ML060960602

2006/04/07-Oyster Creek, License Renewal AMP-AMR Audit Questions AMP-210 Set 3

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D. Ashley_- RE: AuditpQ & A (Question Numbers AMP-141, 210, 356) Page 1 From: <George.Beck@exeloncorp.com>

To: <djal @nrc.gov>, <rkm~nrc.gov>

Date: 04/05/2006 5:15:19 PM

Subject:

RE: Audit Q & A (Question Numbers AMP-141, 210, 356)Attached is third part of AMP-21 0. (third of three)<<PagEzs from AMP-210-2.pdf>>-O-----Original Message-----

> From Beck, George> Sent: Wednesday, April 05, 2006 5:02 PM> To: Donnie Ashley (E-mail);

'Roy Mathew (E-mail)' (E-mail)> Cc: Ouaou, Ahmed; Hufnagel Jr, John G; Warfel Sr, Donald B; Polaski, Frederick W>

Subject:

FW: Audit Q & A (Question Numbers AMP-141, 210, 356)> Note: As originally transmitted this email was undeliverable to the NRC; it exceeded the size limit. It is being retransmitted without the AMP-210.pdf.

This file will be reconstituted and sent in smaller ".pdf"s; the first 11 pages are attached.> Georce-------Original Message-----

> From: Beck, George> Sent: Wednesday, April 05, 2006 4:39 PM> To: Donnie Ashley (E-mail);

'Roy Mathew (E-mail) '(E-mail)> Cc: Ouaou, Ahmed; Hufnagel Jr, John G; Warfel Sr, Donald B; Polaski, Frederick W>

Subject:

Audit 0 & A (Question Numbers AMP-141, 210, 356)> DonniV/Roy,> Attached are the responses to AMP-21 0 and AMP-356 in an updated version of the reports from the AMP/AMR Audit database.

Also included is a revised version of AMP-141. These answers have been reviewed and approved by Technical Lead, Don Warfel.> Regarding AMP-210, please note:> As pointed out in our response to NRC Question AMP-210, (8a)(1), "The 0.806" minimum average thickness verbally discussed with the Staff during the AMP audit was recorded in location 19A in 1994.Additional reviews after the audit noted that lower minimum average thickness values were recorded at the same location in 1991 (0.803") and in September 1992 (0.800").

However, the three values are within the tolerance of +/- 0.010" discussed with the Staff."> Regarding AMP-141, please note:> Our re3ponse to AMP-141 has been revised to reflect additional information developed during the ongoing preparation of RAI responses.

> Please let John Hufnagel or me know if you have any questions.

> George> < F:ile: Pages from AMP-210-2.pdf

>>

D. Ashley -RE: Audit Q & A (Question Numbers AMP-141, 210, 356) Page 2> <c File: Pages from AMP-210.pdf

>>> << File: AMP-141.pdf

>>> << File: AMP-356.pdf

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This e-mail is intended solely for the use of the individual or entity to which it is addressed.

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Creation Date: From: Created By: RE: Audit Q & A (Question Numbers AMP-141, 210, 356)04/05/2006 5:14:15 PM<George.Beck@exeloncorp.com>

George.Beck@exeloncorp.com Recipients nrc.gov OWG-WPO01.HQGWDO01 DJA1 (D. Ashley)nrc.gov TWGWPOO1.HQGWDOO1 RKM (Roy Mathew)exeloncorp.com fred.polaski CC donald.warfel CC john.hufnagel CC ahmed.ouaou CC Post Office OWGVPO010.HQGWDO01 TWGAVPOO1.HQGWDOO1 Files MESSAGE TEXT.htm Pages from AMP-210-2.pdf Mime.:322 Options Expiration Date: Priority: Reply Requested:

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4 S-A OYSTER CREEK NUCLEAR LICENSING REGULATORY CORRESPONDENCE'DISTRIBUTION'SHEET ii OC FILE NO: E93 6 // CORRIESPONDENCE NO: Loor,- ?9 c32/- 9.z _ 2093-) _O~? z LETTER DATE: -9Z Avie ^ W...1. X 0J14 'Y /, /8,>Ws

SUBJECT:

_eysleg Cx4ee/A_ D 4 ey // 6e ;-oO -COGNIZANT Ll'CENSING ENGR: 5>. Z4A , AC1 AI,4?,ION ASSIGNED TO: DISTRIBUTIONI:

TO ED&CC ON: ED&CC DIST'D ON: .' .91a2 ..I ED&CC FILINGi INSTRUCTIONS:

GENERAL (2) = W/O 4,~tcjfA4S 4' se .2-S4c C/ e/'J A'J : 4 4.'A 1 411,7': '?ev1.riaovs Aos 2eAeA n1 00V, !~1 5`Gs /9 f/ 4; 30---- Z- A .1 -.--L -------I- -OFFICE OF THE PRESIDENT NUCLEAR ASSURANCE'" TECH FUNCTIONS P.J.M.K.Clark Wilson A. Kunz Bromery Bldg E Bldg E Bldg 12/FR Bldg 12/FR X 'P.T.X .C.X M.R.3J.R.S.B.X G.X P.X R.G.Fiedler Murphy Mascari Slobodien Tilton Kowalski Markowski Rice Alatary.Cropper Thompson Sullivan Young Bldg E' Bldg E ,Bldg -3 SEB Bldg'12/FR Bldg E Bldg E Bldg E Bldg 12/FR SEB Bldg 12/FR Bldg'12/FR' J.D.B.--J.W.J.Devine'Slear ,Behrle ,Camire Popow ', ,cVAW Bldg F @" Bldg F6 Trl 250'SEB: Bldg F 9 ,.. ) 1/d -'C ... .OYSTER CREEK_/. Coi,/z-..N. yeV-1 J. Barton S. Levin R. Barrett R. Brown N. Chrissotitnos K. Bass M. Budaj P. Cervenka/TSCR J. DeBlasio L. Lammers D. Ranft W. Stewart T. Quintenz T. Dempsey P. Scallon W. Garvey R. Baran R. Booth MOB MOB MOB MOB MOB PEB MOB PEB PEB NMB PEB PEB NMB PEB MOB MOB MOB PEB ADMIN & FINANCE, D. Myers R.C. Cutler Bldg-Bldg F.E X CORPORATE SERVICES R. L. Long Bldg E J. Knubel .Bldg E (M. Laggart Bldg E (K)G. Busch SEB X R. Rogan TMI E. O'Donnell Bldg E C. Gomulka (3) SEB X SITE SERVICES I. Finfrock Bldg E T. Brownridge Trl 300 EXTERNAL DISTRIBUTION NJBNE/K. Tosch SPPT/Bl ake INPO ANI/R. Oliveira BPU/R. Chilton BPU/T.'Gould

'GE/J. Miller..' X X..X_INDEPENDENT FEVIEW OTHER M. Yacabonis M. Fillipone ED&CC B.. Moroney, NSCC(HQ MGO PPY X J. Sullivan E. Roessler Bldg E X Trl 256 X 270 X S. C. 7 Z4tfl wetd~11: Fg WIP51/OISTGEN U GPU Nuclear Corpormtdon E Nu leU0 ar One Upper Pond Road 1 ~Pan§ppany, New Jersey 0:7054-201-316-7000 TELEX 138-482 Wrltera Direct Dil Nunber January 16, 1992 5000-92-2093 C321-92-2008 U. S. Nuclear Regulatory Commission Attn: Document Control Desk Washington, DC 20555 Gentlemen:

Subject:

Oyster Creek Nuclear Generating Station (OCNGS)Docket No. 50-219 -Facility Operating License No. DPR-16 Oyster Creek Drywell Containment

References:

(1) GPU Nuclear Letter dated 12/5/90.-

Oyster Creek Drywell Stress and Stability Analyses (With Sand), GE Reports Index No. 9-1 and 9-2.(2) GPU Nuclear Letter dated 3/4/91 -Oyster.Creek Drywell Stress and Stability Analyses (Without Sand), GE Reports Index No. 9-3 and 9-4.The referenced letters provided GPU Nuclear's ASME Section VIII evaluations of the Oyster Creek Drywell for with and without sand stability analyses, GE Report Indices 9-2 and 9-4. This letter provides you with Revision 1 to these evaluations.

This revision incorporates changes resulting from an internal audit which identified an error in calculating the 7rc factor,-see Figure 2-3 of.Reports 9-2 and 9-4. The calculated stress assumed a cylindrical shape rather than the more appropriate spherical shape. The calculated capacities are still in compliance with all required ASME Code provisions, however, the margins beyond those capacities are reduced. The revisions to the effected pages are indicated by vertical lines in the right hand margin.321920C8.

LET GPU Nuclear Corporation is a subsidiary of General Public Utilities Corporation r C321 2008 Page 2 If you have any questions or comments on this submittal or the overall drywell corrosion program,.please contact Mr. Michael Laggart, Manager, Corporate Nuclear Licensing at (201) 316-7968.Very truly yours,+ .C. DeVine Vice President and Director Technical Functions JCD/RZ/plp cc: Administrator, Region 1 Senior Resident Inspector Dyster Creek NRC Project Manager 32192003.

LET RF# 00664 INDEX 9-2, REV. 0 AN ASME SECTION VIII EVALUATION OF THE OYSTER CREEK DRYWELL PART 2 STABILITY ANALYSIS November 1990 prepared for GPU Nuclear Corporation Parsippany, New Jersey prepared by GE Nuclear Energy San Jose, California RF# 00664 NOEX 9-2, REV. 0 AN ASME SECTION VIII EVALUATION OF THE OYSTER CREEK DRYWELL PART 2 STABILITY ANALYSIS Prepared by: e dA C.D. Frederickson, Senior Engineer Materials Monitoring

&Structural Analysis Services Reviewed by: H..S. Mehta, Principal Engineer Materials Monitoring

&Structural Analysis Services Approved by:: _ _S. Ranganath, Manager Materials Monitoring

&Structural Analysis Services i NE RE:V. 1 TABLE OF CONTENTS.ae 1. INTRODUCTION 1-1 1.1 General 1-1 1.2 Report Outline 1-1 1.3 References 1-1 2. BUCKLING ANALYSIS METHODOLOGY 2-1 2.1 Basic Approach 2-1 2.2 Determination of Capacity Reduction Factor 2-2 2.3 Modification of Capacity Reduction Factor for 2-3 Hoop Stress 2.4 Determination of Plasticity Reduction Factor 2-5 2.5 References 2-5 3. FINITE ELEMENT MODELING AND ANALYSIS 3-1 3.1 Finite Element Buckling Analysis Methodology 3-1 3.2 Finite Element Model 3-2 3.3 Drywell Materials 3-3 3.4 Boundary Conditions 3-3 3.5 Loads 3-4 3.6 Stress Results 3-7 3.7 Theoretical Elastic Buckling Stress Results 3-9 3.8 References 10 4. ALLOWABLE BUCKLING STRESS EVALUATION 4-1 5.

SUMMARY

AND CONCLUSIONS 5-1 iii I r? ORF 8X64 REIlN 0 I 1 li il t A.II~. FEE HUAI'-, REV. 1 LIST OF FIGURES Figure No.Title -Page Nbo.1-2 2- 8 1-1 Drywell Configuration 2-1 Capacity Reduction Factors for Local Buckling of Stiffened and Unstiffened Spherical Shells 2-2 Experimental Data Showing Increase in Compressive Buckling Stress Due to Internal Pressure (Reference 2-6)2-3 Design Curve to Account for Increase in Compressive Buckling Stress Due to Internal Pressure (Reference 2-1.1)2-4 Plasticity Reduction Factors for Inelastic Buckling 3-1 Oyster Creek Drywell Geometry 3-2 Oyster Creek Drywell 3-D Finite Element Model 3-3 Closeup of Lower Drywell Section of FEM (Outside View)3-4 Closeup of Lower Drywell Section of FEM (Inside View)3-5 Boundary Conditions of Finite Element Model 3-6 Application of Loading to Simulate Seismic Bending 3-7 Meridional Stresses -Refueling Case 3-8 Lower Drywell Meridional Stresses -Refueling Case 3-9 Circumferential Stresses -Refueling Case 2-9 2-10 a.2-11 3-21 3-22 3-23 3-24 3-25 3-:26 3-27 3-28 3-29 v U ,YRF# XO624 NDEX 9-, REV. 1 LIST OF FIGURES Figure Paige.No. Title No.3-10 Lower Drywell Circumferential Stresses -'Refueling Case 3-30 3-11 Meridional Stresses -Post-Accident Case 3-31 3-12 Lower Drywell Meridional Stresses -Post-Accident Case 3-32 3-13 Circumferential Stresses -Post-Accident Case '3-33 3-14 Lower Drywell Circumferential Stresses -Post-Accident 3-34 Case 3-15 Symmetric and Anti-Symmetric Buckling Modes 3-35 3-16 Symmetric Buckling Mode Shape -Refueling7 Case 3-36 3-17 Anti-Symmetric Buckling Mode Shape -Refueling Case 3-37 3-18 Buckling Mode Shape -Post-Accident Case '3-38 vi.

?NDEX 9-2, REV. 0 1. INTRODUCTION 1.1 General To address local wall thinning of the Oyster Creek drywell, GPUN his planned to prepare a supplementary report to the Code stress report of record [1-1]. For convenience, the supplementary report is divideid into two parts. Part I of the supplementary report [1-2] includes all of the Code stress analysis results other than the buckling capability for the drywell shell. This report addresses the buckling capability of the drywell shell shown in Figure 1-1 and constitutes the second part of the supplementary report. Buckling of the' entire drywe'll shell is considered in thts analysis with the sandbed region being the area of primary concern.1.2 Report Outline Section 2 of this report outlines the methodology-lused in the buckling capability evaluation.

Finite element modeling, analysis and results are described in section 3. Evaluation of the allowable compressive buckling stresses and comparisons with the calculated compressive stresses for the limiting load combinations are covered in section 4.Section 5 presents the summary of results.and conclusions.

1.3 References 1-1 "Structural Design of the Pressure Suppression Containment Vessels," by Chicago Bridge & Iron Co.',Contract

1. 9-0971, 1965.1-2 'An ASME Section VIII Evaluation of the Oyster Creek Drywel-l," GE Report No. 9-1, DRF# 00664, November 1990, prepared for GPUN.1-1 INDEX 9-2 REV. 0 Mt& a____ro IAvvg --Io 3i-E I..,. is I- 4-- -,a- -9Lc-ll-jtl-A 0 I 44 45: , -- Irz E. % V., 4 -N*Ok -, I Figure 1-1. Drywell Configuration 1-2 N ' N 9- REV. 1 2. BUCKLING-ANALYSIS METHODOLOGY 2.1 Basic Approach The basic approach used in the buckling evaluation follows the methodology outlined in the ASME Code Case N-284 [References 2-1, 2-2]. Following the procedure of this Code Case, the allowable I compressive stress is evaluated in three steps.In the first step, a theoretical elastic buckling stress, aie, is determined.

This value may be calculated either by classical buckling equations or by finite element analysis.

Since the drywell shell geometry is complex, a three dimensional finite element analysis'approach is followed using the eigenvalue extraction technique.

Mare details on the eigenvalue determination are-given in'Section 3.In the second step, the theoretical elastic buckling stress is modified by the appropriate capacity and plasticity reduction factors.The capacity reduction factor, ar, accounts for the difference between classical buckling theory' and actual tested buckling stresses for fabricated shells. This difference is due to imperfections inherent in fabricated shells, not accounted for in classical buckling theory,'which can cause significant reductions in the critical buckling stress. Thus, the elastic buckling stress for fabricated shells is given by the product of the theoretical elastic buckling stress and the capacity reduction factor, i.e., aeai. When the elastic buckling stress exceeds the proportional limit. of the material, a plasticity reduction factor, il, is used to account for non-linear material behavior.

The inelastic buckling stress' for fabricated shells is given by 7iaiaie.In the final step, the allowable compressive stress is obtained by dividing the buckling stress calculated in the second step by the safety factor, FS: Allowable Compressive Stress -nielie/FS 2-1 RF# 0064 YNDEX 9-2, REV. 1 In Reference 2-1, the safety factor for the Design and Level A & B service conditions is specified as 2.0. A safety factor of 1.67 is specified for Level C service conditions (such as the post-accident flooded condition).

The Determination of appropriate values.for capacity and plasticity reduction factors is discussed next.2.2 Determination of Capacity Reduction Factor.The capacity reduction factor, oi, is used to account for reductions in actual buckling strength due to the existence of geometric imperfections.

The capacity reduction factors given in Reference 2-1 are based on extensive-data compiled by Miller [2-3]. The factors appropriate for a spherical shell geometry such as that of the drywell in the sandbed region, are shown in Figure 2-1 (Figure .1512-1 of Reference 2-1). The tail (flat) end of the curves are used for unstiffened shells. The curve marked 'Uniaxial compression' is applicable since the stress state in the sandbed.region is compressive in the meridional direction but tensile in the circumferential direction.

From this curve, Ad is determined to be 0.207.The preceding value of the capacity reduction factor is. very conservative for two reasons. First, it is based on the assumption that the spherical shell has a uniform thickness equal to the reduced thickness.

However, the drywell shell has a greater thickness above the sandbed region which would reinforce the sandbed region. Second, it is assumed that the circumferential stress is zero. The tensile circumferential stress has the effect of rounding the shell arid reducing the effect of imperfections introduced during the fabrication and construction phase. *A modification of the a; value to account for the presence of tensile circumferential stress is discussed -in Subsection 2.3.The capacity reduction factor values given in Reference 2-1 are applicable to shells which meet the tolerance requirements of NE-4220 2-2 TAYREXY 8 REV. 1 of Section III [2-4]. Appendix A of Reference 2-5 compares the tolerance requirements of NE-4220.to the requirements, to which the Oyster Creek drywell shell was fabricated.

The comparison shows that the Oyster Creek drywell shell was erected to the .tolerance requirements of NE-4220. Therefore, although the Oyster Creek drywell is-not a Section III, NE vessel, -it is justified to use the approach outlined in Code Case N-284.2.3 Modification of Capacity Reduction Factor for Hoop Stress The orthogonal tensile stress has the effect of rounding fabricated shells and reducing the effect. of imperfections on the buckling'strength.

The Code Case N-284 [2-1 and 2-2] notes in the last paragraph of Article 1500 that, "The influence of internal pressure on a shell structure may reduce the initial imperfections and therefore higher values of capacity reduction factors may' be acceptable.

Justification for higher values of a, must be.-.given in the Design report." The effect of hoop tensile stress on the buckling strength of cylinders has been extensivelly documented

[2-6 through 2-11].. Since the methods used in accounting for the effect of tensile hoop stress for the cylinders and spheres are similar, the test data and the methods for the cylinders are first reviewed.

Harris, et al [2-6]presented a comprehensive set of test data, including those from References 2-7 and 2-8, which clearly showed-that internal pressure in the form of hoop tension, increases the. axial buckling stress of cylinders.

Figure 2-2 shows a plot of the test data showing the increase in buckling stress as a function of nondimensional pressure.This increase in buckling capacity is accounted for by. defining a separate reduction factor, a. The capacity reduction factor ai can then be modified as follows: ai,mod P i C.p 2-3 O RF#06 4 I NDEX'S'T, REV. 1 The buckling stress in uniaxial compression for a cylinder-or a sphere of uniform thickness with no internal pressure is given by the following:

Sc -(0.605)(ctj)Et/R

-(0.605)(0.207)

Et/R Where, 0.605 is a constant, 0.207 is the capacity reduction factorcti, and .E,t and R are Young's Modulus, wall thickness and- radius, respectively.

In the presence of a tensile stress such as that produced by an internal pressure, the buckling stress is given as follows: Sc,mod ' (0.605)(a 1 + ep)Et/R-(0.605)(0.207

+ tp)Et/R= [(0.605)(0.207)

+ AC]. Et/R Where AC is ap/0.605 and is given for cylindrical geometries in the graphical form in Figure 2-3. As can be seen in Figure 2-3, AC is a function of the parameter X-(p/4E)(2R/t) 2 , where ,ps.is the internal pressure.

Miller [2-12] gives the following equation that fits the graphical relationship between X and AC shown in Figure 2-3: 'AC -tp/0.605 1.25/(5+1/X)

The preceding approach pertains to cylinders.

Along. the similar lines, Miller [2-13] has developed an approach for spheres as described next.The non-dimensional parameter X is essentially (ae/E)(R/t).

Since in the case of a sphere, the hoop stress is one-half of that in the cylinder, the parameter X is redefined for spheres as follows: X(sphere) (p/8E)(2R/t) 2 2-4 UEX06 .RE:V. 1 When the tensile stress magnitude, S, is known, the equivalent internal pressure can be calculated using the expression:

p

• 2tS/R Based on a review, of spherical shell buckling data [2-14, 2-1a), Miller [2-13] proposed the following equation for AC: AC(sphere) 1.06/(3.24

+ 1/X)The modified capacity reduction factor, -i mod' for the drywell geometry was obtained as follows:&i,mod -0.207 + AC(sphere)/0 6 O 5 2.4 Determination of Plasticity Reduction Factor When the elastic buckling stress exceeds the proportional limit of the material, a plasticity reduction factor, 1j, is used to account for the non-linear material behavior.

The inelastic buckling stress for fabricated shells is given by j 7 iaiaie Reference 2-2 gives the.mathematical expressions shown below [Article -1611 (a)] to calculate the plasticity reduction factor for the meridional direction elastic buckling stress. A is equal to ajoie/ay and ay is the material yield strength.

Figure 2-4 shows the relationship in graphical form.1- 1.0 if A S 0.55-(0.45/A) + 0.18 if 0.55 <.A 1.6-1.31/(1+1.15A) if 1.6 < A.S 6.25-1/A if A > 6.25 2.5 References 2-1 ASME Boiler and Pressure Vessel Code Case N-284, "Metal Containment Shell Buckling Design Methods,Section III, Division 1, Class MC", Approved August 25, 1980.2-5 U TIND:X0T66 REV. 1 2-2 Letter (1985) from C.D. Miller to P. Raju;

Subject:

Recommended Revisions to ASME Code Case N-284.2-3 Miller, C.D., "Commentary on the Metal Containment Shell Buckling Design Methods of -the ASME Boiler and Pressure Vessel Code," December 1979.2-4 ASME Boiler & Pressure Vessel Code,Section III, Nuclear Power Plant Components.

2-5 "Justification for Use of Section III, Subsection NE, Guidance in Evaluating the Oyster Creek Drywell,".

Appendix A to .letter. dated December 21, 1990 from H.S. Mehta of GE to S.C. Tumminelli of GPUN.2-6 Harris, L.A., et al, "The Stability of Thin-Walled Unstiffened Circular Cylinders Under Axial Compression Including the Effects of Internal Pressure," Journal of the Aeronautical Sciences, Vol.24, No. 8 (August 1957), pp. 587-596.2-7 Lo, H., Crate, H., and Schwartz, E.B., "Buckling of Thin-Walled Cylinder Under Axial Compression and Internal Pressure,".

NACA.TN 2021, January 1950.2-8 Fung, Y.C., and Sechler, E.E., "Buckling of Thin-Walled Circular Cylinders Under Axial Compression and Internal Pressure," Journal of the Aeronautical Sciences, Vol. 24, No. 5, pp. 351-356, May 1957.2-9 Baker, E.H., et al., "Shell Analysis Manual," NASA, CR-912'(April 1968).2-10 Bushnell, D., "Computerized Buckling Analysis of Shells," Kluwer Academic Publishers, 1989 (Chapter 5).2-11 Johnson, B.G., "Guide to Stability Design Criteria for Metal Structures," Third Edition (1976), John Wiley & Sons.. .Z-b RF# 086624 4 V1 T NDEX -, .REV1 2-12 Miller, C.D., 'Effects of Internal Pressure on Axial Compression Strength of Cylinders," CBI Technical Report No. 022891, Februxry 1991.2-13 Miller, C.D., "Evaluation of Stability Analysis Methods Used for the Oyster Creek Drywell," CBI Technical Report Prepared for GPU Nuclear Corporation, September 1991.2-14 Odland, J., "Theoretical and Experimental Buckling Loads. of Imperfect Spherical Shell Segments," Journal of. Ship Research, Vol. 25, No.3, September 1981, pp. 201-218.2-15 Yao, J.C., "Buckling of a Truncated Hemisphere Under -Axial, Tension," AIAA Journal, Vol. 1, No. 10, October 1963, pp.2316-2319.

2-7 q RDEX T9, REV. 1 OA as 02 OA a 4 12 ' 24 Aia 1 tNV Figure 2-1 Capacity Reduction Factors for Local Buckling of Stiffened and Unstiffened Spherical Shells I 2 RNDEX 8-2, REV. 1 10 WUCY(NAA).

FUNrG a *HE LO, -CRATE S SCHWARTZ-,-*~

NAA bI~cs r TiE t 0-.,I-. I :7-71.-0. ...6-90% PROBAJBILITY CUR -in7 ' 14-.--W I --t-7 7. -I' 'I -, i -I I I I.01 JO LO 10 102 f- (t!,)Figure 2-2 Experimental Data Showing Increase in Compressive Buckling Stress Due to Internal Pressure.(Reference 2-6)2-9 RNEXg-" REV. 1 I o a 6 4 z-, , ..-.I ..i i 1.I -.1.0 I a aC 4 a 0.90 a 64 4 a_ .I ..Wc _ _ _ IICICI_ _ I I I AW_______.I

_ ff I I 6_ _ _ _ _ _J III_.0.0 0.01 C..a OJO.t 4 6 a 4t l 'a 4 6 4 1.0 10 Figure 2-3 Design Curve to Account for Increase in Compressive Buckling Stress Due to Internal Pressure (Reference 2-11)2-10 I RF# 06g E4 YNDEX ~- , REV. 1 7.X& a aq,- Ml /%wa ---4 Figure 2-4 Plasticity Reduction Factors for Inelastic Buckling..I 2-11:

TNDEX 9-2, RE'V. 0 3. FINITE ELEMENT MODELING AND ANALYSIS 3.1 Finite Element Buckling Analysis Methodology This evaluation of the Oyster Creek Drywell buckling capability uses the Finite Element Analysis (FEA) program ANSYS [Reference 3-I]. The ANSYS program uses a two step eigenvalue formulation procedure to perform linear elastic buckling Analysis.

The first step is a static analysis of the structure with all anticipated loads applied. The structural stiffness matrix, [K], the stress stiffness matrix, [S], and the applied stresses, nap' are developed and saved from this static analysis.

A buckling pass is then run to solve for the eigenvalue or load factor, A, for which elastic buckling is predicted using the equation: ([K] + X [S]) (u) -0 where: A is the eigenvalue or load factor.(u) is the eigenvector representing the buckled shape of the structure.

This load factor is a multiplier for the applied stress state at which the onset of elastic buckling will theoretically occur.. All applied loads (pressures, forces, gravity, etc...) are scaled equally.'

For example, a load factor of 4 would indicate that the structure would buckle for a load condition four times that defined in the stress pass. The critical stress, acr, at a certain location of the structure is thus calculated as: Ccr ' ap This theoretical elastic buckling stress is then modified by the capacity and plasticity reduction factors to determine the predicted buckling stress of the fabricated structure as discussed in Section 2.This stress is further reduced by a factor of safety to determine the allowable compressive stress.3-1I RF# 00664 YNDEX 9-2, REV. 0.3.2 Finite Element Model The Oyster Creek drywell has been previously analyzed using a simplified axisymmetric model to evaluate the buckling capability in the sandbed region [Reference 3-2]. This type of analysis conservatively neglects the vents and reinforcements around the vents which significantly increase the stiffness of the shell near the sandbed region. In order to more accurately determine the buckling capability of the drywell, a three dimensional finite element model is developed.

The geometry of the Oyster Creek drywell is shown in Figure 3-1.Taking advantage of symmetry of the drywell with 10 vents, a 36 section is modeled. Figure 3-2 illustrates the finite element model of the drywell. This model includes the'drywell shell from the base of the sandbed region to the top of the elliptical head-and the'vent and vent header. The torus is not'included in this model because the bellows. provide a very flexible connection which does not allow significant structural interaction between the drywell and torus.Figure 3-3 shows a more detailed view of the lower section of the drywell model. The various colors on Figures 3-2 and 3-3 represent the different shell thicknesses of the drywell and vent. Nominal or, as-designed thicknesses, summarized in' Table 3-1, are used for the'drywell shell for all regions other than the sandbed region. The sandbed region shown in blue in Figure 3-3 is considered to have a thickness of 0.700 inch. This is less than the 95% confidence projected thickness for outage 14R. Figure 3-4 shows the view fron the inside of the drywell with the gussets and the vent jet deflector.

The drywell and vent shell is modeled using the 3-dimensional plastic quadrilateral shell (STIF43) element. Although this element has plastic capabilities, this analysis is conducted using only elastic behavior.

This element type was chosen over the elastic quadrilateral shell (STIF63) element because it is better suited for modeling curved surfaces.3-2 VNOEX 9-2, REV. 0 At a distance of 76 inches from the drywell shell, the vent is simplified using beam elements.

The transition from shell to beam elements is made by extending rigid beam elements from a node along.the centerline of the vent radially outward to each of the shell nodes of the vent. ANSYS STIF4 beam elements are then connected to this centerline node to model the axial and bending stiffness of the vent and header. Spring (STIF14) elements are used to model the vertical header supports inside the torus. ANSYS STIF4 beam elements are also used to model the stiffeners in the cylindrical region of the upper drywell. The section properties of these stiffeners are summarized in Table 3-2.The sandbed region at the base of the drywell was designed to provide a smooth transition to reduce thermal and mechanical discontinuities.

The sand provides lateral support to the drywell sphere in this region. The foundation stiffness for the sandbed is considered to be 366 psi/in per Reference 2.4.10 of Reference 3-2. ANSYS STIF14 spring elements are extended radially outward from each node of the shell -in the sandbed region to model the sand support as shown in Figure 3-3.The stiffness for each of these sand spring elements is calculate d by multiplying the foundation stiffness of the sand by the contributory area of-each node in the' sandbed region.3.3 Drywell Materials The..drywell shell is fabricated from. SA-212B FBX steel. The mechanical properties fdr this material at room temperature are shown in Table 3-3. These are the properties used in the finite element analysis.

For the perforated vent jet deflector, the material properties-were modified to account for the reduction in stiffness due to the perforations.

3.4 Boundary Conditions Symmetric boundary conditions are' defined for both edges of the 36'drywell model for the static stress analysis as shown on Figure 3-5.This allows the nodes at this boundary to expand radially outward from 3-3 TR4006 4 NEX 9-,2 RE:V. 0 the drywell centerline and vertically, but not in the circumferential.

direction.

Rotations are also fixed in two directions to prevent the boundary from rotating out of the. plane *of symmetry.

Nodes at the bottom edge of.the drywell are fixed in all directions to simulate the fixity of the shell within the concrete foundation.

Nodes at the ends of the sand spring.elements and the header support spring elements are: also fixed.3.5 Loads The loads are applied to the drywell finite element' model in the manner which most accurately represents the actual loads anticipated on the drywell. Details on the application of loads are discussed in the following paragraphs.

3.5.1 Load Combinations.

All load combinations to be considered on the drywell are summarized on Table 3-4. The most limiting load combinations in terms of possible.

buckling are those which cause the most compressive stresses in the sandbed region. Many of the design basis load combinations include high internal pressures which would create tensile stresses in the shell and help prevent buckling.

The most severe design load combination identified for the buckling analysis of the drywell is the refueling condition (Case IV). This load-combination consists.of the following loads: Dead weight of vessel, penetrations, compressible material, equipment supports and welding pads.Live loads of welding pads and equipment door Weight of refueling water External Pressure of 2 psig Seismic inertia and deflection loads for unflooded condition The normal operation condition with seismic is very similar to *this condition, however, it will be less severe due to the absence of the refueling.water and equipment door weight.3-4.-

RF# 00664*NOEX 9-2, REV. 0 The-most severe load combination for the emergency condition is for the post-accident (Case VI)'load combination including:

Dead weight.of vessel, penetrations, compressible material and equipment supports Live load of personnel lock Hydrostatic Pressure of Water for Drywell Flooded to 74'-6" External Pressure of 2 psig Seismic inertia and deflection loads for flooded condition The application of these loads is described in more detail in the following sections.3.5.2 Gravity Loads The gravity loads include dead weight loads of the drywell shell, weight of the compressible material and penetrations and live loads.The drywell shell loads are imposed on the model by defining the weight density of the shell material and applying a vertical acceleration of 1.0 g to simulate gravity. The ANSYS program automatically distributes the loads consistent with. the mass and acceleration.

The compressible material weight of 10 lb/ft2 is added by adjusting the weight density of the shell to also include the compressible material.

The adjusted weight densities for the.various shell thicknesses are summarized.

on Table 3-5. The compressible material is assumed to cover the entire drywell shell (not including the vent) up to the elevation of the flange.The additional dead weights, penetration weights'and live loads are applied as additional nodal masses to the model. As shown on Tablie 3-6 for the refueling -case, the total additional massis summed for each 5 foot elevation of the drywell. The total is-then divided by ].0 for the 36- section assuming that the mass is evenly distributed around the perimeter of the drywell. The resulting mass is then-applied uniformly to a set of nodes at the desired elevation as shown on Table 3-6. These applied masses automatically impose gravity loads.on the drywell model with the defined acceleration of lg. The same 3-5.

RF# 00664-NDEX 9-2, REV. O0.method is used to apply the additional masses to the model for the post-accident, flooded case as summarized in Table 3-7.3.5.3 Pressure Loads The 2 psi external pressure load for the refueling case is applied to the external faces of all of the drywell and vent shell elements.

The compressive axial stress at the transition from vent shell to beam elements is simulated by applying equivalent axial forces to the nodes of the shell elements.Considering the post-accident, flooded case, the drywell is assumed to be flooded to elevation 74'-6" (894 inches). Using a water density.of 62.3 lb/ft 3 (0.0361 lb/in 3), the pressure gradient versus elevation is.calculated as shown' in Table 3-8. The hydrostatic pressure at the bottom of the sandbed region is calculated to be 28.3 psi. According.e to the elevation of the element centerline, the appropriate pressures are applied to the inside surface of the shell elements.3.5.4 Seismic Loads Seismic stresses have been calculated for the-Oyster Creek Drywell in Part 1 of this report, Reference 3-3. Meridional stresses are imposed on the drywell during a seismic event due to a 0.058" deflection of the reactor building and due to horizontal and vertical inertial loads.on the drywell.The meridional stresses due to a seismic event are imposed on the 3-D drywell model by applying downward forces at four elevations of the model (A: 23'-7",B:

37'-3",C:

50'-11" and D: 88'-9") as shown on Figure 3-6. Using this method, the meridional stresses calculated in Reference 3-3 are duplicated at four sections of the drywell including 1) the mid-elevation of the sandbed region, 2) 17.25' below the equator, 3) 5.75 above the equator. and 4) just above the knuckle region. These four sections were chosen to most accurately represent the load distribution in the lower drywell while also providing a reasonably accurate stress distribution in the upper drywell.3-6 RF# 00664 INDEX 9-2, REXI.-0 To find the correct loads to match the seismic stresses, the total seismic stress (due to reactor building deflection and horizontal arid vertical inertia) are obtained from Reference 3-3 at the'four sections of interest.

The four sections and the corresponding meridional stresses for the refueling (unflooded)

'and post-accident (flooded)seismic cases are summarized in Table 3-9.Unit loads are then applied to the 3-D model in separate load steps at each elevation shown in Figure 3-6. The resulting stresses-at the four sections of interest are then averaged for each of the applied unit loads. By solving four equations with four unknowns, the correct loads are determined to match the stresses shown in Table 3-9 at the four sections.

The calculation for the correct loads are shown' on Tables 3-10 and 3-11 for the refueling and post-accident cases, respectively.

I..3.6 Stress Results The resulting stresses for the two load combinations described in section 3.5 are summarized in this section.3.6.1 Refueling Condition Stress Results The resulting stress distributions for the refueling condition are shown in Figures 3-7 through 3-10. The red colors represent the most tensile stresses and the blue colors, the most compressive.

Figures 3-7 and 3-8 show the meridional stresses for the entire drywell and lower drywell. The circumferential stresses for the same areas are shown on Figures 3-9 and 3-10. The resulting average meridional stress at the mid-elevation of the sandbed region was found to be;rRm 5~-7097 psi 3-7

?RF# 00664 NDEX 9-2, REV. 0 The circumferential stress averaged from the bottom to the top of the sandbed region is;ORc --277 psi 3:6.2 Post-Accident Condition Stress Results The application of all'of the loads described for the post-accident.

condition results in the stress distributions shown in Figures 3-11 through 3-14. The red colors represent the most tensile stresses and the blue colors, the most compressive.

Figures 3-11 and 3-12 show the meridional stresses for the entire drywell and lower drywell. The circumferential stresses for the same areas are shown on Figures 3-13 and 3-14. The resulting average meridional stress at mid-elevation of the sandbed region was found to be;apAm- -9693 psi The circumferential stress averaged from the bottom to the top of the sandbed region is;OPAc -+4049 psi 3-8 TRF#X00664 NDEX 9-2, RE:V. 0 3.7 Theoretical Elastic Buckling Stress Results After completion of the stress runs for the Refueling and Post-Accident load combinations, the eigenvalue buckling runs are made as described in Section 3.1. This analysis determines the theoretical elastic buckling loads and buckling mode shapes.3.7.1 Refueling Condition Buckling Results As shown on Figure 3-15, it is possible for the. drywell to buckle in two different modes. In the case of- symmetric buckling shown on Figure 3-15, each edge of the 36' drywell model experiences radial displacement with no rotation.

This mode is simulated by applying symmetry boundary conditions to the 3-D model the same as used for the stress run. Using these boundary conditions for the refueling case, the critical load factor was found to be 14.32 with the critical buckling occurring in the sandbed region. The critical buckling mode shape is shown in Figure 3-16 for applied symmetry boundary conditions.

The red color indicates sections 'of the shell which.displace radially outward and the blue, those areas which displace inward.The first four buckling modes were solved for in this eigenvalue buckling analysis with no buckling modes- found outside the sandbed region for a load factor as high as 16.32. Therefore, buckling is not a concern outside of the sandbed region.It is also possible for the drywell to buckle in the anti-symmetric manner shown in Figure 3-15. For this mode, the edges of the 3-D model are allowed to rotate but are restrained from expanding radially.

This case is considered by applying anti-symmetric boundary conditions at the edges of the 3-D model. With the two.pass approach used by ANSYS, it is possible to study anti-symmetric buckling of the drywell when the stresses are found based on symmetry boundary conditions.

The resulting load- factor found using anti-symmetric boundary conditions is 16.81. The mode shape for this case is shown on Figure 3-17.3-9 IRF# 006 4 INDEX 9-l, REV. 0 Because the load factor is lower for symmetry boundary conditions with the-same applied stress, the symmetric buckling condition is mor'e limiting.

Multiplying the load factor of 14.32 by the average meridional stress from section 3.6.1, the theoretical elastic buckling stress is found to be;ERie -14.32 x (7097 psi) a 101,650 psi 3.7.2 Post-Accident Condition Buckling Results Considering the post-accident case with symmetry boundary condition!;, the load factor was calculated as 9.91. Multiplying this load factor by the applied stress from section 3.6.2 results in a theoretical elastic buckling stress of 9PAie 9.91 x (9693 psi) 96,060 psi The critical mode shape for this condition-is shown in Figure 3-18.Again, the critical buckling mode is in the sandbed region.3.8 References 3-1 DeSalvo, G.J., Ph.D, and Gorman, R.W., "ANSYS Engineering Analysis System User's Manual, Revision 4.4," Swanson Analysis Systems, Inc., May 1, 1989.3-2 GPUN Specification SP-1302-53-044, Technical Specification for Primary Containment Analysis -Oyster Creek Nuclear Generating Station; Rev. 2, October 1990.3-3 'An ASME Section VIII Evaluation of the Oyster Creek Drywell -Part 1 Stress Analysis," GE Report No. 9-1, DRF # 00664, November 1990, prepared for GPUN.3-10I YNDEX -2, REV. 0 Table 3-1 Oyster Creek Drywell Shell Thicknesses Section Thickness (in.)Sandbed Region -0.700 Lower Sphere 1.154 Mid Sphere .3.770 Upper Sphere 0.722 Knuckle 2.5625 Cylinder 0.640 Reinforcement Below Flange 1.250 Reinforcement Above Flange 1.500 Elliptical Head 1.1875 Ventline Reinforcement 2.875 Gussets 0.875 Vent Jet Deflector 2.500 Ventline Connection 2.500 Upper Ventline 0.4375 Lower Ventline 0.250 3-11 RF# 00664 TNDEX 9-2, REV. 0 Table 3-2 Cylinder Stiffener Locations-and:Section Properties Elevation Height Width Area BendinQ Inertia (in 4)(in) (in) ( in) (in) Horizontal Vertical 966.3 0.75 6.0 4.5 13.5 0.211 -1019.8 0.75 6.0 4.5 13.5 0.211 1064.5 0.50 6.0 3.0 9.0 0.063 1113.0(1) 2.75 7.0 26.6 387.5 12.75 1.00 7.38 1131.0 1.0 12.0 12.0 144.0 1.000 (1) -This stiffener is made up of a 2 beam sections, one 2.75x7" and one 1.x7.375" Table 3-3 Material Properties for FBX-212B Steel Material Pronertv Value -Young's Modulus Yield Strength Poisson's Ratio Density 29.6x10 6 psi 38000 psi 0.3 0.283 lb/in 3 3-12 NDEX 9-2, REV. 0 Table 3-4 Oyster Creek Drywell Load Com6inations CASE I -INITIAL TEST CONDITION Deadweight

+.Design Pressure (62 psi) + Seismic (2 x DBE)CASE II -FINAL TEST CONDITION Deadweight

+ Design Pressure (35 psi) + Seismic (2 x.DBE)CASE III -NORMAL OPERATING CONDITION Deadweight

+ Pressure (2 psi external)

+ Seismic (2 x DBE)CASE IV -REFUELING CONDITION Deadweight

+ Pressure (2 psi external)

+ Water Load +Seismic (2 x DBE) .CASE V : ACCIDENT CONDITION Deadweight

+ Pressure(62 psi @ 175OF or 35 psi @ 281'F) +Seismic (2 x DBE)CASE VI -POST ACCIDENT CONDITION Deadweight

+ Water Load @ 74'6" + Seismic (2 x DBE)3-13

?RF# 00664 NDEX 9-2, REV. 0 Table 3-5 Adjusted Weight Densities of Shell to Account for Compressible Material Weight Shell* Adjusted Weight Density (lb/in3)Thickness (in.)1.154 0.770 0.722 2.563 0.640 1.250 0.343 0.373 0.379 0.310..0.392 I.0.339 3-14 RF# 00664 INDEX 9-2, RE'/. 0.Table 3-6 I Oyster Creek Drywell Additional Weights -Refueling Condition DEAD ELEVATION IWEI6HT (fet) (lbf)15.56 50000 16 20* 15-20-221 556000"21-250 26 30 64100 30.25 105000* 26-30 31 32 33 34 35" 31-35 36 PENETR.WEIGHT (lbf)168100 11200 11100 51500 26500 750 15450 28050 1500 1550 43350 40' 36-40 4S501`~ 45-501 54*1* 51-55 -56 60* 56-60 65-- 61-65 70" 66-70 73" 71-75 82X17*1 81-85 87 90* 86-90 93.75 94.751 95.75*1 91-96 TOTOLS: 41000 1102000 a MISC. TOTAL LOADS LOAD (Ibf) (lbf)50000 168100 11200 556000 11100 115600 100000 205000 16500 750 15450 28050 1500 1550 84350 1102000 7850 24000 80400 20000 115900 20000 72000 5750 8850 21650 1000 15000 .20700 698000 698000 20100 862000 3434350-5 FOOT LOAD 229300 556000*LOAD PER 36 DE6. f OF (lbf) ELEMENTS MooES OF APPLICATION 116-119 161-169 22930 55600 6 a 331700 33170 62250 6225 8 179-187 8 188-196 4146 2073.3 77B 389 LOAD PER FULL NOOE (Ibf).3622 6950 LOAD PER HALF NODE -(lbf)' 1911 3475 7850 56400 95200 52000 700 85900 1102000 7850 196300 72000 5750 8850 21650 8590 110200 785 8 a 5750 8850 19630 7200 575 8as 2165.8 8 a 8 a 197-205 418-426 436-4U 454-462 472-480 508-516 526-534 553-561 1074 13775 913-2454S' 901)72 11.1 27:1 200 923!i 537 6888 49 1227 450 36 55 135 100.I.1.I 21650 1000 15000 16000 1600 20700 20100 8 571-579 8 589-597 21___150 2184150 388200 738800 3____35_3434350 73880 343435._343435 4618 I -LOAD TO BE APPLIEO IN VERTICAL DIRECTION ONLY.& -MISCELLANEOUS LOADS INCLUDE 698000 LB WXiTER WEIGHT AT 94.75 FT. ELEVATION 100000 LB EQUIPMENT DOOR WEIGHT AT 30.25 FT. ELEVATION AND WELD PAD LIVE LOADS OF 24000. 20000 AND 20000 AT 56. f0 AND 65 FT. ELEVATIONS REnhGT.WK1 3-15...ml RF# 00664 YNOEX 9-2, REV~. 0 Table 3-7 Oyster Creek ELEVATION (feet)16 20 15-20 22t 21-25#26 30 30.25* 26-30 31 32 33 34 35 31-35 36 40 36-40 5W'9 65-501 54'951-55 56 60 5 56-60 65 51-65 70 S56-70 73 71-75.12.17'.31-85 87 90*3 .36-90 33.75.35.75 31-96 TOTALS:* DEAD WEIGHT (lbf)50000 PENETR.WEIGHT (lbf)168100 11200 Drywell MISC.LOADS (Tlbf)_______556000 Additional Weights -Post-Accident Condition TOTAL S FOOT LOAD PER LOAD PER LOAD RANGE 36 DEG. J Or N0OES OF FULL NODE (lbf) LOAD (1bf) ELEMENTS APPLICATION (lbf)50000 168100 11200 229300 22930 6 .116-119 3822 556000 586000 55600 a 161-169 6950 11100 115600 105000 231700 23170 8 179-187 2896 16500 750 154S5 28050 1500 62250 6225 8 .188-196 778 1550 11100 51500 64100 105000 LOAD PER HALF NOOE (Ibf)1911.3475..389 537 6888 49 16500 750 15450 28050 1500 1550 43350 41000 1102000 7850 56400 95200 52000 700 84350 1102000 7850 56400 95900 52000 5750 8850 21650 1000 15000 20700 20100 85900 1102000 7850 152300 52000 5750 8850 21650* 8590 110200 785.8 8 8 5750 8850 15230 5200 575.885 2165.8 a a 8 a 197-205 418-426 436-444 454-462 472-480 508-516 526-534 553-561 571-579 1074 13775 98 1904 650 72 lil 271 952 325 36 55 135 21650 1000 15000 16000 1600 8 20700 20100 200 100 510 255 40800 2184150 386200 0 2572350 2572350 f -LOAD TO BE APPLIED IN VERTICAL DIRECTION ONLY.6 -NO MISCELLANEOUS LOADS FOR*THIS CONDITION.

4080__5___35 25J235 8 589-597 FL0OD4GT.WKI 3-16

-TNDEXO-2, REV. 0 Table 3-8 Hydrostatic Pressures for Post-Accident, Flooded Condition WATER DENSITY: FLOODED ELEV: 62.32 lb/ft3 0.03606 lb/in3 74.5 ft 894 inches ELEMENTS AJ3OVE NI)DES 27 40 53 66 79 92 102 108 112 116 120 124 130 138 148 161 170 179 188 197 400 409 418 427 436 445 454 463 472 481 490 499 508 517 526 ANGLE ABOVE EQUATOR (degrees)-53.32-51.97-50.62-49.27-47.50-46.20-44.35-41.89-39.43-36.93-34.40-31.87-29.33-26.80-24.27-20.13-14.38-8.63-2.88 2.88 8.63 14.38 20.13 25.50 30.50 35.50 40.50 45.50 50.50 54.86 ELEVATION (inch)110.2 116.2 122.4 128.8 137.3 143.9 153.4 166.6 180.2 194'.6 209.7 225.2 241.3 257.6 274.4 302.5 342.7 384.0 425.9 468.1 510.0 551.3 591.5 627.8 660.2 690.9 719.8 746.6 771.1 790.5 805.6 820.7 835.7 850.8 885.3 187.3 DEPTH (inch)783.8 777.8 771.6 765.2 756.7 750.1 740.6 727.4 713.8 699.4 684.3 668.8 652.7 636.4 619.6 591.5 551.3 510.0 468.1 425.9 384.0 342.7 302.5 266.2 233.8 203.1 174.2* 147.4 122.9 103.5 88.4 73.3 58.3 43.2 8.7 706.7 PRESSURE (psi)28.3 28.1 27.8 27.6 27.3 27.1 26.7.26.2 25.7 25.2 24.7 24.1 23.5 23.0 22.3 21.3 19.9 18.4 16.9 15.4 13.8 12.4 10.9 9.6 8.4 7.3-6.3 5.3 4.4 3.7 3.2.2.6 2.1 1.6 0.3 25.5 ELEMENTS 1-12 13-24 25-36 37-48 49-51, 61-66 ,55-57 52-54, 138-141 ,58.'60 142-147, 240-242, p257-259 148-151, 243, 256 152-155, 244, 255.156-159i 245, 254 160-165, 246, 253 166-173, 247, 252 174-183, 248-251 184-395 196-207 208-215 216-223 224-231 232-239 430-437 438-445 446-453 454-461 462-469 470-477 478-485 486-493.494-501 502-509 510-517 518-525 526-533 534-541 542-549 550-557 340-399 (Ventline)

FLOODP .WK1 3-17

.ORN 00624 .I NOEX 9- z, REV. 0 Secti.-A) Middle B) 17.25-C) 5.75 , D) Above of'Be'Abo, Knui Table 3-9 Meridional Seismic Stresses 2-D Shell Elevation Model on (inches) Node Sandbed 119 32 low Equator 323 302 ve Equator 489 461:kle 1037 1037 at Four Sections Meridional Stresses _Refueling Post-Accident (psi ) (psi )1258 1288 295 585 214 616; 216 808 3-18 DRFO 00664 INDEX 9-2, REV. 0 Table 3-10 Application of Loads to Match Seismic Stresses -Refudling Case SECTION: 2-0 NODE: ELEV: COMPRESSIVE STRESSES FROM 2-0 ANALYSIS___ __ ___ __ ___ __ ___ __ __ ---- -____-0.058o SEISMIC DEFLECTION:

HORIZ. PLUS VERTICAL SEISMIC INERTIA: TOTAL SEISMIC COMPRESSIVE STRESSES: 2-0 SEISMIC STRESSES AT SECTION (psi)1 2 3

• 4 32 302 461 1037 119.3" 322.S 489.1" 912.3" 785.67 155.54 103.46 85.31 469.55 139.44 110.13 130.21 1258.22 294.98 213.59 .215.52 3-D STRESSES AT SECTION (psi)._______________________-_____..

...I.I.I 3-0 INPUT LOAD SECTION_______A C D SECTION: 3-0 MODES: ELEV: INPUT 3-0 UNIT LOAD DESCRIPTION 1000 lbs at nodes 563 through 569 500 lbs at 427&435. 1000 lbs at 428-434 500 lbs at 197.205. 1000 lbs at 198-204 500 lbs at 161L169. 1000 lbs at 162-168 I 53-65 119.3" 85.43 89.88* 97.64 89.85 1258.22 170-178 322.5'37.94 39.92 43.37.____00 294.98 3 400-408 489.1" 34.94 36.76 0.00* 0.00 213.59_213.59 4 526-531 912.3"____ _..55.23 0.00 0.00 0.00 215__.5.215.52 DESIRED COMPRESSIVE STRESSES (psi): 3-D INPUT LOAD SECTION A B C D LOAD TO BE APPLIED TO MATCH 2-0 STRESSES 3902.2 2101.4 1453.6 6611.6 RESULTING STRESSES AT SECTION (psi)_______ --_----------_-_----_____.._

333.37 188.87 141.93 S94.05 1258.22 125B.22 148.05 83.89 63.04 0.00 294.98_294.98'136.34 77.25 0.00 0.00 213.59 215.52 0.00 0.00 0.00 2.__2.__.215.52 SUM: SEISUNFL.WKI A 3-19 RFx 00664 INDEX 9-2, REV. 0 Table 3-11 Application of Loads to Match Seismic Stresses -Post-Accident Case 2-D SEISMIC STRESSES AT SECTION (psi)SECTION: 2-0 NODE: ELEV: COMPRESSIVE STRESSES FROM 2-D ANALYSIS-___________--_._____________________.__

0.058" SEISMIC DEFLECTION:

HORIZ. PLUS VERTICAL SEISMIC INERTIA:-________________

-TOTAL SEISMIC COMPRESSIVE STRESSES: 1 2 32 302 119.3- 322.5'?78.67 155.54 499.79 429.39 1288.46 54.93 3 461-489.1" 103.46 512.76 616.22 4 1037.912.3"____. --85.31 723.14 808.45 3-D STRESSES AT.SECTIOM (psi)3-0 INPUT LOAD SECTION A 8 C D SECTION: 3-D NODES: ELEY: INPUT 3-D UNIT LOAD DESCRIPTION

-_________________...._.__

_1000 lbs at nodes 563 through 569 500 lbs at 427&U35. 1000 lbs at 426-434 500 lbs at 197&205. 1000 lbs at 198-204 500 lbs at 161&169. 1000 lbs at 162-168 1 53-65 119.3'85.43 89.88 97.64 89.85 1288 __1288.46* 2 170-178 322.5" 37.94 39.92 43.37 0.00 5..____584.93 3 400-408 489.1" 34.94 36.76 0.00: 0.00 616.22 4 526-534 912.3" 55.23 O.CO 808.45 0ESIRED.COMPRESSIVE STRESSES (psi): 3-D INPUT LOAD SECTION_______A 8 C D LOAU TO SE APPLIED TO MATCH 2-0 STRESSES-_____________-

.___________________.__

14637.9 2850.2-1941.7-318.8 RESULTING STRESSES AT SECTION (psi)---------------------------------

1250.51 555.36 511.45 808.45 256.17 113.78 104.77 0.00-189.58 -84.21 0.00 0.CO-28.64 0.00 0.00 0.CO-______ ------ .1288.46 584.93 616.22 808.45 SUN: SEISFL.WKI 3-20 DRYWELL ELEV. SFO ., N- T 7&W R. THK. .676" Figure 3-1.Oyster Creek Drywell Geometry 3-21 1 NOW B. 19 9 L4:33:33 REAL "UKN MU =1 Mr =639.49U CEZ4?ROW HZDDEH w" N4 P4 OYSTER D RM AIMLYSIS -OYC+/-Q CaNs. RFUI)-I I. Fin~rt op 1-4 nvetpr rvooie nvioomi i 4-n ri"ito riamovit mmail..

0 il anvil ah sf HOU X3 1999 14: :17 REAL "U"3 XU =1 vistas l 03 76 XT 429.432 Sr =216.528 ANGZ I99 D CENIhOID HIDDSH W Lab i. V V I % 'S % 1% x %. 1 I.I I E C ...L I .. ..IOWESTZ CRERX 'DRYNY AMALYS.1.-

OVCRtQ. C3AHD, RZ:FUEISNilG)

Figure 3-3..e. .I Closeup of lower Orywell Section of FEN (Outside View).

m 1 fNSYS 4.4 Nov 13 1990 N:56:a3]MAL MM~XU =-1i XT =420.4!3 2 =216.528 ANGZ9O CIENROID HIDDEN w... ......OYSTER CREEX _RWL A"ALYIS -*V'CRI C3As_ REFD[_>;. C~mw 1 ego 1ASpuln of Inw"' neel! Dor.4n"n nf cCm Trnemis Vaswt..

-I Amsys 4.4*OCT 15 1996 0 9:31:36 TYPE NUN BC SYN"DOL DIST71876

.76)(J =363.,031 F2 =639.498--~ac' sok HI'T IDD N OYSTER CREEX DRYWrELL AM~LYS1S c4RASN. NTLA U'

-1 OCT 15 1996 09:33:26 TVP;7 "UNMgT BC SY BOLS D1ST718 .786 XF -a2a .931 ZY =639.498 OYSTR A H I DD OYST_ .RI ASELAmYI YCI S?.UI ODC w'F4nilr 1-0i Amnnlir:%tinn nf flxdinn ton Mimilatp spilmic Rpndina 7.

0 .- -- 1 NOV 14 1999 08: 1a .s STE?=1 IT KRI DWX =6.266376 At =-968 SMX =442.967 XU =1 DIST=710 .786 XJW -393 .931 CYtJOI HI DDPZ 6w -1434 442.967 w-R .X -RL .V .-OQ C OYSTERCRE DRYWIL ANALYSIS -OYCRQ C3AND.RIFUELING)

Fin,," I-7 -Moridjnn1l Strp-cq -Raftiplinrn rxRP 1 ANSYS 4.4 NOV 14 1996 98:14:57 STEW=l ITEB=l Sy c QUG)P DNX =91.19679$mm =-Go=SNX =442. 9G7 XU =1 YU .=-G.s DIST=2698.376 XCF =429.452 CENThOID HIDDEN--7064-1434 442.907 P w 4.I.6--OYSTERl C2REM DRYIIELL ANALYSIS-.OYCRIQ CSAND&REFUELK 1103-tI Finfotrs 'A_Fi.,.il nwpr flrvwpl1 mpi'idinna1 Stptp -Rpfipl inn rasp I aII5VS 4.4 NOV 14 1998 08:13:14 STEP=1 I TERI SX * (UG)DINX =9.26378 SHN =-3363 SiM =2873)CU =1 Wu =-e. a DIST718 .706 CENTROIV HIDPDt XJ-33. 93-2670* :&1?427 , -I.w'a OYSTER CREEX DRYWELL ANALYSIS -OYCRIQ CSAND. REFUE.LIN)

I;, 940flra 1.4 ii,,ormfprpntial qtrto -.Rpfilif~r rate

-I NsYS 4.4 Ii .NOU L4 C9os s .08:13:19 STEP=1 IT]UtI S :X (BUG)DMX -8.166793 CNN =-3363 SMX =2873 XII =1 WV =-S.8 D.ST=208 376 CEIITROZP HIDDEN--3362-2670 2873 ClIOYSTER' CREEX4 DRYHELL ANALYSIS -OVCRIQ CSAND. REFUELING)I Figure 3-10. Lower Drywell Circumferential Stresses -Refueling Case

i. 1 .C4.4 mo l3 L" TER1l"~S X I IA6J,9a DMX =9.486?0 SN =-12326 SHX =2718= 1 DIST=1 ..786 CEWNTOID HIOVKN-12329-19656.*-63o.al1 2719 O.STER CREEX DRYHELL ANALYSIS -4 0. POST-ACCID.

Fimiwrp 1-1 1 Moridional'Strpgcon

.Pnqt-Arridont rare ANSVS4.4 STEP--I SY (AUG')DMX =0.311487 CMn =-12329 SEIX =271S YU, =-e.sa DIST=298*3.76

.: XFr =420.452 CEIITROID HIDDIN E -:22 2710 IOYSTERI CxEMX DRVWELL AM~LVSIS -OVCILUUC3AtW.

POST-ACCID.)I I;.. .Finilop 1-1?I nv.117 fwer Drywell Mpv*idinnal qtata .Pnct~.arr~idfpt raqp

• "Seys 4.4 I. MOU 15 1996.W .c"ts.w t .STABI MMX =.. 4X2?7 SMN =-4594 ShU =12763 YU =-S. 8 DI ST=710 .706 i)X =3- s3.t631-'-45194MI

• -2666 o .2 12763 IOYSTER CREIo DRYMELL AALYSIS t 4S-m FOST-ACCIP.)J

~~***~- 'i :-'p~ 4i ' 'lA ~ , ' ' :

1 ANSYS 4.4.IL mu a:u a99v STEP=1 KTERIl;~~~s A- --dI- Wtl_____ SbX (BUG)<DMX =9.311487 SMH =-4594 SNX =12763..._.Ye =-he.s8 DI t9=29 ..376lCZERQID HIDIDENl.-4594-2666-.... 996 12763 OWSTEJR CMRXE DIRYWILL AtMLYSIS -OVcR+/-U CSAND,. POST-ACCtKP.J

.Fiatlio 1_14 Inwavf nrvwol rfiresmmfowrnantiii ciaceae -Dnet.aereIanf rse..

I-Unbuckled Shapel Buckled Shape Vent f Radial Displacemen:

No Rotation Symmetric Buckling of Drywell Unbuckled Shape Buckled Shape Vent ( Rotation No Radial Disp.4)Anti-symmetric Buckling of Drywell SYM.DRW Figure 3-15. Synnetric and Anti-Synmnmtric Buckling Modes 3-35 ANSYS 4.4 NOU 14 1990 98:25:22 STEP--I STERN tRS_FCT=14. 322 HPXCLOBSIL DM(X = .O618 ZMN =-G.692209 Sit =R. .SB85 XU =1 YU =-su.e DI ST=192. 353 Xr =327.423.. -62;1.6136 U -9. 60254 B. '" _'.OYSTER CRUP DRYWELL ANALYSIS-OYCRIR CSAND. REFUELING)

..L D _ _ _ _; Finairo 1-1f 'Rwvetric Bucklin" Mnfdp shank -Q-oaftiI4nn rcas STZP=l FACE =16 6312 DMX =9. 83214$MM =-G 992837 SMX =0. 9214)MU =1 DISI=192.351

-6.392837-8.892164-9.390492 OYSER RI DRYWIMEL ANALYSIS -8CR+/-H CSOtiM. -R!FUELKNG)

Cinvorn 1.17.8in*4_t4~vmfi ptlevH,~e, IR40i~ chk2"l -oftelMa14,,m 1Ime hNSYS 4.4 NOU la 1990 e8:13:29 STE?=1 ITERtl FACT=9.911.XGLODOL;DPIX =6. 9395.SN -e. 69363 SNX =e.e8=58 X13 =1 YU =-9.6.~D 57=186.* 59 XI =327.422* -9je 3.9128 0.00144.. s3m8o co OYSTER CREEX DRY WLL AHMLYSIS -OYCRLU CSAND. 'POST-CCID.

)c*4_.. 1 le 0-1-V1 Ann Un~ da -n-.4s A -- -- t~ ---

RF 006 DRFu 00664* NDEX 9-2, REV. 0 4. ALLOWABLE BUCKLING STRESS EVALUATION Applying the methodology described in Section 2 for the modification of the theoretical elastic buckling-stress, the allowable compressive stresses.

are now calculated.

Tables 4-1 and 4-2 summarize the calculation of the allowable buckling stresses for the Refueling and Post-Accident conditions, respectively.

The modified capacity reduction factors are first calculated as described in sections 2.2 and 2.3. After reducing the theoretical instability stress by this reduction factor, the plasticity reduction factor is calculated and applied. The resulting inelastic buckling-stresses are then divided by the factor of safety of 2.0 for the Refueling case and 1.67 for the Post-Accident case to obtain the final allowable compressive stresses.The allowable compressive stress for the Refueling case is 10.44 ksi.Since the applied compressive stress is 7.10 ksi, there is a 47%, margin. *The allowable compressive stress for the Post-Accident, flooded case is 14.34 ksi. This results in a margin of 48% for the.applied compressive stress of 9.69 ksi.4-1 RF# 00664 TNDEX 9-2, RE'/. :I Table 4-1 Calculation of Allowable Buckling Stresses -Refueling Case Parameter Value Theoretical Elastic Instability Stress, aie (ksi) 101.65 Capacity Reduction Factor, t 0..207 Circumferential Stress, ac (ksi) -0.28 Equivalent Pressure, p (psi) 0.000"X" Parameter 0.000 AC 0.000 Modified Capacity Reduction Factor, 1imod 0.207 Elastic Buckling Stress, ae ' 4,mod aie (ksi) 21.04 Proportional Limit Ratio, A -aeay. 0.554 Plasticity Reduction Factor, 17 0.993 Inelastic Buckling Stress, a1 -nice (ksi) 20.89 Factor of Safety, FS 2.0 Allowable Compressive Stress, aail -ao/FS (ksi) 10.44 Applied Compressive Meridional Stress, am (ksi) 7.10 Margin -[((all/m) 1] x 100% 47% l 4-2 MEMW DRF# 086~4 INDEX -, REV. 1 Table 4-2 Calculation of Allowable Buckling Stresses -'Post-Accident Case'Parameter Value Theoretical Elastic Instability Stress, ale (ksi) 96.06 Capacity Reduction Factor, ea 0.207 Circumferential Stress, ac (ksi) 4.05 Equivalent Pressure, p.(psi) 13.50"X" Parameter 0.082.AC 0.069 Modified Capacity Reduction Factor, aijmod 0.32 Elastic Buckling Stress, ae -ai,mod aie (ksi) 30.74 Proportional Limit Ratio, A -ae/ay- 0.809 Plasticity Reduction Factor, noj 0.736 Inelastic Buckling Stress, a 1 -17iae (ksi) 22.62 Factor of Safety, FS 1.67 Allowable Compressive Stress, aall -a 1/FS (ksi) 13.55 Applied Compressive Meridional Stress, am (ksi) 9.69 Margin -[(aall/am)

-1] x 100%. 39.7%r 4-3 DRF# 00664 INDEX 9-2, REV. O 5.

SUMMARY

AND CONCLUSIONS The results of this buckling analysis for the refueling and post-accident load combinations are summarized on Table 5-1. The applied and allowable compressive meridional stresses shown in.Table 5-1 are for the sandbed region which is the most limiting region in terms of buckling.

This analysis demonstrates that the Oyster Creek drywell has adequate margin against buckling for an assumed sandbed shell thickness of 0.700 inch. This thickness is less than the 95%confidence projected thickness of 0.736 inches for the 14R outage.5-1 F# 00664 RDEX.9-2, REV. 1 Table 5-1 Buckling.Analysis Summary Service Condition Factor of Safety Applied Applied Compressive Meridional Stress..(ksi)

Allowable Compressive Meridional Stress (ksi)Buckling Margin Load Combination Refueling Post-Accident Design Level C 2.00 1.67 7.10 9.69 10.44 13.55 47 40%47% .40%.E 5-2 IN X 9-4, REV. 0 AN ASME SECTION VIII'EVALUATION OF THE OYSTER CREEK DRYWELL FOR WITHOUT SAND CASE PART 2 STABILITY ANALYSIS February 1991 prepared for GPU Nuclear Corporation Parsippany, New Jersey.,-prepared by GE Nuclear Energy San Jose, California I I I I?RF#X00664 NDEX 9-4, RIV. 0 AN ASME SECTION VIII EVALUATION OF THE OYSTER CREEK DRYWELL FOR WITHOUT SAND.CASE PART 2.STABILITY ANALYSIS Prepared by: ¢C.D. Frederickson, Senior Engineer Materials Monitoring

&Structural .Analysis Services Reviewed by:__ _ ___ ___________

H. S. Mehta, Principal Engineer Materials Monitoring

&Structural Analysis Services Approved by:_-_S. Ranganath, Manager Materials Monitoring

&Structural Analysis Services..-^ .. ..I

?N E6 4 REV. 1 TABLE OF CONTENTS 1.. INTRODUCTION 1.1 General 1.2 Report Outline 1.3 References

2. BUCKLING ANALYSIS METHODOLOGY 1-1 1-1.1-1.*1-2 2-1 2.1 Basic Approach 2.2 Determination of Capacity Reduction Factor 2.3 Modification of Capacity Reduction Factor for Hoop Stress 2.4 Determination of Plasticity Reduction Factor 2.5 References

3. FINITE ELEMENT.MODELING AND ANALYSIS 2-1 2-2: 2-:3 2-.)3-.3-1 3-:2 3-:3 3 -3 3-4S 3 -7 3-9 3-10 3.1 3.2 3.3 3.4 3.5 3.6 3.7 3.8 Finite Element Buckling Analysis Methodology Finite Element Model Drywell Materials Boundary Conditions Loads Stress Results Theoretical Elastic Buckling Stress Results References

4. ALLOWABLE BUCKLING STRESS EVALUATION 4-1 5.

SUMMARY

AND CONCLUSIONS

.5-1 iii DRF# 00664 INDEX 9-4, RE:V. O LIST OF TABLES Table Page No. Title No.3-1 Oyster Creek Drywell ShellThicknesses.

3-11 3-2 Cylinder Stiffener Locations and Section Properties 3-12 3-3 Material Properties for SA-212 Grade B Steel 3-12 3-4 Oyster Creek Drywell Load Combinations 3-13 3-5 Adjusted Weight Densities of Shell to Account for .3-14 Compressible.Material Weight 3-6 Oyster Creek Drywell Additional Weights -Refueling 3-15 3-7 Oyster Creek Drywell Additional Weights -Post-Accident 3-16 3-8 Hydrostatic Pressures for Post-Accident, Flooded Cond. 3-17 3-9 Meridional Seismic Stresses at Four Sections 3-18 3-10 Application-of Loads to Match Seismic Stresses :19 Refueling Case 3-11 Application of Loads to Match Seismic Stresses 20 Post-Accident Case 4-1 Calculation of Allowable Buckling Stresses -Refueling 4-2 4-2 Calculation of Allowable Buckling Stresses -Post-Accident 4-3 5-1 Buckling Analysis Summary 5-2 iv DRF# 00664 INDEX 9-4, REV. 0 LIST OF FIGURES Figure Page No. Title No.1-1 Drywell Configuration 1-3 2-1 Capacity Reduction Factors for Local Buckling of 2-7 Stiffened and Unstiffened Spherical Shells 2-2 Experimental Data Showing Increase in Compressive 2-;3 Buckling Stress Due to Internal Pressure 2-3 Design Curve to Account for Increase in Compressive 2-9 Buckling Stress due to Internal Pressure 2-4 Plasticity Reduction Factors for Inelastic Buckling 2-10 3-1 Oyster Creek Drywell Geometry 3-21 3-2 Oyster Creek Drywell 3-D Finite Element Model 3-.'2 3-3 Closeup of Lower Drywell Section of FEM (Outside View) 3-.3 3-4 Closeup of Lower Drywell Section of FEM (Inside View) 3-.4 3-5 Boundary Conditions of Finite Element Model 3-25 3-6 -Application of Loading to Simulate Seismic Bending 3-26 3-7 Meridional Stresses Refueling Case 3-27 3-8 Lower Drywell Meridional Stresses -Refueling Case 3-28 v--

YRFW O?664 NOEX 9-4, REV. 0 LIST OF FIGURES Figure No.3-9 3-10 3-11 3-12 3-13 3-14 Title Circumferential Stresses -Refueling Case Lower Drywell Circumferential Stresses -Refueling Case Meridional Stresses -Post-Accident Case Lower Drywell Meridional Stresses -Post-Accident Case Circumferential Stresses -Post-Accident Case Lower Drywell Circumferential-Stresses

-Post-Accident Case Page No.3-29 3-.30 3-:31 3-:32 3-:33 3-34 3-15 3-16 3-17 3-18 Symmetric and Asymmetric Buckling Modes.Symmetric Buckling Mode Shape.-.Refueling Case Asymmetric Buckling Mode Shape Refueling Case Buckling Mode Shape -Post-Accident Case 3-:35* 3-:36* :37 3-38 vi ..

RF# 00664 TN0EX 9-4, REV.. 0 1. INTRODUCTION 1.1 General To address local wall thinning of the Oyster Creek drywell, GPUN has prepared a supplementary report to the Code stress report of record[1-1] which is divided into two parts. Part 1 includes all of the Code stress analysis results other than the buckling capability for the drywell shell [1-2]. Part 2 addresses the buckling capability of the drywell shell shown in Figure 1-1 [1-3]. The supplementary report for the degraded drywell -is for the present configuration

'(with sand support in the lower sphere). One option which is being considered by GPUN to mitigate further corrosion in the sandbed region is to remove the sand. Reference 1-4 and this report evaluate the influence of removing the sand on the code stress analysis and buckling evaluation,.

respectively.

Buckling of the entire drywell shell is considered in this analysis with the sandbed region being the area of primary concern.1.2 Report Outline Section 2 of this report outlines the methodology used in the bucklinc capability evaluation.

Finite element modeling, analysis and results;are described in section 3. Evaluation -of the allowable compressive buckling stresses and comparisons with the calculated compressive stresses for the limiting load combinations are covered in section 4.Section 5 presents the summary of results and conclusions.

1-1 DRF# 00664 INDEX 9-4, REV. 0 1.3 References 1-1 "Structural Design of *the Pressure.

Suppression Containment Vessels," by Chicago Bridge St Iron Co.,Contract

1. 9-0971, 1965.1-2 "An ASME Section VIII Evaluation of the Oyster Creek Drywell -Part 1 Stress Analysis," GE Report No. 9-1, DRF# 00664, November 1990, prepared for GPUN.1-3 "An ASME Section VIII Evaluation of the Oyster Creek Drywell -Part 2 Stability Analysis," GE Report No. 9-2, DRF# 00664, November 1990, prepared for GPUN.1-4 "An ASME Section VIII Evaluation of the Oyster Creek Drywell -Part 1 Stress Analysis," GE Report No. 9-3, DRF# 00664, February 1991, prepared for GPUN.1-2 MMMM NDEX -, RE'V. 0.Ad tu l0 9-3-i.t 1 W wEEkv. 8-£LS Vt. M Figure 1-1 Drywell Configuration 1-3 REV. 0-2. BUCKLING ANALYSIS.METHODOLOGY 2.1 Basic Approach The basic approach used in the buckling evaluation follows the methodology outlined in the ASME Code Case N-284 E2-1 and :2-2].Following the procedure of this Code Case,. the allowable compressive stress is evaluated in three steps.In the first step, a theoretical elastic buckling stress,; aie is determined.

This value may be calculated either by classical buckling equations or by finite element analysis.

Since the drywell shell geometry is complex, a three dimensional -finite element, analysis approach is followed-using the eigenvalue extraction technique.

More details on the eigenvalue determination are given in Section 3.In the second step, the theoretical.

elastic buckling stress is modified by the appropriate capacity and plasticity reduction factors.The capacity reduction factor, et, accounts for the difference between classical buckling theory and actual tested buckling stresses for fabricated shells. This difference is due to imperfections inherent in fabricated shells, not accounted for in classical buckling.theory, which can cause significant reductions in the critical 'buckling stress. Thus, the blastic buckling stress for fabricated shells is given by the product of the theoretical elastic buckling stress and, the capacity.reduction factor, i.e., aie i. When the elastic buckling stress exceeds the proportional limit of the material, a plasticity reduction factor, ri 1 , is used 'to account for non-linear material behavior.

The inelastic buckling stress for fabricated shells is given by '7 i~icie-In the final step, the allowable compressive stress is obtained by dividing the' buckling stress calculated in the second step by the safety factor, FS: Allowable Compressive Stress -7iaaie/FS.

2-1 MIMI 8- REV. 0 In Reference 2-1, the safety factor for the- Design and Level A .& B service conditions is specified as 2.0. A safety factor of 1.67 is specified for Level C service conditions (such as the post-accident condition).

The determination of appropriate values for capacity and plasticity reduction factors is discussed next.2.2 Determination of Capacity Reduction Factor The capacity reduction factor, a 1 , is used to account for reductions in actual buckling strength due to the. existence of geometric imperfections.

The capacity reduction factors given in Reference..2-1 are based on extensive data compiled by Miller (2-33. The factors appropriate for a spherical shell geometry such as that of the drywell in the sandbed region, are shown in Figure 2-1 (Figure 1512-1 of'Reference 2-1). The tail (flat) end of the curves are used for unstiffened shells. The curve marked 'Uniaxial compression' is applicable since the stress state in the sandbed region is compressive in the meridional direction but tensile in the circumferential direction.

From this curve, Ad is determined to be 0.207.The preceding value of the capacity reduction factor is very conservative for two reasons. First, it is based: on. the assumption that the spherical shell has a uniform.thickness equal to the reduced thickness.

However, the drywell shell has a greater.thickness above the sandbed region which would reinforce the sandbed region.. Second, it is assumed that the circumferential stress is zero. The tensile circumferential stress has the effect. of rounding the shell and reducing the effect of imperfections introduced during the fabrication and construction phase. A modification of the Ad value to account for the presence of tensile circumferential stress is discussed in'.Subsection 2.3.The capacity reduction factor values given in Reference 2-1 are.applicable to shells which-meet the tolerance requirements of NE-4220 2-2 NUEX 9-4, REE. 1 of Section III [2-4]. Reference 2-5 compares the. tolerance requirements of NE-4220 to the requirements to which the Oyster. Creek drywell shell was 'fabricated.

The comparison shows that the Oyster Creek drywell shell was erected to the tolerance requirements of NE-4220. Therefore, although the Oyster Creek dryweli. is not a Section III, NE vessel, it is justified to use the approach-outlined in Code Case N-284.2.3 Modification of Capacity Reduction Factor for Hoop Stress The orthogonal tensile stress has the effect of rounding fabricated shells' and reducing the effect of imperfections on the buckling strength.

The Code Case N-284 (2-1 and 2-2] notes in the last paragraph of Article 1500 that, "The influence of internal pressure on a shell structure may reduce the initial imperfections and therefore higher values of capacity reduction factors. may be acceptable..-

Justification for higher values of ae must be given in the Design report." The effect of hoop tensile stress on the buckling strength of cylinders has been extensivelly documented (2-6 through 2-11]. Since the methods used in accounting for the effect of tensile hoop stress for the cylinders and. spheres are similar, the test data and the methods for the cylinders are first reviewed.

Harris, et al (2-6]presented a comprehensive set of te'st data, including those from References 2-7 and 2-8, which clearly showed that internal pressure in the form of hoop tension, increases the axial buckling.stress of cylinders.

Figure 2-2 shows a plot of the test data showing the increase in buckling stress as a function of nondimensional.

pressure.This increase in buckling capacity is accounted for by defining a separate reduction factor, ap. The capacity reduction factor -a: can then be modified as follows: 9imod ' ai + Up 2-3

• EX 9- REV. 1 The buckling stress in uniaxial compression for a cylinder or a sphere of uniform thickness with no internal pressure is given by the following:

Sc -(0.605)(aoi)Et/R

-(0.605)(0.207)

Et/R.Where, 0.605 is a constant, 0.207 is the-capacity reduction factor,c:j, and E,t and R are Young's Modulus, wall thickness Land radius, respectively.

In the presence of a tensile stress such as that produced by an internal pressure, the buckling stress is given as follows: Scmod (0.605)(ai

+:r )Et/R-(0.605)(0.207

+ ap)Et/R[(0.605)(0.207).

+ AC] Et/R Where AC is ap/0.605 and is given for cylindrical geometries in the graphical form in Figure 2-3. As can be seen in Figure 2-3, AC -is a function of the parameter X-(p/4E)(2R/t) 2 , where ,p, is the internal pressure.

Miller (2-12] gives the following equation that fits the graphical relationship between X and AC shown in Figure 2-3: AC -ap/0.605 -1.25/(5+1/X).

The preceding approach pertains to cylinders.

Along the similar.lines, Miller (2-13] has developed an. approach for :spheres as described next.The non-dimensional parameter X is essentially (oe/E)(R/t).

Since.in the case of a sphere, the hoop stress is one-half of that in the cylinder, the parameter X is redefined for spheres as follows: X(sphere) (p/8E)(2R/t) 2 2-4

.DEX REV. 1 When the tensile stress magnitude, S, is known, the equivalent internal pressure can be calculated using the expression:.

p -2tS/R Based on a review of spherical shell buckling data [2-14, 2-15], Miller [2-13] proposed the following equation for AC: AC(sphere) 1.06/(3.24

+ 1/X)The modified capacity reduction factor, 'ijmod' for the drywell geometry was obtained as follows:, ai,mod 0.207 + AC(sphere)/O.65-2.4 Determination of Plasticity Reduction Factor.When the elastic buckling stress exceeds the proportional limit of the material, a plasticity reduction factor,.j, is used to account for the non-linear material behavior.

The inelastic buckling stress for fabricated shells is given by 1 7 iciaie* Reference 2-2 gives the mathematical expressions shown below.(Article

-1611 (a)] to calculate the plasticity reduction factor for the meridional direction elastic buckling stress. A is equal to aiaie/oy and ay is the material yield strength.

Figure 2-4 shows the relationship in graphical form.87 -1.0 if A < 0.55-(0.45/A) + 0.18 if 0.55 < A S 1.6-1.31/(1+1.15A) if 1.6 < A 6.25-1/A if A > 6.25 2.5 References 2-1 ASME Boiler and Pressure Vessel Code Case N-284, 'Metal Containment Shell Buckling Design Methods,Section III, Division 1, Class MC", Approved August 25, 1980.2-5 KX0I T, RIv. 1 2-2 Letter (1985) from C.D. Miller to P. Raju;

Subject:

Recommended Revisions to ASME Code Case N-284.2-3 Miller, C.D., "Commentary on the Metal Containment Shell Buckling Design Methods of the ASME Boiler and Pressure Vessel Code," December 1979.2-4 ASME Boiler & Pressure Vessel Code,Section III, Nuclear Power Plant Components.

2-5 "Justification for Use of Section III, Subsection NE, Guidance in Evaluating the Oyster Creek Drywell," Appendix A to letter dated December 21, 1990 from H.S. Mehta of GE to S.C. Tumminelli of GPUN.2-6 Harris, L.A., et al, "The Stability of Thin-Walled Unstiffened Circular Cylinders Under Axial Compression Including the Effects of Internal Pressure," Journal of the Aeronautical Sciences, Vol.24, No. 8 (August 1957), pp. 587-596.2-7 Lo, H., Crate, H., and Schwartz, E.B., "Buckling of Thin-Walled Cylinder Under Axial Compression and Internal Pressure," NACA rN 2021, January 1950.2-8 Fung, Y.C., and Sechler, E.E., "Buckling of Thin-Walled Circular.Cylinders Under Axial Compression and Internal Pressure," Journal of the Aeronautical Sciences, Vol. 24, No. 5, pp. 351-356, May 1957.2-9 Baker, E.H., et al., "Shell Analysis Manual," NASA, CR-912 (April -1968).2-10 Bushnell, D., "Computerized Buckling Analysis of Shells," Kluwer -Academic Publishers, 1989 (Chapter 5).2-11 Johnson, B.G., "Guide to Stability Design Criteria for Metal Structures," Third Edition (1976), John Wiley & Sons.2-6 DRF# 00664 INDEX 9-4, REV. 1 2-12 Miller, C.D., "Effects of Internal Pressure on Axial Compression Strength of Cylinders," CBI Technical Report No. 022891, February.1991.2-13 Miller, C.D., "Evaluation of Stability Analysis Methods Used for the Oyster Creek Drywell," CBI Technical Report Prepared for GPU Nuclear Corporation, September 1991.2-14 Odland, J., "Theoretical and Experimental Buckling Loads Of Imperfect Spherical Shell Segments," Journal of Ship Research, .Vol. 25, No.3, September 1981, pp. 201-218.2-15 Yao, J.C., "Buckling of a Truncated Hemisphere Under Axi'al Tension," AIAA Journal, Vol. 1, No. 10, October 1963, pp.*2316-2319.

2-7 O RF# 08Q6, 4 INDEX 9-4 REV. 1 0.5 0.3 0.2 0.0 a 4 a 12 Is 20 24 ..20 Figure 2-1 Capacity Reduction Factors for Local Buckling of Stiffened and Unstiffened Spherical Shells 2-8 N E l-2, RE.V. 1* "T. ., ,, .SUCHY(NAA)

FUNG & SECHLER iLo CRATE a SCHWARTZ-NAA:11-SEST..FIT CURV N .....-:&Orco r T1 E I i E ( )2 Figure 2-2 Experimental Data Showing Increase in Compressive Buckling Stress Due to Internal Pressure (Reference 2-6)2-9 INDEX'9-4, RE.V. 1 I 0 I a 6 4.z i i ..-1.0 a 6.AC 4 2== _ _ :: : : : =-1--1 11111 1 I I I li__ _ _ _ ___ _ ._ _ _ _ ..._ _ _ _ _ _ .= = : _ _ _ ._ _ _ _ _ _ _ ..............................................

l .._ _ ; = = _ : -.._ _ _ _ _ e .__ _ _ ._ = = _ = = = _ _ 7 .._====-_ _ = = _ _ _ _ _ _ _ _=- _ -== _ _ ._ _ _ _ _ T=__ _ =__ _ _ _ _ _ _ : .__- -----_G--- -------- 7= = z = = I = = = rTo= = _ _ _ i = = = I _ = = I_= _ _ _ Z , ; _ __ _ z _ : : _ = _ , .H_= _ _ t _ .I _ _ _ -T fiI= 2 _= = _ _ I =-_ I I [I Z _ _ _= = _ I = = _ IIII 7-_ _ _ _- l = _ _ XII., _ _ L ____ l --_ S 0.10 a 6 4 2 0.0i 0.05 a 4 OJO a 4..a 1.0 a.a.10 ,, r T Figure 2-3 Design Curve to Account for Increase in Compressive Buckling Stress Due to Internal Pressure (Reference 2-11)2-10

?UESEX YT, REV. I I OC .Ax q,-6 Figure 2-4 Plasticity Reduction Factors for Inelastic Buckling:1 2-11............

DRF# 00664 INDEX 9-4, REV. 0 3. FINITE ELEMENT MODELING AND ANALYSIS 3.1 Finite Element Buckling Analysis Methodology This evaluation of the Oyster Creek Drywell buckling capability uses the Finite Element Analysis (FEA) program ANSYS [Reference 3-1]. The ANSYS program uses a two step eigenvalue formulation procedure to perform linear elastic buckling analysis.

The first step is a static analysis of the structure with all anticipated loads applied. The structural stiffness matrix, [K], the stress stiffness matrix, [S], and the applied stresses, 0 ap are developed and saved from this static analysis.

A buckling pass is then run to solve for the eigenvalue or load factor, A, for which elastic buckling is predicted using the equation: ([K] + X [S] ) (u) 0 0 where: X is the eigenvalue or load factor.(u) is the eigenvector representing the buckled shape of the structure.

This load factor is a multiplier for the applied stress state at which the onset of elastic buckling will theoretically occur. All applied loads (pressures, forces, gravity, etc...) are scaled equally. For example, a load factor of 4 would indicate that the structure would buckle for a load condition four times that defined in the stress pass. The critical stress, acr' at a certain location of the structure is thus calculated as: dcr A Gap This theoretical elastic buckling stress is then modified by the capacity and plasticity reduction factors to determine the predicted buckling stress of the fabricated structure as discussed in Section 2.This stress is further reduced by a factor of safety to determine the allowable compressive stress.3-1_-_

NDEX 9-2 REV'.:0 3.2 Finite Element Model The Oyster Creek drywell has been previously analyzed using a simplified axisymmetric model to evaluate the buckling capability in the sandbed region [Reference 3-2]. This type of analysis conservatively neglects the vents and reinforcements around the vents which significantly increase the stiffness of the shell near the sandbed region. In order to more accurately determine the buckling capability of the drywell, a three dimensional finite element model is developed.

The geometry of the Oyster Creek drywell is shown in Figure 3-1.Taking advantage of symmetry of the drywell with 10 vents,- a 36 section is modeled. Figure 3-2 illustrates the finite element model of the drywell. This model includes the drywell shell from the base of the sandbed region to the top of the elliptical head and the vent and vent header. The torus is not included in this model because the bellows provide a very flexible connection which does not allow significant structural interaction between the drywell and torus.Figure 3-3 shows a more detailed view of the lower section of the drywell model. The various colors on Figures 3-2 and 3-3 represent the different shell thicknesses of the drywell and vent. Nominal or as-designed thicknesses, summarized in Table 3-1, are used for the drywell shell for all regions other than the sandbed region. The sandbed region shown in blue in Figure 3-3 is considered to have a thickness of 0.736 inch. This is the 95% confidence projected thickness to outage 14R. Figure 3-4 shows the view from the inside of the drywell with the gussets and the vent jet deflector.

The drywell and vent shell are modeled using the 3-dimensional plastic quadrilateral shell (STIF43) element. Although this element' has plastic capabilities, this analysis is conducted using only elastic behavior.

This element type was chosen over the elastic quadrilateral shell (STIF63) element because it is better suited for modeling curved surfaces.3-2

-RF#0 REV. 0 At a distance of 76 inches from the drywell shell, the vent is"...simplified using beam elements.

The transition from shell to bea.m elements is made by extending rigid beam elements from a node along the centerline of the vent radially outward to each of the shell nodes of the vent. ANSYS STIF4 beam elements are then connected to this centerline node to model the axial and bending stiffness of the vent and header. Spring (STIF14) elements are used to model the vertical header supports inside the torus. ANSYS STIF4 beam elements are-also used to model the stiffeners in the cylindrical region of the.upper drywell. The section properties of these stiffeners are summarized in Table 3-2.3.3 Drywell Materials.

The drywell shell is fabricated from SA-212, Grade B -high tensile strength carbon-silicon steel plates for boilers and other pressure vessels ordered to SA-300 specifications.

The mechanical properties for this material at room temperature are shown in Table 3-3. These are the properties used in the finite element analysis.

For the perforated vent jet deflector, the material properties were modified to account for the reduction in stiffness due to the perforations.

3.4 Boundary Conditions Symmetric boundary conditions are defined for both edges of the 36'drywell model for the static stress analysis as shown on Figure 3-5.This allows the nodes at this boundary to expand radially outward from.the drywell centerline and vertically, but not. in the circumferential direction.

Rotations are also fixed in two 'directions to prevent the boundary from rotating out of the plane of symmetry.

Nodes at the.bottom edge of the drywell are fixed in all directions to.simulate.the fixity of the.shell within the concrete foundation.

Nodes at the end of the header support spring elements are also fixed.3-3 DRF# 00664 INDEX 9-4, REV. 0 3.5 Loads The loads are applied to the drywell finite element model in the manner which most accurately represents the actual loads anticipated on the drywell. Details on the application of loads are discussed in the following paragraphs.

3.5.1 Load Combinations All load combinations to be considered on the drywell are summarized on Table 3-4. The most limiting load combinations

'in terms, of-possible buckling are those which cause the most compressive stresses in the sandbed region. Many of the design basis load combinations include high internal pressures which would create tensile 'stresses in the shell and help prevent buckling.

The most severe design load combination identified for the buckling analysis of the drywell is the refueling condition (Case IV). This load combination consists of the'following loads: Dead weight of vessel, penetrations, compressible material, equipment supports and welding pads.Live'loads of welding pads and equipment door Weight of refueling water External Pressure of 2 psig Seismic inertia and deflection loads for unflooded condition The normal operation condition with seismic is very similar to.this condition, however, it will be less severe due to the absence of the refueling water and equipment door weight.The most severe load combination for the emergency condition'is f or'the post-accident (Case VI) load combination including:

3-4 DRF# 00664 INDEX 9-4, REV. 0 Dead weight.of vessel, penetrations, compressible material and equipment supports Live load of personnel lock Hydrostatic Pressure of Water for Drywell Flooded to 74'-6" External-Pressure of 2 psig Seismic inertia and deflection loads for flooded condition The application of these loads is described in more detail in the following sections.3.5.2 Gravity.Loads The gravity loads include dead weight loads of the drywell shell, weight of the compressible material and penetrations and live loads.The drywell shell loads are imposed on the model by. defining thI;e weight density of the shell material and applying a vertical acceleration of .'O g to simulate gravity. The. ANSYS -program.automatically distributes the loads consistent with the mass and acceleration.

The compressible material weight of 10 lb/ft2 is added by adjusting the weight density of the shell to also include the compressible material.

The adjusted weight densities for the various shell thicknesses are summarized in Table 3-5. The compressible material is assumed to cover the entire drywell shell (not including the vent) up to the elevation of the flange.The additional dead weights, penetration weights and live loads are applied as additional nodal masses to the model. As shown on Table 3-6 for the refueling case, the total additional mass is summed for each 5 foot elevation of the drywell. The total is then divided by I0.for the 36 section assuming that the mass is evenly distributed around the perimeter of the drywell. The resulting mass is then applied uniformly to a set of nodes at the desired elevation as shown on Table 3-6. These applied masses automatically impose gravity loads on the drywell model with the defined acceleration of Ig. The same method is used to apply the additional masses to the model for the post-accident case as summarized in Table 3-7.3-5 DRF# 00664 INDEX 9-4, REV. 0 3.5.3 Pressure Loads The 2 psi external pressure load for the.refueling case is applied to the external faces of all of the drywell and vent shell elements.

The compressive axial stress at the transition from vent shell. to. beam*eTements is simulated by applying equivalent axial forces-to the nodes of the shell elements.Considering the post-accident case, the drywell is assumed to be flooded to elevation 74'-6" (894 inches). Using a water density of 62.3 lb/ft 3 (0.0361 lb/in 3), the pressure gradient versus elevation is calculated as shown. in Table 3-8. The hydrostatic pressure at the bottom of the sandbed region is calculated to be 28.3.psi.

According to the elevation of the element centerline, the appropriate.pressures are applied to the inside surface of the shell elements.3.5.4 Seismic Loads Seismic stresses have been calculated for the Oyster Creek Drywell in Part I of this report, Reference 3-3. Meridional stresses are imposed on the drywell during a seismic event due to a 0.058" deflection of the reactor building and due to horizontal and vertical inertial loads on the drywell.The meridional stresses due to a seismic event are imposed on the 3-D drywell model by applying downward forces at four elevations of the'model (A: 23'-7',B:

37'-3",C:

50'-11" and D: 88'-9") as shown on Figure 3-6. Using this method, the meridional stresses calculated in Reference 3-3 are duplicated at four sections of the drywell including 1) the mid-elevation of the sandbed region, 2) 17.25* below "the equator, 3) 5.75* above the equator and 4) just above the knuckle region. These four sections were chosen to most accurately represent the load distribution in the lower drywell while also providing a reasonably accurate stress distribution in the upper drywell.3-6.. .

• RF# 00664 INDEX 9-4,RE To find the correct loads to match -the seismic stresses, the total seismic stress (due to reactor building deflection and horizontal and vertical inertia) are obtained from Reference 3-3 at the four sections of interest.

The four sections and the corresponding meridional stresses for the refueling and post-accident seismic cases are summarized in Table 3-9.Unit loads are then applied to the 3-D model in separate load steps at each elevation shown in Figure 3-6. The resulting stresses at the four sections of interest are then averaged for each of the applied unit loads. By solving four equations with four unknowns, the correct loads are determined to match the stresses shown in Table 3-9 at the four sections.

The calculation for the correct loads are shown .on Tables 3-10 and 3-11 for the refueling and. post-accident cases, respectively.

3.6 Stress Results The resulting stresses for the two load combinations described in section 3.5 are summarized in this section.3.6.1 Refueling Condition Stress Results The resulting stress distributions for the refueling condition are shown in Figures 3-7 through 3-10. The red colors represent the most tensile stresses and the blue colors, the most compressive.

Figures 3-7 and 3-8 show the meridional stresses for the entire drywell and lower drywell. The circumferential stresses for the same areas are shown on Figures 3-9 and 3-10. The resulting average meridional stress at the mid-elevation of the sandbed region was found to be;ORmi -7580 psi* 3-7 YNDEX 8-4, REV. 0 The circumferential stress averaged from the bottom to the top of the sandbed region is;°Rc 4490 psi 3.6.2 Post-Accident Condition Stress Results The application of all of the loads described for the post-accident.

condition results in the stress distributions shown in Figures 3-11 through 3-14. The red colors represent the most tensile stresses and.the blue colors, the most compressive.

Figures 3-11 and 3-12 show the meridional stresses for the entire drywell .and lower drywell..

The circumferential stresses for the same areas are shown.on Figures 3-13 and 3-14. The resulting average meridional stress at mid-elevation of the sandbed region was found to be;0 pAm "-11960 psi The circumferential stress averaged from the bottom to the top of the sandbed region is;UPAC s+20080 psi 3-8 RDF# 800664, R NDEX 4, REV.3.7 Theoretical Elastic Buckling Stress Results After completion of the stress runs for the Refueling and Post-Accident load combinations, the eigenvalue buckling runs are made as described in Section 3.1. This analysis determines the theoretical elastic buckling loads and.buckling mode shapes. .3'.7.1 Refueling Condition Buckling Results -As shown on Figure 3-15, it is possible for the drywell to buckle in two different modes. In the case of symmetric buckling shown on Figure 3-15, each edge of the 36' drywell model experiences radial displacement with no rotation.

This mode is simulated by applying symmetric boundary conditions to the 3-D model the same as used for the stress run. Using these boundary conditions for the refueling' case, the critical load factor was found to be 7.67 with the critical buckling occurring in the sandbed region. The critical buckling mode shape is shown in Figure 3-16 for symmetric boundary conditions.

The red color indicates sections of the shell which displace radially outward and the blue, those areas-which displace inward.The first four buckling modes were computed in this eigenvalue' buckling analysis with no buckling modes found outside the sandbed .region for a load factor as high as 9.94. Therefore, buckling is not a concern outside of the sandbed region.It is also possible for the drywell-to buckle in the asymmetric manner shown in Figure 3-15. For this mode, the edges of the.3-D model are, allowed to rotate but are restrained from expanding radially.

This case ts considered by applying asymmetric boundary conditions at the edges of the 3-0 model. With the two pass approach used by ANSYS, it is possible to study asymmetric buckling of -the drywell when the'stresses are found based on symmetric boundary conditions.

The resulting.

load factor found using asymmetric boundary conditions is 10.13. The mode shape for this case is shown.on Figure 3-17...3-9 DRF# 00664 INDEX 9-4, REV. 0 Because the load factor is lower for symmetric boundary conditions with the same applied stress, the symmetric buckling condition is more limiting.

Multiplying the load factor of 7.67 by the average meridional stress from section 3.6.1, the theoretical elastic buckling stress is found to be;ORie = 7.67 x (7580 psi) 58,100 psi 3.7.2 Post-Accident Condition Buckling-Results Considering the post-accident case with symmetric boundary conditions, the load factor was calculated as 5.18. Multiplying this load factor by the applied stress from section 3.6.2 results in a theoretical elastic buckling stress of aPAie = 5.18 x (11960 psi) 61,950 psi The critical mode shape for this condition is shown in Figure 3-18.Again, the critical buckling mode is in the sandbed region..3.8 References 3-1 DeSalvo, G.J., Ph.D, and Gorman, R.W., "ANSYS Engineerincg Analysis System User's Manual, Revision 4.4," Swanson Analysis Systems, Inc., May 1, 1989.3-2 GPUN Specification SP-1302-53-044, Technical Specification for Primary Containment-Analysis

-Oyster Creek Nuclear Generatinc Station; Rev. 2, October 1990.3-3 "An ASME Section VIII Evaluation of the Oyster. Creek Drywell -Part 1 Stress Analysis," GE Report No. 9-1, DRF #-00664, November 1990, prepared for GPUN.3-10 Table 3-1.Oyster Creek Drywell Shell Thicknesses RF# 00664 INDEX 9-4, REV. 0 Section Sandbed Region Lower Sphere Mid Sphere Upper Sphere Knuckle Cylinder Reinforcement Below Flange Reinforcement Above Flange Elliptical Head Ventline Reinforcement Gussets Vent Jet Deflector Ventline Connection Upper Ventline Lower Ventline Thickness (in.)0.736 *1.154 0.770 0.722 2.5625 0.640 1.250 ..1.500 1.1875: 2.875 0.875 2.500 2.500 -0.4375 0.250* 95% confidence projected thickness to 14R.3-11 i DRF# 00664 INDEX 9-4, REV. 0 Cylinder Elevation (in)966.3 1019.8 1064.5 1113.0(1)1131.0 (1) -Table 3-2 Stiffener Locations and Sec Height Width Area (in) (in) (in2)0.75 6.0 4.5 0.75 6.0 4.5 0.50 6.0 3.0 2.75 7.0 26.6 1.00 7.38 1.0 12.0 12.0 This stiffener is made up 2.75x7" and one 1.0x7.375" tion Properties Bending Inertia (in 4 Horizontal -Vertical 13.5 0.211 13.5 0.211 -9.0 0.063 387.5 12.75 144.0 of 2 beam 1.000 sections, one Table 3-3 Material Properties for SA-212 Grade Material Propertv B Steel Val ue 29.6x10 6 psi 38000 psi 0.3 0.283 lb/in 3 Young's Modulus Yield Strength Poisson's Ratio Density 3-12 RF# 00664*NDEX 9-4, REV. 0 Table 3-4 Oyster Creek Drywell Load Combinations CASE I -INITIAL TEST CONDITION Deadweight

+ Design Pressure (62 psi) + Seismic (2 x.DBE).CASE II -FINAL TEST CONDITION' Deadweight

+ Design Pressure (35 psi) + Seismic (2 x DBE)CASE III -NORMAL OPERATING CONDITION Deadweight

+ Pressure (2 psi external)

+ Seismic (2 x DBE)CASE IV -REFUELING CONDITION Deadweight

+ Pressure (2 psi external)

+ Water Load +Seismic (2 x DBE)CASE V -ACCIDENT CONDITION Deadweight.+

Pressure(62 psi @ 1757F or 35 psi @ 281"F) +Seismic (2 x DBE)CASE VI -POST ACCIDENT CONDITION Deadweight

+ Water Load @ 74'6" + Seismic (2 x DBE)3-13 YRF# 00664 NDEX 9-4, RE'l. 0 Table 3-5 Adjusted Weight Densities of Shell to Account for Compressible Material Weight Shell Thickness (in.)Adjusted Weight Density (1 b/in 3)1.154 0.770 0.722 2.563 0.640 1.250 0.343 0.373 0.379 0.310 0.392 0.339 3-14 TRF#X00664 NOEX 9-4, REV. 0 Table 3-6 Oyster Creek Drywell Additional Weights -Refueling Condition ELEVATION (feet)15.56 16 20-- 15-20 229 21-25S 26 30 30.25* 26-30 31 32 33 34 35*- 31-35 36 40* 36-40 501*- 45-SO5 54 51-55 56 60*- 56-60 65 61-65 70* 66-70 73 71-75 82.17* 81-85 87 90* 86-90 93.75 94.759 95.75* 91-96 TOTALS: DEAD YE16HT (lbf)50000 556000 64100 105000 41000 1102000 56400 95200 52000 PEN YE (11___.161 1: I2 1!21 ETR. MISC. TOTAL I6HT LOADS LOAD bf) (lbf) (lbf)50000 5100 168100 1200 11200 556000 1100 11100 l500 115600 100000 205000 6500 16500 750 750 5450 15450 8050 28050 1500 1500 LSSO 1550 3350 84350 1102000 7850 7850 24000 80400 700 20000 115900 20000 72000 5750 5750 8850 8850 21650 1000 1000 5000 15000 20700 698000 698000 20100 5200 862000 3434350 5 FOOT RANGE LOAD 229300 556000 LOAD PER 36 DES.(lbf)________t OF ELEMENTS_ ___ __NODES OF APPLICATION

___________

.116-119 161-169 LOAD PER FULL NODE (lbf)_________L0OD PER HALF NODE (lbf)I PM._22930 55600 6 a 3822 1911 6950 *3475 331700 33170 62250 6225 21650--1!20700 20100 21U4150 381 85900 1102000 7850 196300 72000 5750 8850 21650 16000 738800 3434350_3434350 .8590 110200 785 19630 7200 57S 885 2165 1600 73880 343435__343435 8 179-187 S 188-196 S .197-205 8 418-426 8 436-444 8 454-462 8 472-480 a 508-516 8 526-534 8 553-561 a 571-579 8 -589-597 4146 778 1074 13775 98 2454 900 72 111 271 389 537 6888 49 1227*450* 36' 55 135 2073 .200 10 9235 4618 F -LOAD TO BE APPLIED IN VERTICAL OIRECTION ONLY.& -MISCELLANEOUS LOADS INCLUDE 698000 LB WATER WEIGHT AT 94.75 FT. ELEVATION 100000 LB EQUIPMENT DOOR WEIGHT AT 30.25 FT. ELEVATION AND WELD PAD LIVE LOADS OF 24000. 20000 AND 20000 AT 56, 60 AND 65 FT. ELEVATIONS REFWGT.WK1 3 -15 a RF# 00664 INDEX 9-4, REV. 0 Table 3-7 Oyster Creek Drywell DEAD PENETR. MISC.WEIGHT WEIGHT LOADS (lbf) (Ibf) (lbf)________ --------- --------Additional Weiqhts -Post-Accident Condition..ELEVATION (feet)15.56 16 20 15-20 229 21-25J 26 30 30.25* 26-30 31 32 33 34 35* 31-35 36 40-36-40 509^ 45-50t 54* 51-55 56 60*- 56-60 65* 61-65 70* 66-70 73 e 71-75 82.17* 81-85 87 90* 86-90 93.75 95.75-- 91-96 TOTALS: 50000 168100 11200 556000 64100 105000 41000 1102000 56400 95200 52000 11100 51500 16500 750 15450 28050 1500 1550 43350 7850 700 TOTAL LOAD (lbf)50000 168100 11200 556000 11100 115600 105000 16500 750 15450 28050 1500 1550 84350 1102000 7850 56400 95900 52000 5750 B8S0 21650 1000 15000 20700 20100 5 FOOT RANGE LOAD 229300 556000 LOAD PER 36 DEG.(lbf)________22930 55600 6 a* OF ELEMENTS________NODES OF APPLICATION

• 116-119* 161-169 LOAD PER LOAD PER FULL NODE IALF NODE (lbf) (lbf)_________

-- ---------3822 6950 2896 1911 3475 1448 231700 23170 62250 6225 8 179-187 85900 1102000 7850 152300 52000 5750 8850 21650 8590 110200 785 15230 5200 575 885 2165 8 188-196 8 197-205 8 418-426 8 436-444 8 454-462 8 472-480 8 508-516 8 526-534 8 553-561 8 571-579 8 589-597 1074 13775 98 778 389 537 6888.-49 5750 S850 21650 1000.15000 1904 952 650 325 72

• 36 111 55 271 135 200 100 510 255 16000 1600 20700 20100 40800 2184150 388200 0 2572350 2572350 I -LOAD TO BE APPLIED IN VERTICAL DIRECTION ONLY.& -NO MISCELLANEOUS LOADS FOR THIS CONDITION.

4080 257235__257235 FLOOOGT .W:1 3-16 YRF#X 0664 NDEX 9-4, REV. o Table 3-8 Hydrostatic Pressures for Post-Accident, Flooded Condition WATER DENSITY: FLOODED ELEV:* 62.32 lb/ft3 0.03606 lb/in3 74.5 ft 894 inches ELEMENTS ABOVE NODES 27 40 53 66 79'92 102 lOB 112 116 12D 124 13D 133 14B 161 17D 179 183 197 401)40)413 427 435 445 454 46:3 47;2 481 491)491;5013 517 521'ANGLE ABOVE EQUATOR (degrees)-53.32-51.97-50.62-49.27-47.50-46.20-44.35-41.89-39.43-36.93-34.40-31.87-29.33-26.80-24.27-20.13-14.38-8.63-2.88 2.88 8.63 14.38 20.13 25.50 30.50 35.50 40.50 45.50.50.50 54.86 ELEVATION (inch)110.2 116.2 122.4 128.8 137.3 143.9 153.4 166.6 180.2 194.6 209.7 225.2 241.3 257.6 274.4 302.5 342.7 384:0 425.9 468.1 510.0 551.3 591.5 627.8 660.2 690.9 719.8 746.6 771.1 790.5 805.6 820.7 835.7 850.8 885.3 DEPTH (inch)783.8 777.8 771.6 765.2 756.7 750.1 740.6 727.4 713.8 699.4 684.3 668.8 652.7 636.4 619.6 591.5 551.3 510.0 468.1 425.9 384.0 342.7 302.5 266.2 233.8 203.1 174.2 147.4 122.9 103.5 88.4 73.3 58.3 43.2 8.7 706.7 PRESSURE (psi')28.3 28.1 27.8 27.6 27.3 27.1 26.7 26.2 25.7 25.2 24.7 24.1 23.5 23.0 22.3 21.3 19.9 18.4 16.9 15.4 13.8 12.4 10.9 9.6 8.4 7.3 6.3 5.3 4.4 3.7 3.2 2.6 2.1 1.6 0.3 25.5 ELEMENTS-1-12 13-24 25-36 37-48 49-51, 61-66 ,55-57 52-54, 138-141 ,56-60 142-147, 240-242, 257-259' 148-151, 243, 256 152-155, 244, 255 156-159, 245, 254 160-165, 246, 253 166-173, 247, 252 174-183, 248-251 184-195 196-207 208-215 216-223 224-231 232-239 430-437 438-445 446-453: 454-461 462-469 470-477 478-485 486-493.494-501-502-509 510-517 518-525 526-533 534-541 542-549-550-557 340-399 (Ventl in e)187.3 FLOOOP.WK1 3-17 F DRF#X0664 00 INDEX 9-4, REV. 0 Sectic A) Middle B) 17.25'C) 5.75 /D) Above K of Bel kbov Knuc Table 3-9 Meridional Seismic Stresses at 2-D Shell Elevation Model an (inches) Node Sandbed 119 32 ow Equator 323 302'e Equator 489 461 kle 1037 1037 Refueling (gsi )1258 295 214 216 Four Sections Meridional Stresses..Post-Accident

• (si) ..1288 585 616 808 3-18

• RF# 00664 YNDEX 9-4, REV. 0 Table 3-10 Application of Loads to Match Seismic Stresses -Refueling Case 2-D SEISMIC STRESSES AT SECTION (psi)____._________________--------..-.

-..----_SECTION: 2-0 NOOE: ELEV: COMPRESSIVE STRESSES FROM 2-0 ANALYSIS 0.058" SEISMIC DEFLECTION:

HORIZ. PLUS VERTICAL SEISMIC INERTIA: TOTAL SEISMIC CO0PRESSIVE STRESSES: 1 32 119.3" 788.67 469.55 2 302 322.5" 155.54 139.44 3 461 489.1" 103.46 110.13 4 1037 912.3" 8S.31 130.21 1258.22 294.98 213.59 215.52 3-0 STRESSES AT SECTION (psi)_____.._______________._._____..__-

3-D INPUT LOAD SECTION___;_._A C 0 3-D INPUT LOAD SECTION-------A B C D SECTION: 3-0 NODES: ELEV: INPUT 3-D UNIT LOAD DESCRIPTION 1000 lbs at nodes 563 through 569 500 lbs at 427U435. 1000 lbt at 428-434 500 lbs at 1971205. 1000 lbs at 198-204 500 lbs at 161I169. 1000 lbs at 162-168 DESIRED COMPRESSIVE STRESSES (psi): LOAD TO BE APPLIED TO MATCH 2-0 STRESSES 3902.2 2101.4 1453.6 6611.6 1 53-65 119.3" 85.43 89.88 97.64 89.85I 2 170-178 322.5S-______37.94 39.92 43.37 0.00 3 400-408 489.1'34.94 36.76 0.00 0.00.4.526-534 912.3" 55. Z3 0. DO 0.DO O.DO 1258.22 294.98 213.59 215.52 RESULTIN6 STRESSES AT SECTION (psi)333.37 148.05 136.34 215.52-168.87 83.89 77.25 0.00 141.93 63.04 0.00 0.00 594.05 0.00 0.00 0.00 1258.22 294.98 213.59 215.52 SUN: iEISUNFL.WK1 3-19 DRF# 00664 INDEX 9-4, REV. 0 Table 3-11 Application of Loads to Match Seismic Stresses -Post-Accident Case 2-0 SEISMIC STRESSES AT SECTION (pil).SECTION: 2-D NODE: ELEV: COMPRESSIVE STRESSES FROM 2-D ANALYSIS_________________________________________

0.058" SEISMIC DEFLECTION:

HORIZ. PLUS VERTICAL SEISMIC INERTIA: TOTAL SEISMIC COMPRESSIVE STRESSES:-1 .2 3 4 32 302 461 1037 119.3" 322.3" 489.1" 912.3'788.67 155.54 103.46 85.31 499.79 429.39 512.76 723.14 128B.46 584.93 616.22 808.45 3-D STRESSES AT SECTION (psi)_____________.___________________

_3-0 INPUT I.OAD ST.CTION_.._____-A C 0 SECTION: 3-D NODES: ELEV: INPUT 3-D UNIT LOAD DESCRIPTION 1000 lbs at nodes 563 through 569 500 lbs at 427U435, 1000 lbs at 428-434 500 lbs at 1971205. 1000 lbs at 198-204 500 lbs at 161U169. 1000 lbs at 162-168 I 53-65 119.3" 85.43 89.88 97.64 89.85--2 170-178 322.S5 37.94 39.92 43.37'0.00 3 400-408 489.1" 34.94 36.76 0.00 0.00 4 526-S34 912.3" 55.23-'0.00 0.00 0.00 1288.46 584.93 616.22 808.45 DESIRED COMPRESSIVE STRESSES (psi): 3-D INPUT I.OAD SuCTION A B C D LOAD TO BE APPLIED TO MATCH 2-0 STRESSES* RESULTING STRESSES AT SECTION (psi)_________________

________________

14637.9* 2850.2-1941.7-318.8 1250.51 256.17-189.58-28.64 1288.46 555.36 113.78-84.21 0.00 584.93 511.45 104.77 0.00 0.00 616.22 808.45-0.00 0.00.0.00 808.45 SUN: SEISFL.WK1 3-20_ r __ENM~

RF# 00664 YNDEX 9-4, REV. 0 DRYWELL i ELEV.5 1'I Pt THK .6767 Oyster Creek Drywell Geometry Figure 3-1.3-21 1 AI4SYS 4.4 DEC 4 1990 IM:96: el R~EAL HUM 5u=1 XI' =383.931 21' =639.490 QNG'Z--90 CENITROID HZDVEH W I1.J OYSTER CRMEE DWYWELL AHA~LVSIS-OWCR+/-O CHO SAND, POST-ACC--; ..Fiure 3-2. Oyster Creek Orwell 3-0 FinitelElement Model I ft4SVS 4.4 DEC 4 1990 REAL HUM MU =1 DISTO~s'se.376 XT =420.452 CENRO~lD HIDDEN'I I'J'a IRNN .% ...I.I.OYSTER CREE4 1DRYWELL ANALYSIS -OYCR.LO CHO SAHD, FOST-ACC.I-Figure 3-3. Closeup of Lower Drywell Section of FEM (Outside Viewl 1 DEXC V4 19.940;LOTHNO. 3)REAL HUM)CU =-I.Yu =-1e.DISI=288 .376 X(F =428.432 23 =21.6.328 aNC2,.-90 CEtITROID HIDDENI 4W 1.1 I I OYSTER CREEX DRYWELL ANA~LYSIS

-QYCUJSO CHO SANDSI POST-ACC.Fioure 3-4. Clnoeu of loC:wer Drywel ISection of FFM lTnside View)-.

1 SNSyS 4.4 DEC 4 1990 15:18:37 PLOT NO 6 PREP? iLEMENTS TYPE HUM BC SYMBOLS)(U =1 YU =--.8 DIST=718 .786 Xi' =383.031 2W =639.498 AHGi=--m w t4a ch OYSTER CREEI(DRYWEL]L SC S L ANALY SIS -.OCR{FLO -XHNO SoAND, POST-ACC.;I Fioure 3-5.Boundary Conditions of Finite Element Model 1 I w hNSYS 4.4 OCT 15 1990 89:32:26 PLOT HO. 2 PREP7 XLDIENKTS TYPE HUM BC ScMBOw yuu =--. 8 DIST=718 .786 XF =363.031 ZF =639.498 AHGZ=-90 Crt4ROID HIDDEN SE).OYSTER CREEX DRYWELL ANALYSIS -OYCR+/-A (SAHD. UNIT LOAD CA...I; FWintira tof. -Annlicatinn nf Inmdinn tn Cimolato Seismic R ndinr I 44"""O r.pi #'molts wood W**'As eptso 1001,000 016888#40 f fit%frevorp NOV16~ 1999j;LoT HO.STE?=l ITEP,=l SY (AfUG)o DUIX =0.221779 SHN =-8174 SMX =695.B47 XV =1 YU =-6.8 DI ST=71. 796 X~ 363.931 CENT~OI H IDDEN-8174-7198-3247-1276"95.947-OYSTER CREEX DRYWELL ANALYSIS -. OYCRJS 0NO SAND, BEFMELINGj)

..... I ..........I ...; Figure 3-7.: Meridional Stresses -Refueling Case POTI 'STRESS STEP--I I TER~1* *,SY (oU GI)S* S*~i*g a I 3DMX =9.139473 6MM =-8174 SPIX =695.947 XU =1.YLU =-9. 8 DIST=288.376 ClMIROID HIDDEN-8174-7188-1276 695.947 OYSTER c2lEE1X DRYWELL'ANiALYSIS

-]SCO AD.REFUEL!

4... C;rvttva 1-9.m i niwar nlrvpil' mp'~dinnil

~tptow -cafl~ipiin rat C i NOU 16 1996 1 .O09:. s :39 STEP=1 ITEN=1..SX CSU DIX =0.221779 SHN =-3547 SMX =6754 XU =1 YU =-e.8 DIST=710 .786 XFr =33.931 CZHITROID HIDDEN-l -3547-2463 2176 4465 6754'a OYSTER CREEX DRYWELL: ANALYSIS OYCRS CHO SAND,'.REUELINCs.imirp 3-9. Circ.umferential Stresses Refuelin Case I *Uo- A. A 1NOU 16 199 99:57:15 STEP=ITERIl DMX =9.139473 SMN =-3547_a( 1ii._ .-6754 YU =-e.8 D.ST=20 ,376 XT =29.528 C~tITROID HIDDEN--.3547-2463 2176 4465 6754 OYSTER CREEX DRYNELL AIMLYSIS -OVCH+/-S CHO SAND. REFUEL!Figure 3-10. Lower DryWel Circumferential Stresses -Refueling Case

--- .U-4 1ah.NSHYS 4.4 L6:38:58 STEPil ITEJ01-SY C UC , IDLS DMX =9.479734 SHM =-131355 SMX =3894 XU =1 Yu =-e.s DI ST=78 10706 Xi =363.831 C2 =61&. 498.CDtIROID HIODDK_.-+/-3+/-55 E +/-11260-3683 6 95.136 3894 w OYSTER CREEX DRYWILL AIMLYSIS SAMD.' POST-ACC.:;. Fintirp 1-1 o Mprfidinnal StreapcSp

-pont-Acnident C~aip

[1 I. de I. I B I S £S.. S D I I S S **ANS¶YS 4.4 16:33:46 STEP=i.K TE~t=1l SY c Uc)l DIIX =0.479734 SHM =-13153 SMX =3894 XU =1 YU =-9.8 DIST=289 .376 XT =429.432 CENTROIlD HIDDEN-13+/-55-11260-3683 165.136 3994 k ic~ ~ R IW f%3 OYSTEi CaElM DRYWELL ANALYSIS -OYCRlO CHO SAND. POST-ACC., Figure 3-12. Lower Drywell Meridional Stresses -Post-Accident Case I i I AM-Ap.Vq A-*NWUI9 1993d 16:38:42 STEP=I tTEP.=l DMX =9.479734 SHN =-5205 SMX =2?791-xli =1 Yu =-eG8 DIST=710.*786 XT =303.831 Za2-6a.4 9 8 1 CENTRtOID HIIDDEN-5295-1538 13126 26459 27791 Ca~OYSTER CREEX DRYWXLL AALYS I S MtIMICPOST F rACc r .J .1- -. : Figure 3^13. Circumferential Stresses -Post-Accident Case.. )i 1*,HSYS 4-1.6:32: 33 ST EP=1 VLDOL&S DMX .=G. 479734 CNN =-5295 SMRX =27791 X(U =1 YU =-e.8a DI 87=288 *376 XT =42G.432-5205*-1538 13126 2G459 27791 OVSTZR CREEX DRYMELL AHALYSIS -OYiCR+/-O tHO SAND, POST-ACC.-Fiqure 3-14. Lower Drywell Circumferential Stresses -Post-Accident Case F# 00664 NDEX 9-4,- REV. 0.Unrbuned Shape Budded Shape -Vent (Racial Displacement

.No Rotation-3 Symmetric Buckling of Drywell-UnIbluded Shape' Bucked Shape ( Rotation I kNo Racial Diap.).Vent Asymmetric Buckling of Drywell SYM.DRW Figura 3-15.Symmetric and Asynnetric Buckling Modes 3-35 OkNSVS 4.4 HOU 16 1996£3:23:62 STEP=I ITEN=1 FOCT=7.665 ux D MODAL DNx =90.90381 SMm =-B.609986 SHX =9. 9614)(U =1 YU =-e. 8 DI 5T=106.0369

-0.603147-09.02486 D. 137X-Q3 0.8GIE-933 91.06214 OYSTER CREEX DRYWELL A1ALYSIS -OYCRIT CtUO SANDO REIrUEL 4 Fiqure 3-16. Symmetric Buckling Mode Shape -Refueling Case I AHV 4. i.4.DEC 11 1999 08:21:39 LSOTI LSTmmS STEP=1 ITERN-1 FACT=I. 134 D NODAL DPlX =9 .083744 SMN =-0.99+/-52 SMX =9.093744 XU =1 YU =-es8 DIST=186 .859 XF =327.422-6.359K-Wi 9.912574.90 I3744 L.Ca)(A)OYISTER C3EEX( DRYWEL1L AHALYSIS -OYCR+/-CC (No SAND,. HEPuELjtI-Fioure 3-17. Asynmetric Bucklinq Mode ShaDe -Refuelina Case

• "SYS 4.4 NOU 20 1990.8:27:14 STEP=l I TER-1 FACT=5.181 ux D NODAL DMX =9.62868 SH. =-6.081173 SM~C =8.992867_XU =1 YU =-8.a DI S=186 .859 XJ =327.422 B =z2W3 128 m I.-.724E3-3 .2753-93 w .. _B 6 9.092867 CD 'OYSTER CRERIX DRYWELL AtMALYSIS -OVCRPM CHO SAND. PS-C-I .3-18.. RlIrklina Mnd1 ghmna -Dne+ fiR48an^F rien DRF# 00664 INDEX 9-4, REV. 0 4. ALLOWABLE BUCKLING STRESS EVALUATION Applying the methodology described in Section 2 for the modification of the theoretical elastic buckling stress, the allowable compressive stresses are now calculated.

Tables 4-1 and 4-2 summarize the calculation of the allowable buckling stresses for the Refueling and Post-Accident conditions, respectively.

The modified capacity reduction factors are first calculated as described in sections 2.2 and 2.3. After reducing the.theoretical instability stress by this reduction factor, the plasticity reduction factor is calculated and applied. The-resulting inelastic buckling stresses are then divided by the factor of safety of 2.0 for the Refueling case and 1.67 for the Post-Accident case to obtain the final allowable compressive stresses.The allowable compressive stress for the Refueling case is 10.65 ksi.Since the applied compressive stress is 7.58 ksi, there is a 4]% -margin. The allowable compressive stress for. the Post-Accident, flooded case is 13.77 ksi. This results in a margin of 15% for thee applied compressive stress of 11.96 ksi.4-1

• WEX4 4REX. 1 Table 4-1 Calculation of Allowable Buckling Stresses -Refueling Case Parameter Value Theoretical Elastic Instability Stress, aie (ksi)Capacity Reduction Factor, Ard Circumferential Stress, ac (ksi)Equivalent Pressure, p (psi)"X" Parameter AC Modified Capacity Reduction Factor, ai.,mod Elastic Buckling Stress, a. ' ai,mod ate (ksi)Proportional Limit Ratio, A a Plasticity Reduction Factor, Adi Inelastic Buckling Stress, °; -nie' (ksi)Factor of Safety, FS Allowable Compressive Stress, 0 all -1/FS (ksi)Applied.Compressive Meridional Stress, am (ksi)Margin -[((all/am)

-1] x 100%58.10 0.207 4.49 15.74 0.0865 0.0716 0.325.* 18.88 0.497 1.00 18.88 2.0.9.44 7.58 24.5%4 DRF# 00664 INDEX 9-4, RE:V. 1 Table 4-2 Calculation of Allowable Buckling Stresses -Post-Accident'Case Parameter

'Value Theoretical Elastic Instability Stress, orie (ksi) 61.95 Capacity Reduction Factor, Ad ' 0.207 Circumferential Stress, ac (ksi) 20.08 Equivalent Pressure, p (psi) 70.38"X" Parameter 0.387 AC 0.182 Modified Capacity Reduction Factor, aimod 0.508 Elastic Buckling Stress, ae ' li mod aie (ksi) 31.47 Proportional Limit Ratio, A -aelay 0.828 Plasticity Reduction Factor, Ad 0.724 Inelastic Buckling Stress, ai NMe (ksi) 22.78 Factor of Safety, FS 1.61 Allowable Compressive Stress, call -a/FS (ksi) '13.64 Applied Compressive Meridional Stress, am (ksi)' 11.96 Margin -[(aall/om)

-1] x 100% 14%4-3 DRF 0066 INDEX 9-4, REV. 0 5.

SUMMARY

AND CONCLUSIONS The results of this buckling analysis for the refueling and post-accident load combinations are summarized in Table 5-1. The applied and allowable compressive meridional stresses shown in Table 5-1 are for the sandbed region which is the most limiting region in terms of-buckling.

This analysis demonstrates that the Oyster Creek drywell has adequate margin against buckling with no sand support for an assumed sandbed shell thickness of 0.736 inch. This thickness is the 95% confidence projected thickness for. the 14R outage..5-1 --

E -4, REV. 1 Table 5-1 Buckling Analysis Summary Load Combination Refuel inQ Service Condition Design Factor of Safety Applied 2.00 Applied Compressive Meridional.Stress (ksi) 7.58 Allowable Compressive Meridional Stress (ksi) 9.44 Buckling Margin 24.5%Post-Accident Level C 1.67-11.96 13.64 14.0%5-2