ML20107F213

Rev 1 to Mechanical Design Rept Suppl for Kewaunee High Burnup (49GWd/MTU) Fuel Assemblies

ML20107F213
Person / Time
Site:	Kewaunee
Issue date:	02/28/1985
From:	Hoppe N SIEMENS POWER CORP. (FORMERLY SIEMENS NUCLEAR POWER
To:
Shared Package
ML111750906	List: ML20107F213
References
XN-NF-84-28(NP), XN-NF-84-28(NP)-R01, XN-NF-84-28(NP)-R1, NUDOCS 8502260016
Download: ML20107F213 (46)

Text

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g XN-NF- 84-28 NPJ l REVISION 1 lI lI MECHANICAL DESIGN REPORT SUPPLEMENT FOR

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! KEWAUNEE HIGH BURNUP (49GWd/MTU?

I FUEL ASSEMBLIES I

I I FEBRUARY 1985 I

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ERON NUCLEAR COMPANY, INC.

XN-NF-84-28(NP), Rev. 1 Issue Date: 2/11/85 MECHANICAL DESIGN REPORT SUPPLEMENT FOR KEWAUNEE HIGH BURNUP (49 GWd/MTU) FUEL ASSEMBLIES

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Prepared: ,- /-ch' t")

<h N . E .' fidppe ITitT

'W Project Engineer Approved: /[VN

'G'./J. BusseTman, Manager Date Fuel Design I . , i , D,

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Concur: h li L (,d1 li i n,: ( i-t 8.,B. Stout, Manager 4.1 censing & Safety Engineering Date

'0 < z A//tf H. E.'Williamson, Manager Date Neutroitics & Fuel Management

_ h ,);l lL =l&

I J.1!P' Morgan, Manager Pr6posals & Customer Services Engineering Date G. A. Sofep Manat;er b .

r h. L"J '

Date Fuel Eng7heering & Technical Services i sh NOTE: THERE WAS NO (NP) ISSUE'0F REVISION 0 0F THIS REPORT.

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CUSTOMER DISCLAIMER 8

IMPORTANT NOTICE REGARDING CONTENTS AND USE OF THIS DOCUMENT i

PLl'ASE READ CAREFULLY _

. Exxon Nuclear Company's warranties and representations concoming the subject metter of this document are those set forth in the Agreemer't between Exxon Nucteer Company, Inc. and the Customer pursuant to which this document is issued. Accordingly, except as othervnse expressly provedad in such Agreement, neither Exxon Nuclear Company, Inc. not any person actmg on its behalf makes any warranty or representonon, expressed or implied, with roepect to the accuracy, completeness, or useftsiness of the information contained in this document, or that the use of any information, apparatus, method or process disclosed in this document will not infnnge private 4y owned rights; or assumes any liathlities wt1h respect to the use of any information, apparetus, method or process diecioned in this Jacument.

The informenon contained herem is for the soie use of Customer.

In crder to avoid impeermont of rights of Exxon Nuch Company, Inc.

in potents or invennons which may be included in the information contamed in this document, the recipient, by its acceptone of this document agrees not to publish or make public use (in the potent use of the term) of sudi informecon unal so authorized in wntmg by Exxon Nucteer Company, Inc.

or until after six (6) months followeg termineoon or expiretuwt of the aforcesid Agreement and any exonsson thereof, unises otherwise expressly provided in the Agreement. No rights or licenses in or to any potents are implied by the fumiehing of this domsment. W I

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-i- XN-NF-84-28(NP), Rev. 1 TABLE OF CONTENTS I Section 1.0 Title ~

INTRODUCTION ..........................................

Page 1

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l 2.0

SUMMARY

............................................... 1 3.0 DESIGN BASES .......................................... 3 0 3.1 CLADDING PHYSICAL AND MECHANICAL PROPERTIES ... 3 3.2 CLADDING STRESS LIMITS ........................ 4 3.3 CLADDING STRAIN LIMITS ........................ 5 3.4 STRAIN FATIGUE ................................ 6 3.5 FRETTING CORROSION AND KEAR ................... 6 3.6 CORROSION ..................................... 7 3.7 HYDROGEN ABSORPTION ........................... 7 3.8 CREEP COLLAPSE ................................ 8 3.9 FUEL R0D INTERNAL PRESSURE .................... 9 3.10 CREEP BOW .................................... 10 3.11 OVERHEATING 0F CLADDING ...................... 10 3.12 OVERHEATING 0F FUEL PELLETS .................. 10 3.13 FUEL R00 AND ASSEMBLY GROWTH ................. 11 4.0 DESIGN DESCRIPTION ................................... 13

. 4.1 FUEL ASSEMBLY ................................ 13 4.2 FUEL R00....................................... 13 5.0 FUEL ASSEMBLY MATERIAL PROPERTIES .................... 16-'

8 5.1 ZIRCALOY-4 ................................... 16 5.2 FISSILE. MATERIAL (URANIUM DIOXIDE) ........... 16

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. INCONEL SPRINGS .................. . ........... 17 I

-ii- XN-NF-84-28(NP) Rev. 1 TABLE OF CONTENTS (Continued)

Section Title Page 6.0 CONDITIONS FOR FUEL R00 MECHANICAL DESIGN ............ 18 6.1 REACTOR OPERATING CONDITIONS ................. 18 6.2 R0D DIMENSIONAL DATA ......................... 18 6.3 EXPOSURE HISTORY ............................. 18 6.4 DESIGN CRITERIA .............................. 19 7.0 FUEL R00 MECHANICAL DESIGN ANALYS IS . . . . . . . . . . . . . . . . . . 24 7.1 STEADY-STATE STRESS ANALYSIS ................. 24 7.2 STEADY-STATE STRAIN .......................... 24 7.3 TRANSIENT CLADDING STRESSES AND STRAINS ...... 25 7.4 CYCLING FATIGUE .............................. 25 7.5 COLLAPSE ..................................... 26 7.6 CORROSION AND HYDR 0 GEN P ICKUP . . . . . . . . . . . . . . . . 27 7.7- FUEL R0D ELONGATION . . . . . . . . . . . . . . . . . . . . . . . . . . 27 7.8 INTERNAL PRESSURE ............................ 28 7.9 R00 B0 WING ................................... 28 7.10 FUEL R0D PLENUM SPRING . . . . . . . . . . . . . . . . . . . . . . . 29 7.11 FUEL AND CLADDING TEMPERATURE ................ 30 8.0 FUEL ASSEMBLY EVALUATION ............................. 30 8.1 GENERAL DESCRIPTION .......................... 30.

8.2- DESIGN CRITERIA .............................. 30 8.3 DESIGN ANALYSIS .............................. 31

9.0 REFERENCES

............................................'33 8

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g -iii- XN-NF-84-28(NP) Rev. 1 LIST OF TABLES i

Table No. Title , Page 3.1 STEADY STATE STRESS DESIGN LIMIT .....................

i 12 4.1 FUEL ASSEMBLY DESIGN ................................. 14 6.1 FUEL R00 ATTRIBUTES .................................. 21 6.2 POWER HISTORIES A AND B .............................. 22 6.3 POWER HISTOR IE S C AND D . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 23 8

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i XN-NF-84-28(NP), Rev. 1 I Issue Date: 2/11/85 Page 1 i MECHANICAL DESIGN REPORT SUPPLEMENT FOR KEWAUNEE HIGH BURNUP (49 GWd/MTU) FUEL ASSEMBLIES

1.0 INTRODUCTION

The Kewaunee XN-1 through XN-4 Reload fuel was originally designed for an average fuel assembly burnup of 33 GWd/MTV, then reanalyzed for a peak fuel rod burnup of 43 GWd/MTU. This report describes the mechanical design analyses which show that the fuel from Reloads XN-1 to XN-4 can be irrad-iated to 49 GWd/MTV peak rod burnup, using as-built fuel dimensional characteristics. It also describes the analyses performed to qualify the XN-5 through XN-9 fuel Reloads, which are characterized by design improve-

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ments, for a peak rod burnup of 49 GWd/MTU.

2.0

SUMMARY

The fuel design for the Kewaunee plant has been modified starting with

~ the XN-6 Reload to accommodate to higher burnup. The changes consist of tighter specifications for the cladding characteristics and the applica-tion ccf a fuel resinter density change limit. The as-built 9eload XN-5 satisfied these new specifications.

The existing reload fuel designs and the modified design have been

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reanalyzed to support an increase in peak rod burnup up to 49 GWd/MTU.

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Issues not affected by the increased burnup are covered by the base Design f Report.(1) Mechanical design analyses were performed to evaluate cladding I

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i XN-NF-84-28(NP), Rev. 1 Page 2 steady-state strain, transient stress and strain, fatigue, creep collapse, corrosion, hydrogen absorption, fuel rod internal pressure, elongation, and fuel assembly growth.

Design criteria consistent with current ENC methodology were used in the analyses. Some of the design codes and techniques have been improved since the original mechanical design analysis was performed for Reloads XN-1 and XN-4 which justified up to a peak rod burnup of 43 GWd/MTU. The strain, pressure and collapse analyses have been performed using the RODEX2 code version approved by the NRC in 1983. The ramp stress / strain analysis has been evaluated against both the latest strain criteria and against stress criteria, namely to protect against f ailure by stress corrosion cracking.

The current analyses were performed to a peak rod burnup of 49 .

GWd/MTU, both for the reloads with the new specifications. and for the earlier reloads. Bounding power histories have been used.

The results indicate that all the mechanical. design criteria.are satisfied.

o The maximum end-of-life (E0L) steady-state cladding strain meets the 1.0% design limit.

o- The cladding stress and strain during power ramps, calculated

u. sing different overpower conditions,- do not exceed .the design stress corrosion cracking threshold or the~1.0% strain limit.

o The cladding fatigue usage factor is within the design limit.

o The end-of-life -fuel rod internal pressure :is' less than- the system-pressure..

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a_s XN-NF-84-28(NP), Rev. 1 NdM g Page 3 y n .x .,

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g o The criterion for the prevention of creep collapse is satisfied. -m W

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The maximum calcul ated EOL thickness of the oxide corrosion layer o

A and the maximum calculated concentration of hydrogen in the cladding are within the design limits.

4 3.0 DESIGN BASES g t he design considers ef fects and changes in physical properties of fuel assembly components which result from burnup.

f The integrity of the fuel rods is ensured by analyzing the fuel to

? 1 snow that excessive fuel temperatures, excessive internal rod gas pres-sures, and excessive cladding stresses and strains do not occur. This end -

is achieved by showing the fuel rods to satisfy the design bases for normal operation and anticipated operational occurrences over the fuel lifetime. ,

5_ For each design basis, the performance of the most limiting fuel rod shall

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not exceed the specified limits.

j The functional capability of the fuel assembly is ensured by analyzing  ;

the fuel assembly to show that the fuel system dimensions and properties 3 remain within operational tolerances. This is achieved by showing that the )

1 3= fuel assemblies satisfy the design bases for normal operatica and antici- 1 1

pated operational occurrences over the fuel lifetime. d

_ 3.1 CLADDING PHYSICAL AND MECHANICAL PROPERTIES Zircaloy-4 combines a low neutron absorption cross section, high

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5 corrosion resistance, and high strength and ductility at operating tempera-

= tures. Principal physical and mechanical properties including irradiation 7

effects on Zircaloy-4 are proviced in Section 5.

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I XN-NF-84-28(NP), Rev. 1 Page 4 3.2 CLADDING STRESS LIMITS The design basis for the fuel cladding stress limits is that the fuel system will not be damaged due to fuel cladding stresses exceeding '

material capability. Conservative limits are derived from the ASME Boiler and Pressure Code,Section III, Article III-2000 (Reference 3).

The cladding may also be damaged by the combination of volatile

! fission products and high cladding tensile stresses which may lead to stress corrosion cracking.(4,5) Stress corrosion cracking of fuel rod cladding is considered the principal f ailure mechanism for PCI f ailures encountered during changes in reactor operating conditions.(6,7,8) Even though unanimous agreement has not been reached on which chemical species enhances f ailure, the iodine atmosphere is usually considered the primary attacking media in irradiated fuel. If the stress level is low enough in the cladding, then stress corrosion cracking does not occurs Tests have .

been done under EPRI support (9,10,11) to evaluate a stress threshold "

associated with stress corrosion cracking in an iodine atmosphere. Typical data from those programs show that the time dependence of stress corrosion rupture involves two processes. At lower stresses, time to failure is ,

largely controlled by a time-depen' dent' crack nucleation process. 'Thus, if stress levels remain low enough, a flaw or crack that would ' subsequently. .

propagate will not be nucleated.

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The concept used to' avoid failures from the stress' corrosion cracking failure mechanism from power. ramps is to keep' the Lfuel redsLfrom operating above the. stress threshold associated with the nucleation of a

-propagating' stress corrosion crack. The modelling of the-stress corrosion I

I XN-NF-84-28(NP), Rev. 1 Page 5 I

crack propagation process and methods for predicting the stress levels in I fuel rods operating under prototype exposure histories, incorporate many assumptions. The design procedure used to evaluate ENC fuel rods uses a stress threshold determined from benchmarking studies using the RODEX2(12) and RAMPEX codes. The design criterion for the transient stress limit, resulting from a power ramp, is to keep the predicted stress levels below the stress threshold obtained in the benchmarking studies of test ramp cases.

The benchmarking test results were obtained from the Studsvik Inter-Ramp, Over-Ramp and Super Ramp test series. Conservatism in the design bases is obtained by using a safety factor on the code benchmarked failure stress threshold, by using conservative input values for the f uel rod dimensions in the design analyses, and by assuming wo st case power histories and ramp powers for the analysis.

3.3 CLADDING STRAIN LIMITS Tests (14,15) on irradiated tubing indicate potential for failure at relatively low mean strains. The data on tensile, burst and split ring tests indicate a ductility ranging between 1.2% and 5% at normal reactor

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.I operating temperatures. The failures are usually associated with unstable or localized regions-of h_id deformation af ter some uniform defo*mation. To prevent cladding failure due to plastic instability and localization cf strain, the_ total mean hoop cladding strain for steady-state conditions is limited to 1%,_and the increment of de thermal creep during a transientLis also limited to 1%.

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E XN-NF-84-28(NP), Rev. 1 Page 6 3.4 STRAIN FATIGUE Cyclic PCI loading, colnbined with other cyclic loading associated with relatively large changes in power, can cause cumulative damage which may eventually lead to fatigue failure. Cyclic loading limits are estab-lished to prevent fuel failures due to this mechanism. The design life is based on correlations which give a safety factor of 2 on stress amplitude .

or a safety factor of 20 on the number of cycles, whichever is more conservative.(16) 3.5 FRETTING CORROSION AND WEAR The design basis for fretting corrosion and wear is that fuel rod failures due to fretting shall not occur. Since significant amounts of fretting wear can eventually lead to fuel rod f ailure, the grid spacer assemblies are designed to prevent such wear. The spring dimple system in the spacer grid is designed such that the minimum spring / dimple forces throughout the design life are greater than the maximum fuel rod flow vibration forces. Testing of a wide variety of ENC fuel designs shows fuel rod wear is due primarily to fuel rod loading and unloading, and not due to fuel rod motion during the test. ' There has been little or no difference-between observed wear for' 500 hour0.00579 days <br />0.139 hours <br />8.267196e-4 weeks <br />1.9025e-4 months <br />,1000 hour and 1500 ' hour tests. ' No I-active fretting corrosion has been observed despite spacer spring relaxa .

tion in several test assemblies. - Examination of a 'large number of irra-diated rods. has substantiated the minimal wear observed af ter loop tests. .

Numerous.PWR reload: batches with this typical ENC bimetallic l spacer have.

operated in sixteen reactors with no adverse effec 1s.due to-fretting; I corrosion or wear unrelated to baffle jetting.

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I XN-NF-84-28(NP), Rev. 1 Page 7 I

3.6 CORROSION Cladding oxidation and corrosion product buildup are limited in order to prevent significant degradation of clad strength. A PWR clad external temperature limit is chosen, as corrosion rates are very slow below this temperature, and therefore, overall corrosion is limited. An external corrosion layer limit is also specified, as this amount of corrosion will not significantly affect thermal and mechanical design margins. This decreasse in clad thickness does not increase clad stresses above allowable levels.

Corrosion product buildup, and the resulting temperature in-creases, are calculated directly in the RODEX2 code.

3.7 HYDR 0 GEN ABSORPTION The as-f abricated cladding hydrogen level and the fuel rod

- cladding hydrogen level during life are limited to prevent adverse effects on the mechanical behavior of the cladding due to hydriding. Hydrogen can be absorbed on either the outside or the inside of the cladding. Excessive absorption of hydrogen can result in premature cladding failure due to reduced ductility and the formation of hydride platelets.

The effects of hydrogen on mechanical properties have been investigated at hydrogen concentrations to about 1000 ppm. The effect on strength and ductility depends on such factors as:

o The - tube texture which tends to promote .or' minimize radially orientated hydrides.

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XN-NF-84-28(NP), Rev. 1 Page 8 o Stress and temperature cycling which may promote reorien-tation of hydrides into radial directions. Tensile hoop stress tends to orient hydrides radially.(20) o Distribution of hydrides (hydride case layers on the I.D.

or 0.D. surface tend to promote brittle failures).

o Ratio of cladding wall thickness to average length of hydride platelet.

o The fineness and uniformity in dispersion of the second phase precipitate tend to improve corrosion resistance and decrease hydrogen absorption.

The calculation of hydrogen concentration due to pickup from the coolant is calculated in the RODEX2 code. Hydrogen absorption from inside the clad is minimized by careful moisture control during fuel fabrication. 3 5

3.8 CREEP COLLAPSE The design basis for creep collapse of the cladding is that significant. axial gaps due to fuel densification shall not occur, and therefore, that fuel failure due to creep collapse shall not occur. Creep collapse of the. cladding can increase nuclear peaking, inhibit heat transfer, and cause failure due to localized strain.

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_If significant gaps form in the, pellet column due to fuel .

--densification, the pressure differential between the inside and.outside of N the cladding can act t'o increase cladding-ovality. 0v slity . increase by -

clad creep to the point of plastic instability would result in collapse of r

the cladding. During power changes, such collapse could result in' fuel  :

failure.

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I XN-NF-84-28(NP), Rev. 1 Page 9 Through proper design, the formation of axial gaps and the probabi'lity of creep collapse can be significantly reduced. Typical ENC pellets are stable dimensionally. For high burnup designs, the lot average resinter density change is limited by specification. This specification ensures stable pellets during irradiation, and limits the potential size of fuel column gaps.

An Inconel X-750 plenum spring is included in the ENC fuel rod design, and the rods are pressurized with helium to help prevent the formation of gaps in the pellet column. The plenum spring provides a compressive force on the fuel column throughout the densification phase of the fuel life, and the internal pressure prevents rapid clad creepdown as well as providing a good heat transfer medium for the fuel.

An analysis is performed in order to guard against the unlikely 5 event that sufficient densification occurs to allow pellet column gaps of sufficient size for clad flattening to occur. With this method, creep ovality is calculated with the COLAPX code and cladding uniform creepdown is calculated with the RODEX2 code (12) utilizing conservative design conditions.

3.9 FUEL R0D INTERNAL PRESSURE The internal gas pressure of the fuel rods shall not exceed the:

external coolant pressure. Significant outward circumferential creep which-may cause an increase in pellet-to-cladding gap must be prevented, since it would lead to higher fuel temperature and higher fission gas reiaase. -Fuel rod internal pressure is calculated throughout life with the RODEX2 code.

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XN-NF-84-28(NP), Rev. 1 Page 10 3.10 CREEP B0W I

l Differential expansion between the fuel rods and lateral thermal l W

l and flux gradients can lead to lateral creep bow of the rods in the span l between spacer grids. The design basis for fuel rod bowing is that lateral displacement of the fuel rods shall not be of sufficient magnitude to impact thermal margins. ENC fuel has been designed to minimize creep bow.

Extensive post-irradiation examinations have confirmed that such rod bow i has not reduced spacing between adjacent rods by.more than 50%. The g potential effect on thermal margins is negligible. E 3.11 OVERHEATING 0F CLADDING

! The design basis for fuel rod cladding overheating is that transition boiling shall be prevented. Prevention of potential fuel l

i f ailure from overheating of the cladding is accomplished by minimizing the l

probability that boiling transition occurs on the peak fuel rods during I normal operation and anticipated operational occurrences. Margin to boiling transition is evaluated using applicable DNB correlations, with ENC's XCOBRA-IIIC based PWR thermal-hydraulic methodology.

l 3.12 OVERHEATING 0F FUEL PELLETS Prevention of fuel f ailure from overheating of- the fuel pellets is accomplished by assuring that the peak linear heat generation rate i

'(LHGR) during normal operation and anticipated operational occurrences does not result in fuel centerline melting. ' The melting point of the fuel -is -

' adjusted for burnup in the-centerline _ temperature analysis.

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I XN-NF-84-28(NP), Rev. 1 Page 11 I 3.13. FUEL R0D AND ASSEMBLY GROWTH The design basis for fuel rod and assembly growth is that adequate clearance shall be provided to prevent any interference which might lead to buckling or damage. Cladding and guide tube growth measure-ments of ENC fuel are used in establishing the growth correlations used for calculations.

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I XN-NF-84-28(NP), Rev. 1 I

Page 12 TABLE 3.1 STEADY STATE STRESS DESIGN LIMIT Stress Intensity Limits **

Stress Category

• Ultimate a Yield Tensile g Strangth Strength General Primary Membrane 2/3 1/3 Primary Membrane Plus Primary Bending 1.0 1/2 Primary Plus Secondary 2.0 1.0

• Characteristics of the stress categories are defined as follows:

a) Primary stress is a stress developed by the imposed loading which is j necessary to satisfy the laws of equilibrium between external and 5 internal forces and moments. The basic characteristic of a primary stress is that it is not self-limiting. If a primary stress exceeds 3 the yield strength of the material through the entire thickness, the g prevention of f ailure is entirely dependent on the strain-hardening properties of the material.

b) Secondary stress is a stress developed by the self-constraint of a structure. It must satisfy an imposed strain pattern rather than being in equilibrium with an external load. The basic characteristic 5 cf a secondary stress is that it is self-limiting. Local yielding and 5 minor distortions can satisfy the discontinuity conditions of thermal expansions which cause the stress to occur.

• • - The stress intensity is _ defined as twice the maximum shear stress and is equal to the largest algebraic difference between any two of the three principal stresses.

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XN-NF-84-28(NP), Rev. 1 Page 13 4.0 DESIGN DESCRIPTION 4.1 FUEL ASSEMBLY The 14x14 fuel assembly array includes 16 guide tubes,179 fuel rods and one instrumentation tube. The grid spacers are of standard ENC bi-metallic design, and the fuel assembly tie plates are stainless steel castings with Inconel holddown springs. . Fuel assembly characteristics are summarized in Table 4.1.

4.2 FUEL R0D The fuel rods consist of cylindrical UO2 pellets in Zircaloy-4 tubular cladding.

The Zircaloy-4 fuel rod cladding is cold-worked and lightly stress relieved. Zircaloy-4 plug type end caps are seal welded to each end. The upper end cap has external features to allow remote underwater fuel rod handling. The lower end cap has a truncated cone exterior to aid

_ fuel rod reinsertion into the fuel assembly during inspection t.nd/or reconstitution.

-I Each fuel rod contains a 144.0 inch column of enriched U02 fuel pellets.

The fuel rod upper plenum contains an Inconel X-750 comoression spring to prevent fuel column separation during f abrication and shipping, and during in-core operation.

Fuel rods are pressurized with helium which provides a good heat; -

transfer medium and assists in the prever: tion of clad creep collaose.

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l TABLE 4.1 FUEL ASSEMBLY DESIGN I

FUEL PELLET l

Fuel Material UO2 Sintered Pellets Pellet Diameter, (in.) 0.3565 CLADDING

Clad Material Zircaloy-4 Cold Worked and g l

Stress Relieved 5 l Clad ID, (in.) 0.364 Clad 00, (in.) 0.424 l

l Clad Thickness, Nominal, (in.) 0.030 FUEL R00 Diametral Gap, Cold Nominal, (in.) 0.0075 Active Length, (in.) 144,0

, Total Rod Length, (in.) 152.065 Fill Gas Helium l

SPACER I

Material Zr-4 & Inconel 718 g Rod Pitch 0.556 5 Envelope (in.) 7.763 square GUIDE TUBE Material Zr-4 ID/ID Above Dashpot (in.) 0.541/0.507 TIE PLATES Material ._Styinless Steel -.

110LD00WN SFRINGS g

Material .Inconel 4 I

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l TABLE 4.1 (Continued) iI i

CAP SCREWS

, Materials Inconel and SS FUEL ASSEMBLY I Array 14x14 Assembly Pitch 7.803

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i No. Spacers 7

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No. Fuel Rods 179 No. Guide Tubes 16 No. Instrumentation Tubes 1

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5.0 FUEL ASSEMBLY MATERIAL PROPERTIES The material properties used in the design evaluation are described in this section. The Zircaloy cladding properties and the U02 fuel properties utilized are as incorporated in the RODEX2 and RAMPEX fuel performance codes.

5.1 ZIRCALOY-4 S.1.1 Chemical Properties Zircaloy-4 is used in three forms: (1) Coldworked and stress relieved cladding; (2) Recrystallized annealed tubing; and (3)

Recrystallilzed annealed strip. The chemical properties are in accordance with Grade R60804 (RA-2).

5.1.2 Physical Properties The Zircaloy cladding properties are as incorporated in

-I the R0DEX2 code.

5.2 FISSILE MATERIAL (URANIUM DIOXIDE) 5.2.1 Chemical Composition a) Uranium Content The uranium . content shall be a minimum of 87.7% by I weight of the uranium dioxide on a dry weight basis.

b) Stoichiometry The oxygen-to-uranium ratio of the sin'tered fuel pellets shall. be within the limits of .1.99 and 2.01.

5.2.2 . Thermal Propertiet The thermal properties utilized are as incorocrated in the RODEX2 code.

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Page 17 5.2.3 Mechanical Properties a) Mechanistic Fuel Swelling Model The irradiation environment and fissioning events cause the fuel material to alter its volume and, consequently, its dimen-sions, of the RODEX2 report.

b) Fission Gas Release The evaluation of fission gas release is done by the i RODEX2 code. For design evaluations of end-of-life pressures, pellet-cladding interaction and general thermal mechanical conditions, a physi-cally based two-stage release model is used. First stage fission gas release is to grain boundaries, and then the second stage release is from the grain boundaries to the interconnected free gas volume. This release model is described in detail in Appendix E of the R0DEX2 report.

c) Melting Point The value used for the U02 melting point (unirradi-ated) is 28050C-(50810F). Based on measurements by Christensen, et al(45),

the melting point is reduced linearly with irradiation at the rate of 12.20C (22.00F) per 1022 fiss/cm2 or 320C (57.60F) per 104

- mwd /MTU.

5.3 INCONEL SPRINGS Coil springs are f abricated from Inconel X-750 wire or rod with

.an. alley composition in accordance with Table 5.4 (AMS 5699B).

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I XN-NF-84-28(NP), Rev. 1 Page 18 I 6.0 CONDITIONS FOR FUEL R0D MECHANICAL DESIGN 6.1 REACTOR OPERATING CONDITIONS Core power level (Nominal)

I 1650 MWt Coolant operating pressure (Nominal) 2250 psia Coolant flow rate (at nominal power)

Total 68.2 x 106 lb/hr.

Active Core 65.2 x 106 lb/hr.

g Heat generation in fuel 97.4%

l g Coolant inlet temperature (Nominal) 5340F Number of assemblies in core 121

! Maximum peak pellet LHGR 14.47 kW/ft.

Maximum peak rod burnup 49 GWd/MTU The fuel shall be capable of load-follow operation between 40%

.I and 100% of rated power for at least two months per year, and not preclude the transients set forth in the FSAR.

l 6.2 R0D DIMENSIONAL DATA Some of the cladding and fuel pellet characteristics for each reload are listed in Table 6.1. The characteristics of Reloads XN-6 through XN-9 are based on specifications, while values for Reloads XN-1 to XN-4 are based on as-built measured data. The values for Reload XN-5 are covered by the specifications of Reloads XN-6 through XN-9.

6.3 EXPOSURE HISTORY Multiple power histories were develpped by Wisconsin Public'

-- Service and ENC.

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I XN-NF-84-28(NP), Rev. 1 Page 19 Four bounding histories which simulate various fuel shuffling I

schemes were selected. Tables 6.2 to 6.3 give the power histories and the corresponding fast fluxes. The power histories used were:

Case A: Three high power cycles and a low power fourth cycle.

Case B: Medium-High power during four cycles.

Case C: Four cycles, three high power with low power during the second cycle.

Case D: Four cycles, three high power with lower power during the third cycle.

All power histories are such that the rod average burnup is 49 GWd/MTU.

6.4 DESIGN CRITERIA The mechanical design criteria are: ,

1. The maximum steady-state primary and secondary stresses shall meet the' ASME Boiler and Pressure Vessel Code,Section III require-ments,(3) as defined in Table 3.1.

2. The maximum cladding hoop stress at. pellet ends during power ramping is limited to avoid failure by stress corrosion cracking.

3. The cumulative usage factor for cyclic stresses shall not -

exceed 0.67.

4 The net cladding mean hoop strain shall not increase by more than 1% for steady-state operation. The-increment of the cladding hoop thermal creep strain at pellet ends during a ramp is also limited'to

.1%.

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XN-NF-84-28(NP), Rev. 1 Page 20 I 5. Cladding creep collapse shall not occur.

6. The hydrogen absorption of the cladding and the thickness of the corrosion layer shall not exceed design limits.

7. The internal pressure in the fuel rod at end-of-life shall not exceed the system pressure.

8. The fuel elongation must be accommodated by the clearance between fuel rods and tie plates.

9. The fuel assembly growth must be accommodated by the clearance between the fuel assembly and the core plates.

10. Fuel rod creep bow throughout the design life of the assemblies shall be limited so as to maintain licensing and operational limit restraints.

11. The fuel rod plenum spring shall maintain a positive

_ compression on the fuel column during shipping and during the fuel densifi -

. cation stage.

12. Cladding temperatures shall not exceed the design limits.

13. Pellet temperatures shall' not exceed the melting tempera-ture during normal operation and anticipated transients.

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! XN-NF-84-28(NP), Rev. 1 I

Page 21 I

TABLE 6.1 l

FUEL R0D ATTRIBUTES g

l 5 XN-6 Reload Through Cladding XN-9 Units Clad ID Avg. 0.3640 inch OD Avg. 0.4240 inch Pellet Density Avg. 94.0 %TD Nominal Enrichment 3.2-3.4 %U-235 I

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l XN-NF-84-28(NP), Rev.1 Page 22 I TABLE 6.2 POWER HISTORIES A AND 8 B0UNDING CASE A - HIGH, HIGH, HIGH, LOW POWER Peak P e a'<

I Cycle Burnup Irradiation Time Hrs.

Assembly Burnup mwd /MTU Peak Rod Burnup mwd /MTV Rod Average LHGR kW/ft Rod Fast Flux (>1MeV) 1013n/cm2sec I 11000 0 0 7340 15500 0 0 17050 9.839 9.839 8.86 9.44 0 0 15500 17050 8.252 8.84 5 11000 7340 28500 31350 8.252 9.05 0 0 28500 31350 7.646 9.70 5 11000 7340 41000 44600 7.646 9.60 0 0 41000 44600 2.539 3.00 11000 7340 45000 49000 2.539' 3.00 B0UNDING CASE 8 - MEDIUM POWER & HIGH BURNUP Irradiation Assembly Peak Rod Rod Average Rod Fast I Cycle Burnup Time Hrs.

Burnup mwd /MTU Burnup mwd /MTU LHGR kW/ft Flyx (>1McV) 1013n/cm2sec I 11000 0 0 7340 12775 0 0 14000 8.079 8.079 7.28 7.75 0 0 12775 14000 6.925 8.70

, -l- 11000 7340 23700 26000 6.925 8.90 0 0 23700 26000 6.925 9.00 11000 7340 35000 38000 6.925 9.15

'O O 35000 38000 6.348- 8.70

11000 7340 45000_ 49000 6.348' -8.80 I #

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XN-NF-84-28(NP), Rev. 1 Page 23 B0UNDING CASE C - HIGH, LOW, HIGH, HIGH POWER I

Irradiation Assembly Peak Rod Rod Average Rod Fast Cycle Time Burnup Burnup LHGR- Flyx (>1MeV)

Burnup Hrs. mwd /MTU mwd /MTV kW/ft 1013n/cm2sec 0 0 0 0 9.839 8.86 11000 7340 15500 17050 9.839 9.44 0 0 15500 17050 2.539 2.90 11000 7340 19500 21450 2.539 3.00 0 0 19500 21450 8.252 9.60 11000 7340 32500 35750 8.252 9.80 0 0 32500 35750 7.646 9.80 11000 7340 45000 49000 7.646 9.80 ROUNDING CASE D - HIGH, HIGH, LOW, HIGH POWER Irradiation Assembly. Peak Rod Rod Average Rod Fast a Cycle Burnup Time Hrs.

Burnup mwd /MTU Ruenup mwd /MTU LHGR kW/ft Flyx (>1MeV) 1013n/cm2sec 5

0 0 0 0 '9.839 _8.86 11000 7340 15500 17050 9.839 9.44 0 0 15500 17050 8.252 8.84 11000 7340 28500 31350 8.252 9.05 0 0 28500 31350 2.539 2.95 m 11000 7340 32500 35750 2.539 _3.00 g.

0 0 32500 35750 7.646 9.80 11000 7340 45000 49000 7.646 9.80 .

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I XN-NF-84-28(NP), Rev. I I Page 24 7.0 FUEL R00 MECHANICAL DESIGN ANALYSIS 7.1 STEADY-STATE STRESS ANALYSIS The stresses were calculated at BOL hot and cold conditions, and B0C4 hot and cold conditions when the pellet / clad mechanical interaction is a maximum (History D). The' calculations were performed using the long term cladding behavior -from the RODEX2(12) calculations done for the collapse determinations, and using l spacer-induced stresses calculated by an ANSYS(46) calculation. The collapse calculation cladding conditions were chosen because the minimum cladding thickness is used; and therefore, the

_ stresses would be higher. The ANSYS analysis was a finite element stress analysis done for 0.424 OD cladding. The maximum stress intensities have

'been determined using the. same technique as in the original design i, report.(1)

The results indicate that the calculated stresses are well_ below the design limits.

7.2 STEADY-STATE STRAIN The cladding steady-state strain was evaluated with the RODEX2(12) code, latest version, as approved by the_NRC in 1983. The code calculatbs the thermal, mechanical and' compositional state of'the fuel,-and cladding ~for-a given duty history. Conservative-input salues were used in the strain analysis. Bounding dimension values covering all reloads were

. selected for.the calculat' ions. On the basis of. previous experience, the

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calculations were'per^fo m d for pcwer history "D".

g.Thecriterionof1%maximumatE0Lissatisfied.

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XN-NF-84-28(NP), Rev. 1 Page 25 7.3 TRANSIENT CLADDING STRESSES AND STRAINS I

The stresses and strains during operating transients were evaluated with the RAMPEX code, on the basis of long term fuel conditions evaluated with the RODEX2 code.(12)

The benchmarking of the 1981 versions of these codes has deter-mined a failure threshold stress for ENC cladding. This threshold applies to the cladding hoop stress calculated at pellet ends assuming pellet hourglassing and a chip filling part of the pe.llet-to-clad gap in cold condition as a result of fuel handling. Using that chipping assumption, the stresses are evaluated at the beginning of each cycle, in order to obtain the maximum stress.

The RAMPEX code applies only to one axial location of the rod so that a complete analysis requires many RAMPEX runs. The power ramps are from 0% power to the BOL Fg limit (14.47 kW/f t) for the highest power I assembly, ~or to a level.that would be consistent with the highest powe'r assembly reaching the Fg limit. The ramp rates have been selected on tha basis of current plant ramp rates.

These results are within the design criteria limits, and by using .

the'most conservative dimensional inputs, as in the steady-state strain- -

analysis, they cover all.the reloads. .

7.4 CYCLING FATIGUE lhe cl' adding stresses calculated with RAMPEX at the b= ginning of each cycle were used to determine a stress amplitude during 'each type of transient. The frequency of occurrence of por:er changes was the same. as the duty cycles presented in Reference 1. The stres3_ amplitudes have bean' .

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I I XN-NF-84-28(NP), Rev. 1 Page 26 I

taken as such, since they already include local effects and the RAMPEX code predicts conservative stresses. Moreover, the calculated stress at the beginning of a cycle has been taken as such for the rest of the cycle without taking into account the fuel rod conditioning, which would relax the stresses.

The damage for each type of transient is determined by dividing the expected number of occurrences in each cycle by the allowed frequency for the corresponding stress amplitude and by accumulating the damage for each cycle. The allowed frequency is determined by conservative relations deduced from the fatigue curves of 0'Donnel and Langer.(16) There is substantial margin compared to the design limit of 0.67.

7.5 COLLAPSE The collapse calculation is done using the procedure described in I the extended burnup report, and since approved by the NRC.(52) RODEX2 is run first, assuming nominal pellet dimensions, nominal gap, minimum wall cladding, minimum fill gas pressure, fill gas absorption, and no gas release to determine the temperature and pressure conditions through'out the fuel rod lifetime, and to determine.the clad creepdown. These conditions 4 are used as input for'COLAPX. The COLAPX code is run with a conservative flux history. The code then predicts the time dependent creep ovality deformations in an-infinite length tube subjected to external pressure',

internal pressure, and linearly varying temperature gradients through the .

thickness of the cylinder.

If significant gaps are not allowed to form, then tube ovality, as predicted by~the COLAPX evaluation, cannot ' occur beyond the point of

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I XN-NF-84-28(NP), Rev. 1 I

Page 27 fuel support. The ENC fuel rod design uses an Inconel X-750 plenum spring to maintain an axial load on the pellet column well beyond the time when pellet densification is complete. This assists in the prevention of axial gaps. The pellet maximum resinter densification criteria also assures the presence of stable fuel so that the formation of significant gaps is prevented, and so that clad support is available during the life of the fuel.

In order to guard against the highly unlikely event that enough densification occurs to form pellet column gaps of sufficient size to allow clad flattening, an evaluation was performed. The cladding ovality increase was calculated with COLAPX, and the creepdown was calculated with RODEX2. The combined creepdown at the cladding minor axis was determined not to exceed the minimum level to allow the fuel column to relocate axially without the formation of axial gaps. I-7.6 -CORROSION AND HYDR 0 GEN PICKUP .

R00EX2 -includes the MATPRO corrosion and hydrogen pickup model.(12) -This model considers temperature, exposure time, irradiation enhancement, and original oxide film thickness as parameters. The most-oxide thickness formed in any of the power histories is_less than allowed.

- The maximum hydrogen level in either cladding for all of the power histories is well below the limit for hydrogen content.

7.7_ FUEL R0D ELONGATION The fuel rod elongation must be less'than the clearance between the tie plates. This was an'alyzed using the worst case tolerances for_ the fuel assembly and fuel rod, and using the maximum rod average f ast' fluence I

I I XN-NF-84-28(NP), Rev.1 Page 28

,I in the four power histories. The rod elongation model for fuel rods is based on ENC measured data for PWR rods. A design limit, which very conservatively bounds the data, is used.

The fuel assembly grows also from the f ast i~adiation damage to the guide tubes. However, because the guide tubes are innealed, the growth is not as rapid as the fuel rods. The MATPR0 irrad'.ation growth model(42)

I was used to compute tie plate separation growth. With conservative assumptions, the relative rod / assembly maximum elongation provides a minimum E0L clearance.

7.8 INTERNAL PRESSURE A RODEX2 analysis was performed to evaluate the end-of-life (E0L) internal fuel rod pressure for extended burnup. To prevent cladding instability, the rod internal pressure cannot exceed the system pressure or else the cladding may creep away from the pellet, which increases the fuel rod pellet temperatures. Higher fuel temperatures result in increased fission gas release, and therefore, higher internal rod pressures. The results of this analysis show the E0L internal rod pressure does not exceed the system pressure of 2250 psia. The-fuel rod will, therefore, remain stable throughout the expected power history.

. 7.9 R0D B0 WING Fuel rod bow is determined throughout the life of the fuel assembly so that the reactor operating thermal _ limits can be established.

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These limits -include the minimum critical heat flux ratio associated with protection against boiling transition and the maximum fuel- rod LHGR

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I XN-NF-84-28(NP), Rev. 1 I

Page 29 associated with protection of metal-water reaction, and peak cladding temperature limits for a postulated loss-of-coolant accident (LOCA).

The effect of rod bow on boiling transition and LHGR limits was evaluated to a peak average rod burnup of 49000 mwd /NTU. This evaluation was performed in accordance with the approved methodology. The evaluation of the impact on the most limiting DNB FSAR transient showed that, includ-ing the effects of rod bow, the DNBR is not reduced below the DNB limit.

L The evaluation of impact on limiting LHGR showed that the effects of rod bow are conservatively bounded by .the ECCS analysis and nuclear uncertainty factors. Therefore, no penalty for either effect due to rod bow need be included in operating limits for fuel operating at the 2.28 FT and 1.55 Fg limits to a burnup of 49000 mwd /MTM.

I 7.10 FUEL ROD PLENUM SPRING ,

The major functional requirements on the plenum spring are during shipments and during the densification phase 'of the fuel. Since both of these situations occur relatively early in the, life of the fuel, no reanalysis is required for extended burnup'. However, as the design of the -

rod end cap has been changed since Reload XN-5, the design of the spring has been reevaluated.

The spring force is proportional to the fuel stack weight for -

shipping, the spring never becomes solid, and it compresses the fuel-column during the densification phase. The results,. applicable to all reloads,.

show-that the plenum spring' serves its' intended purpose.  :

A new spring has been designed for Reload XN-7. That spring also satisfies all the d'esign criteria. .

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I XN-NF-84-28(NP), Rev. 1 Page 30 7.11 FUEL AND CLADDING TEMPERATURE As the peak LHGR is reached during the first cycles of the fuel life, no reanalysis should be required for the fuel and cladding tempera-ture at extended burnup. The fuel and cladding temperature were, however, reevaluated with RODEX2, and the results indicate that the design limits

, are not reached.

8.0 FUEL ASSEMBLY EVALUATION 8.1 GENERAL DESCRIPTION The fuel assemblies consist of a 14x14 array occupied by 179 fuel rods, 16 guide tubes, and one instrument tube. Seven Zircaloy-4 spacers with Inconel springs are positioned along the length of the assembly to locate the fuel rods and tubes, and are attached to the guide tubes by resistance spot welds. The guide tubes are mechanically attached to the upper and lower tie plates to form the structural skeleton of the fuel assembly.

8.2- DESIGN CRITERIA The mechanical design criteria for the fuel assembly are to provide for:

o Dimensional Compatibility o Differential Thermal Expansion and Irradiation ' Growth o Fuel Rod Support o Fuel Assembly Holddown o' Upper Tie Plate Removability-

.o Handling and Storage I

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I XN-NF-84-28(NP), Rev. 1 Page 31 Since the design of the fuel assembly structure is unchanged, I

only the irradiation growth and the fuel rod support behavior are affected by the extended burnup.

Specifically, the criteria require the design to provide adequate clearance between the tie plates to accommodate fuel rod growth, and adequate clearance between the fuel assembly and core plates to accomodate fuel assembly growth. The criteria for fuel rod support is to provide sufficient spring force at E0L to minimize flow-induced vibrations and to prevent fretting corrosion at the spacer-fuel rod contact points, consider-ing the effects of irradiation-induced spring force relaxation.

8.3 DESIGN ANALYSIS 8.3.1 Fuel Assembly Growth The fuel assembly growth must be accommodated by the core plate separation in cold conditions. This is more restrictive because of the differential thermal expansion between the core barrel and the fuel -

assembly. The MATPRO growth model(42) is used to calculate the fuel assembly growth using the worst case dimensions,'and conservatively using the peak rod average fast fluence.

The MATPRO correlation is compared with ENC PWR fuel

-assembly growth data for four designs with identical guide tube character-istics. The designs. vary'in' guide tube length, number of tubes per assembly, and holddown force. Growth results for thel lowest stressed design are greater than predicted by the MATPRO model. Growth of the three .

more highly loaded designs is less than the MATPRO curves with less growth

. occurring as the stress is increased. The nominal hot BOL holddown stress I

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i XN-NF-84-28(NP), Rev. 1 Page 32 _=

in the Kewaunee fuel is comparable to the stress of other fuel assemblies, 9 l which were measured to have less growth than the MATPRO model predicted. I" Using MATPR0 and conservative assumptions listed above, the maximum fuel assembly growth, at a burnup of 49 GWd/MTV, gives a fuel -;

assembly / core plate clearan'ce. _

' 8.3.2 Spacer Spring Relaxation

{

The Inconel spacer springs are known to relax during irradiation and the fuel rod cladding tends to creepdown. Together, these K

two characteristics combine to reduce the spacer spring force on a fuel rod b during its lifetime. These characteristics have been considered in the design of the spring to assure an adequate holding force when the assembly ~

has completed its design operating life. -:

Spacer spring relaxation and rod creepdown characteristics I have been monitored in relation to burnup on both BWR and PWR fuel rods by measuring the force required to pull a fuel rod through a spacer. Data _f have been obtained on fuel rods of several . reactor types, including ENC I 15x15 rods for Westinghouse reactors, which have attained an assembly

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burnup of 47700 mwd /MTU. Inspection of the 15x15 rods showed no evidence of significant fretting or wear damage at the contact points.

The spacer spring relaxation, based on this and other I data, follow an~ asymptotic relationship with burnup. For the rod and spacer spring type incorporated in Kewaunee, the average spring force at-47700 mwd /MTU is adequate to prevent fretting wear.

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XN-NF-84-28(NP), Rev. 1 Page 33 I

9.0 REFERENCES

I 1. Generic Mechanical and Thermal Hydraulic Design, XN-NF-78-34(NP),

Jan. 1979, for Exxon Nuclear 14x14 Reload Assemblies with Zircaloy Guide Tubes for Westinghouse 2 Loop Pressurized Water Reactors.

2. Not Used.

3. ASME Boiler and Pressure Vessel Code,Section III,1971 Edition, ASME, New York, NY.

s I 4. H. S. Rosenbaum, "The Interaction of Iodine with Zr-2",

Electrochemical Technology, . Volume 4, Number 3-4 (March-April 1966).

5. A. Garlick, Stress Corrosion Cracking of Zirconium Alloys in Iodine Vapour, British Energy Conference, London, July 1971.

6. R. A. Lorenz, J. L. Collings, S. R. Manning, Fission Product Release From Simulated LWR Fuel, NUREG/CR-0274, ORNL/NUREG/

TM M4, October 1978.

7. .1 . Peels, H. Stehle, and E. Steinberg, Out-of-Pile Testing of Iodine Stress Corrosion Cracking in Zircaloy Tubing in Relation I to the PCI-Phenomenon, Fourth Internationai Conference on Zirconium in the Nuclear Industry, 1979.

8. Nukleare Sicherheit, Halbjahresbericht 1977/1, KFK 2500, Kern-forschungszentrum, Karlsruhe, December 19//.

9. A. K. Miller, et al, Zircaloy Cladding Deformation and Fracture Analysis EPRI NP-856, August 1978.

10. Stress Corrosion Cracking of Zircaloys, SRI International, EPRI NP-1329, March 1980.

11. EPRI-NASA Cooperative Project on Stress Corrosion Cracking of Zircaloy, SRI International, EPRI NP-717, March 1978.

12. K. R. Merckx, RODEX2 Fuel Rod Thermal-Mechanical Response Evaluation Model, XN-NF-81-58(NP), November 1981.

13. Not Used.

14. A. A. Bauer, L. M. Lowry, and J. S. Perrin, Process on Evaluating-LI ' Strength and Ductility of Irradiated Zircaloy During July through Teptember 1975. BMI-1938, September 1975.

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XN-NF-84-28(NP), Rev. 1 Page 34

15. A. A. Bauer, L. M. Lowry, W. J. Gallaugher, and A. J. Markworth, Evaluating Strength and Ductility of Irradiated Zircaloy -

Quarterly Progress Report January throuah March 1978, NUREG/CR-0085, BMI-2000, June 1978.

16. W. J. O'Donnel and B. F. Langer, " Fatigue Design Bases for g Zircaloy Components," Nuclear Science and Enaineering, Volume 20, January 1964.

g

17. Not Used.

18. Not Used.

19. Not Used.

20. H. W. Wilson, K. K. Yoon, and D. L. Baty, "The Effect of Fuel Rod Design on SCC Susceptibility," ANS Liaht Water Reactor Fuel Performance Conferences, Portland, OR, April 29-May 3, 1979.

21. Not Used.

22. Not Used.

23. Not Used.

24. Not Used.

25. Not Used.

26. Not Used.

27. Not Used.

28. Not Used.

29. E. Duncombe, Westinghouse (USA) Report, WAPD-TM-794, (1968).

30. .H. R. Higgy and F. H. Hammad, "Effect of Neutron Irradiation on the Tensile Properties of Zircaloy-2 and Zircaloy-4," J. of Nuclear Materials, 44, August 1972.

31. B. Watkins and D. S. Wood, The Significance of Irradiation Induced Creep on Reactor Performance of a Zircaloy-2 Pressure Tube, Applications Related Phenomena for Zirconium and Its l.

Alloys, ASTM-STP-458, American Society for Testing and Materials, W 1969, pp. 226-240.

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I XN-NF-84-28(NP), Rev. 1 Page 35

32. D. Lee, C. F. Shih, F. Zaverl, Jr. and M. D. German, Plastic

.I Theories and Structure Analysis of Anisotropic Metals -Zircaloys, EPRI NP-500, RP 456-2, May 1977.

I 33. D. G. Hardy, High Temperature Expansion and Rupture Behavior of Zircaloy Tubing, CONF-730304 USAEC/ TIC, Water Reactor Safety, March 26-28, 1973, p. 254-273.

34. D. A. Powers, R. D. Meyer, Cladding Swelling and Rupture Models for LOCA Analyses, NUREG-0630, 1981.

35. E. Duncombe, F. A. Nichols, S. H. Leiden and W. F. Bourgeois, Prediction of the In-Reactor Deformation of Zircaloy Cladding for Oxide Fuel Rods, WAPD-TM-80C(L), December 1969.

36. R. V. Hesketh, J. E. Harbottle, N. A. Waterman and R. C. Lof f,

" Irradiation Growth and Creep in Zircaloy-2," Radiation Damage in I Reactor Materials, Volume 1, Proc. of Vienna Synposium, IAEA-SM-120/D-3, 1969.

37. E. R. Gilbert, "In-Reactor Creep of Reactor Materials," Reactor I 38.

Technology, Volume 14, 1971.

R. C. Daniel, "In-Pile Dimensional Changes of Zircaloy-4 Tubing I Having Low Hoop Stresses," Nuclear Technology, Volume 14, May 1972.

J. Pankaskie, Irradiation Effects on the Mechanical I 39. P. -

Properties of Zirconium and Dilute Zirconium Alloys, BN-5A-618, July 1976.

40. R. V. Hesketh, "Non-Linear Growth in Zircaloy-4," Journal of

Nuclear Materials, 30, (1969), Pages 219-221.

41. Not Used.

42. D. L. Hagrman, G. A. Reymann and R. E. Mason, "MATPRO Version II (Revision 2), A Handbook of Materials Properties for Use-in the I Analysis of Light Water Reactor Fuel Rod Behavior, NUREG/CR-0497 TREE-1280, Rev. 2, August 1981. ,

I 43. J. B. Conway and R. M. Fincel, "The Thermal Expansion and Heat Capacity of UO2 to 20000C," Trans. Am. Nucl. Soc., 6, June 1963.

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I I 44. Lyons, et al, U02 Properties Affectin Engineering Design, 21, pp 184-185 (1972)g Performance," Nuclear lI I .

I XN-NF-84-28(NP), Rev. 1 I

Page 36

45. J. A. Christensen, et al, " Melting Point of Irradiated U02 "

WCAP-6065, February 1965.

46. ANSYS - Engineering Analysis System Theoretical Manual, P. C.

Kohnke, 1977. ANSYS - User's Guide, 1979. Swanson Analysis System, Houston, PA.

47. Not Used.

48. Not Used.

49. Not Used.

50. Not Used.

51. Not Used.

52. Safety Evaluation by the Office of Nuclear Reactor Regulation

'Related to Amendment No. 54 to Facility Operating License No.

OPR Wisconsin Power and Light Company, Madison Gas and Electric Company, Kewaunee Nuclear Power Plant, Docket No.

50-305.

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l XN-NF-84-28 (NP), Rev.1 Issue Date: 2/11/85 1

I MECHANICAL DESIGN REPORT SUPPLEMENT FOR KEWAUNEE HIGH BURNUP (49 GWd/MTU) FUEL ASSEMBLIES 1

I DISTRIBUTION CA Brown

! JC Chandler

!E RA Copeland iE NE Hoppe

Customer /LC O'Malley (20)

Document Control (5) 4

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