ML19305E312

() LIST OF ILLUSTRATIONS Figure Title Page 1-1 Typical Core Cell 1-5 1-2 Schematic of Four Bundle Cell Arrangement 1-6 2-1 P8x8R Fuel Assembly 2-45 2-2.1 Enrichment Distribution P8DRB239 Fuel Bundle 2-46 2-2.2 Enrichment Distribution P8DRB265L Fuel Bundle 2-47 2-2.3 Enrichment Distribution P8DRB265H Fuel Bundle 2-48 -

2-2.4 Enrichment Distribution P8DRB282 Fuel Bundle 2-49 2-2.5 Enrichment Distribution, P8DN3277 Fuel Bundle 2-50 2-3 Cladding Temperature Versus Heat Flux at Beginning of Life (BOL), P8x8R Fuel 2-51 2-4 Cladding Temperature Versus Heat Flux at End of Life (EOL) P8x8R Fuel 2-52 2-5 Cladding Average Temperature at a Fuel Column Axial Gap 2-53 2-6 Definition of Stress Analysis Subscripts 2-54 s 3-1 Calculated Range of Hot, Uncontrolled k gVersus Exposure 3-22 3-2 Calculated Range of k Versus In-Channel Void Fraction 3-23 3-3 Calculated Range of k Versus In-Channel Void Fraction 3-24 3-4 Calculated Range of Hot Uncontrolled Maximum Local Peaking Versus Exposure 3-25 3-5 Envelope of Doppier Coefficient Versus Temperature, E = 200 mwd /t 3-26

3-6 Envelope of Doppler Coefficient Versus Temperature, E = 15,000 MRd/t 3-27 3-7 Envelope of Ak, (From 0.4 Void to Other Voids) Versus Exposure 3-28

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4-1 Schematic of Reactor Assembly Showing the Bypass Flow Paths 4-11 5-1 Relative Bundle Power Histogram for Power Distribution Used in Statistical Analysis for P8x8R Reloads 5-60 5-2 CPR Histogram for Distribution Used in Statistical Analysis for P8x8R Reload (bundles with CFR ratio >1.52 not included) 5-61 5-3 Effects of Initial MCPR on ACPR Typical BWR Core 5 5-4 Flow Factor, Kg 5-63

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(~% 5-5 Turbine Trip Without Bypass 5-64 d

l vii L

NEDO-24195 LIST OF ILLUSTRATIONS (Continued)

Figure Title Page 5-6 Increase in Heat Generated as a Function of Distance 5-65 from the Gap 5-7 Idealization of Flux Spike 5-66 5-8 T' as a Function of a General Profile 5-67 5-9 Central Peak Heat-Flux Distribution (TS No. 65) 5-67 5-10 Central Peak Heat-Flux Distrubiton (TS No. 76) 5-68 5-11 Critical Quality Versus Boiling Length for Tests 65 and 76 5-68 5-12 Damping Coefficient Versus Decay Ratio (Second Order Systems) 5-69 5-13 Accident Reactivity Shape Functions at 20*C 5-70 5-14 Accident Reactivity Shape Functions at 286'c 5-71 5-15 Doppler Reactivity Coefficient vs Average Fuel Temperature as a Function of Exposure and Moderator Condition 5-72 p 5-16 Scram Reactivity Function for Cold Startup 5-73 5-17 Scram Reactivity Function for Hot Startup 5-74 5-18a Water Level Inside the Shroud and Reactor Vessel Pressure Following a 4.66 ft2 Recirculation Suction R Line Break, Emergency Condenser Failure 5-75a ,g 5-18b Water Level Inside the Shroud and Reactor Vessel Pressure Following a 1.0 ft2 Recirculation Suction R Line Break, Emergenc$f Condenser Failure 5-75b _g 5-18e Water Level Inside the Shroud and Reactor Vessel .,

Pressure Following a 0.13 ft2 Recirculation Suction q Line Break, Emergency Condenser Failure 5-75c _cs 5-19a Peak Cladding Temperature Following a 4.66 ft2 Recirculation Line Suction Break, Emergency Condenser Failure 5-76a _e 5-19b Peak Cladding Temperature Following a 1.0 ft2 Recirculation Line Suction Break, Emergency Condenser Failure 5-76b o 5-19c Peak Cladding Temperature Following a 0.13 ft Recirculation g Line Suction Break, Emergency Condenser Failure 5-76c ,g 3-20a Fuel Rod Convective Heat Transfer Coefficient at the Highest Power Axial Node for a 4.66 ft2 Recirculation Line 1/80 Suction Break. Fuel Type - P8DRB239 5-77a}9/79 A

l viii i

j NEDO-24195

() LIST OF ILLUSTRATIONS (Continued)

Figure Title Page 5-20b Fuel Rod Convective Heat Transfer Coefficient at the Highest Power Axial Node for a 1.0 ft2 Recirculation Line "1/80 Suction Break. Fuel Type - P8DRB239 5-77b '9/79, l 5-20c Fuel Rod Convective Heat Transfer Coefficient at the !

Highest Power Ax'ial Node for a 0.13 ft2 Recirculation -

- 1/80 Line Suction Break. Fuel Type - P8DRB239 5-77c 9/79 5-21 Normalized Power Versus Time 5-78 as 5-22 Peak Cladding Temperature Ver;us Break Area 5-79 C;

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%J LIST OF TABLES Table Title Page 1-1 Oyster Creek Reactor Fuel Parameter 1-4 1-2 Lattice and Bundle Average Enrichments for Oyster Creek Reload Fuel 1-4 2-1 Oyster Creek Fuel Assembly Design Specification 2-41 2-2 Zircaloy-2 Material Properties 2-43 2-3 Linear Heat Generation Rate of Calculated 1% Plastic Diametral Strain for P8x8R Fuel 2-43 2-4 Conditions of Design Resulting from In-Reactor Process Conditions Combined with Earthquake Loading 2-44

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3-1 Maximum Cold Clean k, All P8x8R Fuel Designs in Use 3-19 3-2 Definitions of Some Peaking Factors 3-20 3-3 Relationship of Technical Specifications and Peaking 27 lr x Factors 3-21

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4-1 Bypass Flow Paths 4-10 5-1 Uncertainties Used in Statistical Analysis 5-52 5-2 Nominal Values of Statistical Analysis Parameters 5-53 5-3 Oyster Creek Pressure Relief Systems 5-54 5-4 Conservatism Factors 5-55 5-5 Licensing Basis Values for Transient Operating Parameters - Oyster Creek 5-55 5-6 Axial Power Factors for the Thermal-Hydraulic Program 5-55

, 5-7 Nonvarying Plant GETAB Analysis Initial Conditions -

i Oyster Creex 5-56 5-8 Control Rod Withdrawal Error Analysis 5-56 5-9 Sensitivity of CPR to Various Thermal-Hydraulic Parameters 5-57 l 5-10 Significant Input Parameters to the loss-of-Coolant l Accident Analysis 5-58

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5-11 Summary of Break Spectrum Results 5-58 l

ix

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'J NEDO-24195 I

l LIST OF TABLES (Continued)

Table Title Page 5-12 LOCA Analysis Figure Summary 5-59 5-13 Single Failures Considered in the Oyster Creek LOCA <

Analysis 5-59 e

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> 1. INTRODUCTION This report presents information relative to the General Electric reload fuel design and analyses for the Oyster Creek Nuclear Power Plant, Unit 1 (Oyster Creek). The reactor fuel parameters (reactor power, total number of fuel bundles, fuel length, and power density) for Oyster Creek are given in Table 1-1. This report is intended to be a comprehensive reference document for use of General Electric reload fuel in Oyster Creek. Application of the report is limited to this plant. Plant-unique information which does not vary with each reload is also listed within this docusent. The information con-tained in this report represents information that is independent of a cycle-by-cycle reload application. Thus, efficiency is gained by not repeating information in individual reload submittals which is subjected to individual review.

The purpose and applicability of each section is contained in the initial para-graphs of that section. Wherever possible, results of bounding analyses are h given. Bounding analyses are performed where sufficient margins exist to allow analyses to be performed using parameters that are not expected to be exceeded in a group of reloads. Bounding analyses are not performed where they would result in restricted operation. If any parameter ;n a given reload exceeds a parameter used in a bounding analysis, a new evaluation is required. This May take the form of a new analysis or a sensitivity study which demonstrates that the amount that a parameter exceeds the bounding parameter is compensated for by another parameter being more conservative than the bounding value. Where a bounding analysis is not applicable, a cycle-specific analysis will be performed.

A reference cycle has been established to provide a basis for future General Electric reloads. This reference cycle is based on a single General Electric reload batch and current core conditions and plant performance data. Results of the reference cycle analyses are documented in Appendix A.

The fuel design covered in this report is Prepressurized 8x8 Retrofit (P8x8R)

) design. A detailed description of this fuel design is given in Section 2.

1-1

NEDO-24195 m

s_) Additional description and evaluation of the Prepressurized 8x8 Retrofit fuel design is also presented in Reference 1-1. The models and results of the generic analyses performed to assure the structural adequacy of the reload fuel is also described in Section 2. Plant-unique nuclear, hydraulic and limit analyses which are performed for each individual reload are documented in Sec-tions 3, 4 and 5, respectively.

The Oyster Creek reactor core is comprised of numerous core cells. Every core cell consists of a control rod and four fuel assemblies which immediately surround it (Figure 1-1). Each core cell is associated with a four-lobed fuel support piece. Around the outer edge of the core, certain fuel assemblies are not immediately adjacent to a control rod and are supported by individual peripheral fuel support pieces. The four fuel assemblies are lowered into the core cell and when seated, springs mounted at the tops of the chennels force the channels into the corners of the cell such that the sides of the channels contact the grid beams (Figure 1-2).

) Core lattice designation is based upon relative water gap size between adjacent fuel assemblies. In Oyster Creek D lattice cores the gap between assemblies on the sides adjacent to the control blade is greater than the gap on the sides away from the control blade.

The P8x8R reload bundle average enrichments that can be used in the Oyster Creek core are 2.39 wt%, 2.65 wt%, 2.77 wt%, and 2.82 wt% for the 145.24-in. active fuel length rods where two separate bundle designs with an average U-235 enrich-ment of 2.65 wt% are included. Correspcnding bundle designations are P8DRB239, P8DRB265L, P8DRB265H, P8DNB277 and P8DRB282. The P8DNB277 bundle was specifi-cally designed to improve MAPLHGR performance in non-jet pump plants.

It should be noted that two separate enrichment designations will be used throughout the report. ..The bundle designations, given above,' refer to the averagevolumetricbundleU-233' enrichment. Volumetric enrichments are indi-cated by a letter "B" as the fifth character in the prepressurized 8x8 retrofit l bundle designations. The second fuel enrichment distribution used in this m

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- report is the lattice average enrichment. . Lattice distinctions refer to the

i. fuel bundle average planar enrichment.at a.given axial position along the fuel bundle and are. indicated by a letter "L" as'the fifth character of the P8x8R lattice designation. : A list of bundle designations with corresponding lattice j designations is given in Table 1-2 for the P8x8R fuel.

i 1.1 -REFERENCES

! l-1 R.B. Elkins, " Fuel Rod Propressurization - Amendment 1", NEDE-23786-1-P, May 1978.

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O Table 1-1 OYSTER CREEK REACTOR FUEL PARAMETERS l (CE P8x8R FUEL)

Reactor Power (MWt) 1930 i

Number of Fuel Bundles 560 Fuel Length (in.) 145.24

, -o Power Density (kW/f,) 40.22 E l _d i

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Table 1-2 LATTICE AND BUNDLE AVERAGE ENRICHMENTS FOR v h OYSTER CREEK RELOAD FUEL l

l Bundle 145.24-in. Active Lattice Fuel Length P8DRL254 P8DRB239 P8DRL282L P8DRB265L P8DRL282H P8DRB265H-P8DNL282 P8DNB277 l

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2. FUEL MECHANICAL DESIGN 2.1 FUEL ASSEMBLY DESCRIPTION The General Electric fuel assembly (Figure 2-1) consists of a fuel bundle and a channel which surrounds it. Fuel assembly parameters are given in Table 2-1.

The P8x8R fuel bundle contains 62 fuel rods and 2 water rods. The rods of all bundle types are spaced and supported in a square (8x8) array by the upper and lower tieplates and seven spacers. The lower tieplate has a nose piece which has the function of supporting the fuel assembly in the reactor. The upper tieplate has a handle for transferring the fuel bundle from one location to another. The identifying assembly serial number is engraved on the top of the handle. No two assemblics bear the same serial number. A boss projects from one side of the handle to aid in ensuring proper fuel assembly orienta-tion. Both upper and lower tieplates are fabricated from Type-304 stainless steel castings. Zircaloy-4 fuel rod spacers equipped with Inconel-X springs are employed to maintain rod-to-rod spacing. Finger springs located between I the lower tieplate and the channel are utilized on some fuel assemblies to etntrol the bypass flow through that flow path.

2.1.1 Fuel Rods Each fuel rod consists of high density ceramic uranium dioxide fuel pellets stacked within Zircaloy-2 cladding which is evacuated, backfilled with helium and sealed with Zircaloy end plugs welded in each end. The heliun backfill pressure is 3 ATM for the P8x8R' fuel design. The fuel pellets are manufactured by compacting and sintering uranium dioxide powder into right cylindrical pel-lets with flat ends and chamfered edges. The average pollet immersion density is approximately 95% of the theoretical density of UO . Ceramic uranium 2

dioxide is chemically inert to the cladding at operating temperatures and is resistant to attack by water. Several U-235 enrichments are used in the fuel assemblies to redece the local peak-to-average fuel rod power ratios (Figure 2-2). Selected fuel rods within each reload bundle also incorporate small amounts of gadolinium as burnable poison. Gd 03 '" "" #* 7 8 # "

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NED0-24195

] in the UO pellet and forms a solid solution. Details of UO ~cd 0 2

given in Reference 2-1.

2 23 The fuel rod cladding thickness is adequate to be essentially free-standing in the 1000 psi range. Adequate free volume is provided within

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each fuel rod in the form of a pellet-to-cladding gap and a plenum region at the top of the fuel rod to accommodate thermal and irradiation expansion of the UO 2 and the internal pressures resulting from the helium fill gas, impurities, and gaseous fission products liberated over the design life of the fuel. A plenum spring, or retainer, is provided in the plenum space to minimize move-ment of the fuel column inside the fuel rod during fuel shipping and handling.

A hydrogen getter is also provided in the plenum space as assurance against chemical attack from the inadvertent admission of moisture or hydrogenous impurities into a fuel rod during manufacturing. The content and subsequent reaction of the getter is described in Reference 2-32.

Two types of fuel rods are used in a fuel bundle: tie rods and standard rods .

(Figure 2-1). The eight tie rods in each bundle have lower end plugs which v -

thread into the lower tieplate casting and threaded upper end plugs which extend through the upper tieplate casting. A stainless steel hexagonal nut and locking tab are installed on the upper end plug to hold the fuel bundle _

cs together. These tie rods support the weight of the bundle only during fuel j handling operations when the assembly hangs by the handle. During operation, the fuel assembly is supported by the lower tieplate. Fifty-four rods in the /

P8x8R bundle are standard rods. The end plugs of the standard rods have shanks which fit into bosses in the tieplates. An Inconel-X expansion spring is located over the upper end plug shank of each rod in the assembly to keep the rods seated in the lower tieplate, while allowing independent axial expansion by sliding within the holes of the upper.tieplate. Additional expansion spring design information is given in Section VII of Reference 2-2.

2.1.2 Water Rods The P8x8R fuel bundle contains two water rods. These-rods are hollow Zircaloy :

tubes with several holes punched around the circumference near each end to 'l

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allow coolant to flow through. Dimensions and locations of these holes are l

described in Reference 2-32. . !

2-2 l

1

NEDO-24195 One water rod in each bundle positions the seven Zircaloy-4 fuel spacers axially. This spacer-positioning water rod is equipped with a square bottom end plug and with 14 tabs which are welded to the exterior (Diagrams are available in Reference 2-32.). The rod and spacers are assembled by sliding the rod through the appropriate spacer cell with the welded tabs oriented in the direction of the corner of the spacer cell. It' is then rotated so that the tabs are above and below the spacer structure. Once in position, the water rod is prevented from rotating by the engagement of its square lower end plug with the lower tie plate hole.

2.1.3 Other Fuel Assembly Components The primary function of the fuel spacer is to provide lateral support and spacing of the fuel rods, with consideration of thermal-hydraulic performance, fretting wear, strength, neutron economy, and producibility. The spacer represents an optimization of all these considerations. Details of the mechanical design of the P8x8R spacers are shown in Refereners 2-32.

Finger springs (details shown in Reference 2-32) nay be employed to control the bypass flow through the channel-to-lower tieplate flow path for some fuel assemblies. These finger spring seals, located between the lower tieplate j and the channel, provide control over the flow through this path due to channel wall deflections by maintaining a nearly constant flow area as the channel wall deforms.

l The upper and lower tieplates serve the functions of supporting the weight of the fuel and position the rod ends during all phases of operation and handling, g Both the upper and lower tieplates are shown in Figure 2-1. Reload fuel I

bundles with alternate path bypass holes in the lower tieplate may be elected for use at Oyster Creek. These holes are drilled to augment flow in the bypass region. The method by which this bypass flow is taken into account is .o-given in Section 4.

1 D l 2-3 i

NEDO-24195 2.1.4 Channels A separate licensing topical report (Reference 2-3) provides a complete description and analytical results for channels supplied by the General Electric Company and used in conjunction with the reload fuel described herein. However, the following functional description is included in this report for completeness. The description and evaluation supplied in Refer-r ence 2-3 and the information given below applies only to channels supplied by General Electric.

The General Electric BWR Zircaloy-4 fuel channel performs the following functions:

(1) forms the fuel bundle flow path outer periphery for bundle coolant flow; (2) provides surfaces for control rod guidance in the reactor core; H

r (3) provides structural stiffness to the fuel bundle during lateral loadings applied from fuel rods through the fuel spacers; (4) minimizes, in conjunction with the fingersprings and bundle lower tieplate, bypass flow at the channel / lower tieplate interface; q (5) transmits fuel assembly seismic loadings to the top guide and ~m

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fuel support of the core internal structures; (6) provides a heat sink during loss-of-coolant accident (LOCA); and (7) provides a stagnatica envelope for in-core fuel sipping.

The channel is open at the bottom and makes a sliding seal fit on the lower tieplate surface.- The upper end of the fuel assemblies in a four-bundle cell l is positioned in the corners of the cell against the top guide beams by the l channel fastener springs. At the top of the channel, two diagonally opposite corners have welded cabs, one of which supports the weight of the channel from a threaded raised post on the upper tieplate. One of these raised posts has a threaded hole. The channel is attached using the threaded channel fastener essembly, which also includes the fuel assembly positioning spring. Channel-

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co-channel ~ spacing is provided for by means of spacer buttons located on the upper portion of the channel adjacent to the control rod passage area. _

2-4 L

./T x-) The fuel channel enclosing the fuel bundle has a square cross section with a 5.278-in. (nominal)' inside width and round corners, each having a 0.38-in.

-(nominal) inside radius. The nominal length of the fuel channel is 162.156 in.

and the wall thickness is 0.080 in.

2.2 FUNCTIONAL REQUIREMINTS The fuel assembly must be designed to ensure that possible fuel damage would not result in .the release of radioactive materials in excess of applicable regulations. Evaluations 'are made in conjunction with the core nuclear characteristics in Section 3, the core hydraulic characteristics in Section 4, the safety evaluations in Section 5, the plant equipment characteristics and the instrumentation and protaction system to assure that this requirement is met. Adequacy of the fuel assembly is demonstrated if it is shown to provide substantial fission product retention capability during all potential opera-tional modes and sufficient structural integrity to prevent operational impair-ment of any reactor safety equipment. The fuel assembly and its components b) g, are designed to withstand:

(1) the predicted thermal, pressure and mechanical interaction loadings occurring during startup testing, normal operation and abnormal operational transients without impairment of operational capability; (2) loading predicted to occur during handling without impairment of operational capability; and (3) in-core loading predicted to occur from an operating basis earth-quake (OBE), occurring during normal operating conditions, without impairment of operational capability; and are evaluated for their capability to withstand:

(1) in-core loading predicted to occur from a Safe Shutdown Earthquake (SSE) when occurring during normal operation; and' (2) control rod drop, pipe breaks inside and outside containment, fuel handling and one recirculation pump seizure accidents.

Determination of the capability to withstand accidents is by analysis'of the specific event. A description of these analyses is given in Subsection.5.5.

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-NEDO-24195- l 2.3.' MATERIAL PROPERTIES The basic materials used in fuel assemblies are -Zircaloy-2 and Zircaloy-4, Type-304. stainless steel Inconel-X and ceramic uranium dioxide and gadolinia.

These materials have been shown from earlier reactor experience to be com-patible with BWR conditions and.to retain their design function capability during reactor operation. Material properties used in the analysis of the cladding, UO2 pellets, UO 2-Gd 0 Pellets,'Zircaloy-4-spacer, Inconel X-750 23 spacer springs and stainless steel tieplates are given below. Differences.

between irradiated and unirra'diated values are noted.

2.3.1 Cladding The metallurgical state of the Zircaloy-2 cladding:for fuel rods, tie rodd

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and water rods is given in Reference 2-32. Values of thermal conductivity, coefficient of thermal expansion, and plastic strain capability for Zircaloy-2 cladding are given in Table 2-2.

The effect of temperature on the ultimate and yield strengths of unirradiated and irradiated Zircaloy-2 tubing is given in Reference 2-32.

A 1% plastic strain safety limit was established based on General Electric dats on the strain capability of irradiated Zircaloy cladding segments from fuel rods operated in several BWRs (Reference 2-4). None of.the data obtained-fall below the 1% plastic strain value; however, a statistical distribution.

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fit to the available' data indicates che.1% plastic strain value to be approxi-mately the 95% point.in the total population. This' distribution implies,,

therefore,.a small (less than 5%). probability that'some cladding segments may have plastic elongation:less than 1% ac failure.

t The~ stress / cycle relationship used in'the-fatigueievaluations'given in Sub-

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: section.2.5.3.1.4 for cladding (presented-in Reference 2-32)'is based on.a mathematical. fit'to'the data given.in Reference 2-5.-:Other. fatigue analyses

- use the criteria documented in -Reference.2-6.

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_q Additional information with respect to cladding corrosion, hydriding and fretting wear is contained in Section 2.6.

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'2.3.2 Uranium Dioxide Fuel Pellets Thermal conductivity, thermal expansion and melting temperature for UO2 "#*

given in Section 3 of Reference 2-1.

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2.3.3 Urania-Godolinia Fuel Pellets The addition of small amounts of gadolinia to UO results in a reduction in 2

the fuel thermal conductivity and melting temperature. The resulting equa ns for UO - 0 thermal conductivity and melting temperature are 2 23 given in Reference 2-1. Other UO -Gd 0 material properties are also given 2 23

] in Reference 2-1.

2.3.4 Zircaloy-4 Spacer O

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}!aterial properties for Zircaloy-4 are documented in Section 2.2 of Refer-ence 2-3.

4 2.3.5 Inconel X-750 Spacer Springs s

Yield strengths for irradiated and unirradiated inconel at 550*F are given-in Reference 2-32.

2.3.6 Stainless Steel Tieplate The average strength values for type-304 stainless steel tieplates at room temperature and 550*F are given in Reference 2-32.

2.A FUEL' ROD THERMAL ANALYSES Safety l evaluations are performed and measured against establistied safety criteria. The' consequence of calculating values which exceed such criteria 2-7

NEDo-24195 .

\/ is that fuel failure must be assumed to occur. For plant normal and abnormal operation, this is not permissible. The established safety criteria for the fuel are documented in Subsection 2.4.1.1. The evaluation of the linear heat generation rate (LHGR) associated with one of these limits (1%. plastic strain) follows in Subsection 2.4.1.2.

Design evaluations are also performed and measured against established criteria.

The criteria are not intended to predict failure of the fuel if.they are-exceeded, but are rather considered prudent design practice. These design evaluations are included in Subsection 2.5. Thermal analyses performed to provide input into these design analyses are given in Subsection 2.4.2.

4 The models used for fuel rod thermal analyses have been described in Refer-ences 2-7 and 2-8. The latter model is used only for the evaluation of fuel rod initial conditions at the initiation of a loss-of-ccolant accident (LOCA).

Design and safety evaluations performed using the model described in Refer-ence 2-7 assume the most limiting combination of tolerances for all critical y,) dimensions. Assumptions used in conjunction with the model described in Reference 2-8 are stated in the reference.

A power spike allowance due to fuel densification is added to the 13.4 kW/ft LHGR for analyses sensitive to localized power increases (i.e., cladding temperature and stress and strain evaluations). This assures with 95% con-fidence that <1 fuel rod in the core will exceed the maximum LHGR for which the fuel has been designed and precludes technical specification requirements to reduce the 13.4 kW/ft LHGR by a power spike peaking penalty as a function of axial core location. A discussion of the effects of fuel densification is given in Subsection 2.4.2.

2.4.1 Safety Evaluation

. This subsection provides the basis and the.results of a generic safety evalua-tion of linear heat generation rate -(LHGR) associated with the 1% plastic -

strain safety limit. -Results provided in this subsection.are used in'

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conjunction with results from analyses of plant transients which are described in Section 5. The peak LHGR calculated for normal and abnormal operating transients must be less than or equal to the LHGR at which 1% plastic strain is calculated to occur.

2.4.1.1 Basis for Safety Evaluation Fuel rod damage is defined as a perforation of the cladding which would per-mit the release of fission products to the reactor coolant. The mechanisms which could cause fuel damage in reactor abnormal operational transients are:

(1) rupture of the fuel rod cladding due to strain caused by relative expan-sion of the UO pellet, and (2) severe overheating of the fuel rod cladding 2

caused by inadequate cooling.

A value of 1% plastic strain of the Zircaloy cladding has been established as the safety limit below which fuel damage due to overstraining of the fuel cledding is not expected to occur (Subsection 2.3.1). The model used in the U evaluatien of the 1% plastic strain limit is described in detail in Refer-ence 2-7. Dimensions used in conjunction with the model for this evaluation are the most limiting combination of tolerances.

The Fuel Cladding Integrity Safety Limit ensures that fuel damage resulting from severe overheating of the fuel rod cladding (caused by inadequate cooling) is avoided. This limit is discussed in Subsection 5.1.

2.4.1.2 Safety Evaluation Results The equation used to derive the 1%-plastic strain safety limit from test data is given in Reference 2-32.

The linear heat generation rate (LHGR) corresponding to 1% plastic strain has been calculated and results are-presented in Table 2-3. Because gadolinia-urania fuel has a lower thermal' conductivity and melting. temperature'(Sub-section 2.3) than UO fuel,'there is a reduction in the LHGR calculated.to 2

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'NEDO-24195 i '

I . cause 1% plastic diametral strain for'gadolinia-urania fuel rods. However, I to compensate for this, the gadolinia-urania fuel rods are designed to pro-vide margins similar to standard UO r ds.

2 The values calculated as resulting in 1% plastic strain in the cladding are' j

! used during specific plant evaluations of transients due to single operator error or equipment malfunctions to ensure that the safety limit is not exceeded (see Section 5).

Based on these results, it has been determined that the power required to produce 1% plastic strain in the cladding is greater than 160% of the design l

, marinun steady-state power, throughout life, for all. rod types in the assembly for the P8x8R fuel design. This ratio considers the presence of a calculated i power spiking penalty being added to the MLHGR.

~i 2.4.2 Design Evaluations 4

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() The integral design model used in the thermal analyses treats a fuel rod as a number of distinct axial nodes. Analyses at each axial node are based on a solid right circular cylinder pellet geometry divided into' annular ring nodes I with concentric cladding.

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Specific analytical results for the P8x8R fuel design are given below. No I criteria are specified, as results of thefevaluations documented in Sub-I sections 2.4.2.2 through 2.4.2.6 are used as input to the mechanical evalua-tions given in Subsection.:2.5. -

Parameters reviewed by General Electric when it.provides the' reload batch to ,

ensure that these fuel analyses are' applicable for General Electric-supplied reload fuel are'given~1n Subsection'2.7.

2.4.2.1 Effects of Fuel Densification 2

Evaluations to determine the effects of fuel.densification are made.using;the

models described lLn References 2-9.and 2-10. Possible. effects duetto. fuel b; '

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'NEDO-24195 O

densification are: (1) power spikes due to axial gap formation; (2) increase in linear heat generation ute (LHGR) because of pellet length shortening; (3) creep collapse of the cladding due to axial gap formation; and (4) changes in stored energy due to decreased pellet-cladding thermal conductance resulting from increased radial gap size.

2.4.2.1.1 Power Spiking The power spiking allowance, calculated using the referenced methods, resulted in a power spiking penalty at the top of the core which is less than or equal to 2.2% for the P8x8R design. The power spiking penalty as a function of ele-vation from the bottom of the core is conservatively expressed by:

x

=-

L f (2-1)

O where

/fP\

g)x = power spiking penalty at elevation x from bottom of core; (p L AP = power spiking penalty at top of core; F = elevation from bottom of core; and L =- fuel column length.

2.4.2.1.2 Increased Linear. Heat Generation Rate No' power' increase is calculated due to densification. A fuel pellet expands 1.2% in going from the cold to hot condition at 13.4 kW/ft. . While this increase 1

-- -in length from the cold to hot condition is not taken credit'for either:in

.2 l

{:

~-

NEDO-24195 design calculations or in the process of core performance analysis during reactor operation, the expansion more than offsets the decrease in pellet length due to densification. The following expression is employed to calcu-late the decrease in fuel column length due to densification in calculation of an increase in linese heat generation rate:

f=f (2-2) where ao = the average change in density as measured by thermal simulation for 24 hours2.777778e-4 days 0.00667 hours 3.968254e-5 weeks 9.132e-6 months at 1700*C, and 2 = anisotropic factor applied to densification.

Using this equation, the pellet decrease in length due to densification is less

) then the increase in length due to thermal expansion of the pellet in going from cold to hot condition. Therefore, no power increase is calculated due to densification.

2.4.2.1.3 Cladding Creep Collapse Cladding creep collapse is not predicted to occur in the FSx8R fuel design.

Details of this analysis are given in Subsection 2.5.3.1.1.

2.4.2.1.4 Stored Energy The effects of densification on stored energy are considered in the LOCA evalu-ation. Stored energy in the fuel pellet at the initiation of the LOCA is calculated using the model and assumptions described in Reference 2-8.

Analysis of the LOCA is presented in Subsection 5.5.

p 2-12

r NED0-24195

-q

- k/ - m 2.4.2.2 Fuel Cladding Temperature

~

The cladding surface temperature is calculated using the cladding surface heat flux at a given axial position on a fuel element is conjunction with the overall cladding-to-coolant film coefficient-representing the combined' effects of crud and oxide resistances and the liquid fibn resistance based on the Jens-Lottes (Reference 2-11) wall superheat equation.

. The impact.of high cladding temperature, such as decreased yield strength and reduced cladding thickness due to. oxidation, is considered in the design evaluation. Details of considerations when determining cladding surface temperature are given in Reference 2-32.

The model used to calculate'the fuel cladding temperature is documented in Reference 2-7.

O k-s Fuel cladding temperature as a function of heat flux for the P8x8R design le shown in Figure 2-3 for the beginning-of-life conditions and in Figure 2-4 for late-in-life conditions.

2.4.2.3 Fission Gas Release The fuel rod internal pressure is calculated using the perfect gas law. Fuel rod internal pressure is due to the helium, which is backfilled at three atmospheres pressure for_the P8x8R design during rod fabrication,-the volatile content of the UO2_ , and the fraction of gaseous fission products which are

~

released from the UO2 . A c nservative_ combination of dimensional tolerances-is assumed in defining the hot plenum volume used to compute fuel rod internal gas pressure. ' The available fission gas retention . volume is determined based upon the following assumptions:

-(l) Minim 6m (per allowable tolerances) as-built plenum lengthfand

' cladding 11nside 'diaceter'.

+- %

.s i:

2-13

2 NEDO-24195

'(2) Maximum expected fuel'eladding differential expansion.

(3) No credit for fuel-cladding annulus.(gap).

[ ~'(4) The " net" volume is corrected. for the volume of the: components _

contained within the fuel rod plenum.

j -3 I The amount of fission gas released during a time' increment is calculated! based-on the fission gas generated and fission gas release fraction. The temperature-dependent fission gas release fraction model currently employed for fuel rod

! design is a two-region model.

Release Fraction (%) Temperature Range (*F) l 4 <3000 1 100 >3000 I This basis has been demonstrated by experiment to be conservative over the i

complete range of design temperature and exposure conditions (Reference 2-4). .

O For reload fuel, the calculated maximum fission gas release fraction in the l

l' highest design power density rod with the most limiting peaking factors.is:less than 25%. (The percentage of total fuel rod radioactivityfreleased to the rod

' plenum is much less than 25% because of radioactive. decay during-diffusion from i

the UO2 *) "

I 1-

The internal pressure is used in conjunction with.other loads on the fuel rod' cladding when calculating cladding stresses and coteparing these _ stresses 'to-k' the design criteria. Details of thi,s evaluation are described in Sub -

!' section 2.5.3.1.2.

i The fuel rod internal. pressure is calculated using the perfect gas _ law and'the

~

assumptions detailed la Subsection 2.4.2.3. _ The temperature of- the ' gases in' the':

fuel rod plenum is, calculated.-using the assumptions given;in Reference 2-32..

6 b

L 2 '

3

+ . -

~

c t* _

c

_g , .

-w 1,.

, ,, % Pv d + . - ~ . , , .

NEDO-24195 2.4.2.4 Fuel and Cladding Expansion The fuel rod is designed to accommodate predicted fuel and cladding differ-

+

ential expansion. Assuming no cladding restraint, the total relative axial expansion of fuel and cladding is calculated as the difference in.the cladding axial thermal expansion and the fuel axial thermal exps.nsion and irradiation swelling.

The axial thermal expansion of the fuel cladding is calculated based on clad-ding average temperature and cladding longitudinal thermal expansion coefficient given in Subsection 2.3.1. Similarly, the axial fuel thermal expansion is a function of the average fuel temperature and the fuel thermal expansion coefficient.

The cladding temperature is calculated as the sum of inside and outside surface temperature divided by two at an axial elevation. Based on this cladding average temperature, the cladding axial thermal expansions for the separate axial nodes are summed to determine the total fuel rod cladding axial thermal expansion.

The average fuel temperature in this context is the volumetric average cempera-ture of the fuel pellet cross-section at an axial elevation. Again, this calculation is repeated for several axial nodes and,the fuel thermal expansions are summed.

The volume expansions of the fuel attributable to the formation and accumula-tion of fission products are separated into two categories: internal and external' swelling. The internal swelling is postulated to result from the accumulation of both gaseous and solid fission products within the fuel matrix and can be accommodated by the internal porosity of the fuel. This porosity.

volume must.be filled before the. internal swellir.g can contribute to external volume change. The external swelling is postulatd to result from the accumu-lation of solid fission products and results in fuel external dimension changes as soon as irradiation commences. Once the-internal porosity is calculated to

, 2-15

NEDO-24195 O de <i11e467c8 i=ter 1 e111==.ta 1cer 1 111==*esi tece trid te to an external fuel dimension change in addition to that due to the solid external' swelling. The irradiation swelling inodel is based on data reported in References 2-12 and 2-13, as well as an evaluation of all applicable high exposure data (Reference 2-4).

Axial ratcheting of fuel cladding is not considered in BWR fuel rod design.

O eretec791 ca1 <=e1 red aave been everated in the a 1 den te t reaccer wita i 1 elongation transducers. No significant axial ratcheting has been observed (Reference 2-14).

2.4.2.5 Incipient Center Melting

The fuel melting temperature employed is a function of exposure and gadolinia concentration as shown in Subsections 2.3.2 and 2.3.3. The model used for calculating fuel temperatures is described in Reference 2-7. . The fuel is designed so that fuel melting is not expected to occur during normal steady-O t ce <=11 9e er ever tie -

cipient center melting is expected to occur in fresh UO fuel r ds at a 2

linear heat generation rata of approximately 20.7 kW/ft. The effect of gadolinia concentration and fuel exposure on LHGR's at calculated incipient center melting is shown in Reference 2-32.

3.e 2.4.2.6 Pellet-to-Cladding Radial Differential Expansion Pellet-to-cladding radial differential expansion is determined by subtracting O cae i=1ci 1 > i=e e=d c1 dd1 caer t e 9 ie erom cae incre e 1 <=e1 pellet diameter due to thermal expansion and irradiation swelling.

2.5 FUEL ASSEMBLY MECHANICAL EVALUATIONS The fuel assembly is evaluated to assure that the functional requirements given in Subsection 2.2 are maintained. The evaluations consist of analyses, 2-16 k~ _.,

_NEDO-24195 o tests and experience which demonstrate the required fuel assembly structural integrity. When analyses are used to demonstrate structural integrity, resulting stress and/or strain levels are compared to the associatedLmech-anistic limits to ensure that the undesired mechanism being analyzed will not result. These mechanistic limits are documented in Subsection 2.5.1. Differ-ent design limits are applied-to certain elastic fuel rod analyses. These limits are described in Subsection 2.5.1.1.

The evaluations performed address three primary conditions: (1) shipping and handling (Subsection 2.5.2); (2) Normal and Transient-Operational Condition (Subsection 2.5.3); and (3) combined LOCA and seismic loading (Subsection 2.5.4).

i 2.5.1 Analytical Criteria for Assurance of Mechanical Integrity Mechanistic criteria have been established to assure that the functional requirements given in Subsection 2.2 are met.

I Limiting stress values are determined for the normal and upset conditions. .

defined in Table 2-4. The limiting value for stress resulting from mean i value or steady-state. loading is determined by consideration of the material 0.2% offset yield strength or the equivalent strain as established at operating temperature. For stress resulting from load cycling, limiting parameter values are determined from fatigue limits given in Subsection 2.3.

For stress resulting from loading of significant duration, the. limiting parameter is determined from considerations of stress rupture as defined by the Larson-Miller. parameter. If metal' temperatures are below the-level of applicability of stress rupture for the material or if the yield strength.

, is more limiting, then the limiting value of stress is determined from con-sideration of'the material 0.2% offset yield strength or the equivalent strain,.as established at operating temperatures. Where stress rupture and

' l%,] I

~2-17 LW '*

v fatigue cycling are both significant, the following limiting condition is applied:

actual time at stress . actual number of cycles at stress

-<l.0 (2-3)

( allowable time at stress allowable cycles at stress i=1 to n i=1 to m Critical instability loads are derived from test data when available or-from analytical methods when applicable test data are not available. A deflection limit of 0.060-in, fuel rod-to-fuel rod spacing has been established to ensure that the fuel rod-to-fuel rod and 0.030-in. fuel rod-to-channel clearances are sufficient to allow free passage of collant water to all heat transfer sur-faces. The deflection limit was derived from thermal hydraulic testing which demonstrated that allowing a statistical minimum clearance of 0.060-in. rod-to-rod and 0.030-in. rod-to-channel at two standard deviations away from the nominal clearance is sufficient to assure a very low probability of local rod overheating due to boiling transition. More.recent testing to clearances below these values

- would indicate that less clearance is required. Details of these tests are given in References 2-15 through 2-17.

Limiting parameter values for the emergency and faulted conditions defined by Table 2-4 are determined in the following manner. Stress limits are determined i from consideration of the ultimate tensile strength or. equivalent strain of the material, as established at operating temperatures. Critical instability loads are determined from test data, when available, or from analytical methods when applicable test data are not available. -Deflection limits are those values of deformation which preclude serious consequences.

Actual parameter' values are determined from the following considerations. Effec-tive stresses are determined at each point of interest using the theory of con-t

! stant elastic strain energy of distortion (Re.ference 2-18):

2a e

=

(o x -_o y ) + (oy

-o)z

+ (oz -o) x 2

l

( + 6(t xy 2,T yz 2,tzx )1 (2-4)

E2-18

F:. 3. .-

V where 4

e = effective Von Mises stress (psi);

o = longitudinal component:of stress (psi);

o = tangential component .of stress (psi);

a = radial component of stress (psi);

T = tangentially-directed shear stress (psi) on longitudinal face:

y T = radially-directed shear stress on tangential face (psi); and yg T = longitudinally-directed shear stress on radial face (psi) .

3 Stress concentration may be applied'only to the alternating stress component.

Design values of instability loads are scaled up to allow for uncertainty in-manner of load application, variation in modulus of elasticity and difference 1

between the actual and theoretical cases. Calculated values of. deflection for-comparison with deflection limits are based ~on the resulting permanent set after load removal if load removal occurs before damage results.

This procedure is used to evaluate fuel assembly capacity to withstand loadings

due to shipping and handling, normal and transient operation and combined'LOCA and seismic.

As stated in Subsection 2.5, different limits are currently applied to certain elastic analyses which are performed for the-fuel rod. These elastic stress

limits are described in Subsection 2.5.l.1.

2.5.1.1 Stress Limits for Fuel Rod Elastic Analyses-LThe strength theory, terminology and' stress' categories presented in the'ASME

~

Boiler-and Pressure Vessel Code,Section III,'are used'as a guide for.

. v i

l

~2-19' -!

l

d NEDO-24195

/~y

~ i -

~

determining these elastic stress limits. The stress intensity 1imits are given in Reference 2-32. There are different stress intensity limits for normal and abnormal conditions. Each stress intensity limit is based on both yield strength and ultimate tensile strength, and the lower value is always applied.

These criteria are applied to the analysis described in Subsection 2.5.3.1.2.

2.5.2 Evaluation of Fuel Shipping and Handling Loads The structural adequacy of the fuel assembly is determined by demonstrating that each major fuel assembly component can withstand the peak loading which

, occurs due to shipping or handling at room temperature. The functional ade-quacy of each structural fuel assembly component part is established either:

(1) by analyses in which the resulting stresses are compared to' material properties to assure that significant permanent deformations will not occur, i or (2) by testing to demonstrate that significant permanent deformations do not occur.

O k .)

m The bases upon which the fuel assembly components are evaluated to demonstrate their adequacy for sustaining loading from shipping and handling events are given in Reference 2-19. This reference presents the detailed methods.for determining fuel assembly adequacy to withstand shipping and handling loadings, as well as the detailed evaluations of fuel assembly component structural integrity.

Each major fuel assembly component part is shown in Reference 2-19 to be functionally adequate to withstand the peak loadings from assembly shipping and handling at room temperature.

2.5.3 Fuel Assembly Normal and Transient Load Evaluations I

The fuel assembly and its component parts are analyzed to ensure mechanical integrity after exposure to normal and transient operational loadings. The

! evaluations-performed for the fuel rod, water rod, spacer and upper and lower I

(*

x; e

2-20

.l l

l

NEDO-24195 l

tieplate are given below. Fuel rod evaluations are performed for dimensional stability, creep collapse, stress, deflection, f atigue, water-log rupture and flow-induced vibrations.

The fuel assembly and fuel components are designed to assure in-service dimen-j sional stability. The fuel cladding and channel specifications include pro- -

visions to preclude dimensional changes due to residual stresses. In addition, the fuel assembly is designed to accommodate dimensional changes that occur in service due to thermal diff erential expansion and irradiation eff ects. For example, the fuel rods are free to expand axially independent of each other.

Mechanical analyses have been performed to assess the ef f ects of the diff er-ential thermal expansion between the tieplates and spacer grids. The dif f er-ential thermal expansion introduces a bending stress of less tLan 400 psi at the end of the fuel tube. Additional information regarding the model employed -

in this calculation is presented in Section 4 of Ref erence 2-20.

2.5.3.1 Fuel Rod Evaluations 2.5.3.1.1 Cladding Creep Collapse A cladding creep collapse analysis has been performed employing the finite ele-ment model documented in Reference 2-21. Figure 2-5 presents the cladding midwall temperature versus time employed in the analysis. The temperature vs time shown in the figure reflects the assumptions that, at a location of the j fuel rod, the fuel is operating at 13.4 kW/f t plus the calculated power spiking penalty, up to an exposure of 4000 mwd /t. At that point in time, a gap is assumed to form as a result of densification, and the cladding temperature therefore decreases. No' credit was taken for internal gas pressure due to released fission gas or volatiles. The internal pressure due to helium back-

, fill during fabrication was considered. Based on the analysis results, cladding l

l collapse was not calculated to occur for the P8x8R fuel design.

i K._/

2-21

I NEDO-24195 l

-D 1.

2.5.3.1.2 Stress Evaluations The computer program used for the stress analysis of fuel rods is linear and elastic; plasticity, creep or relaxation are not included. The maximum shear stress theory of failure is used, with all radial stresses assumed to be zero.

The radial, longitudinal and circumf erential directions are assumed to be the principal stress directions. Tensile stresses are considered positive, and compressive stresses are considered negative. It is assumed that stress super-position is valid and that all stresses are within the range of Hooke's Law .

deformation, except the pellet-cladding interaction stress during transient overpower. Stress subscripts are defined in Figure 2-6.

An example of fuel description input parameters used in the 8x8R analyses is given in Ref erence 2-32. These input parameters are based on worst tolerance d imensions. Thermal analysis inputs are given in Subsection 2.4. Cladding material properties used in the analyses are given in Subsection 2.3. At beginning of life, the unirradiated mechanical properties given in Subsec-O tion 2.3 are used. At subsequent times in lif e, saturated irradiated mechanical properties are assumed unless the temperature is high enough that the irradia-tion ef f ects on cladding mechanical properties are assumed to be annealed out.

For the latter condition, unirradiated mechanical properties are used.

Three regions of the cladding (at the spacer, between the spacers and at the end plug weld) are analyzed for three different conditions at the beginning of.

lif e and at the end of chosen exposure steps. The three conditions considered are: (1) rated pouer and steady-state design pressures; (2) rated power-and transient design pressures; and (3) transient power.

The loads applied to the fuel rods include pressure diff erentials, flow-induced .

vibration, spacer contact, thermal mismatch (of cladding and relative to lower end plug), radial and circumf erential thermal gradients, end plug misalignment, and pellet-cladding interaction. These loads result in stresses which are categorized and defined in Ref erence 2-32. - There are two analyses performed -

for- each time point and condition analyzed.

2 f)/

y .

2-22

NED0-24195 (D

- V. The first analysis uses the cladding pressure dif ferential resulting from

. maximum coolant pressure and zero psia internal gas pressure; the second analy-sis uses the cladding pressure. diff erential resulting from minimum coolant pressure and maximum internal gas pressure.

The equivalent stress intensity is calculated for each stress category at each of the three regions of cladding, for the inside and outside surfaces, and for each time point analyzed. The stress combinations at each region of cladding are given in Reference 2-32. The equivalent stress intensity is defined as the diff erence between the most positive and least positive principal stresses in a triaxial field. The resulting stress intensity for each stress combination is divided by the appropriate normal or abnormal stress intensity limit to obtain a design ratio. The strengths in the design limits are evaluated at the temperature of the location where stresses are being summed. This is either the mean cladding temperature or the temperature of the applicable cladding surface, depending on the stress category. No design ratio is allowed to exceed unity.

1 This large set of numbers is summarized by showing the maximum design ratio at each rod location for each time point analyzed. Example analytical results are given in Ref erence 2-32.

2.5.3.1.3 Deflection Evaluation The operational- fuel rod deflections considered are a result of manufacturing tolerances, flow-induced . vibration, thermal eff ects and axial load. The deflec-tion equations and nomenclature for these equations are given in Reference 2-32.

The deflections were combined and compared with the fuel rod-to-fuel rod and fuel rod-to-channel spacing deflection limits given in Subsection 2.5.1. This comparison demonstrated that the fuel rod clearance criterion was met.

2.5.3.1.4 Fatigue Evaluation .

The fatigue analysis utilizes the linear cumulative damage rule (Miner's hypothesis-- Reference 2-22.) In the analysis, thermal stresses are assumed to-

- b,a

~2-23 i

NED0-24195 O be proportional to power, based on reference values at 100% power. Because fatigue damage is calculated as a state function, and not a path function, explicit time histories of pressure and temperature are not required. The cyclic condition relating to overpower transients would result from an operator i

error or equipment malfunction, and would therefore be expected to be of short duration (less than 8 hours9.259259e-5 days 0.00222 hours 1.322751e-5 weeks 3.044e-6 months ).

Details of the analysis of areas most subject to fatigue damage are given in Reference 2-32.

The cyclic loads considered in cladding fatigue analysis are coolant pressure and thermal gradients. The analysis is based on the cycles and maximum and minimum coolant pressures shown in Reference 2-32 and the stresses determined in Subsection 2.5.3.1.2. Beginning-of-life thermal gradients are used because they are maximum values. Reference 2-32 shows the load cycles considered, along with the coolant pressures, predicted number of cycles, alternating stress

(g amplitude, fatigue stress, allowable cycles and fatigue damage for each type of

\--

load cycle for fuel with a four-year residence time. The cumulative fatigue damage is less than the allowable fatigue damage.

2.5.3.1.5 Water-Log Rupture

-o As indicated in Subsection 2.7, waterlogging of a fuel rod is not expected 'o because of manuf acturing requirements. Because of this, and because the phenomenon has not been observed in commercial power BWR fuel, no specific analysis of the consequences is performed.

In the unlikely event that a waterlogged fuel element does exist in a BWR core, '

it should not have a significant potential for clad burst (due to internal pressure) during a transient power increase unless the transient started from a !

( cold or very low power condition. Normal reactor heatup rates are sufficiently f slow (t100*F/hr increase in coolant temperature) that water-vapor formed inside a water logged fuel rod would be expected to evacuate the rod through the same

, <s passage it entered, allowing internal and external pressures to equilibriate as

- (,,)

1 2-24

1 NED0-24195 1

the coolant temperature and. pressure rise to ~ the rated conditions. 'Once the internal and external pressures are at equilibrium, at' rated coolant pressure and temperature, transient power increases should, in general, have the effect of only slightly reducing the internal fuel rod plenum volume due to diff er-ential thermal expansion between fuel and clad, thus eff ecting a small short-term increase in internal fuel rod pressure. The potential short-term increase in pressure due to this eff ect would, in general, be small (e.g., a power increase from the cold condition to peak rated power would increase internal pressure less than 15% in the peak power fuel rod). For the range of antici-pated transients, the cladding primary membrane stress resulting from the temporary increase in internal pressure above the coolant pressure would not be expected to exceed the cladding stress design limits.

Experiments have been performed to show that waterlogged fuel elements can fail at a lower damage threshold than nonwaterlogged fuel during rapid reac-tivity excursion from the cold condition (References 2-23 and 2-24) (i.e.,

-~ 60 cal /gm as compared to >300 cal /gm). No analysis of clad stress has been

\/ performed by GE for such conditions. One can postulate that, if such a failure occurred, the resultant energy release and pressure pulse would be much less than for a nonwaterlogged fuel rod which exceeded its damage threshold, since ,

the energy level required for damage is apparently much lower in the water-logged fuel ehement. Any fuel dispersion that might result in such case would s

u further reduce the severity of such a transient.

2.5.3.1.6 Flow-Induced Vibrations Flow-induced fuel rod vibrations depend primarily on flow velocity and fuel

-1 rod geometry. The vibrational model used and maximum calculated vibration amplitude is described in Ref erence 2-32. -

[ 2.5.3.2 Water Rod-l The water rods experience less axial growth than the fueled rods during thermal I' cycling. Therefore, loads can be induced in the spacer positioning-water rod, -

. O.

V l

2-25

NED0-24195 r~5 I) if the positioning device (tab) on the water rod engages the spacer grid. Any additional growth af ter engagement requires the fueled rods to slide through the spacer grid and thereby creates friction forces which must be -reacted by

. the spacer positioning rod. The 14-tab water rod has sufficient margin against the spacer positioning rod sustaining any significant load due to spacer friction forces. This is accomplished by providing additional space between the spacer positioning tabs above that required to clear the divider.

The stret.tth of the spacer grid positioning device was tested for the P8x8R 14-tab water rod design. Uncoupling of the positioning device occurred when the loads induced on the spacer divider were of such magnitude that the divider rotated and allowed the tab to pass. This could result in minor damage to the divider and smearing of the edge of the water rod tab. Both the tab and divider re=ain intact. The estimated hot load carrying capability of the spacer grid positioning device is about three tLnes the load which can be imparted by friction between the fuel rods and spacers.

O

, (s/ In the P8x8R design, the spacer grid positioning device is not located at the geometric center of the spacer grid. The location of the geometric center of the spacer grid has been identified in Ref erence 2-32. The short distance from the geometric center to the tab for the 14-tab water rod reduces the maximum force required to slide the spacer over fuel rods with a load being

, applied to the water rod. This was also substantiated by testing which is

, described in Reference 2-32.

I Another benefit derived from the short distance between the load reaction f

point and the geometric center of the grid is the 14-tab design is that spacer cocking experienced is less than 0.5 degree. This is about the minimum achievable with this spacer grid positioning device in an 8x8 lat-t l tic e. Cocking measurements were obtained by the dial indicator readings shown la Reference 2-32.

1 m

2 - _.

.~ .

'~'

2.5.3.3 cSpacers The. mechanical loadings on the spacer structure during normal operation and transients result from the rod positioning spacer spring forces and from local

, loadings at the water rod-spacer positioning device and a small pressure drop loading.

The spacer represents an optimization of a number of considerations. Thermal-hydraulic development effort has gone into designing the particular configu-ration of the spacer parts. The resultant configurations give enhanced hydraulic performance. Extensive flow testing has been performed employing prototypical spacers to define single-phase and two-phase flow characteristics.

2.5.3.4 Upper and Lower Tieplates The loading on the lower tieplate during operation and transients comprise p-- the fuel weight, the weight of the channel and the forces from the expansion

(_) springs at the top of the fuel rods. The loading on the upper tieplate during operation is due to the expansion spring force, channel weight and tie rod forces. The expansion springs permit diff erential expansion between the fuel 4

rods without introducing high axial forces into the rods. Most of the loading on the lower tieplate is due to the weight of the fuel rods and the channel, which are not cyclic loadings. Operational stresses due to these loadings are -

i less than the material yield strength.

2.5.4 Combined LOCA and Seismic Evaluation Evaluations of the eff ect of combined LOCA and seismic loads upon the com-ponents of the fuel assembly have been performed for BWR/6 reactors (Refer-

~

ence 2-5). The applicability of these evaluations to reload fuel assemblies -

is described below.

1 b)

I

, L 2-27

NED0-24195 2.5.4.1 Dynamic Analysis and Component Seismic Loads The seismic accelerations determined for the BWR/6 fuel assembly are greater 4

than those for reload fuel assemblies. Thus, the results of the BWR/6 dynamic analysis form an upper bound for reload . fuel accelerations, and the seismic loads determined for the BWR/6 fuel assembly components exceed those of the reload fuel assemblies. The procedure for determining these loads for reload fuel is similar to that used for BWR/6 fuel assemblies.

2.5.4.2 LOCA Loads i

The pressure differentials on the BWR/6 lower tieplate, upper tieplate, and spacer resulting from a recirculation line break or from a steamline break are larger than the corresponding pressure differentials for reload fuel.

assembly components. Water rod pressure differentials are insignificantly small. The methodology for evaluating the LOCA pressure diff erentials for reload fuel is similar to that used for BWR/6 fuel assemblies.

2.5.4.3 Component Evaluat::ns The results of the evaluations on the BWR/6 fuel assembly are conservative for the upper tieplate, fuel rods, water rods, and spacers. Thus, the demon-strated adequacy of these components is applicable to reload fuel assem-blies. The methodology for evaluating each BWR/6 fuel assembly component is applicable to the evaluation of reload' fuel assembly components.

2.6 FUEL ROD CORROSION, HYDRIDING AND FRETTING WEAR CONSIDERATIONS' i

2.6.1 Potential for Hydriding In general, the Zircaloy-2 cladding performance in the very early plants was good; .however, some fuel failure mechanisms were exposed and corrected and are not significantly affecting current fuel performance. Details-of this experience are-provided in References 2-4, 2-25, 2-26, and 2 .

t v

2-28

NEDO-24195 One of these fuel failure mechanisms was localized clad hydriding. The hydriding was the result of excessive amounts of hydrogeneous impurities inadvertently introduced into the rod during the fuel fabrication process.

j Hydriding defects were identified by significant numbers of rods perforating at relatively low power and exposure. Hydriding has been eliminated by a combination of improved engineering specifications and stricter manufacturing controls, which ensures that production fuel will not contain excessive quan-tities of hydrogeneous impurities.

An engineering specification limit on moisture content in a loaded fuel rod is defined in Reference 2-32 which is well below the threshold of fuel

, failure.

Procedural controls are utilized in manufacturing to prevent introduction of hydrogenous impurities such as oils, plastics, etc. , to the fuel rod. Hot vacuum outgassing (drying) of each loaded fuel rod just prior to final end I

\

plug welding is employed to assure that the level of moisture is well below the specification limit. As a further assurance against chemical attack j from the inadvertent admission of moisture or hydrogenous impurities into a !

fuel rod during manufacture, a hydrogen getter material is employed in the . j upper plenum of all fuel rods. The getter material is in the form of small chips. These getter chips are loosely packed in a stainless steel tube of 'lE which one end is capped, and the other end is covered by wire screening.

Calculations of getter material performance are described in Reference 2-32.

2.6.2 Fuel Element Energy Release The metal-water chemical reaction between zirconium and water is given by:

Zr + 2H O + Zr0 + 2H - AH ~(2-5) 2 2-29 1

l l

i

e NEDO-24195 where -aH = 140 cal /g-mole.. The reaction rate is conservatively given by the familiar Baker-Just rate equation:

2 6 /-

W = 33.3 x 10 ,t exp ( 45'500T \j (2-6) vhere W =

milligrams of zirconium reacted per cm of surface area; i = time see; R =

the gas constant (cal /mol 'K); and T =

the temperature of zirconium (*K) .

This rate equation has been shown to be conservatively high by a factor of 2 (Reference 2-28). Tha above equation can be differentiated to give the rate

/ at which the thickness of the clad is oxidized. This yields:

th =

1 exp

^2 7 (2-7)

AX where th =

rate at which the clad thickness is oxidizing:-

= oxidized clad thickness; AX A,A = appropriate constants; and g

T = reaction temperature.

.D l

The reaction rate is inversely proportional to the oxide buildup; therefore, at a given cladding temperature, the reaction rate is self-limiting as the oxide O

2-30

'NEDO-24195 O ' build s up . The total energy release from this chemical reaction over a time period is given by:

T QT = .

N rods X& (2 4 where

=

number of rods experiencing boiling transition (at temperature T);

rods i

-AH = heat of reaction; C = cladding circumferences; L = axial length of rod experiencing boiling transition; and p = density of zirconium.

This equation can be integrated and compared to the normal bundle energy release if the following conservative assumptions are made:

(1) At an axial plane, all the rods experience boiling transition and are at the same temperature. This is highly conservative since, if boiling transition occurs, it will normally occur on the high power rod (a) .

(2) Boiling transition is assumed- to occur uniformly around the circum-ference of a rod. This generally occurs only at one spot.

(e) The rods are assumed to reach some temperature T instantaneously-and stay at this temperature for an indefinite amount of' time.

This integration has been performed per axial foot of bundle, and the total energy release as a function of time has been compared to the total energy release of a high power bundle (N6 MW) over an equal amount of time. The results are shown in Reference 2-32. For example, if the temperature of all the rods along a 1-ft section of the bundle were instantly increased to 1500*F, the total amount of energy that has been released.at 0.1 sec is 0.4%-

2-31

. ~ .

NEDO-24195 O' of the total energy that has been released by the bundle (6 MW x 0.1 sec) .

Note that the fractional e nergy release decreases rapidly with time even though a constant temperature is maintained. This is because the reaction is self-limiting.

The amount of energy released is dependent on the temperature-transient and the surface area that has experienced heatup. This, of course, is dependent _,

on the initiating transient. For example, if boiling transition were to f

_k occur during steady-state operating conditions, the cladding surface tempera-ture would range from 1000-1500*F, depending on the heat fluxes and heat transfer coefficient. Even assuming all rods experience boiling transition instantaneously, the magnitude of the energy release is seen to be insignifi-cant. Significant boiling transition is not possible at normal operating conditions or under conditions of abnormal operational transients because of the thermal margins at which the fuel is operated (see Subsections 5.1 and 5.2) .

It can therefore be concluded that the energy release and potential for a f' chemical reaction is not an important consideration during normal operation or

\- abnormal transients.

2.6.3 Fretting Wear and Corrosion As described in Ref erence 2-32, no significant fretting wear or corrosion has been observed throughout a continuing surveillance program.

2.7 APPLICATION OF GENERIC DESIGNS TO RELOAD PROJECTS The fuel bundle designs presented in this section were designed for use in the

-e operating BWR/2 plants. The reference cycle supplement documents the ao number and designation of new and irradiated bundles of the cycle. Conditions which rary from one design to another or from plant to plant, were taken into account in design evaluations. All dimensional, enrichnent or gadolinia varia-l tions among the designs were either analyzed separately or the most limiting case l

was analyzed. Operating conditions which vary from plant to plant are parameters

^3 such as core pressure and the expected maximum power vs exposure for the peak duty

)

2-32

,. , . ~

.NED0-24195 -

b fuel rod. In the generic analysis, the highest core operating pressure was used which results in the highest coolant temperatures. With respect to power vs expo-sure, the limiting fuel rod is assumed to operate at its maximum permitted power over its entire lifetime, based on thermal-mechanical limitations such as the stress limits presented in Reference 2-32.

The power level and exposure' range considered in the generic analyscs, there-fore, provides a basis for review of these same parameters for a particular cycle on an individual project. General Electric performs the following evalua-tions whenever it provides the reload batch to ensure the applicability of the generic analyses to the General Electric-supplied fuel in the core:

1. The performance of the reload fuel batch and other General Electric-supplied fuel already present in the core is proj ected for the reload cycle.

2. The combinations of power, exposure and residence time from this projected cycle performance are compared with those used in the

() respective generic analyses. The application to General Electric fuel is acceptable only if the power levels and exposures analyzed in the generic analyses equal or exceed those projected for the cycle.

In addition, other Oyster Creek operating conditions such as core pressure -

are reviewed and compared with.those used in the generic analyses to ensure that limiting conditions. have not changed. Adherence to these procedures i

ensures the applicability of the generic design to Oyster Creek whenever.

General Electric provides the reload batch.

2.8 INSPECTION AND TESTING 2.8.1 Fuel Manufacturing Rigid quality control requirements are enforced at every stage of General Electric' fuel rod. manufacturing to ensure 'that the design; specifications are

(T met. Written manuf acturing procedures and quality control plans define the' V

2-33

NEDO-24195 Q) steps in the manufacturing process. Fuel cladding is subj ected to 100% dimen-sional inspection and ultrasonic testing to reveal defects in the cladding wall. Destructive tests are performed on representative samples from each lot of tubing, including chemical analysis, tensile and burst tests. Integrity of end plug welds is controlled by standardization of weld processes based on radiographic and metallographic inspection of welds. Sample tests are performed for qualification of weld stations, weld parameters and weld operators prior to application. Production samples are tested as a check on the process and process controls.

UO powder characteristics and pellet densities, composition and surface 2

finish are controlled by regular sampling inspections. UO weights are 2

recorded at every stage in manufacturing. Each separate pellet group is char-acterized by a single stamp. Because individual rods may contain segments of dif f erent fuel compositions, physical and administrative controls are utilized during fuel rod assembly. These controls are overchecked during fuel rod inspection (e.g., scanning to verify pellet enrichment and proper assembly).

\-- Fuel rods are individually serialized prior to fuel loading to: (1) identify which pellet group (s) are to be loaded in each fuel rod: (2) identify which

! position in the fuel assembly each fuel rod is to be loaded; and (3) f acilitate total fuel material accountability. Each finished fuel rod is gamma scanned to detect any enrichment or rod pellet loading deviations which exceed design specification.

The fuel rods within each assembly are designed with characteristic upper end plugs each of which is common to a discrete enrichment range. Gadolinia-urania fuel rods are designed with characteristic extended end plugs. These extended end plugs permit a positive visual check on the location of each gadolinium-bearing rod after bundle assembly. Further identification of individual fuel rods is accomplished by symbolization on the upper end plug. Each upper end plug is ensured proper placement on a fuel rod. by reference to the specific fuel rod type. Proper assembly is verified prior to bundle assembly. Each fuel rod is ensured of proper placement within a fuel bundle by_ inspection of O

V 2-34

g it NEDO-24195 the fuel rod serial number on the lower and plug, by upper end plug-tieplate diameter control and by visual verification of enrichment unique symbols on the upper and plug assembly.

Fuel rod inspection includes metallographic examination of weld samples and radiographic examination of fuel rod welds on a sampling basis. Completed fuel bundles are helium leak tested to detect the escape of helium through O the t se a d e=d 9 1 er e1ded re 1en . Thi test grect de water-te==1==

which could result from cladding which has a small pinhold. The leak detector system consists of a high vacuum system and a mass spectrometer which can detect leaks of rod helium fill gas smaller than the design limit (1 x 10- std. cc/sec).

Each bundle is given a complete dimensional inspection prior to shipment.

Dimensional measurements and visual inspections of critical areas are verified ~

before shipment and again at the reactor site on a planned basis. The sampling rate, method and tools of the post-shipment fuel inspection are outlined in Reference 2-32.

O The general procedures and acceptance criteria used to evaluate fuel assembly, subassembly, component parts and materials which deviate from applicable manu-facturing drawings and specifications are defined in Reference 2-29. Methods defined in this document have previously been accepted by the NRC (Ref ar-ence 2-30).

2.8.2 Enrichment Control Program .

The incoming UF6 *= a re8- 8- groduction UO 2 powder are wrmed by emission O s> ectr <>ceer fer ime ritie d sr mm ce t1== for U-235 e=ricamene-l The sintered pellet is also sampled for impuricles by emission spectroscopy.

Chemical verification of impurities is also performed including gravimetric analysis for 0/U determination.

O 1

2-35 i

-.__--a --w ._w---w--..--w . ne e.,o w w - w.- q m a-.w4 y e- * 'W'

NEDO-24195 O '

-O The enrichment blended material is verified by gamma counting. Each rod is gamma scanned to detect enrichment deviations which exceed design criteria.

Three levels of enrichment control are associated with bundle assembly:

(1) Lower end plug serialization - each rod is uniquely serialized for enrichment which facilitates proj ect traceability.

(2) Upper end plug diameter control - used such that a discrete range of enrichments are associated with one upper end plug diameter. For fuel rods with higber enrichments, the upper end plugs are larger.

The upper tieplate has a distinct hole pattern with varying hole size which precludes the positioning of a fuel rod with higher enrichment into a location for a significantly lower enrichment rod.

(3) Upper end plug symbolization - allows visual inspection ef ter assembly and permits making a permanent photographic assembly record for each bundle. Symbols are uniquely related to specific U-235 and gadolinia configurations.

O Cr 2.8.3 Gadolinia Inspections The same rigid quality control requirements observed for standard UO f"*L *#*

2 employed in manuf acturing gadolinia-urania fuel. Gadolinia-bearing UO fuel 2

pellets of a given enrichment and gadolinia concentration are maintained in separate groups throughout the manufacturing process. The percent enrichment and gadolinia concentration characterizing a pellet group are identified by a stamp on the pellet. The amount of Gd 0 in fuel powder and pellets is verified 23 by testing samples using X-ray tiuorescence. Individual rods, which may con-tain more than one pellet composition, are assembled based on rigid physical and administrative controls, normally including a computer interactive inventory and assembly system with overchecks to assure proper assembly. The Gd 0 content of 23 each fuel rod is verified based on X-ray fluorescence scanning of pellets or fuel rod scanning for Gd 0 content and proper assembly. Correct placement of 3

gadolinia-bearing rods within the fuel assembly is further assured by the extended upper end plug shanks and enrichment control symbols for these rods.

(D 2-36~

NEDO-24195 The following quality control inspections are made:

(1) Gadolinia concentration in the gadolinia-urania powder blend is verified.

(2) Sintered pellet UO -Gd 0 s lid solution homogeneity across a fuel 2 22 pellet is verified by examination of ceramographic specimens.

(3) Gadolinia-urania pellet identification is verified.

(4) Gadolinia-urania fuel rod identification is checked.

(5) Each gadolinia-urania fuel rod is scanned to assure proper assembly.

(6) Gadolinia content is verified by .X-ray fluorescence measurements of each pellet or scanning the assembled rod.

All assemblies and rods for a given project are inspected to assure overall accountability of fuel quantity and rod placement for the proj ect.

2.8.4 Surveillance Inspection and Testing of Irradiated Fuel Rods.

General Electric has an active program of surveillance of both prodection and developmental BWR fuel. The sche.dule of inspection is contingent on the avail-ability of the fuel as influenced by plant operation.

The lead experience fuel rods (with respect to exposure, linear heat generation rate, and the combination of both) are selectively inspected Inspection i techniques used include:

1 (1) Leak detection tests, such as " sipping."

(2) Visual inspection with various aids such as binoculars, borescope, periscope and/or underwater TV with a~ photographic record of observations as appropriate.

(3) Nondestructive testing of selected fuel rods by ultrasonic and eddy current test techniques.

(4) Dimensional measurements of selected fuel rods.

Unexpected conditions or abnormalities which may arise, such as dis'tortions, .

cladding perforation, or surface disturbances are analyzed. -Resolution of

.(A

'v'

\

2 .

l

V specific technical questions indicated by site examinations may require examination of selected fuel rods in Radioactive Material Laboratory (RML) facilities. Details of this surveillance' program are documented in Refer-ences 2-4, 2-25 and 2-27.

A prepressurized test assembly of the 8x8R design was placed in operation in April 1977. For a description of this fuel assembly, see Ref erence 2-31.

2.9 REFERENCES

2-1 G. A. Potts, "Urania-Gadolinia Nuclear Fuel Physical and Irradiation Characteristics and Material Properties," NEDE-20943-P (PROPRIETARY) and NED0-20943, January 1977.

2-2 "8 x 8 Fuel Bundle Development Support," NED0-20377, . February 1975.

2-3 "BWR Fuel Channel Mechanical Design and Deflection", NEDE-21354-P (PROPRIETARY) and NEDO-21354, September 1976.

2-4 H. E. Williamson and D. C. Ditmore, " Experience with BWR Fuel Through q September 1971." NEDO-10505, May 1972.

U 2-5 W. J. O'Donnell and B. F. Langer, " Fatigue . Design Basis for Zircaloy Components," Nuclear Science and Engineering: 20, 1-12 (1964).

2-6 "BWR/6 Fuel Assembly Evaluation of Combined Safe Shutdown Earthquake (SSE) and Loss-of-Coolant Accident (LOCA) Loadings," NEDE-21175-P (PROPRIETARY) and NED0-21175, November 1976.

2-7 " General Electric Standard Safety Analysis Report," Proprietary Supple-ment to Amendment 14, Docket No. STN-50-447, May 1974.

2-8 "GEGAP-III, A Model for the Prediction of Pellet-Cladding Thermal Conductance in BWR Fuel Rods," NED0-20181, Revision 1, November 1973.

2-9 Regulatory Staff (U.S. Atomic Energy Commision), Supplement 1 to the Technical Report.on Densification of General Electric Reactor Fuels, December, 1973.

2-10 V. A. Moore, Letter to I. S. Mitchell, " Modified-GE Model for Fuel Densification," Docket 50-321, March 22, 1974.

2-11 Jens, W. R., and Lottes, P. A., " Analysis.of Heat Transfer, Burnout, Pressure Drop, and Density Data for High Pressure Water, USAEC Report-4627, May, 1951.

. v) ~

2-38

NED0-24195 2-12 UAPD-TM-283, " Effects of High Burnup on Zircaloy-Clad, Bulk UO Plate 2

Fuel Element Samples," September 1962.

2-13 WAPD-TM-629. " Irradiation Behavior of Zircaloy-Clad Fuel Rods Con-taining Dished End UO Pellets," July 1967.

2 2-14 D. C. Ditmore and R. B. Elkins, "Densification Considerations in BWR Fuel Design and Performance," NEDM-10735, December,1972.

2-15 "BWR/6 Fuel Design Amendment No.1," NEDE-20948-lP (Proprietary) and NEDO-20948-1, November 1976.

2-16 "BWR/4 and BWR/5 Fuel Design Amendment 1," NEDE-20944-lP (Proprietary) and NEDO-20944-1, January 1977.

2-17 Attachment to Letter MFN-114-77-050, G. G. Sherwood (GE) to D. G. Eisenhut (NRC), "NRC Questions on Rod Bowing," March 29, 1977.

2-18 S. Timoshenko and J. Goodier, Theory of Elasticity, McGraw-Hill,1951.

1' 2-19 W. G. Jameson Jr., "BWR 2-5 Fuel Assembly Evaluation of Shipping and Handling Loadings," NEDE-23542-P (PROPRIETARY), March 1977, 2-20 "BWR/6 Fuel Design," NEDE-20948-P (PROPRIETARY), June 1976.

2-21 " Creep Collapse Analysis of BWR Fuel Using SAFE-COLAPS Model" NED0-20606A and NEDE-20606-PA (PROPRIETARY), August 1976.

2-22 M. A. Miner, " Cumulative Damage in Fatigue," Journal of Applied Mechanics,12, Transactions of the ASME, 67, 1945.

2-23 Stephan, L. A., "The Response of Waterlogged UO Fuel Rods to Power Bursts," IDO-ITR-105, April 1969.

2-24 Stephan, L. A. , "The Ef fects of Cladding Material and Heat Treatment on the Response of Waterlogged UO Fuel Rods to Power Bursts," IN-ITR-lli, 2

January 1970.

2-25 R. B. Elkins, " Experience with BWR Fuel Through September 1974,"

NEDE-20922 (PROPRIETARY) and NEDO-20922, June 1975.

2-26 H. E. Williamson and D. C. Ditmore, " Current State of Knowledge High Performance BWR Zircaloy-Clad UO Fuel, NED0-10173, May 1970.

2 2-27 R. B. Elkins, " Experience with BWR Fuel through December 1976,"

NEDE-21660-P (Proprietary) and NED0-21660, July 1977.

2-28 " Thermal Response and Cladding Performance of an Internally Pres-surized, Zircaloy-Clad, Simulated ' BWR Fuel Bundle Cooled by Spray- Under Loas-of-Coolant Conditions," GEAP-13112, April 1971.

O .

2-39.

l -NEDO-24195

j. ( ) 2-29 " Nuclear Energy Divisions BWR' Quality- Assurance Program Description,"

l' NEDO-il209,03A, pg. 69, November 1976.

f 2-30 ' Letter, C. J. Heltemes (NRC) to 'J. F. Quirk (GE), "NRC Acceptance of General Electric QA Topical Report," October 27, 1976.

2-31 R. B. Elkins, " Fuel Rod Prepressurization - Amendment 1", NEDE-23786-1-P,

May 1978. ,

! 2-32 " General Electric Reload Fuel Application," NEDE-240ll-P-A*

l l

i i

J t

i

!O l

1 r

4 i

_I.

4

~

Reference refers to revision of NEDE-240ll-P-A which'is approved by the NRC" as of.the date a. specific analysis is. initiated.

.c .

+

40 rE, w = , -- - _ , - e 'u-+ -

~.r y e*--- y

Table 2-1 OYSTER CREEK FUEL ASSEMBLY DESIGN SPECIFICATIONS Fuel Assembly Fuel Bundle P8x8R Geometry 8x8 i

Rod Pitch (in.) 0.64 0 j Fuel Rods Fill Gas helium

! Fill Pressure (atm) 3 i Cetter yes-Number of Fuel Rods 62 Fuel Material sintered UO2

! Pellet Diameter (in.) 0.410 Pellet Length (in.) 0.410 Pellet Immersion Density (%TD) 95.0 1

Cladding Material Zr-2 Outside Diameter (in.) 0.483 O. Thickness 0.032 Water Rod '

Material Zr-2

Outside Diameter (in.). 0.591 i Thickness 0.030 Number of Water Rods 2 1

SP acers Zr-4 with Material Inconel X-570-

' Springs Number per bundle 7 a Fuel Channel j

. Material Zr-4 Inside Dimension (in.) '

5.278 d

Wall Thickness (in.) 0. 08 0-i

! ~ 2-41 l

l l

NEDO-24195 Table 2-1 OYSTER CREEK FUEL ASSmBLY DESIGN SPECIFICATIONS (Continued)

Active Fuel Length (in.) 145.24 Heat Transfer Area (f t ) 94.9 Fission Gas Plenum Length (in.) 9.48 Spacer Pitch (in.) 19.55 Finger Springs

Lower Tieplate Bypass

Flow Holes

Nominal Fuel Weights: Weight of UO 2 ,Kg Weight of U.Kg

^

, FUEL BUNDLE 145.24 AFL 145.24 AFL l

! 8DRB239 200.6 176.9 8DRB265L 200.5 176.8

, 8DRB265H 200.3 176.6 3

~

8DRB282 200.2 176.5 8DNB277 200.2 176.4 i

i

^'""""'""""'"'""""'"""'"""'"

O 2-42

NEDO-24195 VO Table 2-2 1 ZIRCALOY-2 !Le.TERIAL PROPERTIES .

a Parameter Value Thermal Conductivity, 9-10 i Btu /hr-ft *F 3 (600 to 800*F) 4 J

Coefficient of Thermal j Expansion, in./in. *F ,

-5 f Radial 3.2 x 10

~

Longitudinal 2.9 x 10 i l

E Total Plastic Strain > 1.0 3

Capability, %

3

I i

Table 2-3 LINEAR HEAT GENERATION RATE 4 0F CALCULATED 1% PLASTIC DIAMETRAL STRAIN-FOR P8x8R FUEL

!. l i

Exposure LHCR at Calculated 1% Plastic (mud /t) Strain (kW/ft)*

i I

UO2 Gd**

O 224.7 ~221.9-

, 20,000 2 23.0 220.4-40,000 2 19.6 2 17.3 y I

.i

'*The values reported tuve been reduced by an amount equal to_the calculated.

_ power spiking penalty-(%).

- Results for gadolinia are applicable for maximum' concentration used,in j g reload fuel: design.

(' J 2-43 ;

L I. -

Table 2-4 i CONDITIONS OF DESIGN RESULTING FROM IN-REACTOR PROCESS CONDITIONS COMBINED WITH EARTHQUAKE LOADING Conditions of Design Reactor Initial Percent of Safe Shutdown Earthquake Imposed Conditions 0% 50% 100%

Startup Testing Upsat --- ---

Normal Normal Upset Faulted Abnormal Upset --- ---

Acc ident --- ---

Faulted (LOCA Loads)

G 2-44 l

NED0-24195 s . . ,

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2-45

TOP !

W10E141DE CORNE R T

6 N 4 4 4 N 4 4 N 5 N 5 N N N lN 8*6in N N N N Nl lNl N N I

5 N 4 N 3 N 3 N 3 N 2lN 2 N 4 N l to N N N N N Nf N 4 N 3lN 3 N 3 N, 2lN 1 fN 3 N 3lN N lNl N N lN fN N lN 4 N 3lN.

3 N 2lNI W

I 1 N 1 N 2lN N lN' f4 lN 3 N N lN 4 N N 3lN-2 N i WS 1 N 1 N 1 N 2 N lN N, l N N N N 4 N 2lN' 1 N 1 Nl1 N 3 N 2 N 3 N N IN N N, 'N N N N c,

4 Ni t ID 2# '^

lN' 2 N 3lN 1 fN 2 N 2lN 3 N A=

lN N lN N lN N lNl N 5 N 4 NI3 N 2 2lN 3lN 3 N 4 N 4l lN Nj Nf fN lN N, N v

ENRICHYENT E N R ICHYE r.T l IDENTtF tC AllCN at .U 235 l

l 1 l 33 2 3. 0 l l t

3 24 - Y l

4 20 O=6m l y 5 l 17 ENRICHYENT IDE NTIF IC ATION phEL ROD 6 1.3

!' By FUEL RCO mRCHmT 0711 ZONES ZONES N l l

WS - $ PACE R PostTIONING W ATE A RCD W - WATER ROO

's Figure 2-2.1. Enrichment Distribution P8DRB239 Fuel Bundle (b

2-46

NEDO-24195 W10E WlOE CORNER TOP l 6 lN 5 N 4 N 4 N 4 N 4 N 4 N 5 N g.6m N N N N N N N N d k 5 N 4lN 2 N 2 N 2 N 2 N 2 N 4 N

, *! !N N N N N N N 4 N 2 N 2 N 2 N 1 N 1 N 2 N 2 N N N N N N N N N 4

lN 2 N 2 N 2 N 1 N 1 N 2 N g

W N lN N N N N N 4 N 2 N 1 N 1 N 1 N 2 N 2 N WS N N N N N ,

N N 4 N 2 N 1 N 1 N 1 N 1 N 1 N 2 N N N N N N N N N C=

4 N 2 N 2 N 1 N 2 N 1 N 2 N 3 N 139 24 in.

A*

N N N N N N N N I33 24 i"-

5 N 4 N 2 N 2 N 2 N 2 N 3 N 4 lN N N N N N N N lN ENRfCHMENT ENRICHMENT IDENTIFICATION wt '4 U 235 1

l 38 2 3.0 3 2.4 Y l

C 4 20 O=6m l 7

' N E N RICHME N T IDE NTIF IC A T10N 6 13 ,

ENRICHMENT N 0.711 ZONES ZONES WS - SFACE R POSITIONING WATER ROD W- WATER ROD (r^)

v Figure 2-2.2. Enrichment Distribution P8DRB265L Fuel Bundle 2-47 l

NEDO-24195 WlOE WlOE CORNER TOP 6 N 5 N 4 N 4 N 4 N 4 N 4 N 5 N 8

6 en N N N N N N N N 5 N 4 N 2 N 2 N 2 N 2 N 2 N 4 N N N N l N N N N N 4 N 2 N 2 N 2 N 1 N 1 lN 2 N 2 N N N N N N N N N 4 N 2 N 2 N 2 N l 'N 1 N 2fN W

N N N N N N lN 4 N 2 N 1 N 1 N 1 N 2 N 2 N WS N N N N N N N 4 N 2lN 1 N 1 N 1 N 1 N 1 N 2 N N lN N N N N N N C=

139.24 in 4 N 2lN -

2 N 1 N 2 N 1 N 2 N 3 N A=

N lN N N N N N N 133.24 in.

5 N 4 N 2 N 2 N 2 N 2 N 3 N 4 N N N N N N N N N m

u ENRICHMENT E N R IC H M E N T IDE NTIF IC ATION wt % U 235 1

l 38 A 8 2 3.0 3 2.4 O = 6 in 4 l 20 y I'

ENRICHVENT 1.3 IDENNIC ATION FUEL ROO 6

BY FUEL RCD E N RICHME N T ZONES ZONES N l 0 711 WS - SPACER POSITIONING W ATE R ROD W - WATE R ROD s Figure 2-2.3. Enrichment Distribution P8DRB265H Fuel Bundle (G

2-48

NEDO-24195 cs j

> WlOE-WlOE CORNE R TCP 7 N 6 N 5 N 5 N 4 N 4 N 4 N 5 N 86m N N N N N N N N 6 4 2 4 N N 4 N 3fN 2 N 3 N N N N N N lN N N N N 5 N 4 N 3 N 1 N 1 N 1 N 3 N 2 N N N N N N N N N 5lN 3 N 1 N 1 N W

1 N 1 N 2 N lN N N N N N N 4 N 2 N 1 N t N 1 N 1 N 2 N WS N N N N N N N 4 N 3 N 1 N 1 N 1 N 1 N 3 N 2 N N N N N N N N lN ,

4 N 2 N 3 N 1 N 1 N 3 N 1 N 3 N 13 24'"

g A=

lN N N N N N lN N 5 N 4 N 2 N 2 N 2 N 2 N 3 N 4 N N N N N N N N( N O

ENRICHVENT ENRICHVENT IDENTIFICATION wt % U-235 1 38 A 8 2 3.3 3 30 4

2.4 i D =6in f

5 20 E N RICHM E N T IDENTIFICATION FUEL ROD 77 BY FUEL ROO ENRICHMENT ONES ZONES 7 1.3 N 0.711 i WS - SP ACER POSITIONING WATER ROD l W - WATER ROD f; Figure 2-2.4. Enrichment Distribution P8DRB282 Fuel Bundle s.

2-49

O NED0-24195 WlOE WlOE CORNER TOP p 8 N N 5 N 6 N N JL 6]N 5 7 8

6 en N ,

hN N 5 5 N N N g

6 N 4 N l l

N N 2l 2l 2 2 3 5 5 N 3 N 3 N 3 N N 2 N 4 4 4 N N W

5 2 4 4l 4 1 2 3 N WS 5 2 4 N 3 1 2 5 N 3 N 2 N N 2 4 4 3 3 N N C=

6 N 3 N 3 N 146 20~

g, N 3 N 1 1 N 1 3 7 N 3 N , 2 N N 5 N 2 2 lN 3 5 E N RICHY E NT E N R IC M ME N T IDE NTIF IC ATION wt % U 235 1 38 A 8 2 ,

33 V

3g 4 26 1 D = 6 in

[

5 2.4 E N R ICH Y E NT IDE NTinC A TION FUELACO 6 2.0 8v FUEL ROD ENRICHVENT 1.7 ZONES 2ONES 7

8 1.3 N 0.711 WS - SPACE R POSITIONING WATE R ROD W - WATER ROD p Figure 2-2.5. Enrichment Distribution, P8DNB277 Fuel Bundle V

l 2-50 l

l l

NEDO-24195 t

l 800 0

750 -

l INSIDE 700 3

E

?

650 -

j w AVERAGE l e O

! /

SURFACE 550 -

l I f I I l l I l g

0 50 100 150 200 250 300 350 400 460 500 HE AT FLUX (Stu/h ft2) x 10-3 Figure 2-3. Cladd'ing Temperature Versus Heat Flux at Beginning of Life (BOL), P8x8R Fuel 2-51

NEDO-24195 j 1000 l

l 950 -

l I

900 -

850 -

E e,

INSIDE y 800 -

?

(

5 E AVERAGE w

750 -

o O

5 u

SURFACE 700 -

650 -

600 -

! I I I I I 550 O 50 100 150 200 250 300 350 HE AT FLUX (Bruth-f t2)=10-3 Figure 2-4. Cladding Temperature Versus Heat Flux at End of Life O- (EOL) P8x8R Fuel i

2-52

670 660 -

650 -

,' 640 -

m

$ 630 -

Q :

E 620 -

S g 610 -

o

? 600 -

o e

590 -

% sao -

o

$70 --

560 -

sso l l I I O 1 2 4 3 5 IRR AOI ATiON TIME tyd

.i 1

Figure 2-5. Cl. adding Average Temperature at a Fuel Column Axial Gap O

2-53 l 1

1

I NEDO-24195 0

0 1

sunF Act 2 Cy 9

a '

> sunrActt g 1

.l i

1 Figure ~.2-6. .-Definition of Stress-Analysis; Subscripts-2 '

g l

ni -

o I

y z

b i;;

3 l 2

s s

I z O

x 0

8

- -- _ ------ - /

3. NUCLEAR EVALUATION METHODS i

3.1 INTRODUCTION

The nuclear evalutions for Oyster Creek are performed using the analytical tools and methods described in this section. The nuclear evaluation procedure is best addressed as two parts: lattice analysis and core analysis.

9 Most of the lattice analysis is performed during the bundle design process and utilizes methods described in References 3-1 and 3-2. The results of these single bundle calculations are reduced to " libraries" of lattice reactivities, relative rod powers, and few group cross-sections as functions of instantaneous void, exposure, exposure-void history, control state, and fuel and moderator temperature for use in the core analysis. These analyses are dependent upon fuel lattice parameters only and are, therefore, valid for all cycles to which they are applied.

9 The procedure used in modeling non-General Electric fuel is to use the characteristics of GE reload bundles. Whenever possible and appropriate, vendor, operator, and/or GE-calculated specific data are used in place of the above characteristics.

The source of the characteristics used for the non-General Electric-supplied fuel present in the reference cycle is specified in Appendix A.

The core analysis is unique for each reload. It is performed in the months preceding the reload to demonstrate that the core meets all applic-able safety limits. The principal tool used in the core analysis is the

~

three-dimensional boiling water reaE[ simulator code, which computes power distributions, exposure, and redctor thermal-hydraulic characteristics, with spatially varying voids, control rods, burnable poisons and other variables. The BWR simulator code is described in References 3-3 and 3-4.

O 3-1

~

l , . NEDO-24195 l

3.2 LATTICE STCLEAR CHARACTERISTICS DESCRIPTION O -

The mechanical description and physica1 parameters of the e1oad fuel are t given in Section 2.

This section describes the calculated nuc1 ear parameters of the retoad lattice.

There are fav reat 11mits on the lattice design itself. The re.1 11mits are k generally expressed in terms of core parameters, which are discur.eed in -

Subsection 3.3. The intent of Subsection 3.2 is primarily to document the nuclear parameters of the various retoad designs. The choices of parameters for retoad fuel are dependent upon the environment in which the reload fuel bund 1e wi1t be used; that is,'the characteristics of the reload bundle itself are not so important as the characteristics it wi11 exhibit when 1oaded into a specific core. This is demonstrated by the manner in which the analysis is performed; that is, the tattice parameters are reduced to 11braries of information that the simulator code used to predict the behavior of tha core as a who1a.

O th 1 = tic ch = et =1 cic 1 c= a f d=e== == c1 =

r :

1. co1d, c1ean k, data;

2. hot k, as a function of exposure and void fraction;

3. loca1 peaking factor; 4 Dopp1er coefficient; and O 5. veid-rea==1 1=7 effect.

3.2.1 Reactivity Traditionally, bundle reactivities have been expressed in terms of k, (i.e.,

the neutron multip11 cation of an infinite array of 11ke bundles). Genera 11y, the enrichments and reactivities of the bund 1es for reload fue1 wilt be higher

, O than for inicia1 cores. 5,ecifica117 . a hisher va1ee of reactiviev is 3-2

NEDO-24195 i allowable for the low exposure reload bundle than is allowable for the .

initial core bundle (at the same low exposure), because the reload fuel bundle is loaded into an environment of highly exposed bundles of generally lower average reactivit.y. The maximum calculated central section cold, clean, controlled and uncontrolled k, are presented in Table 3-1. Maximum and mini-mum hot reactivity characteristics of the P8x8R reload lattice (axial mid-section) are presentad in Figures 3-1 through 3-3. The first plot depicts the calculated ranga of k., hot, uncontrolled versus exposure for P8x8R central -

O 1attice - The econ 4 and third >1ot de,1ct the cate 1ated ra 8e ef x-. wet. -

uncontrolled and hot, controlled (respectively) for the P8x8R central lattices [

as functions of in-channel void fraction. -

3.2.1.1 Factors Affecting Lattice Reactivity The lattice reactivity is a function of lattice average enrichment, gadolinia loading, void fraction, hydrogen-to-U ratio and exposure. To delineate all these functional dependences, it is necessary to note the following.

O rer siven 2actice. it 1 eb erved chae:

(1) At zero exposure, the reactivity is highest at 0.0 void fraction (VF), followed by 0.4 VF and has least reactivity at 0.7 VF. This is because enriched U-235 fuel lattice is of higher reactivity for a softer neutron spectrum. The softer neutron spectrum is most abundant in the highly thermalized medium of 0.0 VF.

-a (2) The term, " void history" refers to the fact that the k, curves were q e

obtained for the respective lattices based on a nuclear depletion i O co ve re) ai ter7 in t* water de it7 ef o o. o 4. er o 7 vr-In other words, assuming a 0.0, 0.4 or 0.7 VF water density for the whole exposure range, the illustrated k, curves were obtained.

(3) Gadolinia dependence is described in Reference 3-11.

O 3-3 l

NEDO-24195 i

f i .

l (4) The in-situ production of the Pu-nuclides increases with increasing void fractions.- For example 0.7 VF is better for Pu production than 0.4 and 0.0 VF. This is because the first has a harder (less thermalized) neutron flux spectrum than the last two. As the. lattice

[

attains higher exposure, the U-235 is increasingly depleted and the

(

I worth of the bred fissile Pu-nuclides becomes more significant.

. At about 15 GWd/t and thereafter .the fractional fission rate of the bred Pu nuclides (Pu-238 and Pu-241) in the 0.7 VF case is a

! few percent greater than those of the 0.0 and 0.4 VF cases. The net result is that the 0.7 VF case is of higher reactivity than the

, low void fraction cases.

E i 3 l An exception to this % behavior and crossover phenomena at high exposure is the natural U lattice. Here, for all exposures, the higher the void fraction, the greater the reactivity. This is due to the fact that the natural U lattice, having no enriched U-235 and gadolinia, is primarily a U-238 system. It is most favored for I

Pu fissile nuclides generation under a hard neutron flux spectrum, namely that of the 0.7 VF.

It is also noted that the maximum reactivity for the natural U lattice occurs earlier than the other lattices with enriched U and gadolinia. It peaks at about 2.0 GWd/t for the former and 6.0 GWd/t for the latter. This is because the enriched lattice is designed j

such that depletion of the gadolinia occurs around the end of the l 'first cycle, usually about 6.0 GWd/t.

i The primary driving force in fission for the natural U lattices is j from the: bred fissile Pu nuclides. Since-the Pu production is neutronically favored for a hard spectrum, the high void case always displays a higher reactivity. This reactivity difference f- ' increases at the higher exposure range with~the increasing frac-tional fission by Pu-nuclides. .

1

-(5) h, calculational considerations are given in Reference'3-11.

, 3 4

NEDO-24195 3.2.2 Local Peaking Factors

O The local peaking, gross, radial, axial and total peaking factors are design parameters related to reload core analysis. Their respective definitions are shown in Table 3-2. These peaking factors determine, directly or indirectly, the thermal performance parameters of General Electric-supplied fuel such as maximum linear heat generation rate (MLHGR), maximum average planar linear heat generation rate (MAPLHGR) and minimum critical power ratio (MCPR). The relations between the various peaking factors and the core thermal perform-ance parameters are detailed in Table 3-3. The peaking factor, by itself, does not constitute a limiting condition. The thermal performance param-eters such as MLHGR, MAPHLCR and MCPR do limit unacceptable combinations of these peaking factors.

i For a given lattice at a given void fraction, the maximum local peaking factor will occur at different fuel rods as the exposure increases. This is due to the differant depletion and generation rate of the various fissile nuclides in each fuel rod. Figure 3-4 shows the highest and lowest calculated maximum local peaking factor for the P8x8R central lattices at different exposures.

Further characteristics of these curves are given in Reference 3-11.

It has been observed that local peaking factors from infinite lattice calcu-lations exhibit close agreement with experimental results (Reference 3-2);

hence, their use is appropriate.

The axial power shapes and axial peaking factors are dependent on the fuel bundle types and exposures, the in-core locations, the control rod pattern, the specific reload cycle and the plant. These axial power shapes and sxial t

peaking factors are calculated by the three-dimensional BWR simulator, which takes all'of the effects into account. Therefore, while it is possible to j

provide curves showing the variation of the local peaking factor as a function of exposure, it is not possible to do so for the axial peaking factor because it depends on plant-unique features and operating characteristics.

O 1

3-5

--_ = - -_.

NEDO-24195 O The local peaking factors are generated for the exposure range of 0-35 GWd/t. This covers the expected range of exposures at any axial location.

O If a particular location should exceed 35 GWd/t, the power at that location will be sufficiently low that there is no danger of exceeding any performance limits because of its low reactivity at high exposure.

The effect of the enrichment tolerance on local peaking factors has been

( investigated for several types of fuel bundles by intercomparison studies 1

between calculated values and experimental measurements documented in Reference 3-4. The scandard analytical method to evaluate the effect of &

H

> enrichment tolerance on the local power peaking factor is the two-dimensional ,

lattice physics code (Reference 3-1 of the report). It has been found chat

..a E

the standard deviation of the manufactured fuel pellets is about 0.015 wt%

U-235. The corresponding effect on the local power peaking factor is 0.7%. The axial power factors are listed in Table 5-6. Further deviational k I effects on the local power peaking factor are discussed in Reference 3-11. _

The local peaking factor does vary with void fraction, and this dependence is taken into account in the calculations used to assign local peaking factors to each axial segment of the fuel. Figure 3-4 shows the values at 0.40 voids,

~

as this is the typical average BWR bundle voA' caction. This compares to an e eo Oyster Creek avera;;e bundle void faction of pac 0.32 voids.

The curves of local peaking factor in Figure 3-4 are the nominally calculated values. The word " maximum" is used to denote the fact that the maximum pin power at each exposure is used to construct the curves. As' stated above, the nominally calculated values are appropriate, as they agree well with experimental results.

Figures showing the local pin power distributions as a function of void fraction and exposure for typical General Electric fuel lattices are provided 'k*

in Reference 3-10. .

i I

3-6

NEDO-24195 3.2.3 Doppler Reactivity The Doppler coefficient is of prime importance in reactor safety. The Doppler-coefficient is a measure of the reactivity change associated with an increase in the absorption of resonance-energy neutrons caused by a change in the temperature of the material in question. The Doppler reactivity coefficient provides instantaneous negative reactivity feedback to any rise in fuel ~ tem-perature, on either a gross or local basis. The magnitude of the Doppler coefficient is inherent in the fuel design and does not vary significantly among BWR designs having low fuel enrichment. For most structural and moderator materials, this effect is not significant; but in U-238 and Pu-240 an increase in temperature produces a comparatively large increase in the absorption cross-section. The resulting nonparasitic fission absorption of neutrons causes a significant loss in reactivity. In BWR fuel, in which approximately 98% of the uranium in UO 2 is U-238, the Doppler coefficient provides an immediate reactivity response that opposes fuel fission rate changes.

O Although the reactivity change caused by the Doppler effect is small com-pared to other power-related reactivity changes during normal operation, it becomes very important during postulated rapid power excursions in which large fuel temperature changes occur. The most severe power excursions are those associated with rod drop accidents. A local Doppler feedback associated with a 3000'F to 5000*F temperature rise is available for terminating the initial excursion.

The Doppler reactivity decrement is derived directly from the lattice calcula-tions by performing two calculations at different fuel temperatures. The results are used to determine the proportionality constant, CDOP, from:

=

CDOP ( 6 - 69 ). (3-1)

_O-0 3-7

NEDO-24195 The Doppler coefficient is then generated using CDOP (3-2)

(1 dkDOP dT (1 + CDOP ( 8 - 6o ))2 d The theory and methods used are described in detail in Reference 3-5.

Maximum and minimum calculated Doppler coefficients for P8x8R fuel central lattices as a function of fuel temperature and exposure are shown in Figures 3-5 and 3-6. Uncertainty in the nominal Doppler coefficient applica-tion of the point model to three-dimensional analyses and effect of exposure on the Doppler coefficient are discussed in detail in Reference 3-5 and in response to NRC questions in Reference 3-7.

3.2.4 Void Effect o

One of the most important considerations for reactor safety is void reactivity. E The void coefficient must be large enough to prevent power oscillation due to spatial xenon changes yet small enough that pressurization transient do not unduly limit plant operation. In addition, the void coefficient in a BWR has the ability to flatten the radial power distribution and to provide ease of reactor control due to the void feedback mechanism. The overall' void coefficient is always negative over the complete operating range, since the BWR design is undermoderated. The reactivity change due to the formation of voids results from the reduction in neutron slowing down to the decrease in the water-to-fuel ratio.

A detailed discussion of the methods used to calculate void reactivity coefficients, their accuracy and their application to plant transient analysis is presented in Reference 3-5.

The maximum and minimum calculated reactivity effect of a change in void for -

o P8x8R central lattices is shown in Figure 3-7. The specific data are used

~

in the determination of core void coefficients as described in l

()

V Subsection 3.3.2.

l l

3-8

NEDO-24195 c,

3.3 REFERENCE LOADING PATTERN DETERMINATION 3.3.1 Bases

-o The reference loading pattern, documented in each supplement reload R submittal, is the basis for all reload licensing and operational planning and

-o is comprised of the fuel bundles designated in the supplemental

~*

R submittal. It, in turn, is based on the best possible prediction of the core condition at the end of the present cycle and on the desired core energy capability for the reload cycle. It is designed with the intent that it will represent, as closely as possible, the actual core loading pattern.

3.3.2 Core Characteristics The reference core is analyzed in detail to assure compliance with all safety limits. This section discusses the results of core calculations on shutdown margin (including the liquid poison system) and core average reactivity coefficients.

3.3.2.1 Steady-State Core Characteristics 3.3.2.1.1 Core Effective Multiplication and Control System Worth These values are unique to each reload. They are calculated using the BWR simu-lator code to determine the core reactivity with all rods withdrawn and with all rods inserted. A tabulation of the values for the reference cycle is provided in Appendix A.

These three eigenvalues (effective multiplication of the core, uncontrolled, fully controlled and with the strongest rod out)~are calculated at an exposure corresponding to the minimum expected exposure of the previous cycle. The core is assumed to be in a xenon-free condition.

This procedure ensures that the calculated values are conservative. The value 1 (d of R includes equilibrium Sm. Further discussion of the uncertainty cf these calec'4tions is in References 3-4 and 3-8. l l

l 3-9

As exposure accumulates and burnable poison depletes in the least burned fuel bundles, an increase in core reactivity may occur. The nature of this increase depends on specifics of fuel loading and control state. For example, if one control cell is loaded with two fresh bundles, then the core reactivity calculated with that single control rod removed will probably increase as exposure accumulates. However, the core reactivity calculated with a different rod out may decrease with increased exposure.,

Cold k,gg is calculated with the strongest control rod cut at various exposures through the cycle. At each exposure point, a search is made for the single strongest rod, the location of which may change with exposure. The value R is the difference of the strongest rod out k,ff at BOC and the maximum calculated strongest rod out kegg at any exposure point. The strongest rod out k,gg at any exposure point if equal to or less than:

k,gg = (Fully controlled k,fg)BOC + (Str ng Rod Worth)BOC + 1 (~}

3.3.2.1.2 Reactor Shutdown Margin 4

Technical Specifications require that the refueled core must be capable of being made suberitical with margin in the most reactive ccndition throughout the subsequent operating cycle with the most reactive control rod in its full-out position and all other rods fully inserted.

The shutdown margin is determined by using the BWR simulator code to calculate the core multiplication at selected exposure points with the strongest rod fully withdrawn. The shutdown margin for the reference cycle core is obtained by subtracting the BOC k,ff given in the reference cycle supplement from the g critical k,ff of 1.0.

Items which may affect the evaluation of a reactor shutdown margin for a given cycle are:

(1) an early termination or an extension of the previous cycle, and r \

l (2) the inability _to insert a control rod to the fully inserted position. l l

3-10

NEDO-24195 b

3.3.2.1.3 Standby Liquid Control System The standby liquid control system (SLCS) is designed to provide the capability of bringing the reactor, at any time in a cycle, from a full power and minimum control rod inventory (which is defined to be at the peak of the xenon transient) to a subcritical condition with the reactor in the most reactive xenon-free state. The requirements of this system are dependent primarily on the reactor power level and on the reactivity effects of voids and tem-perature between full-power and the cold, xenon-free condition. The shutdown _

capability for the reference cycle is given in the reference cycle supplement. E "D

3.3.2.1.4 Reactivity Coefficients Reactivity ccefficients (the differential changes in reactivity produced by differential changes in core conditions) are useful in calculating relative stability and evaluating response of the core to extcrnal disturbances. The base initial condition of the system and the postulated initiating event l determine which of the several defined coefficients are significant in evaluat-ing the response of the reactor.

The coefficients of interest, relative to BWR systems, are discussed herein individually with references to the types of events in which they significantly

affect the response.

There are two primary reactivity coefficients that characterize the dynamic

^

behavior of boiling water reactors over all operating states: (1) Doppler reactiv1ty coefficient and (2) moderator void reactivity coefficient. Also associated with the BWR is a power reactivity coefficient and a fuel tem-perature coefficient. However, the power ceofficient is just a combination-of the Doppler and void reactivity coefficients in the power operating range and the fuel temperature coefficient is merely a combination of the Doppler and moderator temperature coefficients. Because these two quantities provide only redundan't information, they are no longer calculated for the reload cores. The Doppler and void coefficients are unique for each cora and reload.

-C f

O For the reference cycle, they are given in the reference cycle supplement.

i 3-11 l l

l i

NEDO-24195 3.3.2.1.5 Moderator Void Coefficient The void coefficient as a function of voids is calculated using a point model approach. Equation 3-4.below simulates the total reactor, which is made up -

of a' number of different bundle types.(lattice designs) at characteristic exposures:

N - -

E,-

~

k77 (E g,V) VF g -

(1-CF)+(i(E,N*

g g G

~

k'" (E,V) = I"1 2 (3-4)-

1+bh(V)B(V)+Ak g f where E = average exposure of the ith fuel type; t

V = average void of the core; O k", = k, uncontrolled; k[=k, controlled; 2

th = M volume weighted by control fractionand fuel types; 2

B = geometric buckling; g

ak g = correction factor to preserve criticality; CF = control fraction-(controlled-bundles / total bundles);

N = number of fuel types for the void coefficient calculation (different' batches of- the same lattice design may be considered as separate fuel types for the void coefficient calculation if their variations ir, k. with void are different); and:

VFy . = volume fraction.of the--thL i fuel type..

L3-12,

NED0-24195 .

Equation 3-4 assumes in the volume weighting of k,'s that each fuel type, controlled and uncontrolled, is the core average void. Only volume weighting of k,'s is assumed. A1.,o, each fuel type is assumed to have the same control fraction.

Equation 3-4 is differentiated with respect to void to obtain the void coefficient (1/kgg) (dk gg/d%V). The resulting equation is:

dk"C dk" VF CF dk eff d i (1 - CF) + d, i=1 dV 1 + M (V)B (V) + ak g (3-5)

N c

+ k[VF(1-CF)+k VF CF

g i -

1+MB22+ah g Since k gg = 1, the (1/k gg) term is disregarded.

The void reactivity associated with hanges in void (V) from a critical condi-tion (VC) is also calculated where:

ak(V) =

kd f (V) - k gg(VC). (3-6) kg f(V) and k,gg(VC) are calculated using Equation 3-4 holding the control fraction constant for each critical condition (VC) and parametrically varying VC. The derivative of the void reactivity at the core average void fraction for which the reactor is just critical is the void coef ficient.

Lattice data for the point model simulation are obtained by performing exposure accumulation steps in increments small enough that bundle power distributions do not vary significantly from the beginning to the end of the step. Exposure accumulation is performed in the uncontrolled state at a 3-13

() fixed in-channel void fraction (normally 40%) with no void fraction in the water gaps and water rods (T saturated = 286*C) at constant pressure.

At selected exposures, lattice calculations are performed in which the in-channel void fraction is changed to 0% and 70%. The result of the three o

calculations is k, at three instantaneous void values: 0, 40%, and 70%. 2

-m The function ak (40% V+V) = k (V) - k g(V = 40%) is graphically demonstrated in Subsection 3.2.4. Therefore, for any exposure E, k, (E,V) = k (E,V = 40%)

+ Akg (E, 40% V+V), k is least squares fit to a quadratic in void; derivatives are calculated using term-by-term differentiation. These deriv-atives are then applied in Equation 3-5.

Controlled k,'s are determined by reinserting the control blade into the uncontrolled lattice at the desired exposure intervals.

3.3.2.1.6 Doppler Coefficient T

,-f

( The Doppler coefficients, as a function of void and exposure, are determined as described in Subsection 3.2.3. ~@

27 The core Doppler coefficient is determined as follows. First, the average proportionality constant, (CDOP) is calculated:

N CDOP = [ CDOP (E,UH) (3-7) i=1 where U, E, and UH are the number of nodes in the core, exposure, and a history dependent void function, respectively. Equation 3-8 is then used to determine the core average Doppler coefficient:

1 dk CDOP (kdT DOP 3-8) 1 + CDOP ( d - 60 ) 2 6 l

2.

u/

3-14

NEDO-24195-

. %p) 3.3.2.1.7 Moderator Temperature Coefficient The moderator temperature coefficient is not a significant reactivity coef-ficient because'its effect is limited to primarily the reactor startup range.

Once the reactor reaches the power producing range, boiling begins and the moderator temperature remains essentially constant. As with the void coef-ficient, the moderator temperature coefficient is associated with a change in the moderating power of-the water. The temperature coefficient is negative during power operation.

The range of values of moderator temperature coefficients encountered in current BWR lattices does not include any that are significant from the safety

-5 point of view. Typically, the temperature coefficient may range from +4x10 Ak/k'F to -14x'.0-5 Ak/k'F, depending on base temperature and core exposure.

The small magnitude of this coefficient (relative to that associated with steam voids) and combined with the long time-constant associated with transfer of heat from the fuel to the coolant, makes the reactivity contribution of hv moderator temperature of small importance, Because of its relative insignificance, current core design criteria do not impose limits on the value of the temperature coefficient. A measure of design control over the temperature coefficient is exercised by applying a design limit to the void coefficient. This constraint implies control over the water-to-fuel ratio of the lattice; this, in turn, controls :he temperature coef-ficient. Thus, imposing a quantitative limit on the void coefficient effectively limits the temperature coefficient.

3.4 FINAL LOADING PATTERN COMPARISON 3.4'.1 Introduction and Bases Because the reload licensing process requires an assumptien as to.the condition of the core at the:end'of the previous cycle, it is possible that.the "as-loaded". core may not be identical:to the. reference-design core. To assure that licensing calculations performed on the reference core are applicable to h

L/

3-15

()

1 the "as-loaded" core, certain key parameters, which affect the licensing claculations, have been identified. Conservative bounds on the amount and restrictions on the manner by which these key parameters may vary from the reference are imposed on the "as-loaded" core. When a reload core cannot meet these restrictions, all the affected licensing parameters must be re-examined to assure that there is no adverse impact; only when this re-examination has been completed and it has been established that the as-loaded core satisfies the licensing basis will the core be operated. This re-examination procedure is described in Subsection 3.4.2, '

o 3.4.2 Re-Examination of Basis R w

A re-examination of the parameters that determine the operating limits must be performed. Based on results of the sensitivity studies of the operating limits to these parameters, conservative bounds have been set on the allowable change from the reference. These parameters are:

() (1) scram reactivity insertion; (2) dynamic void coefficient; (3) peak fuel enthalpy for General Electric-supplied fuel during rod drop accident; (4) cold shutdown margin; (5) change in critical power ratio due to a misloaded General Electric-supplied fuel assembly; and (6) rod block monitor response to rod withdrawal error for a General Electric-supplied bundle.

i l

I

! 3-16 l

l

NEDO-24195 These parameters were chosen by one of the following two criteria:

(1) It is a parameter whose magnitude or behavior is explicitly reported in the reload licensing analysis.

Examples: Cold Shutdown Margin, Peak Fuel Enthalpy in Rod Drop Accident, Change in CPR due to a Misloaded Assembly,

~

k and Rod Block Monitor Response.

(2) It is a parameter important to the quantification of an operating limit.

4 Examples: Scram Reactivity Insertion and Dynamic Void Coefficient Affect MCPR limit.

Other parameters were excluded for the following reasons:

(1) SLCS effectiveness - with no significant change in cold shutdown margin, SLCS effectiveness will not change.

(2) Doppler Coefficient and Delayed Neutron Fraction - There are slowly varying functions of exposure which do not change significantly over the expected range of exposure deviations.

3.5 REFERENCES

3-1. C. L. Martin, " Lattice Physics Methods", General Electric Co.,

February 1977, (NED0-20913A) (NEDE-20913-P-A).

3-2. C. L. Martin, " Lattice Physics Methods Verification", General Electric Co., January 1977, (NEDO-20939A).

3-3. J. A. Wooley, "Three-Dimensional BWR Core Simulator," General Electric Co., January 1977,'(NEDO-20953A).

3-4. G. R. Parkos, "BWR Simulator Methods Verification," General Electric Co., January 1977, (NEDO-20946A).

A 3 3-17

7 NEDO-24195 3-5. R. C. Stirn, " Generation of Void and Doppler Reactivity Feedback for :

Application to BWR Design," General Electric Co., December 1975, !

(NEDO-20964). l 3-6. " General Electric Standard Safety Analysis Report," General Electric Co.,

April 14, 1973 (NEDO-10741).

3-7. " Generation of Void and Doppler Reactivity Feedback for Application to BWR Design (Amendment No. 1)," NEDO-20964-1, July 1977.

3-8. "BWR 4/5/6 Standard Safety Analysis Report Rev. 2," Chapter 4, June 1977.

3-9. " Process Computer Performance Evaluation Accuracy Amendment'1,"

NEDO-23040-1, December 1974.

3-10. "BWR/4 and BWR/5 Fuel Design," NEDE-20944-P (Proprietary) and NEDO-20944, October 1976.

3-11. " General Electric Reload Fuel Application," NEDE-24011-P-A*

Reference refers to the revision of NEDE-240ll-P-A, which is approved by the NRC as of the date a specific analysis is initiated.

3-18

NEDO-24195 Table 3-1 MAXIMUM COLD CLEAN k, ALL P8x8R FUEL DESIGNS IN USE

'o Uncontrolled Controlled

. k, 1.2360 1.0656

)

,) .

l e

Q i

3-19 i

I

NEDE-24195

, O u

Table 3-2 DEFINITIONS OF SOME PEAKING FACTORS Local Peaking Factor - Local peaking factor is the ratio of the heat flux in the highest powered rod at a given plane to that of the average rod in the plane.

Radial Peaking Factor (RPF) - The ratio of the fuel assembly power in a particular assembly to the power of the average fuel assembly.

Axial Peaking Factor (APF) - The ratio of the heat flux at the axial plane of interest to the heat flux averaged over the active length of the fuel (assembly or rod) of interest.

Gross Peaking Factor (GPF) - The product of the radial peaking factor for a fuel assembly and the naximum axial peaking factor for the same fuel assembly.

O - "

Total Peaking Factor (TPF) - The total peaking factor is that peaking factor which, when multiplied by the average linear heat generation rate of a specific bundle type, yields the technical specification limit on MLHGR for that bundle type. This definition of the total peaking factor is presented as an equation in the response to the request on Subsec-tion 5.2.1.5. This response also discusses the relationship of the total peaking factor to the APRM rod block and scram setpoints.

I_h s l 1

3-20 l

1

NEDO-24195 r'N Table 3-3 RELATIONSHIP OF TECHNICAL SPECIFICATIONS AND PEAKING FACTORS Maximum Linear Heat Generation Rate (MLHCR)

The MLHGR is the maximum linear heat generation rate (expressed in kW/ft) in any fuel rod allowed by the technical specifications (tech specs) for a given fuel type. The MLHGR is attained when the product of the local, radial and axial peaking factors in an axial segment of a fuel bundle equals the total peaking factor for that fuel type.

Maximum Average Planar Linear Heat Generation Rate (MAPLHCR)

The MAPLHGR is the maximum average linear heat generation rate (expressed in kW/ft) in any plane of a fuel bundle allowed by the tech specs for that fuel type. This parameter is obtained by averaging the LHCR over each fuel rod in the plane and its lisiting value is selected such that:

(1) the peak clad temperature during the design basis loss-of-coolant

\ accident will not exceed 2200*F, and (2) the LEGR will not exceed the MLEGR in the plane of interest. The MAPLHGR is attained when the product of the gross peaking factor and the average rod peaking factor (1.0) equals the tech spec value.

Minimum Critical Power Ratio (MCPR)

The MCPR operating limit is the minimum CPR allowed by the tech specs for a given bundle type. The CPR is a function of several parameters, the most important of which are bundle power, bundle flow and bundle R-factor. The R-factor is dependent upon the local power distribution, but is only indirectly related to the local peaking factor. The limiting value of CPR is selected for each bundle type such that during the most liniting event of moderate frequency, the CPR in that bundle will not be less than the safety limit CPR.

The MCPR is attained when the bundle power, R-factor, flow, and other relevant parameters combine to yield the tech spec value. Therefore, MCPR is not 7'N directly related to any of the peaking factors described in Table 3-2.

b 3-21

NEDO-24195 0 1.18 -

1.14 -

REGION OF CALCULATED HOT, UNCONTROLLEO 4

1.10 3

1.06 s

z 8

5

~-. h 1.02 i

0.98 -

0.94 -

0.90 -

0.86 I I I I I O 5 10 15 20 25 30 35 EXPOSURE (Gwett)

O Figure 3-1. Calculated Range of Hot, Uncontrolled k ,Versus Exposure l

l l 3-22

NEDO-21495 i

) 1.13 1.12 1.11 1.10 1.09 3

8 m REGION OF

$ 1.08 CALCULATED HOT

. ~'

8 UNCONTROLLED ko.0.0 GWd/t 1.07 l

~

/

1.06 -

1.05 -

1.04 -

3,o3 I I I I I I I O 01 02 0.3 04 0.S 0.6 0.7 0.8 e3, IN-CHANNEL VOID FRACTION

'w )

Figure 3-2. Calculated Range of k ,Versus In-Channel Void Fraction 3-23

NEDO-24195 0.90 O-0.89 0.88 0.87 0.86 REGION OF ,

CALCULATED 0.85 HOT. CONTR OLLE D

_ k.00GWWt i /

0.84 t

0.83 -

t w

0.82 -

0.81 -

0.80 -

0.79 -

0.73 _

I 0.77 -

I l i I I I I O 0.1 0.2 0.3 0.4 0.5 06 0.7 0.8 IN. CHANNEL VolO FRACTION Figure 3-3. Calculated Range of k ,Versus In-Channel Void Fraction 3-24

l l

l NEDO-24195 1.25 1.20 e 1.15 -

3 x ,

3 E REGION OF CALCULATED MAXIMUM LOCAL PEAKING 1.10 -

/

1.05 -

g l l l l I I 1.00 0 5 to 15 20 25 30 35 EXPOSURE (GWdit)

Figure 3-4. Calculated Range of Hot Uncontrolled Maximum Local Peaking Versus Exposure 3-25

l NEDO-24195 0 --

l i

-(L4 -

-0.5 -

-o.6 - ,

-o.7 -

REGION OF CALCULATED DOPPLER COEFFICIENTS

- s 3

~ -o.s -

x Y -o.9 -

4 %

O -=12

' -1.0 -

- t .1 -

-1.2 -

\

~

- 1.3

\

l

-1.4 l ;

- 1.5 O 1000 2UOO 3000 4000 0 TEMPER ATURE (OF)

Figure 3-5. Envelope of Doppler Coefficient Versus Temperature, E = 200 mwd /t 3-26

-0.s

-0.7 -

l I

-02 -

l

-0.9 -

REGION OF CALCULATED DOPPLER

- 1.0 - COEFFICt ENTS N

-i.: -

-~

\

=

g - 1.2 -

Ox J

" F?

-1.3 -

s

.. l }

- 1.4 -

i s

- 1.s -

l

- 1.s -

- 1.7 -

_jj f f I I O 1000 m 3000 4000 l TEMPER ATURE (OF)

Figure 3-6. Envelope of Doppler Coeffic!+ t Versus Temperature, E = 15,000 E'd/t 3-27

NEDO-24195 O _

0.03 -

///////////ll

[

0.0 v0so

//////////// ///////

/ce"cN0TEYA /

/

0.01 3

4 0 -

[

G

- 0.01 -

l l

-0.02 -

l 1

~' ~

Lc TE ag 0.7 VOID UNCONTROLLED (WITH XENON) 0 5 10 15 25 30 EXPOSURE (GWd/t Figure 3-7. Envelope of ak, (From 0.4 Void to Other Voids) Versus Exposure l

3-28 l

D. I U -0.3

-0.4 -

-0.5 -

I

-0.6 -

\

-0.7 REGION OF CALCULATEO 7 DOPPLER COEFFICIENTS

~

-0.8 -

o 2 l x 1 l

8

.= * -0,9 -

el ~ 1.0 -

- 1.1 -

- 1.2 -

- 1.3 -

- 1.4 -

- 1.5 0 1000 2000 3000 4000 TEMPER ATURE (OF)

Figure 3-5. Envelope of Doppler Coefficient Versus Temperature, E = 200 mwd /t i

l 3-29

NED0-24195

-0.6

-0.7 - *

-0.8 - I

- 0.9 -

'N AEGION OF.

CALCULATED DOPPLER

- 1.0 - COEFFICtENTS

-1.1 -

o - 1.2 -

'N T X

L - 1.3 -

em.

- 1.4 -

-1.5 -

- 1.6 -

N

- 1.7 -

_gg I I I I O 1000 2000 3000 4000 f3 TEMPER ATURE (OF)

%) Figure 3-6. Envelope of Doppler Coef ficient Versus Temperature, E = 15,000 E'd/t 3-30 0

0.u 0.03 -

/////////// 0 bdBB1n$////$g.0 VOID 0.01 9

l 0 -

D

- 0.01 -

/. /

-0.02 -

ty y o ag 0.7 VOID UNCONTROLLED (WITH XENON) 0 5 10 15 30 EXPOSURE (GWdh Figure 3-7. Envelope of Ak,(From 0.4 Void to Other Voids) Versus Exposure 3-31/3-32

M 9

0 I E h

m Z

0 m

c C

o E

O O

h

. u, e

NEDO-24195 l 4. STEADY-STATE HYDRAULIC MODELS Core steady-state thermal-hydraulic analyses are performed using a model of the reactor core. For thermal-hydraulic analysis, General Electric uses one of the fuel assembly designs documented in Reference 4-9 in place of any non-GE assembly present in the core. The General Electric-supplied bundle which has .,

e.

a fuel rod geometry and enrichment most like that for a non-General Electric-supplied bundle is used in place of that non-General Electric-supplied bundle.

Chosen replacement bundle designations for the reference cycle are given in the reference cycle supplement. This model includes hydraulic descriptions of orifices, 27 lower tieplates, fuel rods, fuel rod spacers, upper tieplates, the fuel channel and core bypass flow paths. The orifice, lower tieplate, fuel rod spacers, upper tieplate and, where applicable, holes in the lower tieplate are hydraulically represented as being separate, distinct local losses of zero thickness. The fuel channel cross section is represented by a square section with enclosed area equal to the unrodded cross sectional area of the actual fuel channel.

-m The flow distribution to the fuel assemblies and bypass flow paths is cciculated

2f on the assumption that the pressure drop across all fuel assemblies and bypass flow paths is the same. This assumption has been confirmed by measuring the flow distribution 1,n boiling water reactors (References 4-1, 4-2, 4-3). The components of bundle pressure drop considered are friction, local, elevation and acceleration (see Subsections 4.1 through 4.4). Pressure drop measurements made in operating reactors confirm that the total measured core pressure drop and calculated core pressure drop are in good agreement. There is reasonable assurance, therefore, that the calculated flow distribution throughout the core is in close agreement with the actual flow distribution of an operating reactor.

An iteration is performed on flow through each flow path (fuel assemblies and bypass paths), which equates the total differential pressure (plenum to plenum) across each path and matches the sum of the flows through each path to the total core flow. The total core flow less the control rod cooling flow enters the lower plenum. A fraction of this passes through various bypass paths

(^]

documented in Subsection 4.5. The remainder passes through the orifice in the fuel support (experiencing a pressure loss), where more flow is lost through 4-1

I NEDO-24195

(

')

7' the fit-up between the fuel support and the lower tieplate and through the lower tieplate holes, if present, and into the bypass region. The nuwoer of fuel assemblies which have lover tieplate holes in the reference cycle will be o

indicated in the reference cycle supplement. The majority of the ao

~

flow continues through the lower tieplate (experiencing a pressure loss),

where some flow is lost through the flow path defined by the fuel channel and lower tieplate into the bypass region. This bypass flow is restricted on those fuel assemblies with finger springs.

Within the fuel assembly, heat balances on the active coolant are performed nodally. Fluid properties are expressed as the bundle average at the particular node of interest and are based on 1967 International Standard Steam-Water Properties. In evaluating fluid propertias, a constant pressure model is used.

1 l The relative axial / power (given in Table 5-6) and radial power distributions g3 are used with the bundle flow to determine the axial coolant property distribu-

~

tion, which gives sufficient information to calculate the pressure drop com-

) ponents within each fuel assembly type. When the equal pressure drop criterion Y described above is satisfied, the flow distributions are established, i

l 4.1 FRICTION PRESSURE DROP 1

Friction pressure drop is calculated with a basic model similar to that used throughout the nuclear power industry. This model is:

2 w fL 2

(~

~

f 2gp 2 TPF A

H ch 1

where I

aP g = friction pressure drop (psi);

w = mass flow rate; g = acceleration of gravity; l("'j c = average nodal liquid density; j I

D = channel hydraulic diameter; H

l 4-2 l l

l

Ach = channel flow area; I

L = incremental length; f = friction factor; and 2

$ = two-phase friction multiplier.

TPF The formation for the two-phase multiplier is based on data that compares closely to that found in Reference 4-4.

Significant amounts of friction pressure-drop data in multirod geometries representative of modern BWR plant fuel bundles have been taken. Both the

~

friction factor and two-phase multipliers were correlated on a best-fit basis using the above pressure drop formulation. Checks are made on a con-tinuing basis to ensure the best models are used over the full range of interest to boiling water reactors.

4.2 LOCAL PRESSURE DROP ib The local pressure drop is defined as the irreversible pressure loss associ-ated with an area change, such as the orifice, lower tieplate, and spacers of a fuel assembly.

The general local pressure drop model'is similar to the friction pressure l

j drop and is:

q AP =

4 p (4-2) where

~

j. AP = local pressure drop (psi);

l K ,=' local pressure drop loss coefficient; J

A = reference area'for local loss coefficient; 2 ~

$ = two-phase local multiplier; TPL.

~

'4-3

NEDO-24195 and w, g, and o are defined above. This basic model is also similar to that used throughout the nuclear power industry. The formulation for the two-phase multiplier is similar to that reported in Reference 4-5. Empirical constants were added to fit the results to data taken for the specific designs of the .

boiling water reactor fuel assembly. These data were obtained from tests per-

.e k

formed in single-phase water to calibrate the orifice, the lower tieplate, and -

the holes in the lower tie place, and in both single- and two-phase flow, to derive the best-fit design values for spacer and upper tieplate pressure drop.

(

The range of test variables was specified to include the range of interest for boiling water reactors. New test data are obtained whenever there is a signifi-cant design changes to ensure the most applicable methods are used.

4.3 ELEVATION PRESSURE DROP The elevation pressure drop is based on the relation:

AP "

-obi c E (4~3) n V o =

A g(1 - a) + o ga *

(4-4) where APg = elevation pressure drop (psi);

L = incremental length; o = average water density; a = nodal average void fraction; o,o g = saturated water and vapor density (resp); and g = acceleration due to gravity.

The voj iraction model used is an extension of the Zuber-Findlay (Reference 4-6) model a uses an empirical fit constant to predict a.large block of steam void l fraction data. Checks against new data are made on a continuing basis to I ensure the best models are used over the full range of interest to boiling water reactors.

4-4

NEDO-24195 4.4 ACCELERATION PRESSURE DROP A reversible pressure cbange occurs when an area change is encountered, and an irreversible loss occurs when the fluid is accelerated through the boiling process. Thu basic formulation for the reversible pressure change resulting from a flow area change in the case of single-phase flow is given by:

AP ACC

1~# 2 (4-5) 2gc pg 2A -

^2 A

~

final flow area

(~

{ ~ initial flow area where AP = acceleration pressure drop; CC

""d A2 " fi"*1 fl " ##*"I d A = initial flow area.

7 In the case of two-phase flow, the liquid density is replaced by a density ratio so that the reversible pressure change is given by:

2 AP ACC

" 1~#

2 (4-7) 2g P e KE ^2 2

i.

where

~

+ , honogeneous density;

P #

H g 1

~

=

pa 2+ o 2 (1 - a)2 , kinetic energy density; e.g g g O

4-5 4

NED0-24195 a = void fraction at A ;

2 x = steam quality at A 3 2

and other terms are as previously defined. The basic formulation for the accelerstion pressure change due to density change is:

" ~

ACC 2 gA OUT IN-ch -

where 2 2 1

OUT (1 ~ *0UT PM OUT P g "OUT # (

1 ~ "OUT}

n , ,

f v

) _1_ ,

1N

+

( ~ *IN D

M IN D

g "IN D 1 1 "IN and is evaluated at the inlet and outlet of each axial node. Other terms are as previously defined. The total acceleration pressure drop in boiling water reactors is on the order of a few percent of the total pressure drop.

4.5 BYPASS FLOW The bypass flow paths considered in the analyses are described in Table 4-1 and are shown schematically in Figure 4-1. Typical values of the fraction of bypass flow through each flow path are given in Reference 4-7.

Due to the low flow velocity, the pressure drop in the bypass region above the core support plate is essentially all elevation head. Thus, the sum of the core support plate differential pressure and the bypass region elevation head is 7

equal to the core differential pressure.

1 4-6

NED0-24195

~-

The flow through the bypass flow paths is expressed by the form:

W = C aP 1

+C a? +C 3

AP . (4-9) 2 Full-scale tests, documented in Reference 4-8, have been performed to estab-lish the flow coefficients for the major flow paths. These tests simulate actual plant configurations, which have several parallel flow paths and, therefore, the flow coefficients for the individual paths could not be separated. However, analytical models of the individual flow paths were developed as an independent check of the tests. The models were derived for actual BWR design dimensions and considered the effects of dimensional '

variations. These models predicted the test results when the as-built dimen-sions were applied. When using these models for hydraulic design calculations, nominal drawing dimensions are used. This is done to yield the most accurate prediction of the expected bypass flow. With the large number of components in a typical BWR core, deviations from the nominal dimensions will tend to j statistically cancel, resulting in a total bypass flow best represented by that calculated using nominal dimensions.

Increases in channel wall permanent deflection at the lower tieplate result in increased bypass flow through the channel to lower tieplate flow path.

For plants without finger springs, this flow path contributes a significant portion (-80%) of the total core bypass flow. Changes in the flow through this path affect the total core bypass flow, which, in turn, affects the active coolant flow, void coefficient and operational transients. To provide control over this flow path, optional finger spring seals, which control the flow rate through the channel to lower tieplate flow path over a wide range of channel wall deflections, may be added to most reload fuel assemblies.

Finger springs are located between the lower tieplate and the channel for the purpose of controlling the flow through that path. The finger springs provide control of the flow through this path by maintaining a nearly constant flow area as the channel wall deforms. A mechanical description of finger spring seals is contained in Section 2.

4-7

NED0-24195 q

Full-scale tests have been performed on. finger spring seals to determine the flow characteristics through the lower-tieplate and fuel channel flow path as a function of pressure drop and channel deformation. The pressure drop and channel deflections used in the tests covered the range expected during reactor service. The fuel bundle components used for the tests were standard lower tieplates, finger springs and a standard channel section. The flow i characteristics as determined by these tests are used in design calculations J

(see Figure 4-19 in Reference 4-7).

]

In cases where finger spring seals are installed on some reload fuel bundles, the hydraulic analysis takes into account the differences between finger .,

spring and nonfinger spring fuel.

Because finger spring seals reduce the bypass flow through the lower tieplate-channel path, the active coolant flow in a bundle with finger spring seals will normally be larger than the flow in a bundle without the seals. This is true, assuming that both bundles are operating at the same power and have identical channel-to-lower-tiephte deflections. To accommodate this varia-tion, for otherwise identical fuel assemblies, the bypass flow characteristics are individually modeled in the analysis. .The model yields lower-tieplate and channel bypass flow as a function of pressure drop for defined channel deflection at the lower tieplate, and the existence (or absence) of finger spring seals.

Variations in bundle flow rates for fuel assemblies with and without finger spring seals are handled by supplying separate hydraulic constants for the finger spring seal and nonfinger spring seal bundles. This ensures that the core flow distribution is properly . calculated.

4.6 REFERENCES

4-l' Core Flav Distriktion in a Modern Boiling Water Reactor as Measured in Monticello, NED0-10299A, October 1976.

4 H. T.' Kim and H. S. Smith, Core Flou Distriktion in a General Electric -

.p Boiling Water Reactor as Measured in Quad-Cities Unit 1, NEDO-10722A,

.t/ August 1976.

4-8'

3 4

4A,y

/

g$*%

1 p++/ _ eEEm e T,em %*%*

TEST TARGET (MT-3) l 1.0 gmsn S B3 E=n =

na l,l h.," lll!b l.8

\

1.25 1.4 1.6

< 6" --+

MICROCOPY RESOLUTION TEST CHART

/f%

t?'%s%g ~

' / 4

?Rhs,,,,, .

, '4A>(b n/fl,l.l .

-_q g m_

-: .gh.

, t '

l, 'Y

L.-.. . ..---..- wL _._ n :.- -

NED0-24195 4-3 " Brunswick Steam Eleccric Plant Unit 1 Safety Analysis Report for Pl Tt i Modifications to Eliminate Significant In-Core Vibrations," HED0-21215,

March 1976.

4-4 R. C. Martinelli and D. E. Nelson, " Prediction of Pressure Drops During 1

Forced Convection Boiling of Water," AS.VF Trans., 70, 695-702, 1948.

l 4-5 C. J. Baroozy, A Systematic Correlation for Bx-Phase Pressure Drop, j

Heat Transfer Conference (Los Angeles), AICLE, Preprint No. 37, 1966.

1 1

4-6 N. Zuber and J. A. Findlay, h>erage Volumetric Concentration in Tuc-

Phase Flou Systems, Transactions of the ASME Journal of Heat Transfer, ,

November 1965.

4-7 "BWR Fuel Channel Mechanical Design and Deflection," NEDE-21354-P (Proprietary) and NED0-21354, September 1976.

I 4-8 " Supplemental Information for Plant Modification to Eliminate Signif-icant In-Core Vibration," NEDE-21156 (Proprietary), February 1976.

i

4-9 " General Electric Reload Fuel Application," NEDE-240ll-P-A*

O i

l 1

I i

l i

r

Reference refers to the revision of NEDE-24011-P-A, which has been

() approved by the NRC as of the data a specific analysis is initiated.

4-9

NED0-24195 s

) Table 4-1 BYPASS FLOW PATHS Flow Path Description Driving Pressure Number of Paths la. Between Fuel Support Core Support Plate One/ Control Rod and the Control Rod Differential Guide Tube (Upper Path) lb. Between Fuel Support Core Support Plate One/ Control Rod and the Control Rod Differential utide Tube (Lower Path)

2. Between Core Support Core Support Plate One/ Control Rod Plate and the Control Differential Rod Guide Tube

3. Between Core Support Core Support Plate One/ Instrument Plate and the In-Core Differential Support Instrument Guide Tube

, 4. Between Core Support Core Support Plate One Plate and Shroud Differential 5 Between Control Rod Core Support Plate One/ Control Rod Guide Tube and Control Differential Rod Drive Housing

6. Between Fuel Support Channel Wall One/ Channel and Lower Tieplate Differential Plus Lower Tieplate Differential

7. Control Rod Drive Independent of Core One/ Control Rod Coolant

8. Between Fuel Channel Channel Wall One/ Channel and Lower Tieplate Differential

9. Holes in Lower Tieplate Channel Wall Two/ Fuel Assembly Differential Plus (where applicable)

Lower Tieplate Grid Differential 4

4-10

NEDO-24195 NOTE: PERIPHERAL FUEL SUPPORTS ARE WELDED INTO THE CORE SUPPORT p PLATE. FOR THESE SUNOLES.

PATH NUMBERS 1. 2. 5 AND 7 00 NOT EXIST.

. CHANNEL i

, _ _ _ _ _ _ _ 3 l

/ -

LOWER TIE PLATE CORE SUPPORT PLATE I '

FUE L SUPPOR T 1a 3 4 1

r h\

IN-COR E GulOE TUBE I

/

T-g ===

tb . CONTROL ROO SHROUD GutOE TU8E p

I CONTROL ROD GutOE TusE-FUEL SUPPORT PIECE

2. CONTROL ROO GuiOE TUGE CORE SUPPORT PLATE l 3. CORE $UPPORT PLATE INCORE GulOE TUBE 4 CORE SUPPORT PLATE SHROUO 5 CONTROL ROO GulOE TUBE ORIVE HOUSING O 7 6 FUEL SUPPORT PIECE LOWER TIE PLATE

7. CONTROL ROO ORIVE COOLING WATER

8. CHANNEL LOWER TIE PLATE A ,

D.

LOWER TIE PLATE HOLES (WHERE APPLICA8LE) 5

. CONTROL ROD ORIVE HOU$tNG lI Figure 4-1.

Schematic of Reactor Assembly Showing the Bypass Flow Paths 4-11 n

N

. m

1 l

l k

1 3 !

l 1

r

?:EDO-24195 h

5. REACTOR LIMITS DETERMINATION Limits on plant operation are established to assure that the plant can be sa-fely operated and not pose any undue risk to the health and safety of the public. This is accomplished by demonstrating that radioactive release f rom plants for normal operation, abnormal operational transients and postulated accidents meet applicable regulations in which conservative limits are docu-mented. This conservatism is augmented by using conservative evaluation models and observing limits which are more restrictive than those documented in the applicable regulations. These observed operating limits are given below.

The obj ective for normal operation and transient events is to maintain nucleate boiling and thus avoid a transition to film boiling. Operating limits are specified to maintain adequate margin to the onset of the boiling transition.

For General Electric-supplied fuel, the figure of merit utilized for plant operation is the critical power ratio. This is defined as the ratio of the critical power (bundle power at which some point within the assembly experience

) onset of boiling transition) to the operating bundle power. The critical power is determined at the same mass flux, inlet temperature and pressure which exists at the specified reactor condition. Thermal margin is stated in terms of the minimum value of the critical power ratio (MCPR), which corresponds to the most limiting General Electric-supplied fuel assembly in the core. To ensure that adequate margin is maintained, a design requirement based on a statistical analysis was selected as follows:

Moderate frequency transients caused by a single operator error or equipment malfunction shall be limited such that, considering uncer-tainties in manufacturing and monitoring the core operating state, more than 99.9% of the fuel rods would be expected to avoid boiling transition (Ref erence 5-1) .

Both the transient (safety) and normal operating thermal limits for General Electric-supplied fuel in terms of MCPR are derived from this basis. A dis-cussion of these limits is given in Subsection 5.1 and 5.2, respectively.

D 5-1

NED0-24195 l

The ASME boiler and pressure vessel code and other codes and standards require that the pressure relief system prevent excessive overpressurization of the primary system process barrier and the pressure vessel. The allowable pressure and prescribed evaluations are determined by these requirements. The analysis performed to demonstrate conformance to the requirements for General Electric-supplied fuel is documented in Subsection 5.3.

Three types of stability are considered by General Electric in the design of boiling water reactors: (1) reactor core (reactivity) stability; (2) channel hydrodynamic stability; and (3) total system stability. A stable system is analytically demonstrated if no inherent limit cycle or divergent oscillation develops within the system as a result of calculated step disturbances of any critical variable, such as steam flow, pressure, neutron flux or recirculation flow. The criteria for evaluating reactor dynamic performance and stability are stated in terms of two compatible parameters. First is the decay ratio (x /x ), which is the ratio of the magnitude of the second overshoot to the first overshoot resulting from a step perturbation. A plot of the decay ratio is a r

graphic representation of the physical responsiveness of the system which, is readily evaluated in a time-domain analysis. Second is the damping coef ficient (c ), the definition of which corresponds to the dominant pole pair closest to the imaginary axis in the e-plane for the system closed-loop transfer function.

As c decreases, the closed-loop roots approach the imaginary axis and the response becomes increasingly oscillatory. This parameter also applies to the f requency-domain interpretation. Limits for these parameters are given in Subsection 5.4.

The effects of the various postulated accidents on General Electric-supplied fuel are investigated for a variety of plant conditions in Subsection 5.5.

Acc ident limits are specified such that:

(1) calculated radioactive material release do not exceed the guide-line value of 10CFR100; (2) catastrophic f ailure of fuel cladding, including f ragmentation of

! fuel cladding and excessive fuel enthalpy is not predicted; 5-2

NEDO-2'.195 O (3) nuclear system or containment (when required) stresses in excess of those allowed for accidents by applicable codes will not result; and (4) overexposure of plant operating personnel due to radiation in the control room will not occur.

5.1 FUEL CLADDING INTEGRITY SAFETY LIMIT The generation of the minimum critical power ratio (MCPR) limit requires a sta-tistical analysis of the core near the limiting MCPR condition. The statistical d

analysis is used to determine the MCPR corresponding to the transient design requirement given in Section 5 for General Electric-supplied fuel. This MCPR established fuel cladding integrity safety limit applies not only for core-wide transients, but is also conservatively applied to the localized rod with-drawal error transient.

5.1.1 Statistical Model The statistical analysis utilizes a generalized model of the BWR core which simulates the process computer function. This code produces a critical power ratio (CPR) map of the cora based on inputs of power distribution and flow and on heat balance information.

A nominal reactor case CPR calculation for each fuel bundle in the core is made by adjusting the design power distribution so that as many fuel assemblies as possible are at or near the MCPR limit. Details of this procedure are docu-mented in Appendix IV of Reference 5-1. Random Monte Carlo selections of all j operating parameters based on the uncertainty ranges of manufacturing toler-I ances, uncertainties in measurement of core operating parameters, calcula-

~

i tional uncertainties, and statistical uncertainty associated with the CEXL boiling transition correlation are imposed upon the analytical representation of the core, and the resulting bundle critical power ratios are calculated.

Sensitivity analyses of the input measurements have permitted elimination of A

N] '

5-3

NEDO-24195 O '

other variables that do not contribute significantly to the uncertainty in the variable of primary interest. In this analysis, any probability less than 10 of the maximum was disregarded.

The rod-by-rod R-factor (used in GEXL) is employed to predict the critical power ratio and ultimately the probability of a boiling transition for each rod. The number of rods expected to experience boiling transition in each trial is then ,

equal to the summation of all the fuel rod probabilities for that trial.

The minimum allowable critical power ratio is set to correspond to the criterion that 99.9% of the rods are expected to avoid boiling transition by interpolation among the means of the distributions formed by all the trials.

For the purposes of determining the core-wide safety limit, an entire core g

of General Electric fuel was assumed. )

) 5.1.2 ?ounding BWR Statistical Analvsis d

Bounding statistical analyses have been perfomed which provide conservative safety limit MCPR's applicable to reload cycles for BWR/2 plants including Oyster Creek. The reactor core selected for these analyses was a 251/764 reload core. The large reload core analysis results conservatively applies to Oyster Creek for all General Electric-supplied reload cycles.

The histograms of relative bundle powers and corresponding CPRs used in the i

statistical analysis for P8x8R reloads are shown in Figures 5-1 and ':-2, ;

respectively. The design basis CPR distribution used in the statistical l

analysis is skewed more to the low MCPR side than actual operating data, thus yielding a conservative value of the 99.9% statistical limit MCPR.

The uncertainty inputs used in the bounding statistical analyses are listed in Table 5-1. Although some of the plant-unique uncertainties may be greater for Oyster Creek, other uncertainties are smaller and, thus, the bounding analysis is conservative. Nominal values of parameters used in the bounding b statistical analyses are listed in Table 5-2.

5-4 L.

.- s '

NEDO-24195 l

The rod-by-rod R-factor distributions used for the bounding statistical analysis are summarized in Reference 5-25. The basis for the additive con- :

stants used to determine the P8x8R R-factor is documenced in Reference 5-2. ?

Results of the analyses show that at least 99.9% of the fuel rods in the ,

core are expected to avoid boiling transition if the MCPR for P8x8R reloads is 1.07 or greater.

O 5.2 MCPR OPERATING LIMIT CALCULATIONAL PROCEDURE t

~

~ ~

A plant-uii1q6e~ MCPR~bp'er'ai1'nilimit is est'ablished for General Electric-supplied R

_m :

fuel to ensure that the fuel cladding integrity safety limit for that fuel is -

not violated for any moderate frequency transient. This operating requirement is obtained by addition of the absolute, maximum CPR value for the most limit-ing transient from rpted conditions postulated to occur at the plant to the fuel [

cladding integrity safety limit.

. . _ _ _ . _ . _ . _ . ___-__ i O c r - 14 c 1 == 1c Pr aiccea e r 1== ca r t ea 1 aece- :

mented in References 5-3, 5-4, and 5-5. This model is based upon a point.

i reactor kinetics model, multinoded thermal-hydraulic and heat transfer rela- i tionships, and mechanical kinetic equations of the equipment. A worse, usually [

m maximum, power condition is assumed with thermally limited fuel conditions. 6 Q j 9

The philosophy with respect to using the equipment performance components of -

l this model for design and safety evaluations is to consider the performance i

of key components at their adverse tolerances. Circuitry delays in the reactor !

protection system, as well as other key equipment circuit delays, are assumed !

at the maximum specified values. The speed of all the control rod drives

[

O re11e 1== cr 1 a ce = ew P #c eec8=1c 1 P cieic cie 1 1= - i Field data have shown considerable conservatism in this key component per- [

formance. The setpoints for the safety / relief valves both in the safety and l

relief function and for pressure scram are assumed at their specified limits. t The assumed setpoints used in the analysis are given in Table 5-3. Other equipment performance such as relief and safety valve opening characteristics, !

recirculation pump drive train inertia and main steamline isolation valve closing ,

times are all assumed to be a adverse tolerances.

5-5

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

o NEDO-24195 End-of-cycle (EOC) conditions for nuclear input data are used (except where specific exposure-dependent evaluations are performed, Subsection 5.2.2) to provide a varying level of conservatism associated with core exposure aspects.

The only significant new input data which are used in reload evaluations are the scram reactivity function, void reactivity function and Doppler reactivity f unc t ion. The basis for using E0C Doppler and void reactivity functions in the analyses is outlined in Reference 5-6. A discussion of the above signi-ficant parameters follows.

Scram reactivity is the worth of control rods as a function of time or position following the scram signal. The scram reactivity insertion is normally lowest at the end of cycle (all-rods-out condition) because there are no stubbed rods to insert negative reactivity more quickly than the remaining blades of the con-trol rod bank. The scram reactivity function is related to dynamic performance when expressed (as plotted) in the form of ak/k3.

The void reactivity coef ficient is an important parameter, not only in trans-k' ient analysis, but also in core stability. The core average void coefficient must be negative; however, it must not be so negative as to yield such a strong positive reactivity f eedback during void collapse events that core and vessel limits are threatened. Conversely, events with void increase must produce suf-ficient negative f eedback to maintain operation within saf ety limits. A trans-ient index which is used to assess the void reactivity characteristics is the ,

dynamic void coefficient. This parameter is defined as the core physics void coefficient multiplied by the average full power void fraction and divided by i the delayed neutron fraction.

The presence of U-238 and, ultimately, Pu-240 contributes to yield a strong negative Doppler coef ficient. This coefficient provides instantaneous nega-tive reactivity f eedback to any fuel temperature rise, either gross or local. !

The magnitude of the Doppler coefficient is not dependent on gadolinium posi-tion or concentration in any bundle because gadolinium has very little ef fect l

l on the resonance group flux or on U-238 content of the core. The core physics O

5-6

NEDO-24195 h

Doppler coefficient is divided by the delayed neutron fraction to define a dynamic Doppler coefficient (g), which most effectively correlates the dynamic response of the plant to the Doppler reactivity feedback.

Additional conservatism is introduced into the analysis to account for biases and uncertainties in the derivation of the nuclear data and its application to the transient model. The conservatism factors used to account for these biases and uncertainties are summarized in Table 5-4. These values, in effect, represent a maximum level of conservatism u. the transient analysis upon which compliance to existing safety limits is generally based.

The reactor core behavior for General Electric-supplied fuel during the rod withdrawal error transient is calculated by doing a series of steady-state three-dimensional coupled nuclear thermal-hydraulic calculations using the three-dimensional BWR simulator (Reference 5-7). This approach assumes that the transient is very slow such that there is sufficient time for heat c transfer and void redistribution to equilibrate and also that the neutron L

k(D J flux and heat flux are in equilibrium with each other. This calculation is achieved by maintaining a constant eigenvalue via increasing core average power as a control rod is withdrawn incrementally.

Descriptions of the transient events are given in Subsection 5.2.1. Inputs to these transients which are specific to the Oyster Creek plant are given in Tables 5-3 and 5-5. Transient inputs which are specific to the reference cycle are~ o x

given in the reference cycle supplement. Because the transient model establishes 2

~

operating conditions, only licensing basis values are given in Table 5-5.

Actual values used in the analyses will be within the tolerances shown in the !

table. The transient descriptions given in Subsection 5.2.1 are used as a basis for the typical analyses performed for plant reloads. Oyster Creek analyses will differ in certain aspects from the typical calculational pro-cedure due to the util ty selected margin improvement option of exposure-dependent limits. A description of this option and its effect upon the transient analysis is given in Subsection 5.2.2. ATWS pump trip is also assumed in the analysis of Oyster Creek.

.)

5-7

4 NEDO-24195 The operating limit MCPR for General Electric fuel for rapid transients is calcu-laced by using the SCAT computer program (Reference 5-8). Inputs to this pro-gram consist of the transient analysis results, steady-state flow distribution (from Section 4), bundle power, axial distribution, and gap conductance. The axial power distribution used in the analyses is given in Table 5-6. Nonvarying plant initial conditions for the GETAB analysis are given in Table 5-7. Reload-dependent plant initial conditions for the GETAB analysis and the resulting reload g MCPR operating 1 %it for General Electric fuel in the reference cycle are given in ,S the reference cycle supplement. The initial MCPR assumed for transient analyses is usually greater than or equal to the GETAB

]s*E operating limit. Figure 5-3 illustrates the effect of the initial MCPR on transient ACPR for a typical BWR core. This figure indicates that the change in ACPR is approximately 0.01 for a 0.05 change in initial MCPR. Therefore, nonlimiting GETAB transient analyses may be initiated from a MCPR below the operating limit because the higher operating limit MCPR more than off sets the increase in ACPR for the event. This may also be applied to limiting transients if the difference between the operating limit and the initial MCPR is small h? (0.01 or 0.02). Densification power spiking is not considered in establishing the MCPR operating limit. Justification for this is presented in Sub-section 5.2.3.

The operational limit MCPR must be increased for low flow conditions. This is because, in the BWR, power increases as core flow increases, which results in a corresponding lower MCPR. If the MCPR at a reduced flow condition were at the 100% power and flow MCPR operating limit, a sufficiently lart,e ipsdvertent flow increase could cause the MCPR to decrease below the Safety Limit MCPR given in Subsection 5.2.1. Therefore, the required operating limit MCPR is increased at reduced core flow rates by a flow factor, K g, such that:

Required MCPR ,g , MCPR Operating Limit Operating Limit f @ 100% core flow The flow f actor (K g ), as a function of the core flow rate, the flow control mode, and (in the manual flow control mode) the maximum possible core flow 5-8

(

rate is given in Figure 5-4. Kg was derived by considering the potential range of safety limit MCPR's and applying a generic set of lines constructed by the i

above procedure. Therefore, the generic lines become more conservative as the operating limit MCPR increases, t

h Events such as loss of feedwater heating and the turbine trip without bypass l become less severe when initiated from power levels less than the design value. !

As power is reduced, total steam flow decreases. In the case of the loss in t O feedwater heating, the reduced steam flow results in a decrease in both feed- '

water flow and the maximum temperature rise across a given heater. The core !

t subcooling change associated with the loss in feedwater heating will be less, >

as will the positive reactivity insertion. Therefore, the event becomes less severe. If the turbine trip without bypass events were initiated from a reduced power state, the pressure rise and the resulting void collapse caused

by closure of the turbine stopvalve would be less because of the decreased steam flow, and the event is less severe. ' Figure 5-5 shows the effect of a turbine trip without bypass at reduced initial power level upon the resultant dome pressure and surface heat flux of an operating BWR. Results for Oyster O' Creek are similar. At lower power states, events such as inadvertent startup j

of an idle recirculation pump, recirculation flow controller failure (increas-

{

ing flow), feedwater flow controller failure (maximum demand) and rod withdrawal i error can become more severe than transients which are limiting at design con-ditions. However, the increased reduction is less than the increase in K at

. r g

the lower power and flow of the initial operating limit MCPR. An example of this i is given in Table 5-8 for rod withdrawal error. !

Transient analyses are performed at the full power, EOC, all-rods-out condition.

Once an individual plant reaches this condition, it may shutdown for refueling or it may be placed in a coastdown mode of operation. In this type of opera-tion, the control rods are held in the all-rods-out position and the plant is allowed to coastdown to a lower percent of rated power while maintaining rated '

core flow. The power profile d ing this period is assumed to be a linear function with respect to exposura. It is expected that the actual profile O

5-9 l

NEDO-24195 O will be a slow, exponential curve. An analysis tc the linear approximation, however, will be conservative, since it overpredicts the power level for any given exposure.

In Reference 5-9, evaluations were made to 40% power level points on the linear curve. The results show that the pressure margins from the limiting pressurization transient and the MCPR operating limits exhibit a larger margin for each of these points than the E0C full power, full flow case. MAPLHGR limits for the full power, rated flow case is conserva tive for the coastdown period, since the power will be decreasing and rated core flow will be main-tained (Subsection 5.5) . Therefore, it can be concluded that the coastdown operation beyond full power operation is conservatively bounded by the analysis at the EOC conditions.

5.2.1 Transient Descriptions

} Eight nuclear system parameter variations can pose potential deleterious V effects to the Nuclear Steam Supply System. The parameter variations are as follows:

(1) Nuclear system pressure increase - threatens to rupture the reactor coolant pressure boundary from internal pressure. Also, a pressure increase collapses the voids in the moderator. This causes an inser-tion of positive reactivity which may result in exceeding the fuel cladding safety limits.

(2) Reactor vessel water (moderator) temperature decreases - results in an insertion of positive reactivity as density increases. Positive reactivity insertions threaten the fuel cladding safety. limits because of higher power.

(3) Positive reactivity insertion - ic possible from causes other than nuclear system pressure or moderator temperature changes. Such reactivity insertions threaten the fuel cladding safety limits because of higher power.

5-10

i ,

NEDO-24195 j (4) Reactor vessel coolant inrentory decrease - threatens the fuel as the coolant becomes unable to maintain nucleate boiling. :

i (5) Reactor core coolant flow decrease - threatens the fuel cladding [

i safety limits as the coolant becomes unable to maintain nucleate boiling.

(6) Reactor core coolant flow increase _ - reduces the void content of the moderator, resulting in a positive reactivity insertion. The result- ,

ing high power may exceed fuel cladding safety limits.

\

(7) Core coolant temperature increase - could 4xceed fuel cladding {

t safety limits.

sl l

I (8) Excess of coolant inventory - could result in damage resulting from !

I excessive cartyover. !

Although the FSAR contains many transient descriptions and analyses, only a few of these transients would result in a significant reduction of MCPR.

r To determine the limiting transient events, the relative dependency of CPR upon various thermal-hydraulic parameters was examined. A sensitivity study was performed to determine the effect of changes in bundle power, bundle flow. -

m subcooling, R-factor and pressure on CPR. R

.e Results of the study are given in Table 5-9. As can be seen from this table, '

CPR is most responsive to fluctuations in the R-factor and bundle power. A (

slight sensitivity to pressure and flow changes and relative independence to !

changes in inlet subcooling was also shown. The R-factor is a function of l bundle geometry and local power distribution and is assumed to be constant :

throughout a transient. Therefore, transients which would be limiting because O

5-11

NEDO-24195 O .

\i e

of MCPR would primarily involve significant changes in power. Based on this, the transients most likely to limit operation because of MCPR considerations are:

(1) turbine trips or generator load rejection without bypass; (2) loss of feedwater heating; (3) feedwater controller failure (maximum demand); and (4) control rod withdrawal error.

Descriptions of the above limiting events are given below. The analytical results of the most limiting transient in each of the above types of events $,

is provided in the reference cycle supplement.

5.2.1.1 Generator Load Rej ection Without Bypass Fast closure of the turbine control valves is initiated whenever electrical grid disturbances occur which result in significant loss of load on the generator. The turbine control valves are required to close as rapidly as possible to prevent overspeed of the turbine generator rotor. The closing causes a sudden reduction of steam flow, which results in a nuclear system pressure increase. The reactor is scrammed by the fast closure of the turbine control valves.

Starting Conditions and Assumptions. The following plant operating conditions and assumptions form the principal bases for which reactor behavior is analyzed during a load rejection:

(1) reactor and turbine generator are initially operating at full power when the load rej ection occurs; (2) all of the plant control systems continue normal operation;

("t (3) auxiliary power is continuously supplied at rated frequency;

()

5-12

NEDO-24195 t

~%

Q(

(4) the reactor is operating in the manual flow control mode when load rejection occurs, although the results do not differ significantly for operation in the automatic flow control mode; and (5) the turbine bypass valve system is failed in the closed position.

Event Description. Complete loss of the generator load produces the following sequence of events:

(1) The acceleration relay detects the load and initiates fast closure of the turbine control valves. The turbine accelerates at a maxi-mum rate until the valves start to close. Oyster Creek has a MHC control system; thus, turbine control valve has a nonlinear closure signature that is a function of the MHC settings.

(2) Reactor scram is initiated upon sensing control valve fast closure, 0, (3) If the pressure rises to the pressure relief setpoint, part, or all, of the relief valves opens, discharging steam to the suppression pool.

(4) If the pressure rises to > 1060 psig, trip of recirculation pump drive motors occurs.

=

U 5-13

()

E 4

s 5.2.1.2 Turbine Trip Without Bypass A variety of turbine at nuclear system malfunctions will initiate a turbine trip. Some examples are: moisture separator and heater drain tank high levels, large vibrations, loss of control fluid pressure, low condensor vacuum and reactor high water level. The turbine stopvalve closes, causing a sudden reduction in steam flow, which results in a nuclear system pressure increase and the shutdown of the reactor.

Starting Conditions and Assumptions. The plant operating conditions and assumptions are identical to those of the generator load rejection.

5-14

NED0-24195

/ Event Description. The sequence of events for a turbine trip is similar to that for a generator load rejection. Stopvalve closure occurs over a period of 0.10 sec.

Position switches at the stopvalves sense the turbine trip and initiate reactor scram. If the pressure rises to the pressure relief setpoints, relief valves open, discharging steam to the suppression pool.

S 2

5.2.1.3 Loss of Feedwater Heating Loss of feedwater heating results in a core power increase due to the increase in core inlet subcooling.

O. Starting Conditions and Assumptions. The following plant operating conditions and assumptions form the principal basis for which reactor behavior is analyzed during the. loss of feedwater heating transient:

(1) The plant is operating at full power.

(2) The plant is operating in the manual flow control mode.

Feedwater heating can be lost when steam extraction line to heater is closed.

E

=

4

/

a 5-15

(-)

5/

T1.is produces a gradual cooling of the feedwater, allowing the reactor vessel E!

to receive cooler feedwater. The maximum number of feedwater heaters which can be tripped or bypassed by a single event represents the most severe transient for

, analysis considerations. An instantaneous loss of the feedwater heating capability of the plant causes an increase in core inlet subcooling. This increases core power due to the negative void reactivity coefficient. In automatic control, some compensation of core power is realized by modulation of core flow.

~

E!

?

'\ 5.2.1.4 Rod Withdrawal Error Starting Conditions and Assumptions. The reactor is operating at a power level above 75% of rated power at the time the control rod withdrawal error occurs. The reactor operator has followed procedures and, up to the point of the withdrawal error, is in a normal mode of operation (i.e., the control rod pattern, flow setpoints, etc. are all within normal operating lbnits) . For these conditions, it is assumed that the withdrawal error occurs with the maximum worth control rod. Therefore, the maximum positive reactivity inser-tion will occur.

Event Description While operating in the power range in a normal mode of operation, the reactor operator makes a procedural error and withdraws the maximum worth control rod to its fully withdrawn position. Due to the positive reactivity insertion, the core average power will increase. More importantly, the local power in

. 5^g the vicinity of the withdrawn control rod will increase and could potentially d

5-16

NED0-24195 O cause cladding damage due to either overheating which may accompany the occur-

~

rence of boili,ng transition or by exceeding the 1% plastic strain limit imposed on the cladding, which are the assumed transient failure thresholds. The following list depicts the sequence of events for this transient:

1 l

l Approximate Elapsed Time from Start l Event of Rod Motion l (1) Event begins, operator i selects the maximum worth control rod, acknowledges any alarms and withdraws the rod at the maximum rod speed. 0 (2) Core average power and local power increase causing LPRM alarm. < 5 sec (3) Event ends - rod block by APRM rod block syst'em. < 30 see D -

v 5-17

V Results and Consequences The analysis considers the continuous withdrawal of the maximum worth control rod at its maximum drive speed from the reactor, which is operating at rated power with a control rod pattern which results in the core being placed on thermal design limits (i.e., MCPR = Tech Spec Value and a peak LHGR of 13.4 kW/ft for the P8x8R fuel). A worst-case condition is analyzed to ensure that the results obtained are conservative; this approach also serves to demonstrate the function of the APRM rod block system.

The worst-case situation is established for the most reactive reactor state and assumes that no xenon is present. This ensures that the maximum amount of excess reactivity which must be controlled with the movable control rods is present. During a normal startup, sufficient time would be available to achieve some xenon and samarium buildup, and, af ter some short period of operation, samarium will always be present. This assumption makes it possible to obtain a worst-case situation in which the maximum worth control rod is fully inserted;

} and the remaining control rod pattern is selected in such a way as to achieve design thermal limits in the fuel bundles in the vicinity of the maximum worth control rod which is to be withdrawn. It should be noted that this control rod configuration would be highly abnormal and could only be achieved by deliberate operator action or by numerous operator errors during rod pattern manipulation prior to the selection and complete withdrawal of the msximum worth rod. '

E sults for this worst-case condition for the reference cycle are given in the g s

reference cycle supplement. The calculated MLHGR during the RWE is compared to -

-4 that associated with 1% plastic strain in the cladding (Subsection 2.4.1.2) to ensure that fuel damage to any General Electric bundle is not expected during the event.

5.2.1.5 Feedwater Controller Failure - Maximum Demand Identification of Causes. This event is postulated on the basis of a single f'T d

failure of a control device, specifically one which can directly cause an increase in coolant inventory by increasing the feedwater flow. The most 5-18

w NEDO-24195

() severe applicable event is a f eedwater controller failure during maximum flow demand. The feedwater controller is forced to its upper limit at the beginning of the event.

Operating Conditions and Assumptions. The operating conditions and assumptions considered in this analysis are as follows:

(1) Feedwater controller fails during maximum flow demand.

(2) Maximum feedwater pump run out is assumed.

(3) The reactor is operating in a manual flow control mode which provides for the most severe transir.nt.

Event Description. A feedwater controller failure during maximum demand produces the following sequence of events:

(1) The reactor vessel receives an excess of feedwater flow.

(2) This excess flow results in an increase in core subcooling, which results in a core power rise, and in the reactor vessel water level.

(3) The rise in the reactor vessel water level eventually leads to high water level turbine trip and reactor scram trip. ~@

7 Results and Consequences. The influx of excess feedwater flow results in an increase in core subcooling, which reduces the void fraction and thus induces an increase in reactor power. The excess feedwater flow also results in a rise in the reactor water level, which eventually leads to high water level; @

main turbine trip and turbine bypass valves are actuated.

m Reactor scram trip is actuated from main turbine stop valve position switches. C; Relief valves open as steamline pressures reach relief valve setpoints. -

j) 5-19

NEDO-24195 l 1

O LJ 5.2.2 Exposure-Dependent Limits The severity of any plant transient is worst at the end of the cyc7e primarily because the EOC, all-rods-out scram curve gives the worst possible Jcram response. It follows that some limits relief may be obtained by analyzing (2e transients at other interim points in the cycle and administering the resulting limits on an exposure-dependent basis.

This technique is straightforward and consists merely of repeating the transient analyses for selected midcycle exposures. Because the scram reactivity function monotonically deteriorates with exposure (af ter the reactivity peak), the limit determined for an exposure E is administered for all exposures in the interval E

g <E<E where E g is the next lower exposure point for which a limit was determined. This results in a table of MCPR limits to be applied through diff erent exposure intervals of the cycle.

~

O v

The reference cycle supplement will designate exposure-dependent limits j*

as a margin improvement option selected for use at Oyster Creek, and will reflect the~resulting effect'of these option on the plant analysis in )**

Appendix A.

5.2.3 Ef f ect of Fuel Densification on MCPR Operating Limit '$

S 5.2.3.1 Effect on Local Power Associated with densification and reduction in volume of the pellets in a given fuel rod, is the possible occurrence of a gap between two adjacent pellets at some point along that rod. In the immediate region of the gap (a roughly spherical region 1 or 2 inches in radius), there is a slight reduction in power, and a local distortion of the heat flux profile along the affected rod. The heat generated is zero at the gap, but peaks sharply at the pellet faces on O

5-20 a

'+ either side,* as shown in Figure 5-6. These flux " spikes" are only a few per-cent above the average flux for the rod in' the affected region, the amount depending on the size of the gap. For a gap of 1.3 in. (3.30 cm), the maxi-mum calculated gap size for General Electric-supplied fuel, the spike is %5% -'g above the local average on the face of the pellet. _S An idealization of the spike, given in Figure 5-6 for purpose of analysis, is shown in Figure 5-7. Note that the spike height and area are the same as in Figure 5-6. For ease of calculation, the axial flux profile, except for the spike, is created as uniform over the boiling length. The following heat flux profile is assumed at critical power:

q" = q ", Z = 0 to L ~

B Z is distance from begin- (5-1) ning of boiling (inches)

= 1.05 q ", Z = L ~

B B, O

where q" = heat flux (energy rate per unit area), and q" = base heat flux at spike location (i.e. , does not include the incremental additional heat flux due to the spike itself). It is defined by Figure 5-7.

Using the Tong F factor method (Reference 5-10), the equivalent uniform heat flux q " over same boiling length L s ghen by:

B L*

LB

[ -CLh B

B

-C (L B

B

~

q" l-e l= C q "e dZ + C 1.05q1 "e dZ o

j o 1

LB-0.4 (3-2)

_a

??

go" =1+0.05(1-e-0.g

-CL B

91 1-e g *The ' heat generated in the rods adjacent to the af f ectec' rod are also affected,

( 5 being increased the order of 1 or 2% in the immediate region. This slight increase in heat generated in neighboring rods will have negligible effect on the occurrence er nonoccurrence of boiling transition on the affected rod.

5-21

/ where q" = the value of the heat flux, if it were uniform in the axial direction, which would result in boiling transition for the particular boiling length. (In practice, a correlation would be used to determine go".)

The two limiting values of this expression are seen to be:

n q

C - = (Local conditions hypothesis): ,,

= 1.05 (5-3) 41 n

q C + 0 (integral hypothesis: ,, = 1 + 0.05 (0.4/LB) = 1.0004 (5-4)

It was shown in Reference 5-1 that the integral hypothesis is applicable to BWR operating conditions, and, thus, Equation 5-4 is applicable. Note that q "/q " is' largest when L is 3 a minimum. The minimum value of L B # """ "

to us in the reactor situation is on the order of 50 inches. Tong suggests a R value of C = 0.135/D. For D = 0.5 in. (typical for BWR's), C = 0.27 in." . S This gives:

n q

= 1.00512. (5-5)

,'l The critical power without the spike is q " L . The critical power with the B

spike is q" L ' " *#*

B E #*

~

* *#*##E" *** "** ***#*E" "*

respect to distance.

z l

1 f q" =

,, q" dz (5-6)

,1 o ,

o e

Solving this equation for the flux profile in Figure 5-7 gives:

0 q" = q"1 1 + 0.05 =

1.0004 q"7 (5-7) o 5-22

The relationship of T[" to a more general type of profile is shown in Figure 5-8.

Per the integral hypothesis (Equation 5-4) q" = 1.0004 q"7, and q" = q".

Thus, the two critical powers are equal. There is no reduction in critical power due to the presence of the spike, according to the integral hypothesis.

1 Using the value of C suggested by Tong, q"y = q". The critical power O 1.00512 j v with the spike in this case is: '

i q'L B "

1. 12 S B" ' S B

i i

The critical power is reduced only 0.47% due to the presence of the spike, according to the Tong F-factor method. , ,

Recourse can be made to data to resolve the question concerning the effect of a flux spike on critical power. There are very few good flux spike data, and ,

nearly all that exist are for flux spikes at the exit end. Exceptions to this are some data reported by B&W personnel several years ago (Reference 5-11), i using an 0.4 in. I.D. x 6 f t long test section with a symmetrical cosine profile (their TS No. 63), and a second test section with the same dimensions and symmetrical cosine profile; but in addition, a flux spike 3/4 of the way '

along the length of the test section (TS No. 76). The profiles are shown in ;

Figures 5-9 and 5-10. Boiling transition occurred a few inches past the ppike at the lower subcoolings, but close to and on the spike at the higher subcoolings. f i

The critical quality versus boiling length for both these test sections is shown f in Figure 5-11 and was derived as follows.

!O l. The axial profiles for Test Section Nos. 65 and 76 (Figures 3 and 4 l

of Reference 5-11) were put in the normalized forms: R i I 1/2 !

TS65 $ = 0.124 + 1.115 4 -4 -

(5-9) !

I

-=

o ,

= 1.239 (5-10) 5-23 l

l l

-- -- , . - . - - . . - . - n - , , - , - - , . . , - - . ,

NEDO-24195 TS76 9 = 1.40 sin 1,j45(0.0725+hf (5-11)

J

$""* = 1.40 (5-12)

av 1

Spike at 53.5 <_ z <_ 54.5 in.

O +*** = 1.40

av ~$

._ __ _2

2. The Reference 5-11 text gives the test condition P = 2000 psia.

Table 1 of this reference gives the inside diameter, D, from which !

flow area A can be calculated and heated length, L, for both test sections. Table 3 of the reference gives X (local quality at loca-tion of ooiling transition), z, (location of boiling transition, measured from beginning of heated length), Ti (inlet temperatur.e from .

which inlet enthalpy h 1and inlet subcooling as can be determined),

G (mass flux), and Q/ATS a (fr m which total power can be determined).

Note that there are sometimes two values of x and z, meaning that BT occurred at two locations simultaneously. Table 3 of the reference also gives Q/ , local heat flux; Q/A,y, average heat flux up to location of BT of the reference, X,, exit quality. These were not used in the preparation of Figure 5-11.

3. The location.of z , the position at which equilibrium balk boiling begins, is calculated from:

o

=

Ah, GA (Q/qg ,y) $ dz (5-13) '

o

4. Boiling Length = L = (z-z )

3 O

N 5-24

NEDO ~.4195 f

It is clear that all the TS No. 76 data correlate well with the TS No. 65 data :

in the X versus LB Plane. The critical quality-boiling length concept is

{

identically the integral hypothesis (Equation 5-4). See rege 3-3 of Refer-

~

ence 5-1 for a discussion of this. It has already been shown in accordance k [

t with the integral hypothesis that the critical power will not be affected by i the presence of the spike. Thus, the B&W data show that the critical power is not affected by the presence of a spike, in particular a spike which is !

considerably more severe than the spike associated with densification (compare Figure 5-10 with Figure 5-6). .

The analysis in accordance with the integral concept (critical quality versus boiling length arrelation) shows no effect of a densifications-type flux spike on critical power.

The data show that, even for a more severe spike, the integral concept holds and there is no effect of' spike on crirical power. The effect on the R-factor ,

is given below. ,

5.2.3.2 Effect of Densification on the R-Factor [

i The R-factor for a given bundle is a parameter those evaluation assumes prior !

knowledge s

of the individual rod powers in that bundle. Densification can affect ;

i individual rod. powers, but prior knowledge of'which rods are affected, gap :

loc.ation, eU:., does not exist. Hence,-the R-factor, as calculated for critical

power predicrion, is not changed by densification. However, we will attempt

.to answer a slightly different question, viz. "What would be the change in {

P.-factor if densification effects could be taken into account, and, hence, whd[inuld'betheeffectonpredictedbundlecriticaipower?" '

O l

Fuel densification can cause gaps in the fuel column which result in local power epikes in the adjacent fuel pellets and in the surrounding fuel rods.

She effect of these power spikes.on. critical power of a single rod was evaluated

,fa the preceding section. It was showa that there is no reduction in the crit-ical power due to the presence of the power spikes according to the integral l

O' .

$ E-w 5-25 !

g.

\, .te b l_+

NEDO-24195 hypothesis. Even using the Tong F-factor method, the critical power is reduced only slightly (<0.5%) due to the present of a power spike at the end of the boiling length.

In the GETM method, using the GEIL correlation (Reference 5-1), the effect of variation in rod-to-rod power on bundle critical power is considered in the R-fac tor. This factor accounts for the details of flow and enthalpy distribu-tion in the bundle.

O The maximum fuel densification gap size expected is 1.3 inches. The rod with a gap will experience a loss in active fuel length (equal to the gap size) and a corresponding increase in the average fuel density in the shorter fueled region.

It will be assumed that a 1% increase in density results in a 0.7% increase in linear heat generation rate and that the gap occurs at a point where the local heat flux is equal to the average heat flux for that rod. Neglecting the effect of local power spikes at the pellet ends at the gap, the rod with the maximum gap will experience a rod power reduction of 0.276% (an extremely unlikely case, since all surrounding rods must have zero densification). Conservatively assuming that the maximum gap case represents a 3-sigma value, the 1-sigma uncertainty in average rod power due to a gap is less than 0.1%.

A fuel rod without a gap but surrounded by gapped rods will experience local power increases. M auming that these power spikes are additions, and all the rods sur-rounding the center rod of a P8x8R bundle have a maximum gap (which is extremely unlikely), the average power in the center (ungapped) rod will increase by 0.305%.

Conservatively assuming that the maximum gap case represent a 3-sigma value, the 1-sigma uncertainty on rod power is <0.1%.

Saveral parameters which contribute to local rod-to-rod power uncertainty are statistically considered,to determine the uncertain y in the R-factor (Section IV 3.4.7 of Reference 5-1). The 1-sigma uncertainties of these parameters range from 0.2% to 2.6% for an overall rod power (1 sigma) uncertainty of 2.868%.

The parameter with a 0.2% l-sigma uncertainty on rod power contributes only 0.014% to the overall rod power uncertainty. The effect of densification and power spikes is seen to be even smaller and is not considered important to the production of rod power and R-factor. (The uncertainty in the R-factor is con-s:rvatively derived in Reference 5-1 as o g

= 0.5o,, where o, is the uncertainty

'in rod average power.) -

5-26

NED0-24195 In summary of the effect of densification en rod power, the uncertainty on rod 4

power to densification increases the overall rod power uncertainty by less than 0.014%. Therefore, densification has negligible effect on R-factor and on predicted bundle critical power, and is currently not considered in the application of GEXL to critical power predictions.

5.3 VESSEL PRESSURE ASME CODE COMPLIANCE MODEL The pressure relief system was designed to prevent excessive overpressurization of the p.imary system process barrier and the pressure vessel and thereby pre-i cludes an uncontrolled release of fission products.

The design capacities of the safety valves for Oyster Creek were determined according to the requirements of Section I, Power Boilers, of the ASME Boiler and Pressure Vessel Code. Under the provisions of this code and Special Code Case 1270N, safety valve capacities were established to prevent a vessel ~

g pressure rise greater than 10% above the design pressure. At least one D

, h safety valve was to be set at or below the vessel design pressure; the highest l

safety valve setting could not exceed 3% of the vessel design pressure. No

credit was allowed for reactor scram as a complimentary pressure protection w device. Thus, the required safety valve capacities were sized assuming essen-tially instantaneous isolation of the pressure vessel with no pressure relief other than that from the safety valves. Results of the overpressurization analysis for the reference cycle are given in the supplemental submittal. ~S D

5.4 STABILITY ANALYSIS METHOD To demonstrate analytically no divergent oscillation or limit cycle oscilla-tion occurs in the system, a stability criterion, referred to as the " Acceptable Performance Limit", is established for General Electric-supplied fuel. This limit corresponds exactly to the analytical threshold of instability. The stability criterion is stated in terms of two compatible parameters representing both the frequency- and time-domain: damping coefficient (( ) and decay ratio es

~,-

(x,/x ).

~ 0 As stated in'Section 5, the definition of the damping coefficient E, t, corresponds to the pole pair closest to the ju in the s plane for the system 5-27

NEDO-24195 closed-loop transfer function. This parameter also applies to the frequency- !

domain. The dscay ratio x2 /*0 s defined as the ratio of the magnitude of the 1] second overshoot to the first overshoot resulting from a step perturbation.

This is a graphic representation of the physical responsiveness of the system' which is readily evaluated in a time domain analysis. There is a direct relationship between the decay ratio and the damping coefficient for any dominant mode response (Figure 5-12).

For reload cores, two types of stability are examined utilizing a linearized analytical model. First is the hydrodynamic channel stability of one or morc types of channels operating in parallel with other channels in the core. This i

is considered because flow esci11ations may impede heat transfer to the mod-l erator and drive the reactor into power oscillations. Second is the reactivity feedback stability of the entire reactor core, which also involves power oscil-lations. Total system dynamics are comprised of the dynamics of the control system comoined with those of the basic process and is, therefore, not reanalyzed for reload cores.

'm The assurance that the total plant is stable and, therefore, has significant O design margin is demonstrated analytically when the acceptable performance limit El e

~~

of a decay ratio less than 1.0 or a damping coefficient greater than 0.0 is met for each type of stability discussed. These criteria are established at the threshold of sustained oscillatory response.

These criteria must be satisfied for both usual and unusual operating conditions of the reactor that may be encountered in the course of BWR plant operation.

Therefore, stability analyses are performed at the natural circulation intercept

, of the rod line corresponding to the highest power level for which safety analyses are performed as documented in plant technical specifications. i l

The mathematical model representing the core evamines the linearized reactivity response of a reactor system with density-dependent reactivity feedback caused by boiling. The hydrodynamics of various hydraulically coupled reactor channels or regians are examined separately on an axially multinoded basis by grouping various channels that are thermodynamica11y and/or hydraulically similar. This O interchannel hydrodynamic interaction or coupling exists through pressure i

l 5-28 4

NEDO-24195 variations in the inlet plenum, such as can be caused by a disturbance in the d flow distribution between regions or channels. This approach provides a i

reasonably accurate, three-dimensional representation of the reactor's hydro-dynamics. Details of the model are given in Reference 5-12.

Also given in Reference 5-12 is a significant base of experimental confirmation which verifies the suitability of the analytical models used in both the linear-j ized frequency domain and the nonlinear time domain. Sensitivity of the transient behavior to key system parameter variations is established and adequate design margins are applied to accommodate such variations. The method and models combined with the established design limits are considered to provide a sufficiently conservative analytical result.

The analysis is performed at the most limiting condition, which usually occurs near the end of cycle, with power peaking toward the bortom of the core.

Because of the decrease in delayed neutron fraction, the value of the density reactivity coefficient, Ak/kS,gg/ao, increases. The most sensitive reactor p operating condition is that corresponding to natural circulation flow intercept of the rod line corresponding to the highest power level, as documented in plant technical specifications, for which safety analyses are performed. A o e

specific analysis for the eference cycle will be provided in the supplemental D submittal.

5.5 ACCIDENT EVALUATION METHODOLOGY As stated in Subsection 5.2, abnormal operating transients are evaluated to determine the normal operating MCPR limit for General Electric-supplied fuel.

In addition to these analyses, evaluations of less frequent postulated events are made for this fuel to assure an even greater depth of safety. Accidents are events which have a projected frequency of occurrence of less than once in every one hundred years for every operating BWR. The broad spectrum of postulated accidents is covered by six categories of design basis events. These events are the control rod drop, main steamline break,-loss of coolant, refuel-

i. ing, recirculation pump seizure, and fuel assembly loading accidents. A description of each of these events follows.

5-29

NEDO-24195 ., = ,

O I

.R 5.5.2 Loss-of-Coolant Accident -.

This analysis of the Oyster Creek Nuclear Generating Station loss-of-coolant accident (LOCA) is provided to demonstrate conformance with the ECCS accept-ance criteria of 10CYR50.46. The objective of the LOCA analysis contained herein is to provide assurance that the most limiting break size, break loca- i tion and single failure combination has been considered for the plant. The documentation contained in this section is intended to satisfy these require-ments. I' I

The general description of the General Electric (GE) LOCA evaluation models is '

contained in Reference 5-18. Applicability and approval for pre-pressurized reload fuel are given in Reference 5-25. Model changes are described in References 5-20 and 5-21, which were recently approved by the USNRC (Refer-ence 5-19). These model changes are employed in the new CHASTE computer code t

(

(CHASTE-05) which has been used in this analysis. !

l 5.5.2.1 Input to Analysis t O !

A list of the significant input parameters to the LOCA analysis is presented in Table 5-10.

5.5.2.2 !

SAFE i

This code is used primarily to track the vessel inventory and to model ECCS performance during the LOCA. The application of SAFE is identical for all ;

break sizes. This code predicts the entire course of the postulated accident.

SAFE calculates reactor system pressura, ECCS flows (which are pressure !

dependent), and hot fuel node uncovery time. (Note: To ' simplify the analyt-ical procedure, no reflooding times are calculated. All peak cladding tem-peratures are turned over by action of core spray heat transfer only.)

O 5-39 t _

~

The SAFE results presented are: ,

O (1) Water Level inside the Shroud (2) Reactor Vessel Pressure ;

l 5.5.2.3 CHASTE t This code is used, with inputs from the SAFE code, to calculate the fuel cladding heatup rate, peak cladding temperature, peak local cladding oxidation,

{

and core-wide metal-water reaction. The fuel model in CHASTE considers tran-sient gap conductance, clad swelling and rupture, and metal-water reaction. '

The duration of nucleate boiling following the recirculation line break is calet:. lated by CHASTE using the GE dryout correlation, which is based on no-flow conditions. This correlation relies on experimental data from many different test assemblies with various axial power shapes and different geomet-ries whose purpose was.to investigate the no-flow dryout phenomenon. The Ellion pool boiling heat transfer correlation is used from the time that the decaying heat transfer from the dryout correlation ceases to apply until the time that the hot fuel core node is uncovered.

The empirical dryout, core spray heat transfer, channel wetting, and rod rewetting correlations are contained in CHASTE, which solves the tratsient heat transfer equations for the entire LOCA at a single axial plane in a single ruel assembly. Iterative applications of CHASTE determine the maximum parmissible planar power to satisfy the 10CFR50.46 acceptance criteria.

The CHASTE results presented are:

(1) peak cladding temperature versus time; (2) heat transfer coefficient versus time; (3) normalized power versus time; and -

(4) peak cladding temperature versus break area.

k*

O -

5-40

9 NEDO-24195 R 2

The CHASTE-05 version of the code is used in thir analyses. This version contains the improved radiation and conduction hea*: transfer models approved by t'ae USNRC in the first quarter of 1977.

A summary of the analytical results is given in Table 5-11. Table 5-12 lists the figures provided for this analysis. The peak cladding temperature e vs. break area curve of Figure 5-22 is based on previous break spectrum analysis.

h 5.5.2.4 Single Failures and Breaks Table 5-13 gives the single failure combinations considered in the LOCA analysis.

A spectrum of break sizes was analyzed for the recirculation line break to determine the worst combination of break size and single failure mode. The k*

break spectrum and analytical results are given in Reference 5.. Full-sized breaks of t'he feedwater, steamline and core spray line were also analyzed.

2 ~m O,(/) Of all the break sizes and locations which were investigated, the 0.13 ft O os recirculation line break was determined to be the most limiting case because this break had the highest peak cladding temperature. The Lime between 2

uncovery and rated spray for the 0.13 ft break is longer than the time for all other breaks, except for the 0.12, 0.11, and 0.10 ft break:; . For these breaks, the uncovery times are 9, 20 and 33 seconds later, respectively, and the times of rated core spray are 13, 30, and 50 seconds later, respectively. Although the maximum time of all the breaks (0.10 to 0.12 ft ) between uncovery and

(

9 rated spray is about 15 seconds longer than for the 0.13 ft" break, the decay heat. generation is also lower. The effect of the lower decay heat outweighs the longer time between uncovery and rated spray, producing a lower peak clad temperature for the 0.12, 0.11, and 0.10 ft breaks.

U 5-41

-' R

_ 3:

() 5.5.2.5 Conclusions ,

The MAPLHGR's given in the reference cycle submittal represent the operating o

MAPLHGR limits determined for Oyster Creek, based on currently USNRC approved

{

GE Evaluation Models. The maximum core-wide metal-water reaction is also given .

for all break locations and sizes considered.

5.5.3 Main Steamline Break Accident Analysis The analysis of the main steamline break accident depends on the operating thermal-hydraulic parameters of the overall reactor (such as pressure) and over-all factors affecting the radiological consequences (such as primary coolant activity). Insertion of reload fuel will not change the radiological conse-

~

quences of this event. Analytical results supporting this are presented in o Reference 5-26. ,

27 5.5.4 Loading Error Accident Calculational Methods lfv t .

One of the events which has been evaluated in BWR safety analysis reports is fuel bundle loading error. A loading error in the core configuration is defined as: (1) a General Electric fuel bundle in an improper location (mislocation),

or (2) a General Electric fuel bundle in an improper orientation (i.e. , mis-oriented - rotated 90* or 180*).

Proper orientation of fuel assemblies in the reactor core is readily verified by visual observation and ossured by verification procedures during core loading.

Five separate visual indications of proper fuel assembly orientation exist:

(1) The channel fastener assemblies, including the spring and guard used to maintain clearances between channels, are located at one corner l of each fuel gasembly adjacent to the center of the control rod.

(2) The identification boss on the fuel assembly handle points toward l

the adjacent control rod.

O O (3) The channel spacing buttons are adjacent to the control rod passage area.

5-42

NED0-24195

[{

_m i

(4) The assembly identification numbers which are located on the fuel

)

C-</

s. assembly handles are all readable from the direction of the center of the cell.

(5)' There is cell-to-cell i u1 cation.

Experience has demonstrated tF-* these design features are clearly visible so that any misloaded fuel assuably would be readily identifiable dur,ing core loading verification.

Proper location of the fuel assembly in the reactor core is readily verified by visual observation and assured by verification procedures during core loading.

Verification procedures include inventory checks, current bundle location logs, serial number verifications, and visual or photographic inspection of the loaded core. The verification procedures are designed to minimize the possibility of the occurrence of the fuel assembly loading error.

Based on the probability assessments given in Reference 5-23, it is seen that the iO probability of a significant fuel assembly loading error is much less than once in a plant lifetime. Additionally, it requires multiple operator errors. Thus, the fuel assembly loading error is classified as an accident, not a transient, so application of LHdR 1Laits is not appropriate.

Analytical Procedures Analysis methods for the misoriented and mislocated fuel assembly are discussed in detail in Reference 5-23. Approval of these methods is-given in Reference 5-24 under the stipulation that a ACPR penalty of 0.02 be added for the tilted mis-

_o oriented bundle. This 0.02 is added on to the calculated ACPR used in determining

((

the operating limit when utilizing this method. General Electric applies the '

fuel cladding integrity safety limit documented in Subsection 5.1 to the accident results. These results are shown in the reference cycle supplement. ,

e -

The utility may elect to substitute an alternative approach such as the

}{

one noted in Reference 5-24 for this limit.

(<)

(.

i 5-43

NEDO-24195 --e O

_ . _e_

5.5.5 One Recirculation Pump Seizure Accident Analysis O l Thid accident is assumed to occur as a consequence of an unspecified, instantaneous stoppage of one recirculation pump shaft while the reactor is operating at full power.

The pump seizure event is a very mild accident in relation to other accidents such as the LOCA. This is easily verified by consideration of the two events. '

'O t deta ecia =c ca c1=c=1 c1== 4:1 1== 1 > 11e- 1 1e = tr rapidly - in the case of the seizure, stoppage of the pump occurs; for the 17

~

s l LOCA, the severance of the line has a similar, but more rapid and severe influ- ;

^

ence. Following a pump seizure event, flow continues, water level is main-tained, the core remains submerged, and this provides a continuous core cooling mechanism. However, for the LOCA, complete flow stoppage occurs and the water ;

level decreases due to loss of coolant resulting in uncovery of the reactor core and subsequent overheating of the fuel rod cladding. In addition, for the pump seizure accident, reactor pressure does not significantly decrease, whereas l

l complete depressurization occurs for the LOCA. Clearly, the increased tempera- '

ture of the cladding and reduced reactor pressure for the LOCA both combine to yield a much more severe stress and potential for cladding perforation for !

the LOCA than for the pump seizure. Therefore, it can be concluded that the I potential effects of the hypothetical pump seizure accident are very conserva-tively bounded by the effects of a LOCA and specific analyses of the pump seizure accident are not required. .

5.5.6 Refueling Accident Analysis l

5.5.6.1 Identification of Causes O

Accidents that result in the release of radioactive materials directly to the :

containment can occur when the drywell is open. A survey of the various condi-tions that could exist when the drywell is open reveals that the greatest i potential for the release of radioactive material occurs when the drywell head and reactor vessel head have been removed. In this case, radioactive material I l released as a result of fuel failure is available for transport directly to the containment.

5-44

l NEDO-24195 'R Various mechanisms for fuel failure under this condition have been investigated.

With the current fuel design, the refueling interlocks, which impose restirctions on the movement of refueling equipment and control rods, prevent an inadvertent criticality during refueling operations. In addition, the reactor protection system can initiate a reactor scram in time to prevent fuel damage for errors or malfunctions occurring during planned criticality tests with the reactor vessel head off. It is concluded that the only accident that could result in the releaae of significant quantities of fission products to the containment O a=r1== **1 =eae er ever cie 1 e re 1=1== <re the cciae=t 1 are>>1== er a fuel bundle onto the top of the core.

5.5.6.2 Effect of Fuel Densification This event occurs under nonoperating conditions for the fuel. The key assump-tion of this postulated occurrence is the inadvertent mechanical damage to the fuel rod cladding as a consequence of the fuel bundle being dropped on the core while in the cold condition.

Fuel densification considerations do not enter into or affect the accident results. l 5.5.6.3 Methods, Assumptions, and Conditions l

The assumptions and analyses applicable to this type of fuel handling accident are described below. '

(1) The fuel assembly is dropped from the maximum height allowed by the ,

fuel handling equipment.

O t (2) The entire amount of potential energy (referenced to the top of the i

reactor core) is available for application to the fuel assemblies ;

j involved in the accident. This assumption neglects the dissipation of some of the mechanical energy of the falling fuel assembly in the water above the core and requires the complete detachment of the assembly from the fuel hoisting equipment. This is only possible if the fuel assembly handle, the fuel grapple, or the grapple cable breaks, or improper grapplings occur.

5-45

~

NEDO-24195- S (3) None of the energy associated with the dropped fuel assembly is absorbed by the fuel material (uranium dioxide).

5.5.6.4 Results and Consequences P

Fuel Damage Dropping a fuel assembly onto the reactor core from the maximum height allowed by the refueling equipment (<30 f t) results in an impact velocity of 40 ft/sec.

(J")

The kinetic energy acquired by the falling fuel assembly is less than -

e, 17,000 ft-lb and is dissipated in one or more impacts.

f[

The first impact is expected to dissipate most of the energy and cause the largest number of cladding failures. To estimate the expected number of failed i fuel rods in each impact, an energy approach is used, i

The fuel assembly is expected to impact on the reactor core at a small angle from the vertical, possibly inducing a bending mode of f ailure on the fuel rods

/"' m U) of the dropped assembly. It is assumed that each fuel rod resists the imposed C; bending load by a couple consisting of two equal, opposite concentrated forces. ~ -

Therefore, fuel' rods are eryected to absorb little energy prior to failure as a result of bending. Actual bending tests with concentrated point-loads show that each fuel rod absorbs approximately 1 ft-lb prior to cladding failure.

Each rod that fails as a result of gross compression distortion is expected to absorb approximately 250 f t-lb before cladding failure (based on 1% uni-form plastic deformation of the reds). The energy of the dropped assembly is conservatively assumed to be absonbed by only the cladding and other core structures.

\(~m

,)

Because a fuel assembly consists of 72% fuel, 11% cladding, and 17% other struc- !

tural material by weight, the assumption that no energy is absorbed by the fuel material results in considerable conservatism in the mass-energy calculations

' hat follow.

6 5-46

NEDO-24195 l

The energy absorption and successive impacts is estimated by considering a plastic impact. Conservation of momentum under a plastic impact shows that the fractional kinetic energy absorbed during impact is:

M,

- M +M where M is the impacting mass and M is the struck mass. Based on the fuel geometry in the reactor core, four fuel assemblies are struck by the impacting assembly. The fractional energy loss on the first impact is approximately 80%.

The second impact is expected to be less direct. The broad side of the dropped assembly impacts approximately 24 more fuel assemblies, so that af ter the second impact only 136 f t-lb (approximately 1% of the original kinetic energy) is available for a third impact. Because a single fuel rod is capable of absorb-ing 250 f t-lb in compression before cladding failure, it is unlikely that any fuel rod will fail on a third impact.

[}

If the dropped fuel assembly strikes only one or two fuel assemblies on the first impact, the energy absorption by the core support structure results in approxi-mately the same energy dissipation on the first impact as in the case where four fuel assemblies are struck. The energy relations on the second and third impacts remain approximately the same as in the original case. Thus, the calculated energy dissipation is as follows:

First impact 80%

Second impact 19%

Third impact 1% (no cladding failures)

The first impact dissipates 0.80 x 17,000 or 13,600 f t-lb of energy. It is assumed that 50% of this energy is absorbed by the dropped fuel assembly and that the remaining 50% is absorbed by the struck fuel assemblies in the core.

Because the fuel rods of the dropped fuel assembly are susceptible to the bending mode of failure and because 1 f t-lb of energy is sufficient to cause L

, 5-47

< O V cladding failure as a result of bending, all 62 rods of the dropped fuel assembly are assumed to fail. Since the tie rods of the struck fuel assemblies

) are more susceptible to bending failure than the other 54 fuel rods, it is assumed that they fail on the first impact. Thus, 4 x 8 = 32 tie rods (total in 4 assemblies) are assumed to fail.

Because the remaining fuel rods of the struck assemblies are held rigidly in place in the core, they are susceptible only to the compression mode of failure. To cause cladding failure of one fuel rod as a result of compression, 250 ft-lb of energy is required. To cause failure of all the remaining rods of the 4 struck assemblies. 250 x 56 x 4 or 56,000 f t-lb of energy would have to be absorbed ir; cladding alone. Thus, it is clear that not all the remain-ing fuel rods of the struck assemblies can fail on the first impact. The number of fuel rod failures caused by compression is computed as follows:

l' O.5 x 13,600 x 11 g 250 O

(/

Thus, during the first impact, fuel rod failures are as follows:

Dropped assembly 62 rods (bending)

Struck assemblies 32 tie rods (bending)

Struck assemblies 11 rods (compression) 105 failed rods Because of the less severe nature of the second impact and the distorted shape of the dropped fuel assembly, it is assumed that in only 2 of the 24 struck assemblies are the tie rods subjected to bending f ailure. Thus, 2 x 8 = 16 tie rods are assumed to fail. The number of fuel rod failures caused by compres - l sion on the second impact is computed as follows: !

0 x 17,000 x 11 g

= 3 250 fi a

l 5-48 l

l

l NEDO-24195 Thus, during the second impact, the fuel rod failures are as follows:

f--

Struck assemblies 16 tie rods (bending)

Struck assemblies 3 rods (compression) 19 failed rods The total number of failed rods resulting from the accident is as follov ::

First impact 105 rods

.m Second impact 19 rods l

-e

!! l Third impact 0 rods \

i 124 total failed rods 5.5.6.5 Radiological Consequences .

~

il Based on the linear heat generation rate applicable to P8x8R fuel, it can bt theoretically predicted that the fractional plenum activity will be approxi-mately one-tenth of that activity contained in the plenum of a 7x7 fuel rod.

For the purpose of this evaluation, it is conservatively assumed that the fraction plenum activity for the P8 x 8R rod is the same as for the 7x7 rod -

Since each P8x8R fuel bundle produces the same power as a 7x7 bundle, th e . $

tverage activity per rod for the P8 x 8R bundle will be 49/62, or 0.79, tin ts ;

the activity in a 7x7 rod. Based on the assumption that 124 P8 x 8R rods fa!.1 L compared to 111 for a 7x7 core, the relative amount of activity released for -

the P8x8R fuel is (L24/111) (0.79) = 0.88 ttnes the activity released for a (( ;

. *1 7x7 core. The activity released to the environment and the radiological expo- !

sures for the P8 x 8R fuel will, therefore, be less than 96% of those values presented in the FSAR for a 7x7 core. As identified in the FSAR, the radiologi-l I

caa axposures for the 7x7 fuel are well below those guidelines set forth on 10CFR100; therefore, it can be concluded that the consequences of this accident for the P8 x 8R fuel will also be well below these guidelines.

l l

l 1

l 5-49 l

NEDO-24195 g,

\

5.6 REFERENCES

5-1 General Electric BWR Thermal Analysis Basis (GETAB): Data, Correlation and Design Application, January 1977 (NEDC-10958-PA and NEDO-10958-A).

52 Basis for 828 Retrofit Fuel Thermal Analysis Application, September 1978 (NEDE-24131).

5-3 R. B. Linford, Analytical Methods of Plant Transient Evaluations for the General Electric Boiling Water Reactor, February 1973 (NED0-10802).

5-4 R. B. Linford, Analytical Methods of Plant Transient Evaluations for the CE BVR Amend ent Nd. 2, June 1975 (NED0-10802-01).

5-5 R. B. Linford, Analytical Methods of Plant Transient Evaluations for the CE BWR Amendnent No. 2, June 1975 (NED0-10802-02) .

5-6 Generation of Void mui Doppler Reactivity Feedback for Application to BWR Design, December 1975 (NEDO-20964).

5-7 J. A. Woolley, "Three Dimensional BWR Core Simulator," January 1977 (NEDO-20953A).

5-8 Analytical Model for Loss-of-Coolant Analyses in Accordance with 10CFR50

(]

V Appendix K, January 1976 (NEDE-20566-P and NED0-20566) .

5-9 R. L. Bolger, Commonwealth Edison Company, letter to E. G. Case, Deputy Director, Office of Nuclear Reactor Regulation, USNRC, Subj ect: Dresden Station Unit No. 2 Proposed Amendment to Facility Operating License No. DPR-19 to Permit Power Coastdown from 70% Power to 40% Power, NRC Docket No. 50-237, dated June 6, 1977.

5-10 0. Glenn Smith, W. M. Rohren, Jr., and L. S. Tong, " Burnout in Steam-Water Flows with Axially Nonuniform Heat Flux," ASME Paper 65-WA/HT-33, November 1965.

5-11 H. S. Swanson, J. R. Carver, and C. R. Kakaral, "The Influence of Axial Heat Flux Distribution on the Departure from Nucleate Boiling in a Water-Cooled Tube," ASME Paper 62-WA-297, November 1962.

5-12 " Stability and Dynamic Performance of the General Electric Boiling Water Reactor," NED0-21506, January 1977.

5-13 C. J. Paone, Banked Position Withdrawal Sequence, January 1977, (NEDO-21231) .

5-14 C. J. Paone and J. A. Woo 11ey, Rod Drop Accident Analysis for Large Boiling Vater Reactors, Licensing Topical Report, March 1972 (NEDO-10527).

k.

5-50

NED0-24195 i \

\ss 5-15 R. C. Stirn, C. J. Paone, and R. M. Young, Rod Drop Accident Analysis for Large Boiling Vater Reactors, Licensing Topical Report, July 1972 (NED0-10527, Supplement 1).

5-16 R. C. Stirn, C. J. Paone, and J. M. Haun, Rod Drop Accident Analysis for Large Boiling Water Reactors, Addendwn No. 2, Exposure Cores, Licensing Topical Report, January 1973 (NEDO-10527, Supplement 2).

5-17 Fuel Densification Effects on General Electric Boiling Water Reactor Fuel, August 1973, Supplement 6 to NEDM-10735.

5-18 Analytical Model for Loss-of-Coolant Analysis in Accordance with 10CFREO Appendi K, January 1976 (NEDE-20566-P and NED0-20566) .

5-19 Letter, K. R. Coller (NRC) to G. G. Sherwood (GE), " Safety Evaluation for General Electric ECCS Evaluation Model Modifications", dated April 12, 1977.

5-20 Letter, A. J. Levine (GE) to D. F. Ross (NRC) dated January 27, 1977,

" General Electric (GE) Loss-of-Coolant Accident (LOCA) Analysis Model Revisions - Core Heat Code CHASTE 05".

5-21 Letter, A. J. Levine (GE) to D. B. Vassallo (NRC), dated March 14, 1977,

" Request for Approval for Use of Loss-of-Coolant Accident (LOCA) Evalu-ation Model Code REFLOOD05".

l 5-22 Letter, N. G. Trikouros (GPU) to J. F. Kilty (GE), dated November 13, 1975, -

R No. S&L-3305, "0yster Creek Nuclear Generating Station Revised Single Failure LOCA Analysis". -

5-23 Letter, R. E. Engel to D. G. Eisenhut (NRC), " Fuel Assembly Loading Error",

November 30, 1977.

5-24 Letter, D. G. Eisenhut (NRC) to R. E. Engel (GE), MFN-200-78, May 8, 1978.

5-25 " General Electric Reload Fuel Application", NEDE-24011-P*.

5-26 Attachment to Letter, S. Bartnoff to A. Giambusso, "0yster Creek Nuclear e Generating Station Docket Number 50-219 Loss-of-Coolant Accident 2 Analysis Re-evaluation Additional Information," April 28, 1975.

g

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Reference refers to the revision of NEDE-24011-P-A which is approved by the NRC as of the date the specific analysis is initiated.

5-51 I

() Table 5-1 UNCERTAINTIES USED IN STATISTICAL ANALYSIS Standard Deviation Quantity (% of Point) Comment Feed %2ter 1.76 This is the largest component of total reactor Flow power uncertainty.

Feedwater 0 76 These are the other significant parameters in Temperature ,

ccre power determination, d

Reactor 0.5 Pressure Core Inlet 0.2 Affect quality and boiling length.

Temperature Core Total 5.0 This uncertainty is for non-jet pump plants and Flow includes the allowance due to bypass flow uncertainty.

Channel Flow 3.0 This accounts for manufacturing and service Area induced variations in the free flow area within the channel.

p\.)

Friction 10.0 Accounts for uncertainty in the correlation Factor representing two-phase pressure losses.

Multiplier Ch tel 5.0 Represents variation in the pressure loss char-Fr .-ion acteristics of individual channels. Flow area Fat

and pressure loss variations affect the core Mul. lier flow distribution, influencing the quality and boiling length in individual channels.

TIP Readings 8.7 These sets of data are the base from which gross power distribution is determined. The assigned uncertainties include all electrical and geometrical compenents plus a contribution from the analytical extrapolation from the chamber location to the adjacent fuel assembly segment. Also included are uncertainties con-

, tributed by the LPRM system. LPRM readings are used to correct the power distribution calcula-tions for changes which have occurred since the last TIP survey. The assigned uncertainty af fects power distribution in the same manner as the base TIP reading uncertainty.

/) R Factor 1.6 This is a function of the uncertainty in local

' '\- ' fuel rod power.

Critical 3.6 Uncertainty in the GEXL correlation in terms Power of critical power.

5-52

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NOMINAL VALUES OF STATISTICAL ANALYSIS PARAMETERS

! Core Thermal Power 3293 MW i Core Flow 102.5 M1b/hr .

i !-

l j Core Bypass Flow Rate 10.25 M1b/hr t .

Core Inlet Temperature 527.68'F

! Dome Prescure 1010.4 psig l R-factor 1.038 (P8x8R) ;

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nus D O Table 5-3 OYSTER CREEK PRESSURE RELIEF SYSTEMS SV RV or S/RV Number Number Lowest Opening of of Number of Set Capacity Lowest Capacity Time Time Safety Relief Safety / Relief Point at Setpoint Setpoint at Setpoint Delay Constant Valves Valves Valves (psig) (No./%) (psig) (No./%) (msec) (msec) 16 5 --

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Table 5-4 ,

l CONSERVATISM FACTORS 2

! Power Increase Power Decrease Nuclear Parameter Transients Transients 1

Doppler Reactivity 0.95 1.05 Void Reactivity 1.25 0.95 Scram Reactivity 0.80 Not Applicable I

The factors are multipliers to the nominal point mcdel nuclear data.

Examples of power increase transients are load rejection without bypass, turbine trip without bypass and feedwater controller failure.

Potier Decrease Transients include events such as recirculation flow con-troller failure (decreasing flow), ene recirculation pump trip, and g five recirculation pump trips.

Q Table 5-5 LICENSING BASIS VALUES FOR TRANSIENT OPERATING PARAMETERS - OYSTER CREEK Rated Steam Thermal Power Dome Pressure Turbine Pressure Flow x 106 Core Flow x 10 (Mut 20.2%) (psig 2 psi) (psig !2 psi) (1b/hr 0.2%) (lb/hr 20.2%)

1930 1020 938 7.25 61 R

3 Table 5-6 AXIAL POWER FACTORS FOR THE THERMt.L-HYDRAULIC PROGRAM Node APF Node APF Node APF (Bottom) 1 0.47 9 1.29 17 1.15 2 0.55 10 1.34 18 1.08 3 0.64 11 1.38 19 1.01 4 0.74 12 1.40 20 0.93 5 0.85 13 1.39 21 0.84 6 0.97 14 1.36 22 0.74 7 1.10 15 1.30 23 0.60 8 1. 21 16 1.23 (Top) 24 0.43 5-55

NEDO-24195 i

Table 5-7

NONVARYING PLANT GETAB ANALYSIS INITIAL CONDITIONS - OYSTER CREEK ,

I 6

Core Flow x 10 Reactor Core Inlet Enthalpy Nonfuel j

, (lb/hr) Pressure (psia) (Btu /lb) Power Fraction 2

61 1050 517.5 0.035 8

i Table 5-8 I

CONTROL R0D WITHDRAWAL ERROR ANALYSIS Example ,

, MCPR Operating Limit Increase with Decreased Flow [

Frem 100% Power, 100% Flow AMCPR Limit to %55%, 40% Flow (on Flow Increase = 0.19 i Control Line)

MCPR Change During Withdrawal Rod In to Rod Out at =

0.33 Decrease in MCPR 100% Power, 100* Flow Rod In to Rod Out at = 0.41 Decrease in MCPR

%55% Power, 40% Flow i 1

AMCPR Increase <:,0.08 l

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5-56

4 NED0-24195 ,

O Table 5-9 SENSITIVITY OF CPR TO VARIOUS THERMAL-HYDRAULIC PARAMETERS Approximate _,

ACPR + Nominal CPR Parameter Nominal Value AParameter + Nominal Parameter P8x8R Fuel Bundle Power 3.3 MWt -1.0 (or Relative Bundle Power)

Bundle Coolant 6 lbm-G = 1.1 x 10 2

+0.2 Flow Rate hr-ft W = 0.12 x 10 lbm/hr Core Coolant 20-27 Stu/lbm +0.1 Inlet Subcooling GEXL R-factor 1.04 -2.1 Core Pressure 1035-1055 psia -0.6 (with constant O coolant subcooling) l O .

5-57

. C3

() Table 5-10 SIGNIFICANT INPUT PARAMETERS TO THE LOSS-OF-COOLANT ACCIDENT ANALYSIS

_ 2{

Plant Parameters:

Core Thermal Power 1969 MWe, which corresponds to 102% of rated power Vessel Steam Output 7.4 x 10 lbm/h, which corresponds to 102% of rated power -*

Vessel Steam Dome Pressure 1020 psig cI Recirculation Line Break Area 4.66 ft2 (DBA), 1.0 ft2 , 0.13 ft 2

C; Number of Drilled candles 'e None Fuel Parameters:

Peak Technical Specification Linear Heat Design Axial Fuel Bundle Generation Rate Peaking Fuel Type Geometry (kW/ft) Factor

__ P8DRB239 8xA 13.4 1.57

) P8DRB265L ftx8 13.4 1.57 es P8DRB265H Sr.E 13.4 1.57 [;

P8DNB277 8x8 13.4 1.57 P8DRB282 8x8 13.4 1.57 _

Table 5-11

SUMMARY

OF BREAK SPECTRUM RESULTS e Break Size Core-Wide e Location Peak Local Metal-Water e Single Failure PCT (*F) Oxidation (%) Reaction (%)

^

e 4.66 ft 2058 *

e Recirculation Suction e Emergency Condenser e 1.0 ft 2082 * *

e Recirculation Suction e Emergency Condenser e 0.13 ft 2200 7.9 0.40 e Recirculation Suction -

s e Emergency Condenser

{d . ~* '

Not limiting case "'m 5-58

I ) Table 5-12 LOCA ANALYSIS FIGURE

SUMMARY

2 Limiting Break (0.13 ft ) DBA 1.0 ft Water Level Inside 5-18e 5-18a 5-18b .

Shroud and Reactor Vessel Pressure Peak Cladding 5-19c 5-19a 5-19b ,

Temperature e-e Heat Transfer 5-20e 5-20a 5-20b Coefficient Normalized Powe.r 5-21 5-21 5-21 S

Peak Cladding 5-22 5-22 5-22 27 Temperature Vs.

Break Area ~

Table 5-13

~

SINGLE FAILURES CONSIDERED IN THE OYSTER CREEK LOCA ANALYSIS (Reference 5-22)

Single 4 Break Location Failure Available Systems Recirculation Line 1 EC 2 CS + 0 EC + 5 ADS 1 ADS 2 CS + 1 EC + 4 ADS '

Feedwater and 1 EC 2 CS + 1 EC + 5 ADS Steam Lines 1 ADS 2 CS + 2 EC + 4 ADS i

Core Spray Line 1 EC 1 CS + 1 EC + 5 ADS 1 ADS 1 CS + 2 EC + 4 ADS EC = Emergency Condenser CS = Core Spray ADS = Automatic Depressurization System

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