. The failure of the core support shield and core barrel upper flanges, or re-lated flanges and other circumferential joints, is not considered credible on the basis of the conservative design criteria and large safety factors employed in the internals design. The final internals design will be capable of with-standing various combinations of forces and loadings resulting from the static weight 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 (0.05 g vertical gives 24,000 lb), coolant flow hydraulic loading (230,000 lb), and other related loadings. The algebraic sum of this simplified loading case is 483,000 lb. This results in a tensile stress of about 700 psi in the core support 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 thermal stress during all tran,sients, and resolution of fabrication details such as shell rolling tolerances and weld joint preparation details will increase the stress leveli listed above. As a final design criterion, the core support components will meet the stress requirements of the ASME Code, Section III. The structural integrity of all core support veld joints in the internals shells [m] V will be insured by compliance with the radiographic inspection requirements in the above code. The seismic analysis will include detailed calculations to de-termine the maximum structural response of the reactor vessel and internals. This analysis will be performed as described in 3 1.2.4.1. In the event of a major loss-of-coolant accident, such as a 36 in. diameter coolant pipe break near the reactor vessel outlet, the fuel assembly and vessel internals would be subjected to dynamic loadings resulting from an oscillating (approximately sinusoidal) differential pressure across the core. A prelimi-nary analysis of this postulated accident indicates that the fuel assemblies wouldmoveupwardlessthan3/8in. Sece deflection of the internals struc-tures would occur, but internals component failure will not occur. The oc-currence of a loss-of-coolant accident and resulting loadings will be evalu-ated during t S detailed design period for the fuel assemblies and related internals str. tural components. The cerlections and movements described above would not prevent CCA insertion because the control poison pins are guided by split tubes throughout their  ; travel, and the guide tube to fuel assembly alignment cannot change regardless 1 of related component deflections. CCA trip could conceivably be delayed = omen- l tarily as a result of the oscillating pressure differential. However, the CCA travel time to full insertion would remain relatively unaffected as transient pressure oscillations are dampened out in approximately 0 5 sec. On this basis, theCCAtriptimeto2/3insertionwillbeapproximately155seeinsteadof the specified 1.40 sec. Also, this possible initial minor delay in trip initia-t'"j tion would not contribute to the severity of the loss-of-coolant accident be-cause at the initiation of CCA trip, the core would be suberitical from voids. m 3-30 00 00$D

Material for the reactor internals bolting vill be subjected to rigid quality , control requirements to insure structural integrity. The bolts will be dye penetrant-inspected for surface flaw indications after all fabrication opera-tions have been completed. Torque values will be specified for the final as-se=bly to develop full-bolting capability. All fasteners will be lock-velded to insure assembly integrity. 3 2.4.1.1 Upper Plenum Assembly The upper plenum assembly is located directly above the reactor core and is removed as a single component prior to refueling. It consists of upper and center grid assemblies, CCA guide tubes, a flanged plenum cylinder,-spcning: -for pt:22ge ^# "2tcrfor 214% and openings for reactor coolant out-let 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. Rectangu-lar flanges on the CCA guide tubes are bolted and lock-velded to the upper grid bars. CCA guide tubes provide CCA guidance and protect the CCA from the effects of coolant cross-flow. Each CCA 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 are slotted for their full length, the brazement provides continuous guidance for the CCA full stroke travel. Design clearances in the guide tube vill ac-com=odate some degree of misalignment between the CCA guide tubes and the fuel assemblies. Final design clearances will be established by tolerance studies and the results of control rod drive assembly prototype tests. The center grid assembly consists of parallel flat bars intersecting to form square lattices at the top and bottom of the center grid assembly. The bars are attached to a flange which is bolted to the plenum cylinder lower flange. The center grid assembly locates the lower end of the individual CCA guide tube relative to the upper end of the corresponding fuel assembly. Incating slots in the upper plenum assembly top flange engage the reactor ves-sel top flange locating devices to align the upper plenum assembly with the reactor vessel, reactor vessel top head control 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, ther=al shield, and surveillance specimen holder tubes. Static loads from the assembled components and fuel assemblies, and dynamic loads from CCA trip, hydraulic flow, thermal expansion, seismic disturbances, and loss-of-coolant accident considerations,are all carried by the core support assembly. Each of the core support assembly components is described as follows: C f f 3-51 ) i

a. Core Support Shield b]

/ The core support shield is a large flanged cylinder which mates with the reactor vessel opening. The top flange rests on a circumferential  : 1 edge in the reactor vessel top closure flange. The core support shield lower flange is bolted to the core barrel. The cylinder wall has two nozzle openings for reactor coolant outlet flow. ??: - i _ficed'n; c' "cre, ;nh containing a n;;;1; op #ng, ar: .:cli;i te th; ;jlic. L ;;11 c# '" :n %rrt chi 12.- Incating blocks on the j inside of the cylinder wall near the bottom guide and align the upper l plenum enamber relative to the core support shield. l The reactor vessel outlet nozzles C tb ocre f1: M'"; - '^~ are l sealed to the mating components of the core support shield by the I differential thbrmal expansion between the carbon steel reactor ves- 1 sel and the stainless steel core support shield. The nozzle seal surfaces are finished and fitted to a predetermined cold gap provid-ing clearance during core support assembly installation and removal. At reactor operating temperature the cating metal surfaces are in contact to make a seal without exceeding allowable stresses in either the reactor vessel or internals.

b. Core Barrel The core barrel supports the fuel assemblies and lower grid and flow O baffle, and directs the reactor coolant flow through the vessel.

O The core barrel consists of a flanged cylinder, a series of internal horizontal spacers bolted to the cylinder, and a series of vertical plates bolted to the inner surfaces of th9 horizontal spacers to form an inner wall enclosing the fuel assemblies. The core barrel construction will be similar to the reactor internals component developed by B&W for the Indian Point Station Unit No.1. Coolant flow is downward along the outside of the core barrel cylin-der, and upward through the fuel assemblies contained in the core barrel. A snall portion of the coolant flows upward through the space between the core barrel outer cylinder and the inner plate wall. Coolant pressure in this space is naintained slightly lower than the core coolant pressure, thus avoiding tension loads on the bolts which attach the plates to the horizontal spacers. The vertical plate inner wall vill be carefully fitted together to reduce reactor coolant leak-age to an acceptable rate. The upper flange of the core barrel outer cylinder is bolted to the mating lower flange of the core support shield, and the lower flange is bolted to the mruing flange of the lower grid and flov baffle. All bolts will be inspected and installed as described in 3 2.l+.1, and will be lock-welded after final assembly. V Spacer bars and support lugs are welded to the core tarrel outer cylinder to position and support the ther=al shield. 3-52 00 66HT

c. Iower Grid and Flow Baffle The lower grid provides alignment and support for the fuel assemblies and aligns the incore instrumnt guide extensions with the fuel as-sembly incore instrument tubes. The 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.

The 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 ther=al shield is installed in the an-nulus between the core barrel outer cylinder and the reactor vessel inner wall. The ther=al shield reduces the neutron and gam =a internal heat generation in the reactor vessel wall, thereby reducing the re-sulting thermal stresses.

The thermal shield is supported on, and bolted to, lugs velded to the core barrel outer cylinder to minimize the possibility of thermal shield vibration. The support lugs are designed to prevent the bolts being loaded in shear. Bolts are lock-welded after final asseubly.

e. Surveillance Specimen Holder Tubes Surveillance specian holder tubes are installed on the core stpport assembly outer wall to contain the surveillance specimen assemb.'.ies.

The tubes extend from the top flange of the core support shield to the lower end of the thermal shield. The tubes will be rigidly at-tached to prevent flow-induced vibration. Slip joints at the lower end of the core support shield will allow the core support shield to be removed from the core support assembly without destructively re-moving the surveillance specimen holder tubes. 3 2.4.1 3 Incore Instrument Guide Extensions The incore instrument guide extensions guide the incore instrument assemblies between the instru=ent penetrations in the reactor vessel bottom head and the instru=ent tubes in the fuel assemblies. A spring-loaded ball joint at the lower end of the instrum nt guide extensions provides for minor misalignmnt between the reactor vessel instrument penetra-tions and the fuel assembly instrum nt tubes. A perforated shroud tube concen-tric with the instrument guide tube adds rigidity to the assembly and reduces the effect of coolant flow forces. Fifty-one incore instrument guide extensions are provided. The incore instrument guide extensions are designed with a slip joint so they will not be affected by the core drop described in 3 2.4.1. O -

          ' '     't 3-53

3 2.4.2 Fuel Assemblies 3 2.4.2.1 Description

a. General Description The fuel for the reactor is sintered pellets of low enrichment ura-nium dioxide clad in Zircaloy-4 tubing. The clad, fuel pellets, end supports, holddown spring, and end caps form a " Fuel Rod Assembly".

Two hundred and eight fuel rod assemblies are mechanically Jcined in a 15 x 15 array to form a " Fuel Assembly". The fuel assembly is shown in Figure 3-49 The center position in the assembly is re-served for instrumentatien. The remaining 16 positions in the array are provided with " Guide Tubes" for use as control pin locations. The complete core consists of 177 fuel assemblies. All, assemblies are identical in mechanical construction, i.e., all are designed to accept the control cluster assemblies (CCA). However, only 69 are provided with CCA's to control the reactivity of the core under oper-ating conditions. In those 108 fuel assemblies which do not contain a CCA during a given core cycle, the guide tices are partially filled at the top by an " Orifice Cluster Assembly" (Figure 3-50) in order to minimize bypass coolant flow. These orifice cluster assemblies also tend to equalize coolant flow between fuel assemblies with CCA's and those with orifice clusters. (7 Fuel assembly components, materials, and dimensions are listed below: V Item Material Dimensions, in. Fuel UO2 Sintered 0 362 diam. l Pellets Fuel Clad Zircaloy-4 0.420 OD x 0 368 ID x 152-7/8 long Fuel Rod Pitch 0 558 Fuel Assembly Pitch 8 587 Active Fuel Iangth 144 Overall Iength 165 Control Pin Guide Tube Zircaloy-4 0 530 OD x 0.015 vall Incore Instrument Zircaloy-4 0 530 OD x 0.075 vall Tubes Spacer Grid Assemblies Stainless Steel, Spaced at 21 in. Tp-304 Structural Can Panel Stainless Steel, 0.031 thick Tp-304 l End Fitting Stainless Steel, Tp-304 O V H 3-Su 00 an1~1L uv.- Mb,

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

The density of the fuel is 95 per cent of theoretical. I Average design burnup of the fuel is 28,200 WD/MIU. Peak burnup is 55,000 wD/MIU. At the peak burnup, the fuel growth is calculated to be 9-1/2 volume per cent by the method given in Reference 49 This growth is accoc=odated by pellet porosity, by the radial clear-ance provided between the pellets and th cladding, and by a small amount of plastic strain in the cladding. Each fuel column is located, at the bottom, by a thin-wall stainless steel 3destal and is held in place during handling by a spring at the top. The spring allows axial differential thereal expansion be-tween fuel and cladding, and axial fuel growth. The bottom pedestal is also collapsible, thus providing a secondary buffer to prevent ex-cess cladding axial strain. Fission gas release from the fuel is accommodated by voids within the fuel, by the radial gap between the pellets and cladding, and by void space at the top and bottom ends of the fuel rod.

c. Fuel Assembly Structur,e_,

(1) General The fuel assembly, shown in FiB ure 3-49, is of the canned type. Eight spacer grid assemblies and four perforated panels form the basic structure. The renela are velded together at the cor-ners for the entire length. The spacer grid assemblies are velded to the panels, and the lower and upper end fitting assen-blies are velded to the panels to conplete the structure. The upper end fitting assembly is not attached until the fuel rods, guide tubes, and instrument tube have been installed. At each spacer grid assembly each fuel rod is supported on four sides by integral leaf-type springs. These springs are designed to pro-vide a radial load on the fuel rod sufficient to restrain it so that flow-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 accoc=odate the differential thermal expansion be-tween the Zircaloy fuel rod and the stainless steel structure. (2) Spacer Grid Assembly These assemblies are comprised of ferrules made of square tubing. The ferrule has a portion of each side formed into spring sec-tions which have hydrodynamically shaped " dimples" that contact the fuel rods. The ferrules are joined to6 ether by bracing to form the spacer grid assemblies. The grid assemblies, which

                                                                              ,J kE                             3-5' o

los

provide the desired pitch spacing between fuel rods, are spot-welded at intervals to the perforated stain 19ss steel can panels. (3) Iower End Fitting Assembly The lower end fitting assembly is constructed from Type 304 stainless steel members which when joined together form a box structure. Four deep cross members serve as the positioning sur-faces for the fuel assembly when it is inserted into the lower 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 Fitting Assembly The upper end fitting assembly is similar to the lower end fit-ting assembly. Tt positions the upper end of the fuel assem-bly and provides coupling between the fuel assembly and the handling equipment. A hollow post,velded in the center of the assembly,is designed to provide an uncoupling means for the CCA-to-drive connection, and to retain the orifice cluster as-sembly. In order to identify a fuel assembly under water, a serial number is milled into a flat, chrome-plated surface which is welded to the box frame. (5) Guide Tubes The guide tubes are Zircaloy tubes which serve to guide the con-trol pins within the fuel assembly during operation. The tubes are restrained axially by the upper and lower end fitting assem-blies in the fuel assembly, and are restrained radially by the spacer grids in the same manner as the fuel rods. 3 2.4.2.2 Evaluation

a. Fuel Rod Assembly (1) General The basis for the design of tk - fuel rod assembly is discussed-in 3 1.2.4. Materials testing and actual operation in reactor service with Zircaloy cladding has demonstrated that Zircaloy-4 material has ample corrosion resistance and sufficient mechani-cal properties to maintain the integrity and serviceability re-quired 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 wall and the greatest ovality permitted by the specifi-(v) cation is operating in the region of the core where performance requirements are the most severe. Fission gas release rates, m 3-56 00 0Q06

fuel Browth, and changes in mechanical properties with irradia- 3 tion are based on a conservative evaluation of currently avail-able data. Thus, it is unlikely that significant failure of the cladding vill result during operation. The actual clad stresses are considerably below the yield strength. Circumferential stresses due to external pressure, calculated us-ing those combinations of clad dimensions, ovality, and eccen-tricity which produce the highest stresses, are shown in Table 3-17 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 pai at design pressure, 27,000 psi at operating pres-sure. At this stress, the =aterial may creep sufficiently to allow an increase in ovality until further creep is restrained by support from the fuel. Contact loads on the order of 20 lb/ in. of length are sufficient to counteract the bending stress. Creep collapse tests have indicated a long time collapse resis-tance in excess of the requirement to prevent collapse in the end void. As the fuel rod internal pressure builds up with tin.0, 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 of short time burst. The possibility of stress-rupture burst has been investigated, using finite-difference methods to esti-mate the long time effects of the increasing pressure on the clad. The predicted pressure-time relationship produced stresses which are less than 1/3 of the stress levels which would produce stress rupture at the end-of-life. Outpile stress-rupture data were used, but the greater than 3:1 margin on stress is more than enough to account for decreased stress-rupture strength due to irradiation. Clad circumferential stresses are listed in Table 3-17 The free 6as content of the fuel rod is calculated by consider-ing (1) initiel helium fill gas, (2) init;ial water vapor and at-mospheric gases adsorbed on the. fuel, and (3) fission product gases. The water vapor present initially is expected to disso-ciate over the life of the fuel and enter into hydriding and oxidizing reactions. The gas remaining at the end-of-life, when the maximum internal pressures exist, consists of the atmospheric gases and helium present initially plus the released fission gases. The fission gas production is evaluated for a range of neutron fluxe d the fissionable material present over the life of the fuel. 5 A design value for gas production of 0.29 atoms of gas per fission has been determined. q !) b ' ' 3-57 l

l Table 3-17 Clad Circumferential Stresses Ultimate Calc. Yield Tensilre Stress, Stress, Stress, psi psi psi Operating Condition ,

1. BOL " - Operating at Design Pre sure Tott.1 Stress (membrane + bending) Due to 2,c00 psig System Design F. 3sure Minus 100 psig Fuel Rod Internal Pressure Aserage clad Temperature - Approximately 635 F (expansion void) , -33,000 46,000

2. EOL - Maximum Overpower l l

System Pressure - 2,185 psig  ; Fuel Rod Intea al Pressure - 3,300 psig Average Temperature Through Clad Thick-ness at Hot Spot - Approximately 725 F Pressure Stress Only ) 9,000

                       - Including 4,000 psi Thermal Stress              13,000          36,000                38,000 3      EOL - Shutdown Immediately After Shutdown System Pressure - 2,200 psig Fuel Roi Tnta nal Pressure - 1,750 psig Average Clad Temperature - Approximately 575 F                                              -4,000          45,000                48,000 3 Hours Later (50 F/hr Pressurizer Cooldown Rate) itel Rod Internal Pressure - 1,050 psig System Pressure - 680 psig Average Clad Tempe.mture - Approximately 425 F                                               3,300          52,000                55,000 (a) Cladding is being ordered with 45,000 psi minimum yield strength and 10 per cent minimna elongation, both at 650 F. Minimum room temperature strengths will be approximately 75,000 psi yield strength (0.2 per cent offset), and 85,000 psi ultimate tensile strength.

Clad stresses due to fuel swelling are discussed further below.

                                                          ,_,,                        000016ff i                                                                                                                        o

The total production of fission gas in the hottest fuel rod as- ' sembly is based on the hot rod average burnup of 38,000 WD/MIU. The corresponding maximum burnup at the hot fuel rod midpoint is 55,000 WD/MIU. The fission gas release is d on temperature versus release fraction experimental data. Fuel temperatures are calcu-lated for small radial and axial increments. The total fission gas release is calculated by integrating the incremental re-leases. The maximum release and gas pressure buildups are determined by evaluating the following factors for the most conservative con-ditions: (a) Gas conductivity at the end-of-life with fission gas pres-ent. (b) Influence of the pellet-to-clad radial gap and contact heat transfer coefficient on fuel temperature and release rate. (c) Unrestrained radial and ax.al thermal growth of the fuel pellets relative to the clad. (d) Hot rod local peaking factors. (e) Radial distribution of fission gas production in the fuel pellets. (f) Fuel temperatures at reactor design overpower. The fuel temperatures used to determine fission gas release and internal gas pressure have been calculated at the reactor over-power condition. Fuel temperatures, total free gas volume, fis-sion gas release, and internal gas pressure have been evaluated for a range of initial diametral clearances. This evaluation shows that the highest internal pressure results when the maxi-mum diametral gap is assumed because of the resulting high aver-age fuel temperature. The release rate increases rapidly with l an increase in fuel temperature, and unrestrained axial growth l reduces the relatively cold gas end plenum volumes. A conserva-tive ideal thermal expansion model is used to calculate fuel temperatures as a function of initial cold diametral clearance. Considerably lover resistance to heat transfer between the fuel and clad is anticipated at the end-of-life due to fuel fracture, swelling, and densification. The resulting maximum fission gas release rate is 43 per cent. (3) Collapse Margins Short time collapse tests have demonstrated a clad collapsing pressure in excess of 4,000 psi, at expansion void maximum tem-perature. Collapse pressure cargin is approximately 1 7 Ex-trapolation to hot spot average clad temperature ( = 725 F)

  !Ob 3-59

indicates a collapse pressure of 3,500 psi and a margin of 1.4, V('] which also greatly exceeds requirem nt. Outpile creep collapse tescs have de=onstrated that the clad meet a the long time (creep collapse) requirement. (k) Fuel Swelling Fuel rod average and hot spot operating conditions and design paramters at 100 per cent power, pertinent to fuel swelling con-siderations,are listed below. Average Maximum 2 HeatFlux, Btu /ft-hr 167,620 543,000 LinearHeatRate,kw/ft 5.4 17 5 Fuel Temperature, F 1,385 4,160 Eurnup (NWD/ M ) at Equilibrium 28,200 55,000 Nominal Values Pellet OD, in. 0 362 Pellet Density, % of Theoretical 95 Pellet-Clad Diamtral Gap at Assy., in. 0.004 - 0.008 Clad Material Cold-Worked Zr-4 Clad Thickness, in. 0.026 The capability of Zircaloy-cled UO2 fuel in solid rod form to perform satisfactorily in PWR service has been amply demonstrated through operation of the CVTP and Shippingport cores, and through l results of their supplemntary development programs, up to ap- i proximately 40,000 WD/M. l l As outlined below, existing experimental information supports l the various individual design paramters and operating conditions up to and perhaps beyond the maximum burnup of 55,000 WD/M, but not in a single experiment. However, the LRD 1rradiation test program, currently in progress, does combine the items of concern in a single experim nt, and the results are expected to be available for final design confirmation by the end of 1966. l l (5) Application of Experimental Data to Design Adequacy of the Clad-Fuel Initial Gap to Accommodate Clad-Fuel Differential Thermal Expansion d Experimental Work 00 0(310 3-60

1 Six rabbit capsules, each containing three, 5 in, fuel length, s Zr-2 clad rods, were irradiated in the Westinghouse Test Reac-tor (41) at power levels up to 24 kw/ft. The 94 per cent theo-retical density (T.n) UO2 pellets (0.430 OD) had initial clad-fuel diametral gaps of 6,12, and 25 mils. No dimensional changes were observed. Central m1 ting occurred at 24 kw/ft only in those rods which had the 25 mil initial Sap. Two additional capsules were tested.(E1} The specia ns were similar to those described above except for length and initial gap. Initial gaps of 2, 6, and 12 mils were used in each cap-sule. In the A-2 capsule three 38 in. long rods were irradiated to 3,450 WD/M at 19 kw/,ft maximum. In the A-4 capsule, four 6 in. long rods were irradiated to 6,250 WD/m at 22.2 kw/ft maximum. No central melting occurred in sny rod, but diameter increases up to 3 mils in the A-2 capsule and up to 1 5 mils in the A-4 capsule were found in rods which had the 2 mil initial gap. Application In addition to demonstrating the adequacy of Zirealoy-clad UO2 pellet rods to operate successfully at the power levels of in-terest (and without central m1 ting), these experiments demon-strate that the design initial clad fuel gap of 4 to 8 mils is adequate to prevent unacceptable clad diamter increases due to differential thermal expansion between the clad and the fuel. A maximum local diametral increase of less than 0.001 in. is in-dicated for fuel rods having the minimum initial gap, operating at the maximum overpower condition. (6) Adequacy of the Available Voids to Accommodate Differential Expansion of Clad and Fuel, Including the Effects of Fuel Swelling Experimntal Work l Zircaloy-clad, UO2 pellet-type rods have performd successfully in the Shippingport reactor up to approximately 40,000 WD/M. Bettis Atomic Power Inboratory(k9) has irradiated plate-type UO2 fuel (96-98 per cent T.n) up to 127,000 WD/m, and at fuel center fuel temperatures between 1,300 and 800 F. 3,20 This work indi-cates swelling rates of 0.16% AV/10 f/cc until fuel in-ternal voids are filled, then 0 7% AV/1020 f voids are filled. This point of " breakaway"/cc appearsafter to beinternal inde-pendent 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 ele-ments tend to reduce these swelling effects by providing greater  ; resistance to radial swelling, and lower resistance to longitu-  ; dinal swelling than was present in the plate-type test speci-  : I

mens.l j ,' -l

This is confirmed in t by the work of Frosc, Bradbury, and Griffiths of Harwell( inwhich1/4in.diameterUO2 pellets, clad in 0.020 in. stainless steel with a 2 mil diametral gap, were irradiated to 53,300 MID/M at a fuel center temperature i of 3,180 F without significant dimensional change. ! In other testing (55) 0.150 in. OD, 82-% per cent T.D. oxide pellets (20 per cent Pu, 80 per cent U), clad with 0.016 in. stainless steel with 6-8 mil diametral gaps, have been irradiated to 77,000 M(D/M at fuel temperatures high enough to approach central melting withmat apparent detrimental results. Compar-able results were oftained on rods svaged to 75 per cent T.D. and irradiated to J 00,000 M(D/M. Application Based on the BAP.', experimental data, swelling of the fuel rods is estimated as outlined below Fuel is assumed to 'twell unifo2mly in all directions. Clad-pellet differential thermal ex]ansion is calculated to be about 0.004 in. at the maximum linear heat rate, so that all of the minicum initial gap of 0.004 in. is filled up by thermal expan-sion. If the initial gap exceeds the minimum, the additional gap volume is assumed available to accommodate swelling. This O additional void volume may initially tend to be filled by pellet thermal expansion because of the low contact pressure and resul-tant low contact coefficient, but as the fuel swells, the con-tact pressure must increase if the clad is to be stretched. Where fuel cracking tends to fill the radial gap, it is assumed that the crack voids are ava m ble to absoro swelling. The external effect of fuel swelling is assumed to occur at 0.16% AV/1020 f/ccuntilthe5percentinitialvoidinthe95 per cent T.D. pellets is filled at about 9 x 1020 f/ce. 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 ffee( ,000 ggp/ m ) is reached. Total fuel volume increase is 4-1 per cent which results in a 1-1/2 per cent diameter increase in a rod with the 0.004 in. minimum initial gap. Clad stress is estimated at 20,000 psi, so that the elastic strain is about O. ' 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. 1 Qualitative information from LSBR(54) suggests that swelling rates for this design may exceed those indicated by the BAPL data, because 9f tche higher fuel temperatures. However, the A.E.R.E. tests (52/ and the General Electric tests (53) do not sup-O. port more than a small increase in post " breakaway" swelling rates at temperatures of interest.

                             ,,                      00 0 @

I Fuel Svelling Studies - IRD Irradiation Program (EE Dimensional stability of.UO2 under inpile conditions simulating large reactor environments is under investigation. This study is currently being carried out under USAEC Contrac' AT(30-1)- 3269, "Iarge Closed-Cycle Water Reactor Research and Develop-ment Program". Parameters which contribute to swelling are burnup, heat rating, j fuel density and grain size, and clad restraint. These are sys-tematica11y being studied by irradiating a series of capsules , containing fuel rods. These experiments were assigned by the i AEC to ETR/MIR. Test variables are shown in Table 3-18. Table 3-18 LRD Fuel Swelling Irradiation Program Initial Goal Heat Rating (b) Capsule (") Enrichment, kv/ vatts/ Fuel Density, Burnup, WAFD-49  % ft em  % T.D. WD/MIU_ A 18 7 12 394 94 and 97 36,000 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 E 13 5 18 591 94 and 97 47,000 F 13 5 18 591 90, 94, and 97 47,000 G 16.0 18 591 90 and 97 47,000 H 17 0 24 788 94 and 97 56,000 I 18 7 24 788 94 and 97 56,000 J 20.0 24 788 94 and 97 56fx0 K 20.0 24 788 90 and 94 56,000 L 20.0 24 788 94 and 97 56,000 (a)Fourrods/ capsule. (b) Fuel center temperatures vary from 1,570 to 4,110 F. O e n

i Effect o( Zircaloy Creep (mu) The effect of Zircaloy creep on the amount of fuel rod growth due to fuel swelling has been investigated. Clad creep has the effect of producing a nearly constant total pressure on the clad ID, by permitting the clad diameter to increase as the fuel diameter increases. Based on out-of-pile data,(56) 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 clad at the hot spot. At the start of this high swelling period (roughly the last 1/3 of the core life), the reactor coolant system pressure vould more or less be balanced by the rod inter-nal pressure, so that the total pressure to produce the clad stress of 22,000 psi vould have to come from the fuel. Contact pressure would be 2,400 psi. At the end-of-life, the rod inter-nal pressure exceeds the system pressure by about 1,100 psi, so

                                                      ~

that the clad-fuel contact pressure vould 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 swelling period, 700 psi at the end-of-life . Since the contact pressure was assumed to be 825 psi in calculating the contact coefficient used to determine the 3 fuel pellet thermal expansion, there is only a short period at the very end-of-life (assuming the 3:1 increase in creep rates due to irradiation) when the pellet is slightly hotter than cal-culated. The effect of this would be a slight increase in pel-let thermal expansion, and therefore in clad strain. Consider-ing the improbability that irradiation will actually increase creep rates by 3:1, no change is at:icipated.

b. Overall Assembly (1) Assurance of Control Cluster Assembly Free Motion The 0.058 in. diametral clearance between the guide tube and the control pin is provided to cool the control pin and to insure ,

adequate freedom to insert the control pin. As indicated below, ' studies have shavn that fuel rode will not bow aufficiently to touch the guide tube. Thus, the guide tube will not undergo de-formation caused by fuel rod bowing effects. Initial lack of straightness of fuel rod and guide tube, plus other adverse tol-erance conditions, conceivably could reduce the 0.083 in. nomi-nal gap between fuel rod and guide tube to a minimum of about 0.045 in., including amplification of bowing due to axial fric-tion loads from the spacer grid. The maximum expected flux gradient across a fuel rod of 1.176 vill produce a temperature difference of 12 F, which will result in a thermal bow of less than 0.002 in. Under these conditions for the fuel rod to touch (p ~j the guide tube, the thermal gradient across the fuel rod diam-eter would have to be on the order of 300 F. c 00 0$ $ 3-eu

The effect of a DNB occurring on the side of a fuel rod adjacent s  ; to a guide tube would result in a large temperature difference. l In this case, however, investigation has shown that the clad - temperature would be so high that insufficient strength would  ; be available to generate a force of sufficient magnitude to j cause a significant deflection of the guide tube. In addition, she guide tube would experience an opposing gradient that would resist fuel rod bowing, and its internal cooling would maintain temperatures much lower than the fuel rod, thus retaining the guide tube strength. (2) Vibration ThesemiempiricalexpressiondevelopedbyBurgreen( was used to calculate the flow-induced vibratory amplitudes for the fuel

                                                       ~

assembly and fuel rod. The calculated amplitude for the fuel assembly is 0.010 in., and less than 0.005 in. for the fuel rod. The fuel rod vibratory amplitude correlates with the masured amplitude obtained from a test on a 3 x 3 fuel rod assembly. In order to substantiate what is believed to be a conservatively calculated amplitude for the fuel assembly, a direct masurement will be obtained for a full size prototype fuel assembly during testing of the assembly in the Control Rod Drive Line Facility (CRDL) at the B&W Research Center, Alliance, Ohio. (3) Demonstration In addition to the specific items discussed above, the overall mechanical performance of the fuel assembly and its i' dividual components is being demonstrated in an extensive experim ntal program in the CRDL. 3 2.4.3 Control System 3 2.4.3 1 Control Rod Drive System Design Criteria The control rod drive system shall be designed to met the following perfor-mance criteria:

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

b. Uncontrolled Withdrawal No single failure or chain of failures shall cause uncontrolled with-drawal of any drive.

O .

        '  ,;s   .) -

Q c. Equipment Removal The disconnection of plug-in type connectors, modules, and subassem-blies from the protective circuits shall be annunciated or cause a reactor trip.

d. Control Cluster Assembly (CCA) Trip The trip connand shall have priority over all other commands. Trip action shall be positive and nonreversible. 'Irip circuitry shall provide the final protective action and shall be direct-acting, in-cur minimum delay, and shall not require external power. Circuit interrupting devices shall not prevent reactor trip. Fuses, where used, shall be provided with blown indicators. Circuit breaker po-sition information shall also be indicated.

e. CCA Insertion Insert command shall have priority over withdraw command. The drive actuator will be capable of overcoming a " stuck rod" condition equiv-alent to a 400 lb weight.

f. Withdrawal The control system allows only two CCA groups out of four regulating

 'O     CCA g oups to withdraw at any time subject to the conditions de-scribed in 7 2.2.1.2.

g. Position Indication Continuous position indication, as well as upper and lower position limit indication, shall be provided for each control rod drive as-sembly. The accuracy of the position indicators shall be consistent with the tolerance sat by reactor safety analysis.

h. System Monitoring The drive controls shall incl 2de provisions for monitoring conditions which are important to safety and reliability. These include such things as detecting power supply troubles and detecting motor or mechanism abnormal conditions.

i. Drive Speed The drive controls shall provide for single uniform speed of the mechanism. The drive controls, or mechanism and motor combination, shall have an inherent speed-limiting feature. The speed of the mechanism shall be 25 in./ min plus or minus 10 per cent of the pre-determined value for both insert and withdraw motion. The withdraw speed shall be limited so as not to exceed 25 per cent overspeed in he event of speed control fault.

l O i l 00 CONG Alto 3-66 1

s J. g hanical Stops The drive mechanisms shall be provided with positive, mechanical stops at both ends of the stroke or travel. The stops shall be capable of receiving the full operating force of the mechanisms without fa ilure . 3.2.4.3.2 Control Rod Drive Assembly The control rod drive assemblies provide for controlled withdrawal or in-sertion of the cluster control assemblies (CCA) out of or into the reactor core to establish and hold the power level required. The drive assemblies are also capable of rapid insertion or trip for emergency reactor condi-tions. The control rod drive assemblies are buffer seal, rack and pinion-type drives under development by Diamond Power Specialty Corporation. The control rod drive assembly data are listed in Table 3-19. Table 3-19 Control Rod Drive Design Data Ite_m Data Nunber of Drives 69 Type Buffer Seal, Rac.t and Pinion Location Top-Mounted / Direction of Trip Down Velocity of Normal Withdrawal and Insertion 25 in./ min Maximum Trip Time for 2/3 Insertion 1.4 see Length of Stroke 139 in. l Design Pressure 2500 psig ' Design Temperature 650 F l l A control rod drive assembly consists of a rack housing, snubber bottoming spring assembly, rack, rack pinion, coupling assembly, drive shaft housing, miter gear set, drive shaft assembly, buffer seal assembly, magnetic clutch, gear reducer, drive motor, position indication transmitters and limit switch system. The spool piece joins the drive assembly to the reactor closure head nozzle as shown in Figure 3-51. The d,ive motor supplies torque through the magnetic clutch to the drive shaf t-gear system to provide vertical positioning of the rack. The control rod drive assembly is shown on Figures 3-51 and 3-52. Sub-assemblies of the control rod drive are described as follows:

           .. .I   .I 3-67 (Revised 4-29-67)               ,

a. Rack Housing Subassembly The rack housing subassembly houses the hydraulic snubber, the bottoming spring assembly, the rack, rack pinion assembly and a rack guide bushing. The lower guide tube is attached to the lower end of the rack housing, and the cap and drive line vent assembly is mounted on the upper end of the rack housing. The hydraulic snubber decelerates the moving elements of the drive at the end of travel by controlled orificing of reactor coolant' water. The bot-toming spring assembly absorbs the bottoming impact in a stack of spring washers. The rack is guided by an upper shoe attached to the upper end of the rack, a rack guide bushing located at the pinion, and a lower guide tube bushing located at the lower end of the lower guide tube. The rack pinion is carried by two ball bearings. The valve on the cap and drive line vent assembly is used to bleed air or gases from the rack housing during reactor start-up. The removal of this assembly provides the access for CCA coupling and uncoupling and for securing the racks in the retracted position when the reactor closure head or individual drives are to be removed.

b. Drive Shaft Housing Subassembly The drive shaft housing subassembly houses the miter gear set, the drive O shafts and their supporting ball bearings. The drive shaft assembly is made up of two shaft assemblies with an intermediate bearing to raise the critical speed of the assembly.

This subassembly is attached to the rack housing subassembly by four through-bolts. All gasketed joints are of the double Conoseal-type with a pressure test-ing tap between the seals. I

c. Buffer Seal Assembly l

A prensure breakdown-type seal is employed to seal the drive shaft pene-tration in the primary pressure container. Seal system water is injected betwe.in the eighth and ninth stages of a 9-stage seal to provide a con-trolled leakage of approximately 3 gal /hr into the reactor coolant system and 12 gal /hr to the letdown storage tank. The seal water is cooled below 120 F, demineralized and specially filtered before injection into the seal. A conventional rotary seal is employed to prevent seal water from entering the drive package.

d. Drive Package The drive package is a synchronous type containing a synchronous motor, a self-locking worm gear reducer, a magnetic clutch, position indication O transmitters and a limit switch system. In conjunction with the magnetic h clutch is a unidirectional mechanical clutch which will allow the motor to drive the rods down following a trip. The motor has inherent braking c 3-68 (Revised 6-16-67)

O, so no separate brake is required. The self-locking worm gear reducer prevents torque feedback to the motor.

o. Position and Limit Switch Transmitters  :

The position transmitters and limit switches are located between the buffer seal and the magnetic clutch and supply redundant position sig-nals and limit switch contacts. There are three separate devices included in the position and limit switch transmitter assembly. A potentiometer generates an analog posi-tion signal; a linear variable differential transformer (LVDT) generates both an analog position signal and limit contacts; and the limit switch mechanism provides limit contacts. Refer to Figure 3-58. The potentiometer is geared directly to the drive shaft and gives a continuous de signal proportional to the CCA position. The LVDT trans-mitter has a core that is moved by means of a ball screw mechanism geared to the drive shaft. A demodulator located within the control cabinet contains the necessary electronic circuitry to gen'erate the analog de signal. 'Ihis demodulator also has relays with adjustable set points for position contacts. The limit switch assembly consists of switches operated by linear cams that are m m d by a ball screw assembly. This is also geared directly to the drive shaft. By using these three transmitters, it is possible to get both redundant e position and redundant limit signals. O . 3-69 (Revised 6-16-67) l[

A h 32.433 Control Rod Drive Assembly Control System The control system for the rod drive mechanism is designed to energize and po-sition the drive, indicate the control cluster assembly (CCA) position in the core, and indicate malfunctions in the system. As shown on Figure 3-57, the control system consists of: Power supplies and monitors. Clock (CCA speed standard).

>              Drive mechanism grouping panel.

Individual CCA control logic. Position indicator system. Travel limit system. Automatic sequence logic. Trip system. The rod mechanism control system provides the reactor operators with the flex-O ibility of CCA grouping, manual or automatic group operation, automatic OCA group sequencing, and infor=ation of CCA position in the core. A total of 12 CCA groups is available through facilities of a drive mecLanism grouping panel which enables up to 12 CCA's to be assigned to each group. In-dividual position indicators are provided for all 69 CCA's and are visible to the operator. The operating control panel includes four group position indi-cators. Associated with each of these four indicators is a switch which se-1ects CCA position data from a single CCA in each group. Three of the indica-tors are assigned to groups A, B, and C. The other is assigned to groups D through L. In addition, individual CCA selection is achieved through these switches for single CCA trim by manual switch action. CCA groups are pro-grammed such that the power peaking values 1kted in Table 3-1 are never ex-ceeded. Automatic sequencing (group overlap) of groups A through D is provided nnd is available for automatic or manual operator CCA motion requirements. It allows a limited overlap of operation of any two groups in a fixed sequence, but no more than two. Inputs from CCA position and travel limits feed this system. Automatic and manual control is provided. In " automatic," the selected drive mechanism group receives an automatic command signal from the reactcr control i system. In " manual," provision is made for operation of any individral CCA or group of CCA's. Manual or automatic operation of four CCA groups in a preset sequence is provided as described above. Grouping is determined at the drive p mechanism grouping panel prior to reactor operation. t ine drive gate is part of the individual CCA control logic circuitry which per-forms the function of selection and gating. It receives inputs from the clock, 3-70 (Revised 4-29-67)

the IN and OLTI control busses, motion " Enable," and travel limits. The drive gate sends pulses to the translator upon receiving (1) clock pulses, (2)

 " Enable" input, and (3) an IN or OUT control signal. End travel limits          and the driver monitor provide inputs to stop CCA motion.

Output signals of the drive gate feed into the translator. This unit pro-duces the proper signals for the drive motor. Direction la determined by the IN and OtTI commands, and speed is determined by the fixed clock fre-quency. The position indication and travel limit systems consist of three different types of transmitters and produce two independent analog position signals and two independent limit signals. One of the devices, the LVDT, produces both position and limit signals. Either source of signals can be used for the position and for the limit signals. Position output jocks are provided for a precision meter and for computer monitoring. Calibration of the potentiometer and the LVDT is accomplished by initial adjustments prior to installing the power package and also by making adjustments within the control cabinet. The limit switches are ad-justed prior to installation of the drive package. A fault detection circuit monitors signals to provide extra protection against unwanted withdrawal and insertion motion. See Figure 3-57. Trip is initiated by de-energizing two series circuit breakers in each of two power sources (Figure 3-59). O Each loss-of-voltage trip coil is fed by a sep-arate two-out-of-four relay circuit powered by four inputs from the reactor protective sys tem. Failure of any two inputs causes trip. The manual trip pushbutton opens all trip circuit breakers. Test pushbuttons are provided to test each circuit breaker action. 3.2.4.3.4 Control Rod Drive Sys tem Evaluation

a. Design Criteria The System will be designed, tested, and analyzed for compliance with the design criteria. A preliminary safety analysis has been made on the control rod drive motor control subsystem for failures of logic functions. It was concluded that no single failure in any CCA control would prevent CCA insertion, nor cause inadvertent l CCA withdrawal of another CCA or CCA group.

l b. Materials Selection Materials are selected to be compatible with, and operate in, the reactor coolant. Certified mill test reports containing chemical analysis and test data of all materials exposed to the reactor system fluid shall be provided and maintained for the control rod s

                     ;D               3-71 (Revised 4-29-67)

• iqd" i+

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

c. Relation to Design Temperature All parts of the control rod drive assembly are designed to operate at 650 F, although it is expected that all parts will 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. It is expected that the more significant temperature changes will be caused by displacement of reactor coolant in and out of the eachanism water space as the drive line is raised and lowered.

d. Design Life The expected life of the control rod drive control system is:

(1) Structural portions, such as flanges and pressure housings, O- have an expected life of 40 years. (2) Moving parts, such as rack, pinions, and other gears, have an expected life of 20 years. (3) Electronic control circuitry has an expected life of 20 years. 3.2.4.3.5 Control Cluster Assembly (CCA) q Each control cluster assembly is made up of 16 control pin assemblies which are coupled to a single Type 304 stainless steel spider cluster assembly (Fig-ure 3-60). Each control pin assembly consists of an absorber section of silver-1 l 3-72 (Revised 4-29-67)

indium-cadmium poison clad with cold-worked, Type 304 stainless steel tubin6 Ci and Type 304 stainless steel upper and lower end pieces. The end pieces are welded to the clad to form a water and pressure-tight container for the poison. The control pin assemblies are loosely coupled to the spider cluster assembly in order to permit maximum conformity with the channels provided by the guide tubes. The CCA is inserted through the upper end fitting assembly of the fuel assembly, each control pin assembly 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 pin assemblies is provided throughout the stroke. With the reactor assembled, the CCA cannot be withdrawn far enough to cause disengagement of the control pin assemblies from these incere guide tube assemblies. Pertinent design data are shown in Table 3-20. Table 3-20 Control Cluster Assembly Design Data Item Data Number of Cluster Assemblies 69 Number of Control Pins per Cluster Assembly 16 Outside Diameter of Control Pin, in. 0.440 Os Cladding Thickness, in. 0.019 Cladding Material Type 304 SS, cold-worked Poison Material 80% Ag, 15% In, 5% Cd Length of Poison Section, in. 134 Stroke of Centrol Pin, in. 139 This type of CCA has been developed under the USAEC Large Reactor Development

Program and offers the following significant advantages:

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

b. Shorter reactor versel and shorter internals due to the elimination of control rod followers.

c. Lower reactor building requirements due to the reduction of reactor coolant inventory.

d. Better core power distribution for a given CCA worth.

A CCA prototype similar to the B8M design has been extensively tested (58) at reactor temperature, pressure, and flow conditions under the LED program. 3-73 00 0H+4@~3

The silver-indiu=-cadmiu= poison caterial is enclosed in stainless steel tubes to provide structural strength to the centrol pin asse=blies. These pins are designed to withstand all operating loads includin; those resulting from hy-draulic forces, themal gradients, and reactor trip deceleration. The cladding of the poison section also prevents corrosion and eliminates pos-sible silver conta=1 nation of the reactor coolant. The ability of the poison clad to resist collapse due to the syste= pressure has been de=onstrated by an extensive collapse test pro;ra= on cold-worked stainless steel rods. The actual collapse carsins are higher than the re-quire =ents. Internal pressure and poison swelling are not expected to cause stressin; or stretching of the clad because the A 3-In-Cd alloy poison does not yield a ;as-ecus product under irradiation. Eecause of their great length and unavoidable lack of straichtness, so=e sli st =echanical interference between control pins and guide tubes must be expected. IIovever, the parts involved, especially the control pins, are so flexible that only very s=all friction dra;s will result. Si=ilarly, thermal distortions of the control pins are expected to be s=all because of the lov heat ceneration and adequate cooling. Consequently, it is not anticipated that the control pin asse'::blies vill encounter significant frictional resistence to their motion in the ;uide tubes. Overall performance of the COA vill be de=:nstrated in the EU Control Red Drive Line Facility (CRDL). 1 01a

             <r,                                                                     l bb    '

3 74 l

33 TESTS AND HISPECTIONS 331 NUCLEAR TESTS AND INSPECTICN 3 3 1.1 Critical Experiments An experimental program is presently underway to verify the control cluster assembly (CCA) design as to reactivity worth, and hence the safety aspect of the CCA. Detailed testing is underway to establish the worth of a CCA and a group of CCA's under various conditions similar to the reference core. These parameters include CCA arrange =ent, fuel enrichments, fuel element metal-to-water ratios, CCA materials, and soluble boron concentration in the moderator. A part of the CCA test program includes the study of local power peaking both between fuel assemblies and around the water holes left by withdrawn CCA's. Additional data on the vorth of soluble boron under various fuel assembly and - water hole arrangements will also be produced by this program. Zero power tests vill be made to establish the worth of soluble boron and the CCA's at the reactor site for both cold and hot conditions. CCA group vorths will be determined for all operating patterns to insure that the various limits placed on a single SCA or CCA pattern worth are not exceeded, and that there is sufficient reactivity always fully withdrawn to provide reactor shutdown vita the l most worthy CCA stuck out. A g 3 3 1.2 Zero Power, Approach to Power, and Power Testing Boron worth and CCA vorth (including stuck CCA vorth) vill be determined by phy-sics tests at the beginning of each core cycle. It is expected that recalibration of boron worth and CCA vorth will be performed at least once during each core cycle. Calculated values of boron vorth and CCA vorth will be adjusted as nec-essary to the test values. The boren worth and CCA vorth at a given time in core life vill be based on CCA position indication and calculated data as adjusted by experimental data. Periodic laboratory analysis of the reactor coolant is performed to determine the boron concentration. The reactivity held in boron is then calculated from the concentration and the reactivity worth of boron. The method of =aintaining the hot shutdown margin (hence stuck CCA margin) is related to operational characteristics (load patterns) and to the power peaking l restrictions upon CCA patterns at power. The CCA pattern restrictions vill be nch that sufficient reactivity is always fully withdrawn to provide adequate shutdown with the stuck CCA margin. Power peaking as related to CCA patterns and shutdown margin vill be monitored by reactivity calculations, and interlocks vill l be provided to prevent CCA patterns which produce excessive power peakinE and/or reduction of shutdown ma gin. l Operation under all power conditions vill be monitored by incore instrumentation, and the resulting data analyzed and compared with multidimensional calculations p in a continuing effort to provide sufficient support for further power escalations. (vl 0 3.,s 00*H59AS

332 THERMAL AND HYDRAULIC TESTS AND INSPECTIoH 3 3 2.1 Reactor Vessel Flow Distribution and Pressure Drop Test A1/6scalemodelofthereactorvesselandinternalsvillbetestedtoachieve the following objectives:

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

b. To measure fluid mixing between the vessel inlet nozzle and the core inlet, and between the inlet and outlet of the core.

c. To measure the overall pressure drop between the vessel inlet and out-let nozzles, and the pressure drop between various points in the reac-tor vessel flow circuit.

The reactor vessel, thermal shield, flow baffle, core barrel, and upper plenum assembly are cade 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 +a-G , surements vill be taken in the inlet and outlet nozzles of the reactor codel and at the inlet and outlet of each of the fuel assemblies. Static pressure taps vill l be located at suitable points along the flow path through the vessel. This in- l' strucentation vill provide the data necessary to accceplish the objectives set forth for tae tests. l 3 3 2.2 Fuel Asse=bly Heat Transfer and Fluid Flow Tests 3W is conducting a continuous research and development progra= for fuel assembly heat transfer and fluid flow applicable to the design of the Duke Power Cc=pany reactors. Single channel tubular and annular test sections and cultiple rod assemblies have been tested at the 3 W Research Center. 3 3 2.2.1 Single Channel Heat Transfer Tests A larse quantity of uniform flux, single channel, critical heat flux data has been obtained. References to uniform flux data are given in 3AW-168 and 3 2 3 1.2 of this report. The effect on the critical heat flux caused by nonuniform axial power generation in a t b .ar test secti on at 2,000 psi pressure was investi-jated as early as 1961. This prosrsm was extended to incit;de pressure 1,000, 1,500, and 2,000 psi and mass velocities up to 2 5 x 106 lb/hr-fta , 5 The effect on the critical heat flux caused by differences in the radial and axial power distribution in an annular test section was recently investigated at reactor desi3n conditions.(60) Data vere obtained at prpssurcs of_l,0Co,1,500, 2,000, and 2,200 psi and at mass velocities up to 2 5 x loc lb/hr-ftc. c A 3-76

3.3.2.2.2 Multiple Rod Fuel Assembly Heat Transfer Tests v Critical heat flux data are being obtained from 4 ft and 6 ft long, 9-rod fuel assemblies, in a 3 x 3 square array. The operating conditions include , pressures up to 2,400 psi and mass velocities up to 3.0 x 106 lb/hr-ft2 The l simulated fuel rod diameter and pitch are identical with the reactor core. l The data obtained to date have verified the design procedure. Data will be obtained from a longer, 9-rod fuel assembly to provide additional confidence  : in the present thermal design and to lead to increased reactor power capability.  ! 3.3.2.2.3 Fuel Assembly Flow Distribution and Pressure Drop Tests l Flow visualization and pressure drop data have been obtained from a 10 times j

 - full scale model of a single rod in a square ficw channel. These data have        -

been used to refine the spacer ferrule designs with respect to mixing turbu-lence and pressure drop. ) l Flow distribution in a square, 4-rod test assembly has been measured. A salt  ! solution injection technique was used to determine the average flow rates in the simulated reactor assembly corner cells, wall cells, and unit cells. Inter-channel 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 pressure drop data will be obtained to improve the core power capability. The following fuel assembly geometries will be tested to provide additional data:

a. A 3 x 3 array identical to that for which critical heat flux data have been obtained to provide additional interchannel mixing data,

b. A 4 x 6 array divided in half by a perforat.ed plate simulating adja- )

cent fuel assemblies to provide data on mixing between assemblies,

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

3.3.2.3 Preoperational Testing and Postoperational Testing Thermocouples are included as part of the incore instrumentation assembly and will enable postoperational temperature measurements to be made at the entrance and exit of all 51 instrumented fuel assemblies. The results of these tests , will be compared to the results of the model tests which were used to design l calculations. 3.3.3 FUEL ASSEMBLY, CONTROL CLUSIER ASSEMBLY, AND CONTROL ROD DRIVE ASSEMBLY MECHANICAL TESTS AND INSPECTION To demonstrate the mechanical adequacy and safety of the fuel assembly, con-trol cluster assembly (CCA), and the control rod drive assembly, a number of functional tests have been performed , are in progress , or are in final stages of preparation. b W 00-00N 3-77 j

3331 prototype Testing A full scale prototype fuel assembly, CCA, and control rod drive assembly is presently being tested in the Control Rod Drive Line (CRDL) Facility located at the B&,1 Research Center, Alliance, Ohio. This full size loop is capable of si=ulating reactor environmental conditions of pressure, temperature, and coolant flow. To verify the mechanical design, operating compatibility, and characteristics of the entire control rod drive assembly-fuel asse=bly system, the drive assembly will be stroked and tripped in excess of expected operating life requiremeV s. A portion of the testing will be perfor=ed with =aximum misalignment c aditions. Equipment is available to record and verify data such as fuel assemb ty pressure drop, vibration characteristics, hydraulic forces, etc., and to di onstrate control drive operation and verify scram times. All prototype compnents will be eynMned periodically for si6ns of material fret-tin 6, wear, and vibration / fatigue to insure that the mechanical design of the equipment meets reactor operating requirements. After the prototype fuel assembly has been tested under simulated reactor op-erating conditions, it will be installed in the full size low pressure loop to verify specific fuel assembly desi 6n data. These data include pressure drop, coolant interchannel mixing, and coolant velocity profiles. 3332 Model Testing Many functional improvements have been incorporated in the desi 6 n of the proto-type fuel assembly as a result of model tests run to date. For example, the spacer grid to fuel rod contact area was fabricated to 10 times reactor size, , and tested in a loop simulating coo 3 ant flow Reynolds numbers of interest. Thus, visually, the shape of the fuel rod support areas was optimized with re-spect to minimizing the severity of flow vortices. Also, a 9-rod (3 x 3) actual size model was fabricated (using production fuel assembly materials) and tested at 640 F, 2,200 psi, and 13 fps coolant flow. Principal objectives of this test weretoevaluatefuelrodcladdingtospacerBridcontactwear,and/orfretting corrosion resulting from flow-induced vibration. A wide range of contact loads (including s=all clearances) was present in this specimen. No significant wear or other flow-induced anmnge was observed after 210 Cays of loop operation. 3333 Componentand/orMaterialTesting 33331 Fuel Rod Cladding Extensive short t6:e collapse testing was perfomed on 1.ircaloy-4 tube speci-mens as part of the B&W overall creep-collapse testing program. Initial test speciuens were 0.436 in. OD with wall thicknesses of 0.020 in., 0.024 in., and 0.T8 in. Ten specimens of each thickness, 8 in. long, were individually tested at 680 F at slowly increasin6 pressure until collapse occurred. Collapse pres-sures for the 0.TO in. wall thickness specimens ran6ed from 1,800 to 2,200 psig, the 0.T4'in. specimens ranged from 2,800 to 3,200 psig, and the 0.T8 in, specimens ran6ed from 4,500 to 4,900 psig. The material yield stren6th of these specimens ranged from 65,000 to 72,000 psi at room temperature, and was 35,800 psi at 680 F. Additional 2.ircaloy 4 short time collapse specimens were l prepared with a material yield stress of 78,000 psi at room temperature and 48,500 psi at 615 F. Fifteen specimens having an OD of 0.410 in. and an ID of l 0 365 in. (0.0225 in, nominal vall thickness) were tested at 615 F at increasin6

          ,,   r

O pressure until collapse occurred. 4,960 psig. Collapse pressures ranged from 4,470 to Creep-collapse testing was performed on the 0.436 in. OD specimens. Twelve specimens of 0.024 in. wall thickness, and 30 specimens of 0.028 in. wall i thickness, were tested in a single autoclave at 680 F and 2,050 psig. During l this test, two 0.024 in. wall thickness specimens collapsed during the first 30 days, and two collapsed between 30 and 60 days. None of the 0.028 iri. wall thickness specimens had collapsed after 60 days. Creep-collapse testing was then performed on thirty 0.410 in. OD by 0.365 in. ID (0.0225 in. nominal wall) , specimens for 60 days at 615 F and 2,140 psig. None of these specimens co?-  ! lapsed, nor was there any significant increase in ovality after 60 days. Results of the 60 day creep-collapse testing on the 0.410 in. OD specimens showed no indication of incipient collapse. The 60 day period for creep-col-lapse testing is used since it exceeds the point of primary creep of the mate-rial, yet is sufficiently long to enter the stage when fuel rod pressure begins to build up during reactor operation, i.e. , past the point of maximum differ-ential pressure that the clad would be subjected to in the reactor. In order to help optimize the final clad thickness, additional clad-collapse testing is scheduled for 1967 using specimens fabricated to the reference de-i sign fuel clad dimensions, material specifications, and operating conditions. 3.3.3.3.2 Fuel Assembly Structural Components The mechanical design of the prototype can panel assembly is the result of an extensive can panel design and structural evaluation program. The full size, simulated loop functional testing as noted in 3.3.3.1 is expected to verify I i e can panel design criteria. Prototype static and dynamic load testing is under-way to verify can panel structural adequacy for vibration, handling, operation and seismic loads. a In the mechanical design of the spacer grids particular attention is given to I the ferrule-to-fuel-rod contact points. Sufficient load must be applied to position the fuel rods and to minimize fuel rod vibration, yet allow axial thermal differential expansion, and not produce fretting wear in the fuel rod cladding. Static load and functional testing of the prototype grids will dem-onstrate their adequacy to perform within the design requirements. 3.3.3.4 Control Rod Drive Assembly Tests and Inspection 3.3.3.4.1 Mechanism Developmental Tests 1 The prototype buffer seal act "ar is under development at the B&W Research l Center. Wear characteristics of critica components such as sleeve bearings, pinion l and rack teeth, snubber piston ,nd sleeve, etc,during' tests to date indicate l

that material compatibility anc structural design of these components will )

I be adequate for the life of the mechanism. j H i 0 0 AUU4it f1 f A 7 i 3-79 (Revised 4-29-67) it9--r- r 'e?.w- y-y vr q.. w p , e-L-.r.s-u y,r 9-.y --

                                                                                                             .+ m       y  - - - , . s-w-.-. -am--c- --- m        -.w.-----..-w,,

Subsequent to completion of the actuator development program, the complete prototype control rod drive assembly will be subjected to environmental testing under simulated reactor conditions (except radiation) in the Control Rod Drive Line (CRDL) Facility at Alliance. Environmental tests will in-clude, but not necessarily be limited to, the following: Operational Tests Operating speeds. Temperature profiles. Trip times for full and partially withdrawn control cluster assem-blies (CCA) for various flow-induced pressure drops across the CCA. Life Tests (With internals assembled to maximum misalignment permitted by drawing dimensions and tolerances-) 2,500 Partial stroke (75 per cent) cycles. 2,500 Full stroke cycles. 25 Partial stroke (40 per cent) trip cycles. 175 Partial stroke (75 per cent) trip cycles. 200 Full stroke trip cycles. Misalignment Tests 100 Full strokes and 100 full stroke trips with internals tolerances altered to 1.5 times maximum allowable mis-alignment. Coupling Tests Complete check of coupling operations af ter testing. The above cycles meet the total test requirements of 5,000 full strokes and 500 trips. Complete disassembly and inspection of the assembly will be accom-plished at various B&W facilities after completion of environmental tests. 3.3.3.4.2 Control Rod Drive Assembly Developmental Tests A control rod drive assembly motor control unit has been built in breadboard I form. Following the testing of this breadboard version, prototype circuits

     % .   "q '

Ik , 3-80 (Revised 4-29-67)

for plu6-in modules will be desi6ned and tested. Testing will censist of bench

     \    testing, life testing, and determining effects of simulated failures. The cim-ulated failure testing will be designed to verify the safety analysis.

The rod drive control system package will be tested in conjunction with the rod drive motor control package to insure proper operation. Simhted failure test-ing vill also be performed on the combined system to insure that protective re-quirements are being met. The position indicator and limit switch subsystem has been built in prototype form and life-te"'ed mechanically under expected environmental conditions. Further testing, both mechanically and electrically, will be done under ex-pected environmental conditions at the B&W Research Center. Characteristics to be determined will include accuracy, repeatability, linearity, short term stability, and long term stability. 3 3 3.4.3 Production Tests The finished control rod drive assembly will be proof-tested as a complete sys-tem; i.e., mechanisms, motor control, and system control working as a system. This proof testing will be above and beyond any developme'2tal testing which is

 ,        performed in the product development stages.

Mechanism production tests will include the following: O a. Ambient Tests V Coupling tests. Operating speeds. Position indication. Trip tests,

b. Operational Tests .

Operating speeds. Position indication. Partial and full stroke cycles. Partial and full stroke trip cycles. Control system assembly production tests will be performed as described in the fo11owin6 Paragraphs. The finished hardware will be systematically operated through all of its oper-ating codes, checked ove c the full range of all set points, and checked for proper operation of all patch plugs. This will check completeness and proper functioning of wiring and components. 3-81 40 00:48 D( I

The operating modes to be checked will include such things as automatic oper-ation, =anual Group operation, trim or single CCA operation, position indica-tion of all CCA's, travel limit on all CCA's, trip circuit operations, IU com-mand, OUI com=and, etc. The trip circuit or circuits will be tested by repeated operation. The overall trip time will be measured. The accuracy and repeatability of the position indication and limit switch sys-tems will be tested. Pe*ser supply tests will be perfomed to determine the upper and lower operating voltage and to prove immmity to switching transients. Fault conditions will be simulated to prove that no unsafe action results from defective components, circuits, or wiring. Ability to detect unsafe fault con-ditions at the operating console will be determined. Typical of faults which will be simulated are:

a. Defective limit switch or circuit.

b. I= proper CCA group patch.

c. Defective patch plugs,

d. Defective Broup sequencer. \

e. Defective clock.

f. Defective automatic control signal.

g. Defective command line.

h. Defective fuses.

i. Defective single CCA control circuit or switch.

j. Defective power supply.

k. Defective :totor translator.

1. Defective motor cable.

m. Defective position transmitter.

l The finished hardware will be visually inspected for quality of workmanship. This inspection will include an examination of the enclosure, cable entrances, dust tightness, maintenance features, drawers and cable retractors, fasteners, stiffeners, module mounts, wire harnesses, and other similar details. l O 9

                    ?S                                                                                                          l
                             )                                                                                                  i Q \ )' -

3-82 (Revised 4-29-67)

334 INTERNAIS TESTS AND INSm;n0NS The internals upper and lower plenum hydraulic design will be evaluated and guided by the results from the 1/6 scale model flow test which is described in detail in 3 3 2.1. These test results will indicate areas of gross flow maldistribution and allow verification of vessel flow-pressure drop computa-tions. In addition, the test results will provide measured pressure pulses at specific locations as an aid in assessing the vibration response character-istics of the internals components. The effects of internals misalignment will be evaluated on the basis of the test results from the CRDL tests described in 3 3 3 4. These test results, when correlated with the internals guide tube final design, will insure that the CCA will have the capability for a reactor trip or fast insertion under all modes of reactor operation in the reactor coolant environment. These tests will not include the effects of neutron flux exposure. After completion of shop fabrication, all internals components will be shop-fitted and assembled to final design requirements. The assembled internals components will be installed in a mockup of the as-built reactor vessel for final shop fitting and alignment of the internals for the mating fit with the reactor vessel. Dummy fuel and CCA's will be used to check out and insure that ample clearances exist between the fuel and internals structures guide tubes to allow free movement of the CCA throughout its full stroke length in various core

locations. Fuel assembly mating fit will be checked at all core locations.

  .q The dummy fuel and CCA's will be identical to the production components, except                                                                j that they will be manufactured to the most adverse tolerance space envelope; and even though the assembly weights will be representative of the production i

units, the dummy components will not contain fissionable or poison materials. Internals shop fabrication quality control tests, inspection, procedures, ard methods will be similar to the pressure vessel tests as described in detail in 4.1.4. With regard to the internals surveillance specimen holder tubes, the material irradiation surveillance program is described in 4.4 3 All internal components can be removed from the reactor vessel to allow in-spection of all vessel interior surfaces (see 4.4.1). Inspection of internals

components surfaces can be perfor=ed when the internals are removed to the canal storage location.

l l 1 l t l 3-88 Pa. _ _ _ . -

l /] V 34 mmalCES (1) Putnam, G. E., TOPIC - A Fortran Program for Calculating Transport of Particles in Cylinders, IDO-16968, April 1964. l (2) Avery, A. F., The Prediction of Neutron Attenuation in Iron-Water Shields, AEEW-R225, April 1962. (3) Bohl, H., Jr. , et al., P3M31, A One-Dimensional Multigroup P-3 Program ) for the Philco-2000 Computer, WAPD-TM-272.  ; (4) Bohl, H., Jr. and Hemphill, A. P., MRT-5, A Fast Neutron Spectrum Pro-gram for the Philco-2000, WAPD-TM-218. (5) Armster, H. J. and Callaghan, J. C., KATE-1, A Program for Calculatin6 Wigner-Wilkins and Maxwellian-Averaged Thermal Constants on the Philco-2000, WAPD-TM-232. (6) Marlowe, 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 Re-actor Inttices, BNL-5826. (8) Cadwell, W. R., Buerger, P. F., and Pfeifer, C. J. , The PDQ-5 and PDQ-6 O Programs for the Solution of the Two-Dimensional Neutron Diffusion-De-pletion Problem, WAPD-TM 477 (9) Marlowe, O. J., Nuclear Reactor Depletion Programs for the Philco-2000 Ccmputer, WAPD-TM-221. (10) Lathrop, K. D., ITfF-IV, A FORERAN-IV Program for Solving the Multigroup Transport Equat, inn With Anisotropic Scattering, IA-3373 (11) Joanou, G. D. and Ddek, J. S. , GAM-1: A Consistent P1 Multigroup Code for the Calculation o.' Fast Neutron Spectra and Multigroup Constants, GA-1850. (12) Baldwin, M. N., Physics Verification Experiments, CORE I, p 28 and Initial Conversion Ratio Measurements, BAW-TM 454. (13) Clark, R. H. and Pitts, T. G., Physics Verification Experiments, Core I, BAW-TM kSS. (14) Clark, R. H. and Pitts, T. G., Physics Verification Experiments, Cores II and III, BAW-TM kS8. (15) Clark, R. H. , Batch, M. L. , and Pitts, T. G. , Lumped Burnable Poison Pro-gram - Final Report, BAW-3492-1. n (16) Neuhold, R. J., Xenon Oscillation, BAW-305, 1966. 00 90150 ";L3C-n

(17) 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 Report BAW-168, November 1961.

(18) U.S.-Euratom Joint IBD Program, Burnout Flow Inside Round Tubes With Non-uniform Heat Fluxes, The Babcock & Wilcox Company, BAW-3238-9, May 1966. (19) Jens, W. H. and Inttes, P. A., Analysis of Heat Transfer Burnout, Pres-sure Drop, and Density Data for High Pressure Water, ANL-4627, May 1951. (20) Owen, D. B., Factors for One-Sided Tolerance Limits and for variable i Sampling Plans, SCR-607, March 1963 (21) 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, Pennsylvenia,1958. (22) USAEC Docket 50-24h, Exhibit D-3, entitled " Rochester Gas and Electric Corporation, Brookwood Nuclear Station Unit No.1", (Tvird Supplement to: Preliminary Facility Description and Safety Analys.'s Report, February 28,1966). (23) Lee, D. H. and Obertelli, J. D., An Experimental Investigation of Forced Convection Burnout in High Pressure Water. Part 1, Round Tubes With Uni-p form Flux Distribution, Am-R-213, August 1963 V (24) Matzner, B. and Griffel, J., Bimonthly Progress Report (BPR-XIII-11 and 12-63), Task XIII of Contract AT(30-3)-187, Basic Experimental Studies of Boiling Fluid Flow and Heat Transfer at Elevated Pressures, for November and December 1963, January 27, 1964. (25) Matzner, B. and Griffel, J., Monthly Progress Report (MPR-XIII 6-63), Task XIII of Contract AT(30-3)-187, Basic Experimental Studies of Boiling Fluid Flow and Heat Transfer at Elevated Pressures, for June 1963, June 28, 1963 (26) Matzner, B., Monthly Progress Report (MPR-XIII-5-63), Task XIII of Con-tract AT(30-3)-187, Basic Experimental Studies of Boiling Fluid Flow and Heat Transfer at Elevated Pressures, for May 1963, May 31, 1963 (27) Internal Memo, Weatherhead, R. J. to Lottes, P. A., critical Heat Flux (Burnout) in Small Diameter Tubes at 2000 psia, December 29, 1950. (28) Swenson, H. W., Carver, J. R., and Kakarals, C. R., The Influence of Axial Heat Flux Distribution on the Departure From Nucleate Boiling in a Water Cooled Tube, ASME Paper 62-WA-297 (29) Nonuniform Heat Generation Experimental Program, Quarterly Progress Re-port No. 7, January - March 1965, BAW-3238-7, Joint U.S.-Euratom B&D Program, AEC Contract No. ATC(30-1)-3236. O y/ (30) Hald, A., Statistical Theory With Engineering Applications, John Wiley

           & Sons, Inc., New York, 1955 e

UU UU r s n 3-@ 2r t-

(31) Worthing, A. G. and Geffner, J., Treatment of Experimental Data, John Wiley & Sons, Inc., New York, 1943 (32) Baker, o., Simultaneous Flow of Oil and Gas, Oil and Gas Journal, Vol. H , pp 185 - 195, 1954. (33) Rose, S. C., Jr. and Griffith, P., Flow Properties of Bubbly Mixtures, ASME Paper No. 6,-IIT-38,1965 5 (34) 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 M1T, June 1964. (35) Bergles, A. E. and Suo, M., Investigation of Boiling Water Flow Regimes at High Pressure, NYO-3304-8, February 1,1966. i (36) Notley, M. J. F., The Thermal Conductivity of Columnar Grains in Irra- I d!ated UO2 Fuel Elements, AECL-1822, July 1963 l l (37) Lyons, M. F., et al., UO2 Fuel Rod Operation With Gross Central Melting, GEAP 4264, October"' 1963 1 (38) Notley, M. J. F., et al., Zircaloy-Sheathed U Fuel Elements Irradiated . at Values of Integ3iIl kd6 between 30 and 83 cm, AECL-1676, December l 1962. l (39) Bain, A. S., Melting of UO2 During Irradiations of Short Duration, AECL-2289, August 1965 (40) Notley, M. J. F., et al., The Iongitudinal and Diametral Expansions of UO2 Fuel Elements, AECL-2143, November 1964. (41) Duncan, R. N., Rabbit Capsule Irradiation of UO ,2CVNA-142, June 1962. (42) Lyces, M. F., et al., UO2 Pellet Thermal Conductivity From Irradiations With Central Melting, GEXP k624, July 1964. (43) 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. (44) Ross, A. M. and Stoute, R. L., Heat Transfer Coefficients Between UO2 and Zircaloy-2, AECL-1552, June 1962. l (45) Hoffman, J. P. and Coplin, D. H., The Release of Fission Gases From Ura-nium Dioxide Pellet Fuel Operated at High Temperatures, GEAP-4596, j

     ~ September 1964.

(46) Spolaris, C. N. and Megerth, F. H., Residual and Fission Gas Release From Uranium Dioxide, GEAP 4314, July 1963 i (47) Robertson, J. A. L. , et al. Behavior of Uranium Dioxide as a Reactor Fuel, AECL-603,1958.- -

                                                                                                             +

l 3-@ 40 00i52 3 %

                                          'l c

l I I O (48) Parker, G. W., et al. , Fission Product Release From UO2 by High Tempera-V ture Diffusion and Melting in Helium and Air, CF-60-22-14, ORNL, February 1961. (49) Daniel, R. C., et al., Effects of High Burnup on Zircaloy-Clad, Bulk UO ,2 Plate Fuel Element Yamples, WAPD-263, September 1962. (50) Blomeke, J. O. and Todd, Mary F. , Uranium Fission Product Production as ) a Function of Thermal Neutron Flux, Irradiation Time, and Decay Time, l ORNL-2127, Part 1, Vol.1 and 2. (51) Duncan, R. N., CVTR Fuel Capsule Irradiations, CVNA-153, August 1962. (52) Frost, Bradbury, and Griffiths (AERE Harwell), Irradiation Effects in Fissile oxides and Carb'. des at Lov and High Burnup Levels, Proceedings of IAEA Symposium on Ra11ation Damage in Solids and Reactor Materials, Venice, Italy, May 1962. (53) Gerhart, J. M., The Pos t-Irradiation Evamination of a Pu0 2-UO2 Fast Re- l actor Fuel, GEAP-3833 1

                                                                                                                            )

(54) Atomic Energy Clearing. House, Vol. 22, No. 3, p 11.  ! (35) Large Closed-Cycle Water Reactor Research and Development Program Progress Report for the Period, January 1 to March 31, 1964, Westing-house Electric Corporation, Pittsburgh, Pa.,1964, WCAP-3269-2. Also WCAP-3269-3 for period from April 1 to June 30, 19 %. 4 (56) Physical and Mechanical Properties of Zircaloy-2 and -4, WCAP-3269-41, Figure 18. (57) Burgreer, D., Byrnes, J. J., and Benforado, D. M., Vibration of Rods Induced by Water in Parallel Flow, T_rans. ASME 80, p 991,1958. (58) Iarge Closed-Cycle Water Reactor P&D Program, Progress Report for the Period January 1 to March 31, 1965, 3 AP-3269-12. (59) Burnout for Flow Inside Round Tubes With Nonuniform Heat Fluxes, BAW-3238-9, May 1966. (60) Nonuniform Heat Generation Experimental Program, BAW-3238-13, July 1966. O - mm vu m m m, gg7

                                                                                    ,vulJJ 3-3P 9/.-                             _. __  ,-             .

O E g 2000 i O, 1800 f 1600 Z W O 1400 3 -m O Z 1200 s O ,

                          \

O m 1000 ,

                                \,

m 8 800 x, x Unit No. 1 H 600 N 3 O 400 Unit I.~o. 2 0 200

                                                      \

U 1 h 0 g 0 30 60 90 120 150 180 210 240 270 300 CORE LIFE, EFFECTIVE FULL POWER DAYS BORON CONCENTRATION VERSUS CORE LIFE OCONEE NUCLEAR STATION FIGURE 3-1 0000l St138 l l

7#/HHH/R E b8S'WN

    /                                   ,    .8
                   /                 /
      /                                                              \

,[ (( Fuel Zone

                             !                       \\

Fuel Zone 3 i;i:i:! Ni \ { 0. 6 iGii:{ .2 'i:i$i

                                          \      \               N

0. 2 Fuel Zone 2 \

                             \
                                            *\             PR hl\

sertion s x s N =

0. 4

   'N                  -

x XENON OVERRIDE ROD NS RTE (CONTOURS TAKEN RELATIVE TO PEAK) utjan OCONEE NUCLEAR STATION

                                               *~

FIGURE 3-2 - i; Oh C'

O O AXIAL POWER PROFILE FOR 55 % INSERTION IS SHOWN ON I FIGURE 3-4

                                                                          -l I

I

1. 7 '

                                                              /~

O # \ 10.

                                             /

h W -

1. 6

                                /
                                    /

3. x / t k

O i x e! 1x n  ;

        -W       1, 4 10       20          30           40            50                                          60          70                                   80 ROD INSERTION, %

AXIAL PEAK TO AVERAGE POWER VERSUS XENON OVERRIDE ROD INSERTION m 2 ourba OCONEE NUCLEAR STATION I FIGURE 3-3 m an nnicc L@ vm mv.ss

O

1. 8

1. 6 # ,

                )

g' / \ s g 1. 0 -'r

               /                         \-~-

x w 0. 8

             /                                    x
           /                                          \

!$ '/

0. 4 X g N

0. 2 g

                                                                   \     3
         ' '                         144" 0

l0 20 30 40 50 60 70 80 90 100 110 120 130 140 150 DISTANCE FROM BOTTOM OF ACTIVE FUEL, INCHES

                                                                              'l AXIAL POWER PROFILE XENON OVERRIDE RODS 55 PER CENT INSERTED uh =   OCONEE NUCLEAR STATION FIGURE 34             1 d

1 i:?'IIL UI p h

O

      +2
                  ,             I                                 m m,

f 53 / HX +1

 <N g)                              580 F            j 62                                           r On                                j                      s     '

H ~ 0 y 9 - Oh >

                                   \

H- L68 F

 <W                     /
 $$   ~'            /
 !!8 AO          j  /

O {

      ~

0 2 4 6 8 10 12 14 16 18 20 22 MODER ATOR BORON CONCENTRATION ppm BORON x 10-2 MODERATOR TEMPER ATURE COEFFICIENTS VERSUS BORON CONCENTRATION i m E Nwie OCONEE NUCLEAR STATION FIGURE 3-6 1 00 00156"3 %

I O-

           +2-l            l We b                                                               #

f N g* c:: s 2000 ppm # s

   $d                                      -

p e,"j

                  #                W                      00 ppm 9   . 0     7     #      -m-                    ,      .'

e$ M g 500 ppDN Od N

   <W dN N

h N WN -1 A o No Boron } AG

           -2                                                          .

100 200 300 400 500 6'00 MODERATOR TEMPER ATURE, F MODERATOR TEMPERATURE COEFFICIENTS VERSUS MODERATOR TEMPERATURE AND VARIOUS BORON LEVELS i OCONEE NUCLEAR STATION {cu.@m FIGURE 3-6 i P

V 100 i 90 i L 80

                         \
  @ 70 5                        \
  $ 60                      i e                          t A 50 b                            \

40 %) \

                                  \ 5. 6% Ak/k

8. 6% A k/k 20 g 4- -

0 1 2 3 4 5 6 TIME, SECONDS PER CENT INITIAL POWER VERSUS TIME FOLLOWING TRIP l

~

( OCONEE NUCLEAR STATION FIGURE 3-7

                                                                            &O 00157 A4.L                 '
                                                                                                            .)

                                                                        ~.

100

1. 90 -

d W H Finite Sample U 99% Confidence f 80 - O M 4 Z b 70 - 3 m L 9 k 60 - 50 , , , ,

1. 0 1. 2 1. 4 1. 6 1. 8 2. 0 DNB RATIO POPULATION INCLUDED IN THE STATISTIC AL STATEMENT VERSUS DNB RATIO OCONEE NUCLEAR STATION FIGURE 3-8  ;

i 1 I

 ,(6
       '?'

l

,n, V

1. 8 g 7 ~ P/ P = 1. 70 (Partial Rod

[ , Ins ertion)

       .5                   r
                             /                  X                 l

(,- I' ' s P/ P = 1. 50

1. 4 f A l g (Modified Cosine) 7 z

1. 3

                        /                     i
                                               /         \       l                \

l ,

                                            /                                         (

1. 2 \

l

                                                                                          \

J

                                       /

I iK ( g

1. 0

                                                                        \,

i , s 0. 9 ,/ l \g

0. 8 f

g

                                                                                                     \

V / I \

                            /                                  ,
                                                                                        \                \              l

o. s / / l \ \ l y- -

                                                                                                       \       s. ..

0. 3 l \ '

                                                                                                                        \

l Fuel Mid-Plane l

0. 2 l '

Core  ; Core h Bottom Top r 144" r

0. 0 0 l 20 40 60 80 ld0 120 l 140 10 30 50 70 90 110 130 DISTANCE FROM BOTTOM OF ACTIVE FUEL, INCHES POWER SHAPE REFLECTING INCREASED '

AX1AL POWER PEAK FOR 144 INCH CORE m

u. \

O]. t E P0wie OCONEE NUCLEAR STATION FIGURE 3-9 y 0L00158 2 Q (o

1.90 1.85-1.70-- 1.60-_ \ Line A (Design)

1. 50-- \

Line B (Nominal - Maximum Calculated) g 1.40-- g Line C (Typical True Distribution) O g y 1.30--

 <                                       N 4g   1.20--                               g CW                                             N 34   1.10--

N g$gO 1.00-- \

 ~-

oa 0.90-- O <1 N

• 0.80-- s W \

( 0.70--

                                                                         \

0.60-- \

                                                                             \

1.85 N 0.50-- (1) Line A = 1. 79 x Line B x 0.40.. (2) Line C for illustration only x (3) Line B is based on detailed data \ from a rod by rod printout of a \ 0.30-- pDQ (two dimensional power and

0. 20-- flux calculation for worst time \

in core life) ' O.10 - i > . . . i i . 0 l'O 26 36 '40 50 60 76 80 90 100 % PERCENTAGE OF FUEL RODS WITH HIGHER PEAKING FACTORS THAN POINT VALUES DISTRIBUTION OF FUEL ROD PEAKING l l OCONEE NUCLEAR STATION

                                                     ,                    FIGURE 3-10                                  m 1
          ,s                ,p
               -.q.
      ?;C'VU              .;

O 1 200 i

 <  180                                      %

N O

 >  160 H                          Line 1 3               Fah Nuclear - 1. 85 140           -

r 8 i / Q. 120

 <                     Line 2                            l    B g            Fah Nuclear - 1. 79
 >  100                                -

b [ Maximum m 6C A / Overpower C 4 40- g J e

                                      /
                                                            ~

M w J / o l 100 102 104 106 108 110 112 114 116 118 120 FULL POWER (2452 MW), % POSSIBLE FUEL ROD DNB'S FOR MAXIMUM DESIGN CONDITIONS - 36,816 ROD CORE I FIGURE 3-11 j 1 Ui

l 1 eJ m Z Q 100

 <                                                                            f
 $   80
 $   70 i

g I y 3 60 h A

 <   50 m                                                         /

5 40 O l g

                                                       /
 <                  Fah Nuclear - 1. 79          %

30 m Maximum g j / Overpower F O c: 10 7"

                                 /                 l W                   /                             l m            /

2 o I

 $      100  102   104   106   108    110    112  114    116 118   120 FULL POWER (2452 MW), %

POSSIBLE FUEL ROD DNB'S FOR MOST PROBABLE CONDITIONS - 36,816 ROD CORE m 1 a OCONEE NUCLEAR STATION FIGURE 3-12 q l l l .

            .c l

U(.  : O

fm i%j) (1 - P) (p) O.1

0. 9 I

0.01 , o,99 I

                      \
                       \
                        \                                                                               ,
                          \
                           \

l 0.001 6T 0.999 i

               \
                                 \        1 114%)
                                   \   a'
 /m             \                    r U                 \
                     \
                                     \
                                        \
                        \                 \
                           \                \

0.0001 \ k 0.9999 s S\ x x

                                                  \

F y y g 100%j g g

                                           \            \
                                               \                                                        l 0.00001                                                                            N 0.99999 0              10        20        30        40 50   60    70    80     90     100 PERCENTAGE OF RODS WITH A LOWER VALUE OF P DISTRIBUTION OF POPULATION PROTECTED P AND 1-P VERSUS NUMBER RODS FOR MOST PROBABLE CONDITIONS O)

(_ surrow OCONEE NUCLEAR STATION FIGURE 3-13 c b bb u h

1. 6 O-e O

Design Overpower 4 1. 5 l 0 4 Z O 1* 4

1.38 \

a _ _ _ . , _A Q W H N O 1. 3 W c: b a W

 <     l. 2 3

lI: O H u) W H H 1.1 O T 3 9 H

 <     1. 0                                 -         -           --

c: CQ Z Q 100 1.L O 120 1-30 140 150 REACTOR POWER,PER CENT OF 2452 MWth DNB RATIOS (BAW-168) VERSUS REACTOR POWER m ovat$wte w OCONEE NUCLEAR STATION FIGURE 3-14 _ l s- Ob i 13

1 O

     +24                                                                           79 l
     +22                                                                           81
     + 20                                                            r             83
                                                                                          ^
     +18                                                                           85 f                        f f                         O
     +16                                                      /                    88     W Design Overpower
                                                            /                           3 4
     +14

[ gW 91 Wp 40

     +12                                                                              O<

94 c:

 #                                                                                    Z; W ";'
     +10 g                                                                                 98 g o, t:
                                                                                      *E N+8 e                            r                                                  102      O W
     +6                 7'                                                       108      [
     +4           7
                   /                                                                  n 116 oD
     +2       ,
               /                                                                 129 ZG W :::

0 O / a ity Subcooled g44 E g 4

     -2
      -4
     -6 100    110        120      130                      140         150 RATED POWER, %

MAXIMUM HOT CHANNEL EXIT QUALITY VERSUS REACTOR OVERPOWER MEbwin OCONEE NUCLEAR STATION J FIGURE 3-15 e 00 00161M2-

O' 4.00 l l UO * *' 2 $ I e l 5 1 7 h 3.00 l c I 2 I H H z 2.00 O O [  ! J < l 2 e I W 's e Data Based on CVNA -142 June 1962 l 1.00  ! - 0 1000 2000 3000 4000 5000 TEMPERATURE, F THERMAL CONDUCTIVITY OF UO2 o DCONEE NUCLEAR STATION FIGURE 3-16 m

O 6000 g UO2 Melting Temperature l 5000 ---- - - - ----- - - --- I I b

 %  4000                                            #
                                                        /                                        l E
                                                  /l                          l 143 %

l

0. Power  !

2 N 3000 , [

 $                                                    I I

l I O

• a 2000 **#

I w 114% Power l l l 1000 N I I I ) 0 0 5 10 15 20 25 LINEAR HEAT R ATE, kw/ft FIGURE 3-17 FUEL CENTER TEMPERATURE AT THE HOT SPOT VERSUS LINEAR POWER u" rom OCONEE NUCLEAR STATION FIGURE 3-17 , _AA A A i z ^, VU U U i V L.

O 70- , Gaussian Distribution ' 60-( il 50- &'O - O g 40- .c - E $ 30-20- - 10-0 L O.60.7 0.8 09 1.0 1.1 12 1.3 1.4 1.5 1.6 1.7 1.8

                                      & E/& C NUMBER OF DAT A POINTS VERSUS $ E/ @ C es u m in    OCONEE NUCLEAR STATION FIGURE 318                _
   +
     )   *
           .i

() t 5. I

O ' 1.025 - 1.020 - F"0 l l l u 1.015 - 3 ,

   %  1.010 -                                                                      l N                                                                               1 1.-005 -                                                       O E

1,000- 60 70 80 90 100 - 995 -

       .990 -

I A l

       .985 -                                                                      l Population Protected, %                                      ,

l HOT CHANNEL FACTOR VERSUS PER CENT POPULATION PROTECTED e OCONEE NUCLEAR STATION FIGURE 3-19 y l nn vu nni e L vuiUJ

O-100 --- Infinite Sample 100% Confidence 90 g Finite Sample 90% Confidence l [v 80 - o 5

        $      70 ~                     Finite Sample
       ]                                99% Confidence
       ~3 C.

60 _ 50 , , 1.0 1.2 1.,4 1.6 1.,8 2.0  ! Burnout Factor (DNB Ratio) )

                                                                                           )

BURNOUT FACTOR VERSUS POPULATION FOR VARIOUS CONFIDENCE LEVELS esu,,4=r e OCONEE NUCLEAR STATION FIGURE 3-20 m 1 m + l

O 200 180 $ I 160 I j l 140 j - Q k 120 r  !

 $ 100                                    /                l l

3 1 8 80-l O 60 - l 7 40 - 20 s l

                            \                    \

1.00 1.05 1.10 1.15 1.20 FRACTION OF RATED POWER RODS IN JEOPARDY VERSUS POWER m on$wa OCONEE NUCLEAR STATION s

                              'Y                 FIGURE 3-21                              e 00          00164 2 5 Tr

m 10- - nn e Tr 0 ' . '2 ' . 's ' . h ' . '8 ' l .'o ' 1.'2 ' l'. 4 ' l ' 6 ' l' 8 ' 2.'0 * '2 . . (1) @EIN C WAPD-188 500 psia Data and BAW-168 20 . 10- - a 0- m j 0 i .2e i .4i i .6e i .8i i 1.0 i i 1.i2 e 1.4 i e i e i i . i 1.6 1.8 2.0>2 (2) $E! C o g WAPD-188 600 psia Data and BAW-168 1 2; G 10_ _ 0 r - -

                                                                                                             ~),

a i e i e i a i e i i e iiie i , ,m, a 0 .2 .4 .6 .8 1.0 1.2 1.4 1.6 1.8 2.0>2 (3) W/ E C WAPD-1881000 psia Data and BAW-168 10- - ei b i 0 F 0 a .2i i e is i i s i i e i a i e i

                           .4        .6      .8                 1.0                 1.2             1.4     1.6     1.8 2.0 >2 (4)                                    9/ E C WAPD- 188 1500 psia Dr.ta and BAW-168 RATIO OF EXPERIMENTAL TO CALCULATED BURNOUT HEAT FLUX i

out,he OCONEE NUCLEAR STATION l ' FIGURE 3-22 m W< L

                                                                                                 & fQ                               .

10-I 0 i i i . . . . . , , i r-i i i i e i i . i 0 .2 .4 .6 .8 1.0 1.2 1.4 1.6 1.8 2.07 2 (5) Wg /CC l WAPD-188 17 50 psia Data and BAW-168 l 70_ _ l 60-1 3 _ 1 d 50- _ t t

 .o E 4o_

b _ i 30-l 20- _ l

10. -

O i i . . . m. I. . . . . . . . > > n. .". . i l 0 .2 .4 .6 .8 1.0 1.2 1.4 1.5 1.8 2.0>2 (6) WE /IC WAPD-188 2000 psia Data and BAW-168 RATIO OF EXPERIMENTAL TO CALCULATED BURNOUT HEAT FLUX sust.hte OCONEE NUCLEAR STATION FIGURE 3-23 c 00 -00165 2foo

l I 10 - 0

                        . . . . . . .           . . . i e i e i .                                i i i .

0 2 .4 .6 .8 1.0 1.2 1.4 1.6 1.8 2.0>2 (7) g jg WAPD-188 2250 psia Data and BAW-168 10 - 0 i i . . . . i i i i i i i .O. . . . . . i 0 .2 .4 .6 .8 1.0 1.2 1.4 1.6 1.8 2.0 >2 0 E C

      $                             WAPD-188 2500 psia Data and BAW-16 3 o

10 g

      .c E

a Z 0 0 i i i i i i . . . . . . i i i i 2 .4 .6 8 1.0 1.2 1.4 1.6 M.i 1.8 2.0>2 (9) Og /0 C WAPD-188 2750 psia Data and BAW-168 20 - 10 - _ _ 0 4 - i i e i i i e 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 1.4 1.8 2.0 >2 (10) g jg G AEEW-R 213 560 psia Data and BAW-168 RATIO OF EXPERIMENTAL TO CALCULATED BURNOUT HEAT FLITX l uhu OCONEE NUCLEAR STATION FIGURE 3-24 c

    > ; I I II. I i

h

O 10-0 i e i : i e i i i i e i e a i i i : i i . 0 .2 .4 .6 .8 1.0 1.2 1.4 1.6 1.8 2.0 >2 (11) #E/8C AEEW-R-213 720 psia Data and BAW-168 50-40-a 30-

 .5                                        -

y - g 20- - o O E 2; - 10-0 i e i i e i i i e i . . . i . l. O. m. U. . i 0 .2 .4 .6 .8 1.0 1.2 1.4 1.6 1.8 2.0>2 (12) 9g/9C AEEW- R-213 1000 psia Data and BAW-168 10-0 M 0'.2 . '4 ' . '6 ' . '8 '1.'O'1$2'1!4' l'. 6 ' l'. 8 ' 2.'0 E2 (13) @E / @C AEE W-R-213 1300 psia Data and BAW-168 RATIO OF EXPERIMENTAL TO CALCULATED BURNOUT HEAT FLUX senbu OCONEE NUCLEAR STATION FIGURE 3-25 c 00 001662fo W

30 - _

                                                                  ~

20 - 10 - 0 1 f-

  .          0     . '2    .4     ' .b ' .b ' 1!0 ' lj2 ' 1!4 ' l'.6 ' 1.'8 ' 2!0-)2 E
 '8     (14)                                                @E! C (1

g AEEW -R-213 1500 psia Data and BAW-168 U

 .c      10-E O

O r, r r, e e, i e e i i e i e i i i i i i i i e i i e i 0 .2 .4 .6 .8 1.0 1.2 1.4 1.6 1.8 2. 0 >2 (15)

                                                            @E/N C Columbia 500 psia Data and BAW-168 10-
                                               -       ~

0 - l ,,

                                                #                i      i     e  i   a  i   ;  i   a  e  i   i 0       .2        4      .6        .8             1.0            1.2    1.4    1.6   1.8    2.0>2 (16)
                                                           $E!'O Columbia 720 psia Data and BAW-168 RATIO OF EXPERIMENTAL TO C ALCUI,ATED l

BURNOUT HEAT FLUX w - OCONEE NUCLEAR STATION FIGURE 3-26 q

    %v a63

D

 'd             50 -
                                                                    ~

40_

30. _

20 -

            .5  10_

o _

            .$   0 CE Z

l 0 . '2 .4 .'6

                                                                    . h ' 1.'0 ' l ' 2 ' 1.' 4 ' 1.'6 ' 1.'8 ' 2.b >h (17)
                                                                            @E! C Columbia 1000 psia Data and BAW-166 10-0 1    I   I       I   I      i     l          3     I       i    i       i   i  l  i  I   l   I I     I I O         .2          .4        .6              .8             1.0       1.2        1.4   1.6    1.8     2.0>2 (18)
                                                                            @E/N C Columbia 1200 psia Data and BAW-168
                                                                                                                                                  ]

RATIO OF EXPERIMENTAL TO C ALCULATED BURNOUT HEAT FLUX OCONEE NUCLEAR STATION FIGURE 3-27 y 1 00-0016f 26 l 1 1

i

                                                                                                        \

20 - 10 _ _ _ 0 I 9 e i i a a i i i i i e i i e i i e i e i i 0 .2 .4 .6 .8 1.0 1.2 1.4 1.6 1.8 2.0>2 (19) g /g Columbia 1500 psia Data and BAW-168 60-B g 50_ U

     .o Z     40-30_                                 -

20-10_ 0 r, r, a i e e i e i i1 i , e i i i i i i i i i 0 .2 .4 .6 .8 1.0 1.2 1.4 1.6 1.8 2.0 > 2 (20) $EINC Argonne 2000 psia Data and BAW-168 RATIO OF EXPERIMENTAL TO CALCULATED BURNOUT HEAT FLUX entrente OCONEE NUCLEAR STATION FIGURE 3-28 - c

   !     I

. a. ,f,.

l Ch NJ 10 - _ 0 F 0 .2 .4 .6 .8 1.0 1.2 1.4 1.6 1.8 2.0>2 (21) @EIO C B& W 2000 psia Data and BAW-168 l 10 -

    =

0 0 i i i i E - T e e e i e i i i i i s i i i i 4 0 .2 .4 .6 .8 1. 0 1. 2 1.4 1.6 1.8 2.0 72 (22) Og /0 C 3  ; Z Euratom 1000 psia Data and BAW-168 ' l l i 10- l l 0 i i i a a i i i i i i e i i i i i i i i 0 .2 .4 .6 .8 1.0 1.2 1.4 1.6 1.8 2.072 (23) $E!N C Euratom 1500 psia Data and BAW-168 RATIO OF EXPERIMENTAL TO CALCULATED BURNOUT HEAT FLUX m nEbu OCONEE NLICLEAR STATION FIGURE 3-29 , 90 00168 2(c(o

                                                     . a .-

20-10_ 0 i i i i i e i i e e i e i i e i e i e i i 0 .2 .4 .6 .8 1.0 1.2 1.4 1.6 1.8 2.0 *2 (24)

                                        $3/6c Euratom 2000 psia Data and BAW-158 50-
 .h S

Y

 }

E 40-o Z _

                                               ~

30- - 20- __ 10- _- " __Y _ _- T Ili,, -, o i i i e iiie i i i i e i i i e i i i i 0 .2 .4 .6 .8 1.0 1.2 1.4 1.6 1.8 2.0>2 (25) Og /9c All 500-720 psia Data and BAW-168 RATIO OF EXPERIMENTAL TO CALCULATED BURNOUT HEAT FLUX W OCONEE NUCLEAR STATION e FIGURE 3-30 -

(3 O 80 - 70-60 -

   ,  50 -
  .E 0  40_

t _ _

  .o E

a 2; I~ 30-20 _ _ _ 10_ -- 0 i i i i r-i i i i i i i i i . .

                                                                                   -I-       l I T ) ,. i 0   .2   .4         .6       .8     1.0       1.2        1.4        1.6        1.8 2.0>2 (26)                                 @EO C All 1000 psia Data and BAW-168 RATIO OF EXPERIMENTAL TO CALCULATED BURNOUT HEAT FLUX l'

(j row OCONEE NUCLEAR STATICN FIGURE 'n 3-31 y nn nn* UV vv .-,

80 - O 70 - 60 - 3 50 - .5 4 u 40 - .E E 30 - 20 - 10 - - 0 i i i i i i I i i i i 79 i i bi i s i rm O i ,

                                                  ,                                        i i

0 .2 .4 .6 .8 1.0 1.2 1.4 1.6 1.8 2.0>2 (27) @E C All 1500 psia Data and BAW-168 RATIO OF EXPERIMENTAL TO CALCULATED BURNOUT HEAT FLUX ovupu OCONEE NUCLEAR STATION FIGURE 3-32  ; 9

       * > Li l

122 - 1004"

• 90 - l 80 -

70 - 60 - B - 3 50 - 2 O t j 40 - 30 - 20 - ' 10 - _

                                                        ~

0 g

             , , , , , , , , i i e i i i br  iia m

a i i 0 .2 .4 .6 .8 1.0 1.2 1.4 1.6 1.8 2.0 72 (28) NEIOC All 2000 psia Data and BAW-168 RATIO OF EXPERIMENTAL TO CALCULATED BURNOUT HEAT FLUX p eastbre V OCONEE NUCLEAR STATION FIGURE 3-33 c 00 00370 i

126 - _ 100 0  :: , 90 - 80 - 70 - 3

     .5 S     60 -

o - U

     .o E

m 50 - Z 40 - _ 30 _ 20 - 10 -

                                                                      ~

0 M TH, H, %, i , i i i i i i , i i . i , i , , , , l , 0 .2 .4 .6 .8 1.0 1.2 1.4 1.6 1.8 2.0 >2 (29) @g/9C All 1750-27 50 psia Data and BAW-168 RATIO OF EXPERIMENTAL TO CALCULATED i BURNOUT HEAT FLUX l j OCONEE NUCLEAR STATION FIGURE 3-34 c

 * <    s'i'       *   '
 ,      k. ,e                                                                               ,

10 - 0 1 i i e e i i i i i i e a i i iie i i : 0 .2 .4 .6 .8 1.0 1.2 1.4 1.6 1.8 2.0>2 (30) $E!NC Euratom Chopped-Cosine 1000 psia Data and BAW-168 10 -

    . 0        i     i i       e i        i e a              , i         i i          i ,     a     i i      a     a    a 3         0         .2      .4        .6          .8       1.0       1.2          1. 4      1.6       1.8       2.0 >2 (31)                                            WE INC

[ Euratom Chopped-Cosine 1500 psia Data j and BAW-168 l 10 - O 0 i i i i i e i a i i i i

                                                                                .           e a      a     i i i i 0         .2       .4       .6         .8       1.0        1.2          1.4        1.6      1.8       2.0>2 (32)                                             9!E C Euratom Chopped-Cosine 2000 psia Data and BAW-168 10 -

0 i i i i i i i i M i i i i i . . i i . . i 0 .2 .4 .6 1.0 1.2 1.4

                                                   .8                                         1.6       1.8 2.0 >2 (33)                                            6E/8C Euratom and B&W Inlet Peak 1000 psia Data and BAW-168 RATIO OF EXPERIMENTAL TO CALCULATED BURNOUT HEAT FLUX b                                                                mtbu                   OCONEE NUCLEAR STATION FIGURE 3-35               _

00 0CQ7iL

10 - O m 0 . '2 ' . '4 '.6 ' . b ' 1.' 0 ' l'. 2 ' l'.4 ' 1[6 ' II8 ' 2.'0 S2 (34) Wg /f C Euratom and B&W Inlet Peak 1500 psia Data and BAW-168 10 - B o & S 8 i ' ' ' ' ' ' ' ' ' ' i ' ' ' ' g 0 .2 ' . '4 .6 .8 1.' 0 1. 2 1. 4 1. 6 1.'8 2.0>2 y (35) 9EINC j Euratom and B&W Inlet Peak 2000 psia uata g and BAW-168 s 10. O O m i . . . . . i i i i e i . . . . . . ,i1 0 .2 .4 .6 .8 1.0 1.2 1.4 1.6 1.8 2.0>2 (36) g/ C Euratom and B&W Outlet Peak 1000 psia Data and B AW- 168 10 - o . . , . i , , . . . , , . . . . . ... 0 .2 .4 .6 .G 1.0 1.2 1.4 1.6 1.8 2.0 >2 (37) $EINC Euratom and B&W Outlet Peak 1500 psia Data and BAW-168 l RATIO OF EXPERIMENTAL TO CALCULATED BURNOUT HEAT FLUX mtpa OCONEE NUCLEAR STATION FIGURE 3-36 933

10-0 i i e i i 4 i a ii6 . 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 > Z (38) NEINC l Euratom and B&W Outlet Peak 2000 psia Data and BAW-168 10-H- M n

                                              ~

o ITT , a 6 i i i e i i e i e i . i i i i i i i i 0 .2 .4 .6 .8 1.0 1.2 1.4 1.6 1.8 2.0 >2 (39)  ; /g 3 All 1000 psia Non-Uniform Data and BAW-168

        .5 E

y 10 - y _ _ s

     ') 2
         =   0 7

i i i i i i i i i i i i i 6 . . . i i e i i 0 .2 .4 .6 .8 1.0 1.2 1.4 1.6 1.8 2.0 >2 (40) @E!0 C l All 1500 psia Non-Uniform Data and BAW-168 20-10- - - 0 m i i i a i i i i i 6 4 i e i i a ie i i i 0 .2 .4 .6 .8 1.0 1.2 1.4 1.6 1.8 2.0 72 (41) (EIN C l All 2000 psia Non-Uniform Data and BAW-168 l RATIO OF EXPERIMENTAL TO CALCULATED BURNOUT HEAT FLUX Ont. net OCONEE NUCLEAR STATION FIGURE 3-37 m l 0000;t74 -

O

   +18                                                                            88      te
   +16                                                            /
                                                                      / //     / 90 1

gDesign [ Z

   +14
                                                                         /

93 k

   +12                                              -
                                                      /              /
                                                                       /         96
                                                                                         ?

0

  +10                                        ,
                                              / //          7                     100
                                                                                      $5 4g g  + 8                                ,
                                       / /p/        7                             104 50 ao N

b +6 # ,/ 1 / g

                                     /

109 g3 . o "'

  +4 j        pj    e      -

jg" g

  +2         r 7                                                    126  g=

j / > 0 / / '

                      /                                       0" IItY            14   P
             /                                               Subcooled                hd 2  //                            2120 psig I                  35 (F Ah- 1. 85)
  -4                                   2185 psig       (F Ah-1. 85)                     ,0 2185 psig       (F Ah-1. 79)                     g
  -6                                                                                    4 100        110              120            130            140          150 RATED POWER, %

MAXIMUM HOT CHANNEL EXIT QUALITY VERSUS REACTOR POWER ja OCONEE NUCLEAR STATION FIGURE 3-38 g roc 215

O

    +14                 l                  g           l      l                        j ll 5% Below Average Assembly Flow
    +12            ---Average Assembly Flow ll        l           l
    +10                          i Fah = 1. 8 5 1
    +8                              Fah = 1. 7 9 m
    +6                                              #
 #                                                         /

N+4 # ' t: k j l , l

                                                        / /
                                                         /

8** /g / l O O

                      /           /

jr

                                              /                 Quality Subcooled
    -2         /     //           /             Fah = 1.79
                    /      /'                         I j  '
                                        - F2.h = 1. 8 5
            /',/
            /                    I      Design Overpower (114%)
    -8 100         110              120                130                               140 R ATED POWER, %

HOTTEST DESIGN AND NOMINAL CHANNEL EXIT QUALITY VERSUS REACTOR POWER (WITHOUT ENGINEERING HOT CHANNEL FACTORS) O f4 - oco uc' FIGURti 3-39 s'ario-0000%76

s 4

  .       3
 '7 '                7 114% POWER

( r O. BAKER o [130% POWER s M y 2 BUBBLY [ j E \ j k ANNULAR 3 \ E m / \ s 2

   <   0. 9       /                         \.

a; s [ S. C. ROSE k y 0. 7 ,

                                                 \

O \ A 0. 6 9

   $   0. 5
                                                    \ -

4 ' 7 SLUG g

0. 4 -

R. D. HABERSTROH \

0. 3 \

                                                                  \
                                                                    \

N N

0. 2 \

0 10 20 30 40 X-QUALITY. % FLOW REGIME MAP AT 2185 PSIG {u$rpwre OCONEE NUCLEAR STATION FIGURE 3-40 - g

      ,1 , ' ij !    '

l

O l

               \               I Design Overpower
                  \

2. 0 -\ \

               \   \                 1. 65 Cosine (W-3)
                 \   \

1. 8 N s 1. 80 Cosine (W-3)

1. 6 \

     <                               \

a: N g BAW-168 Design j 1. 4 -

                                  -- -\-           DNBR (1. 38) o                                       % - -W-3 Design a        _                                        DNBR (1. 30)

W N

     $                      1. 65 Cosine (BAW-168) N
     $  1. 0  -
                                                        \

A y 1. 80 Cosine (BAW-168) \

   \
                                                             \

o 0. 8 -

                                                                \

1. 50 Cosine (B AW-168)

0. 6

              -                                                       \

N

1. 50 Cosine (W-3) \

0. 4 -

0. 2 -

I I I 0 I 100 110 120 130 140 150 REACTOR POWER % OF 2452 MWt HOT CHANNEL DNB RATIO COMPARISON suthu OCONEE NUCLEAR STATION FIGURE 3-41 g i 00 0$78 1

O t 150 Design i e Power I

  ~ 140 2

I I

                                       /

a f 130 . i / '1

  =

I 120 W M [ A l O h 11  ! Y 2300 2400 I 2500 2600 REACTOR CORE POWER MWth REACTOR COOLANT FLOW VERSUS POWER OCONEE NUCLEAR STATION FIGURE 3-42 , m

        .ll*                                I

O 4.00 l l N UO2 Melts g d., " e

                                                                               \,M
           =

3 \ 1 3.00 g i \

           !            \\                                                             i B&W Design Value (CVNA - 142)

E CVNA - 246 o / O. Q g 2.

• h, 7 f i

                                     \                                      /

1 n A

                                           %                         /
                                                                         /             1
                                                                                 /
           ?

K'%L- / - I GEAP 4624 l l 1.000 1000 3000 4000 50b0 2000 1

TEMPERATURE , F THERMAL CONDUCTIVITY OF 95%

DENSE SINTERED UO2 PELLETS OCONEE NUCLEAR STATION FIGURE 3-43 y 0000389

w O 6000 - 5500 Design Overpower

                                                                      /
                                                                    /          ,'
                                                                 /

100% Power / k 5000 A-

• P' ~ ~~~

                                                               ~~ k s

D l / ,

                                                     /     /
 ]

4500 ,, j

                                                  / ,e#

b / n

 ~                                         /
                                         's 4000 O                                     /

J W D 4 3500 3000 B&W Design Value (CVNA-142)

                    ,      ----REF 42 (GEAP-4624) k
                           ----REF 43 (CVNA-246) 6    8    10      12    14   16  18       20       22   24   26     28    30 l

LINEAR HEAT RATE, KW/FT FUEL CENTER TEMPERATURE FOR BEGINNING-OF-LIFE CONDITIONS l e OCONEE NUCLEAR STATION FIGURE 3-44 y 75luO ,I

O s000 , f Design Overpower

                                                                /           /

5500 , 100% Power

                                                                      /

5000 2deMg]egeramre ,

                 ~
                                                   '/ ??

u / / / 5" /'// w W // l// \ N 4000 /lf 0 A 4 3500 B&W Design Value (CVNA-142)

                      ---- REF 42 (GEAP-4624)
                      ---- REF 43 (CVNA-246) 3000 g

6 8 10 12 14 16 18 20 22 24 26 28 30 LINEAR HEAT RATE, KW/FT FUEL CENTER TEMPER ATURE FOR END-OF-LIFE CONDITIONS C) outbre OCONEE NUCLEAR STATION

                                              <'          FIGURE 3-45              ,

1 l

l l l l l l 0 O' CONTROL ROD DRIVE F cnS% ASSE R Y N /,yas;q- -

                                                                         ~

s\ $TUDS l

                       /      .hss
                                                                                            \.t N
                                                                                                   ,x et J m                   aM M %                hM M M                  g---{ %

UPPER PMNUM J, , 't

                                       ]

1 ,-

                                                                                                       ,Y
                                                                                                          '/'

ASSEM8LY - " i j CONTROL CLUSTER GUIDE TU8E 2j- \ , CORE SUPPORT

                          '<4    ,

k. SHIELD p Q WTMT NMZLE g, N fJ

                                                                                                  .             . INLET N0ZZLE I

I , m

                            /\.       dy                        K      '

[ F M.. ' ' i

s. 6 i / REACTOR VESSEL CORE 8ARREL \ $

f ., k ,/ [ FUEL ASSEMBLY E ANCE - f HOLDER TUBE

                                }n;j -                            .,/

THERMAL SHIELD s_. - i ,L - N ,' ',

                                      !           /                            \           i h-        /                                    \\                 '

d ./ '

                          ,          h                           rr r r-r r r n.

M-rrTT-r' ll llrwa lL' il 11 11 i' LOWER GRID s Flow BAFFLE DeCORE INSTRUMENT J I L GUl0E EXTENSION

                                                           /    '
                                                                     \

U "g P - REY: 4-8-67 DELETE 0 htET 00R CCRE

 ,                        j_                                                                                         r eoccm Naz:LE REACTOR VESSEL AND INTERNALS t

GENERAL ARRANGEMENT l seu row OCONEE NUCLEAR STATION FIGURE 3-46 c 6

O

 \s!

FUEL ASSEMBLY

                                         /
                                           ,       '$l
                                                                  //.
                 / ,/'
                                                                           /
                                                                              ,'/f'                              SURVEILLANCE o      o          o~                         /                    SPECIMEN 0;

o; i iO/ HOLDER TUBE e g g 4 l0 ,

            ,/             o           @

• G 3 O O L; ',

          /            O:        e e

G e S O O O O

                                                                                          !                /
                                                                                                            '         CONTROL CLUSTER
         /                                              i
                                                                                         @                            LOCATION g     . g       @       g             @         O m,Vi c e                                                                           j S       O         G

_z y; g g + -.-..g+ @

        /
                        ..       +                           .
                                                                           +    -

O  : . 2 **@ O

                             .

• i 3 @ S I 0; g -

INCORE INSTRUMENT

                        .         . @  .   @        @             2        S

• 8 !U

(, oi . 2 . @ 3 1 @ @ @ f LOCATION

          '                                                                              IO              ,/
           ,                .    .  @      @        @            @         8 o

• O
• 2 -

S * @ 'o 8 *

• REACTOR VESSEL

                /                                                                                     -

f o: _

                        '                           O THERMAL SHIELD
                                      '/
                                           /m ///   i CORE BARREL REACTOR VESSEL AND INTERNALS GENERAL ARRANGEMENT CROSS SECTION O'                                                                                           out rom OCONEE NUCLEAR STATION FIGURE 3-47        ,

00 0 @ f l

O I I ca ta i i ca ry ) I I 1

                                 ~
                                         \

N (. _il \ I u x 3 H s a n xceptrm f I I f) : - i s 3 e CORE FLOODING NOZZLE

                                                        /

r

                                             ! i
                                                   @__  ,/

REV: 4-1-67

: s / CHANGED ENTPANCE s / ARRANGEMENT FOR COPE FLOODING nf '

                                                        /.

i v^ '

                                                          /     -

CORE FLOODING ARRANGEMENT suu p ts OCONEE MUCLEAR STATION FIGURE 3-48 e 1 0-65 .

l l O UPPER END M " l' " I riT m G ASSEMBLY ( 6 j I Il i i W

                                                                                                              ) l1.            .          .         .      i l

s--[ ' ' . - - ' ' ' SMCER GRIO fdJ ~ ~ ~ ?[ ~"Y?I s

                                      .@ ZZz ~ n c r ~' m-                                           E8R8R8g888R8R,::

F A b M M HM ' g W i 8-e- maaqe CAN PANEL i i E > m* P

                                      $h         .h 4              N'                               F TOP VIEW f

l  ! <

                                                                                                   \

U V V V U INSTRUMENTATION TU8E . 6 ' O n , A

                                  ~

o ,it c , aia lo Y '

                                                    'a          ' U_._

l o o o,o O O

                                                                             ,uEL ,c0
                                                                        ,_ ASSEM8LY         ~

s- 1

                                    !                           [ j'                      ,

l  ! CROSS SECTION t _t

                                                                                                                     'l  ?J . lb i LOWER END
                                                                                                             ~

FITTWG

                                                                                                                                        ~

ASSEMBLY J 1

              /N                                FUEL ASSEMBLY

( eut OCONEE NUCLEAR STATION FIGURE 3-49 1 00 OGB6 l

O' i COMlf0G ' N h -lR in 'sn nr

                                                  '-       nBIl
                                                     '//,/,/, f'
                                                             ,p
                                                                     ;q /

gr

                                                                        / ///"

1%

                                                              ~Ni i

t t  ? - e TOP VIEW l uu u u uu ORIFICE CLUSTER ASSEMBLY asAE pta OCONEE NUCLEAR STATION FIGURE 3-50 - b IOb $ hb

C ( !i

                        !ii o   >

fA$a5 i'- L W o a a o e

                             ,:         u <

3

             .l 4

1-s

                       ,,         ,a Y' 5 i'           8 d            *
             .!           ,            u 4             ii           5 r             e L_.c ' g n a'                   o O'8        5e                     e I      W

• H 1

b l'f ll

!i-s ,l' I

1 l i l!. i: i l  !!: I i 1 L  : ii

            .-c    e ! ;'
                         !         !' sL 1        4      z ,-

Y e $ o I.! l s! u  ; CONTROL ROD DRIVE

                   );!!'l* o               z o

GENERAL ARRANGEMENT uwl t e ,, i s' j c! ,l 1

                      -d l                                            e ,

' MO W e $oo 9[o$0I i

                               .,          aw W=
                                               ?

z l aN" sur OCONEE NUCLEAR STATION l l k l FIGURE 3-51 00 00ag$

                                                              @ CAR RE3uCE2                                                                                                                        (.                                                    -- - . w-n~                                                                     . h.-

M AGNETIC CLUTCH SUFFER-S EAL ASSEMBLY = 4 00:lvE MOTOR P (TION INDICATOR TRANSMITTER S .__ _ _ . (.. Y ,5

                                                  ~

I - g 4 - - - + 2;

   -,      --                                0 %. , % -=c.:EH-{[
                                                                            .f i          -:                      _
                                                                                                                           =e                                   7 M'2               '

I 4% gw,.r,qeg4 7mQ m

                                         <C  ..                            - 4
                                                                          ;g - y                  +

83._

                                                                                                                                .+,_g.+_..
                                                                                                                                       -et n

j( M j, 3 fpA --p3 g' I I) ' h [%7g

                                                                                                                                                                                                                                                                              ' ~
                                                                                                                                                                                                    \A m             m                                                        3                                                                                    ' :-                                    A(

w he - ..

                                                                                                                 )
                                                                                                                                        - .. \

h as bh - . . LIMIT SWITCH SYSTEM RACK HOUSING n l RACK PINION _ - . . ~ .f . . I" c4

                                                                   --/

y , . a w;e .f.>,

                                                          . + w5 ~ _ -                                        J .-                      =-i-                                                _

j .- - -_.u--- e

                                                                                                                                                                                                                                          .J x-                        ~ .~ .wm
                               ,                                                                                                           ,..                                        . m.-                                            .

S NUBB ER BOTTOM ING g-j .',$7,',] 9 1 LT, SPRING ASSEMBLY g ~) DRIVE SHAFT ASSEM BLY r

                                                                                                                                                                        .f          .
                                                                                                                                                                                              'S'/f'
                   '-'y n
                                                    -r, ; i
                                                                            - Qfn
                                                                                         ,1__

x-

                                                                                                       ~

Y.7" f"%,

                                                                 ,              7                      y-                            -. - . -y7__ m . ,                          . . --

E ' .Lg. 3 ma i a - bg r% :

                                                                                                                                     , _ _ _ _        _ .3"                                   , gN
                          .2                                                                                                                                               : fli, ~                 3
                                                    / ["h l                                                                                                                    \

h j DRIVE SHAFT HOUSING MITER GEAR SET s' j R1% Q l, I

                                                 ' /b 3           D, !

6e

                                                             - os CPOOL PIECE REACTOR CLOSURE
\r T.L - M - . HEAD x

m j NOZZLE COUPLING ASS EM BLY ,_, _ - l- ::-j

  \           .,

l f i

                                                                                                .                                   -                            r=.2._-- /_-a__      --- % 3;_
                                                                                                                                                                                                -c :            .==

J"

                                                                                                                                                                                                                                                       ==
                                                                               ~

2.y _. -_.-q--M- -*+

                                                                                                                                                               =

a._ L ,.q._, ,u- --- - -w 5. _. - sob s-n %_ . . - . . _ ~

            . f- .             . .k ,,

[~~._~___Z._ 7 3 : 3.~. ~ w- . . . . . - .

                ,     tn,m ,                                                                           g .

s x i_ - , , - a. e es

                !M(                         m y                                                                        \

I \'Og..Y f&.h O,/ n CONTROL ROD DRIVE - VERTICAL SECTION O i 4Rh  ; "

                                        ,y .                                                        --
                                                                                                                                                              ,,m 3'I'. . ike; \ o1 1

s.

                                                                        ' i '. ,                                                                              h,uith' ug    '        m          OCONEE NUCLEAR STATION FIGURE 3-52 kgl   1

CABLE

                            &h L %                     SEAL INTEGRITY VENT GASKETS HOUSING %    dm1                           "                          # .
                             \                                                .                                        RI COVER W \\\\\\\\\2                     9f( * ~ '"

sd DRIVE PINION g g N PINION SHAFT l\ p - . . - -- NUTATING DISC

                                               \  l                               /                                 ROLLER

__ _ x __ ACTUATOR CORE NUTATING DISC ACTUATOR OCONEE NUCLEAR STATION FIGURE 3 53

                                                                                                                                    /j 000040

2 [// / P i /

                                     /'/ v d

v// 4 r

                       '             /
         /                           /
         /                           /
         /                           /
         /                           /
         /                           /
         /            ,
                      '              /
         /                           //
                                     /                  SPOOL PIECE A-
                            -f /4 f

o

           ;,L,

_jp;- g , LOWER GUIDE TUBE 4 I 4 4 ADJUST ABLE SPACER t/ SLEEVE ff ivi II h  : fll'

             '  i g         =

UPPER HARD STOP

                            ,/

k_-_ _!$N ACKRGUIDE RACK SPOOL PIECE AND LOWER GUIDE TUBE ASSEMBLY OCONEE NUCLEAR STATION FIGURE 3-54 m N OE I)b '

i UNCOUPLING LUG \ k@l ,

                                                              \;
                             \                                x   -LOCKING SPRING LET DOWN NUT -g -~
                            > k                               \I g\ m            .        \
                            \

sLaL ,s %-UPPER g PRESSURE HOUSING

                           \'

C CN O <,  % PISTON GUIDE 3 s

                           \     '

s s at FLOW PATH s ' l -i

                           %          s
                                                            ^

SNUBBER SLEEVE g ;0 s l 0 l

                           +l      0  :          9 0  :

g s , , s l0 l y; 0 N ' v  % 0 4 l RACK s s

                           %       0                   0  '

DRILLED HOLES- l

                             "9                . 0     i    g

_ t _ g , 4 g _: _ g { - 0 4 L O , SNUBBER PISTON \ 5  %

                        \%           h                    ,

s kl ., i N  :

                           *     %                            s s l 7-                ?<>;s
                           \

Q'j

                                 /'

f/ g g FLOW PATH

                               .y/                  /                                           1

_ m A r_ - MOUNTING B ELLEVILLE  :# g[ p FLANGE S PRINGS < g , s  : / as s 1 r L l l SNUBBER l mt ro OCONEE NUCLEAR STATION FIGURE 3-55 m 000@h

O' w C n-a-n t/ w

                                   ~

VAPOR BLEED PORT PACKING GLAND NUT JACKING SCREW =G SEAL TEST PORT w$3ps@p gx , SEAL RING CAP l jS hi [ <

                                               -'"c^"

DOG POINT SET SCREWS SEAL RINGS / / j

                                     ~

hk MECHAMCAL LOCKING j/ RETAINER NUT

                           >   n-                  ^

CAP AND DRIVE LINE VENT AJSEMBLY l l 1 l g _. __ __ FIGURE 3-56 el_ i i .

v arv e -gr SOueCE f QQ S A w acorn anastums arvineos now9ea le Teto 4,ecg,T3 CLuTCM rowge b w es m ens 6a4ELES ins rea aussa asur m i.oGsc

                                              *O*te M CLyTLn         om erv eae-er                                             SoveCl e         4             A        eCWie          M 4

ggese to eset censwam ens jg yg,, g,g gy, y g smTwotensWEW C&c CE "" gg ,y g mNg6aw cows

                                                                                                                                          80 3      ~*            I 2*wa%S           SATE N4 I NO. I o
                                                                                       , , , , ,                M..mTi s.t    e                                        i wo i g

aWTOM aTIC SEbENCE -* positsCN Ssakat DCSITION IN o t CATOR

                                                                                                                                                     ==== j w e                             __.                                                                  =-                     l LiMot ti6 sial                                 TW4 web                   l LIMIT       ===_e.-J I                ayyggg                                                                                             tweTCats                   l 0                                                                 .y                  CE                                                l
                              ,        g39g E
                                                                                                                                         ->          ,,,,,, ,p o
                                ;CONTeOL h

4 i l ctocq Q g on:vg l 3YSTEM l TeaNSLATom gg g yg

                                ; PANEL i               eoo ceivE
                  ,  g                                                         co .a aW         aart waa                                  ~a         ~~

McTom e -----s seovrina wo e [D 3 PANEL l oeive CLuTCM I - .kv/ aroup aLanM "0"'T0'

                                                                                                                  ~>                                                  l Teav,6         __

L M,vs positiow soual pos Tion i liI iwo.CATom - -_ _g rov I ggggy _ L eM8T S*6N AL Te4vCL m ,CT _. , , , '~ it

                                                                                                                                      '"'TC =< s
                                                                                                                                                        ---w swiTCM                              i i 1                                                                                   g
                                                                        ' ' '                             Al   DgMomae s o            DC $iTION                 l L%T SGNat                                       TeaN 3      > = - -

2 g j (Lv oT) amoue I

                                  .NomovaL        -

l cows twsTC M l 40C W fiossim0M DeivE '8 NSLAT0e DesvE Ca vaafiON g CM*aW GATE 4 69 ~ McMITom 4M @M ha c,e Snown position OesvE f I up sN0it aTION Mowstoe W 69 ataeM NOG 9 g n00 kJMt? potiTiom siGNat post rion e00 strosCaTION "" sNoitaToe l

                                                                                                                                                    - - -- g
         ,._ _   ntTr                                                                                                                       or                                  ;
         , uwr ,

IW UTEe n coun PANEL

                                                                                        '~T      ea~ L
                                                                                                                                        '"=y        _ _ _ _ a,
                                                                                                                                      ,,,yg,g,                  g               ,

L_____s .o . . . o. ... N u %,,,,,,, ,,o,,,,,, ,

                                                                                        . . . , . . < . . , ~ i,,,,,                    TeaN,       -___                        .

e (Lv of f l F"*"J l 1 l DRIVE MECHANISM CONTROL BLOCK DIAGRAM O

'r    )

seown OCONEE NUCLEAR STATION

 ?';                                                                                                                                      FIGURE 3-57                         .

d 000(9fbi

s MOTOR LIMIT POSITION CLUTCH LI G HTS DIS PLAT UU O I I l l POSITION SIGNAL '" IN TOR - - - - - POTENTIOM ETER I LIMIT SIGN ALS TRAVEL l SHAFT g,7 , _ , , , , _ _ SWITCHES l Position sicwAL I p og,7, o g  ; sysTam LOGIC  : g,g,7 3,cy us DEMoouLATcR TR ANSFO RMER (LVOT) --- - j i I I L SEAL DRIVE REY: 4-23-67

• eE04 AWN M 4EFLECT REVI$.CN Cr CON?90L RCC 09tvC.

LIMIT SIGNAL AND POSITION INDIC ATION SYSTEM OCONEE NUCLEAR STATION FIGURE 3-58 , c

l l l ( l O 1 i SOURCE HQ 4- SINGLE PH A% Soul 4CE Mc E-SiwGLE Pw Ass I l MA40AL TRnR i SWITCH I le g a,: tsT .. __

                                                                                                     . . 337 R.

l

                             -       I                                                                 i
                                    =Rs   i ::Rs e;,                                    ::Rei ::R3 ti l   O - - - -H ***,"                                   0---)-) %T***
                                    " RS3 U RSS,           vuoER VOLTAGE

[RS3 :RS4,i

                                                                                        ~~

UuCER ATAGE l RS RELAYS T l 1 8

                                                     '      coat                         I t

t cog CONT ROLLED l l FROM , i REACTOR f e

         .RQTECTivE                                  I l    i SYSTEM                                    t   TEST ps                                 a e          TESToe (4 CHAuMEL%)

asi :: Res l __Rs i :: as s BREAKEsR I asa ()_ _ _)_) ,, _ _ _)_) BREAWER

      ,                                     :: Rs*         vuoER                        - aS E ;Rs

• um et.R

     /                       L                             voi.TasE                                              VOLTAGE Colt                                                 conc SINGLE PMASE                                          SIMGLE WA%E R ECTI Fi E R                                         R E CTI Ft ER "EV 4-'-8 7                                  (NO FILTER CELAY)                                     (W3 FILTER Ot. LAY)
               .ccc.o  o, ~c 2 ca. rex m.ata                         uo ,                                                  uo. g a                 cm.t
               .E v d .9-4' w:nc=u CLv'OM '*CLL'% O N                                 1*1     I-          -

1 II . I i 2 3

                                                                                  ,1 .I                  ,A   I     I il 61 64 69 t                                                                     j Y

CLUTCH HoLoiNG POWER TO G9 CONTROL ROO ORIVES REACTOR TRIP CIRCUIT OCONEE NUCLEAR STATION f'} FIGURE 3-59 c l 0000%$

                                                               ,1 9-n!p P

i Coueuna . g as m'. . 1s, g, r 4

                                                               ;    il U lj J        l l      ,

1 O O i

                                                -~ s.           '

i

                                                's.g,
                                                            '%.                  ~%~       .
                                                                                ~~- .

N Oe- -. l h l k TOP VIEW w.

                                             'g-                             ~~.     .
                              .Eur.. fi
                              ,0 ISO N
                                                                                          ~ s' '

SECTION CONTROL MN ASSEMBLY h.

                                              $u L         _  _                    _         _     _

l hU U U 00 i CONTROL CLUSTER ASSEMBLY trowt f OCONEE NUCLEAR STATION FIGURE 3-60 c

        +              j
          ~"

< N.}}