ML061020616

Letter Forwarding, ASME Section Viii Evaluation of the Oyster Creek Drywell, Part 2, Stability Analysis, Attachment 3 to Letter Dated 04/07/2006

ML061020616
Person / Time
Site:	Oyster Creek
Issue date:	01/16/1992
From:	Devine J GPU Nuclear Corp
To:	Document Control Desk, Office of Nuclear Reactor Regulation
References
2130-03-20289, 5000-92-2093, C321-92-2008, TAC MC7624 DRF # 00664, Rev 0
Download: ML061020616 (128)

Text

ATTACHMENT 3 (GPU Letter to NRC dated January 16, 1992)

6%A GPU Nuclear Corporation HNulear One Upper Pond Road ParsPny. New Jersey 07054 201-316-7000 TELEX 136-482 Writer's Dkect Del Number 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 WE' 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.

32192008. LET

• GPU Nuclear Corporation is a subsidiary of General Public Utilities Corporation 11

p C321-92-2008 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, J~-. C. DeVine U Vice President and Director Technical Functions JCD/'RZ/pl p cc: Administrator, Region I Senior Resident Inspector Oyster Creek NRC Project Manager 32192008. LET

TRNDEXT 4-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 lNOEX 9-?, REV. 0)

AN ASME SECTION VIII EVALUATION OF THE OYSTER CREEK DRYWELL PART 2 STABILITY ANALYSIS Prepared by: e i?'

C.D. Frederickson, Senior Engineer Materials Monitoring &

Structural Analysis Services a.

Reviewed by: .

H.-S. Mehta, Principal Engineer Materials Monitoring &

Structural Analysis Services Approved by::

S. Ranganath, Manager Materials Monitoring &

Structural Analysis Services i

I

YW8EXo864 REV. I TABLE OF CONTENTS

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 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 3-10

4. ALLOWABLE BUCKLING STRESS EVALUATION 4-1

5.

SUMMARY

AND CONCLUSIONS 5-i iii

• xA I REV. 0 LIST OF TABLES Table Page No. Title Np.

3-1 Oyster Creek Drywell Shell Thickness 3-11 3-2 Cylinder Stiffener Locations and Section Properties 3-12 3-3 Material Properties for FBX-212B Steel 3-12 3-4 Oyster Creek Orywell 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 Case 3-17 3-9 Meridional Seismic Stresses at Four Sections 3-16 3-10 Application of Loads to Match Seismic Stresses - 3-IS Refueling Case' 3-11 Application of Loads to Match Seismic Stresses - 3-20(

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

E REV. 1 LIST OF FIGURES Figure Page No. Title No.

1-1 Drywell Configuration 1-2 2-1 Capacity Reduction Factors for Local Buckling of 2-8 Stiffened and Unstiffened Spherical Shells 2-2 Experimental Data Showing Increase in Compressive 2-9 Buckling Stress Due to Internal Pressure (Reference 2-6) 2-3 Design Curve to Account for Increase in Compressive 2-10 Buckling Stress Due to Internal Pressure (Reference 2-1.1) 2-4 Plasticity Reduction Factors for Inelastic Buckling 2-11 3-1 Oyster Creek Drywell Geometry 3-21 3-2 Oyster Creek Drywell 3-D Finite Element Model 3-22 3-3 Closeup of Lower Drywell Section of FEM (Outside View) 3-23 3-4 Closeup of Lower Drywell Section of FEM (Inside View) .3-24 3-5 Boundary Conditions of Finite Element Model 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 3-9 Circumferential Stresses - Refueling Case 3-29 v

-NEX0 REV. 1 LIST OF FIGURES Figure Page' No. Title N&.

3-10 Lower Drywell Circumferential Stresses - Refueling Case 3-31) 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 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 - Refueling 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

RF# 0066

?NDEX 9-2, REV. 0

1. INTRODUCTION 1.1 General To address local wall thinning of the Oyster Creek drywell, GPUN has planned to prepare a supplementary report to the Code stress report of record [1-1]. For convenience, the supplementary report is divided 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 drywell 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 used 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 # 9-0971, 1965.

1-2 'An ASME Section VIII Evaluation of the Oyster Creek Drywel.l," GE Report No. 9-1, ORF# 00664, November 1990, prepared for GPUN.

1-1

NDEX 00664 YRF# 9-2, REV. 0 x i, ELEVL L S -1L LLI' At, 5*6X, ELTY Figure 1-1 Drywell Configuration 1-2

-NEXT 8 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 compressive stress is evaluated in three steps.

In the first step, a theoretical elastic buckling stress, ale, is determined. This value may be calculated either by classical buckling equations or by finite element analysis. Since the d6ywell shell geometry i-s 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, cti, 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., 0ieal When the elastic buckling stress exceeds the proportional limit of the material, a plasticity reduction factor, ij, is used to account for non-linear material behavior. The inelastic buckling stress' for fabricated shells is given by nictigie-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 - l7iiaie/FS 2-1

DRF# 00664 NDEX 9-Z, 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, ct, 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, a: 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 tensilef circumferential stress has the effect of rounding the shell and reducing the effect of imperfections introduced during the fabricatiori.

and construction phase. 'A modification of the ai 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

HUAIE 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, -itis 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 ac 1 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, ap. The capacity reduction factor 1 can then be modified as follows:

-i,mod Ii+ ap 2-3

pp YAATx° , 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)(e1)Et/R

• (0.605)(0.207) Et/R Where,'0.605 is a constant, 0.207 is the capacity reduction factorai 1, 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 p (0.605)(Qj+ *p)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 (og/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)-

2-4 w

-NUEXS-,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, aimod' for the drywell geometry was obtained as follows:

ci,mod 207 + AC(sphere)/O.6 O5 2.4 Determination of Plasticity Reduction Factor When the elastic buckling stress exceeds the proportional limit of the material, a plasticity reduction factor, ij, is used to account for the non-linear material behavior. The inelastic buckling stress for fabricated shells is given by 7iatiae.i ' 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 eiaie/°y and ay is the material yield strength. Figure 2-4 shows the relationship in graphical form.

n: - 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

YREV. 1 2-2 Letter (1985) from C.D. Miller to P. Raju;

Subject:

Recommended Revisions to ASVIE 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 of1 GPUN.

2-6 Harris, L.A., et al, "The Stability of Thin-Walled Unstiffened e 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 T14 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.,

e e Z-0 MEEENWM__ M

RF# 0Q6J4 YNOEXo-, REV. I 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 Axial Tension," AIAA Journal, Vol. 1, No. 10, October 1963, pp.

2316-2319.

2-7 I

lG7X1'-11 REV. 1 of oC q 0.A 02 0.0 s .U 24 2a a 4 a 12

• e-thI' of.

Figure 2-1 Capacity Reduction Factors for Local Buckling Stiffened and Unstiffened Spherical Shells 2-8 1

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s/03' Figure 2-4 Plasticity Reduction Factors for Inelastic Buckling I

2-11

DRF# 00664 INDEX 9-2, 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, cap? ire developed and saved from this static analysis. A buckling pass is then run to solve for the eigenvalue or load factor, X, for which elastic buckling is predicted using the equation:

([K] + X [S]) (u) 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:

cr ' 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.

3-1

?RF# 00664

?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 dimensionallfinite 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 from 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

RF# 00664 INOEX 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 calculated by multiplying the foundation stiffness of the sand by the contributory area of each node in the sandbed region.

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

.RV . VNRVEXYT-

. REY 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

• DRF# 00664

• INDEX 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 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 10 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 1g. -The same 3-5 M

RF# 064 TNDEXTP Z, REV. 0 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 tcb the external faces of all of the drywell and vent shell elements. The compressive axial stress at the transition from vent shell to beani 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/ft3 (0.0361 lb/in3 ), 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 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 3-1) drywell model by applying downward forces at four elevations of the model (A: 23'-7",B: 37'-3",C: 50'-llP and D: 88-9") as shown on Figure 3-6. Using this method, the meridional stresses calculated 1:1 Reference 3-3 are duplicated at four sections of the drywell including

1) the mid-elevation of the sandbed region, 2) 17.25* below th '

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 I

~lRF# 0066

• NDEX 9-2, REV. 0 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 (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.

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; aRm - -7097 psi 3-7

_M

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

3-11 condition results in the stress distributions shown in Figures 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 The meridional stresses for the entire drywell and lower drywell.

3-13 circumferential stresses for the same areas are shown on Figures of and 3-14. The resulting average meridional stress at mid-elevation the sandbed region was found to be; up - -9693 psi the top of the The circumferential stress averaged from the bottom to sandbed region is; GPAc +4049 psi 3-8

RF# 004 TNDEX 98z REV. 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 1lad factor found using anti-symmetric boundary conditions is 16.81. The mode shape for this case is shown on Figure 3-17.

3-9

TNDEX0 l REV. 0 Because the load factor is lower for symmetry boundary conditions with the same applied stress, the symmetric buckling condition is more 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;

'Rie ' 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 conditions, 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 0PAie ' 9.91 x (9693 psi) a 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.

3-10 I

?NDEX ..2, REV. 0 Table 3 Oyster Creek Drywell Shell Thicknesses Section Thickness (in.)

Sandbed Region 0.700 Lower Sphere 1.154 Mid Sphere 0.770 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

-W

RDEX08 6-4 REV. 0 Table 3-2 Cylinder Stiffener Locations ai nd Section Properties Elevation Height Width Area Bending Inertia (in4)

(in) (in) -inn ) Ulnu iorizontal 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 12.0 112.0 144.0 1.000 (1) - This stiffener is maLde up of a 2 beam sections, one 2.75x7" and one I .0x7.375" Table 3-3 Material Properties for FBX-2128 Steel Material ProDertv Value Young's Modulus 29.6x106 .Psi Yield Strength 38000 psi Poisson's Ratio - 0.3 Density 0.283 lb/in3 3-12

-Emwm

RF# 00664 INOEX 9-2, 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 @ 175*F or 35 psi @ 281'F) +

Seismic (2 x DBE)

CASE VI - POST ACCIDENT CONDITION Deadweight + Water Load e 74'6" + Seismic (2 x DBE) 3-13

RF# 00664 INDEX 9-2, REV. 0 Table 3-5 Adjusted Weight Densities of Shell to Account for Compressible Material Weight Adjusted Shell Weight Density Thickness (in.) Olb/in3i 1.154 0.343 0.770 0.373 0.722 - 0.379 2.563 0.310.

0.640 0.392 4.

1.250 0.339 3-14

N1O4EX 9 , REV. 0

.Table 3-6 Oyster Creek Drywell Additional Weights - Refueling Condition a

DEAD PEIIER. MISC. TOTAL - 5 FOOT LOAD PER LOAD PER LOAD PER ELEVATION VEIGHT WEIGHT LOADS LOAD 36 DEG. # OF IfOOCS OF FULL NODE HALF NODE (fat) (lbf) (lbf) (lbf) . (lbf) LOAD (lbf) ELEMENTS APPLICATION Ilb?) 0Ibf) 15.56 50000 SOOOO 16 168100 168100 20 11200 11200

^ 15-20 229300 22930 6 116-119 3522 1911 2II 556000 SS6000

-- 21-259 SS6000 SS600 a 161-169 6950 3475 26 11100 11100 30 54100 51500 115600 30.25 105000 100000 205000

• 26-30 331700 33170 6 179-167 4146 2073 31 16500 16500 32 750 750 33 15450 15450 34 28050 28050 35 1500 1S00

• 31-35 62250 622S a 168-196 776 389 36 1550 1550 40 41000 43350 U4350 36-40 85900 3590 8 197-205 .1074 .537 SO# 1102000 1102000

• 45-509 1102000 110200 8 418-426 13775 6888 54 7850 7850

• 51-55 7850 785 8 436-444 96 .49 56 56400 24000 80400 60 95200 700 20000 115900 56-60 196300 19630 8 454-462 -2454 1227 6S* 52000 20000 72000

• 6l-65 72000 7200 8 472-480 -900 450 70 5750 5750 O 66-70 5750 575 8 508-516 72 36 73 8850 8850

• 71-75  : 11I 55 8850 88S 8 526-534 62.11 21650 21650

• 081-65 21650 2165 8 553-561 . 271 . 135 87 1000 1000 90 15000 15000

• 66-90 16000 1600 Sil5-S79 200 100 93.75 20700 . 20700 94.75. 698000 698000 95.75 20100 20100

• 91-96 738800 73UO 8 581-597 . 9235 4618 TOTALS: 21U4150 388200 862000 3434350 3434350 34343S

• - LOAD TO BE APPLIED IN VERTICAL DIRECTION ONLY.

& - MISCELLANEOUS LOADS INCLUDE 698000 LU WATER WE16HT AT 94.75 FT. ELEVATION 100000 LB EQUIPMENT OOM WEIGHT'AT 30.25 FT. ELEVATION AND WELD PAD LIVE LOADS OF 24000. 20000 AD 20000 AT 56, 60 AND 65 FT. ELEVATIONS iEFV6T.W.1 3-15

?RF# 00664 NOEX 9-2, REV. 0 Table 3-7 Oyster Creek Drywell Additional Weights - Post-Accident Condition-PEKETR. MISC. TOTAL 3 FOOT LOAD PER LOAD PER ILOAD PER EL.VATION iEIGHT lE16tT LOADS LOAD RANGE 36 DES. L Ot OOMS OF FULL NODE UALF NOOE (Feet) (lbf) (tbf) 0bl) (Tbf) LOAD (lbf) ELEMENTS APPLICATION (lbf) (ibfl

-(----- ________-

15.56 50000 50000 16 168100 168100 20 11200 11200

• 15-20 229300 22930 6 .116-119 3822 1911 221 556000 556000 21-251 556000 55600 8 161-169 6950 3475 26 11100 11100 30 84100 51500 115600 30.25 105000 105000

- 25-30 231700 23170 8 179-187 2896 1448 31 16500 16500 32 750 750 33 35450 15430 34 28050 280S5 35 1500 1500 t 31-35 62250 6225 8 . 188-196 778 389 36 1550 1530 -

40 41000 43350 84350 36-40 85900

• 8590 .a 197-205 1074 537 SO 1102000 1102000

• - 45-50J 1I02000 110200 8 418-426 13775 6888 54 7850 7850 51-55 7850 785 8 436-444 98 49 56 56400 56400 60 95200 700 95900 556-60 I 52300 15230 8 454-462 1904 952 65 52000 52000 61-65 52000 5200 8 472-480 650 325 70 5730 5750

• 66-70 5750 575. S 508-516 72 36 73 8850 8850 71-7S 8850 885 8 S26-S34 111 55 82.17 21650 21650 tl-85 216S0 2165 8 553-561 271 135 1000 1000 90 15000 15000

- 86-90 16000 1600 8 571-579 200 100.

93.75 20700 20700 95.75 20100 20100 91-96 40800 4080 8 589-597 510 255

___-LOA -O E APPLIED

-- -_--- C-- IE-- 2572__50 TOTILS: 2184150 38U200 0 2572350 2572350. 257235 t - LOAD 70 BE APPLIED IN VERTICAL OIRECTlON ONLY.

& - NO MISCELLANEOUS LOADS FOR*THIS CONDITION.

FLOOOOGT.VI 3-16

I mwm.

RFEX08664 TNOEX 9-2, REV. 0 Table 3-8 Hydrostatic Pressures for Post-Accident, Flooded Condition WATER DENSITY: 62.32 lb/ft3 0.03606 lb/in3 FLOODED ELEV: 74.5 ft 894 inches ANGLE ELEMENTS ABOVE ABOVE EQUATOR ELEVATION DEPTH PRESSURE NODES (degrees) (inch) (inch) (psi) ELEMENTS 27 -53.32 110.2 783.8 28.3 1-12 40 -51.97 116.2 777.8 28.1 13-24 53 -50.62 122.4 771.6 27.8 25-36 66 -49.27 128.8 765.2 27.6 37-48 79 -47.50 137.3 756.7 27.3 49-51, 61-66 ,55-57 92 -46.20 143.9 750.1 27.1 52-54, 138-141 ,58'60 102 -44.35 153.4 740.6 26.7 142-147, 240-242, 257-259 108 -41.89 166.6 727.4 26.2 148-151, 243, 256 112 -39.43 180.2 713.8 25.7 152-155, 244, 255 116 -36..93 194'.6 699.4 25.2 156-159i 245, 254 120 -34.40 209.7 684.3 24.7 160-165, 246, 253 124 -31.87 225.2 668.8 24.1 166-173 247, 252 130 -29.33 241.3 652.7 23.5 174-183, 248-251 138 -26.80 257.6 636.4 23.0 184-195 148 -24.27 274.4 619.6 22.3 196-207 161 -20.13 302.5 591.5 21.3 208-215 170 -14.38 342.7 551.3 19.9 216-223 179 -8.63 384.0 510.0 18.4 224-231 188 -2.88 425.9 468.1 16.9 232-239 197 2.88 468.1 425.9 15.4 430-437 400 8.63 510.0 384.0 13.8 438-445 409 14.38 551.3 342.7 12.4 446-453 418 20.13 591.5 302.5 10.9 454-461 427 25.50 627.8 266.2 9.6 462-469 436 30.50 660.2 233.8 8.4 470-477 445 35.50 690.9 203.1 7.3 478-485 454 40.50 719.8 174.2 6.3 486-493 463 45.50 746.6 147.4 5.3 494-501 472 50.50 771.1 122.9 4.4 -502-509 481 54.86 790.5 103.5 3.7 510-517 490 805.6 88.4 3.2 518-525 499 820.7 73.3 .2.6 526-533 508 835.7 58.3 2.1 534-541 517 850.8 43.2 *1.6 542-549 526 885.3 8.7 0.3 550-557 187.3 706.7 25.5 340-399 (Ventline)

FLOODP.WK1 3-17 mmmmmmmmm

RNOEF YNOEX 9- 4 REV. 0 006 Table 3-9 Meridional Seismic Stresses at Four Sections 2-D Shell Meridional Stresses Elevation Model Refueling Post-Accident SectU on (inches) Node NW~ )-

119 32 1258 1288 A) Middle of Sandbed 323 302 295 585 B) 17.250 Be'low Equator 489 461 214 616 C) 5.75 Abo,ve Equator 1037 1037 216 - 808 D) Above Knu,_.5 tkle.

3-18

?NDEX M4 REV. 0 Table 3-10 Application of Loads to Match Seismic Stresses - Refueling Case 2-0 SEISMIC STRESSES AT SECTION (psi)

SECTION: 1 2 .3 4 2-0 MOOE: 32 302 J461 1037 COMPRESSIVE STRESSES FROM 2-0 ANALYSIS ELEY: 111.3" 322.5" 489.1" 912.3'

. I 0.058' SEISMIC DEFLECTION: 78.67 155.54 103.46 85.31 HORIZ. PLUS VERTICAL SEISMIC INERTIA: 469.55 139.44 110.13 130.21 TOTAL SEISMIC COMPRESSIVE STRESSES: 1258.22 294.98 213.59 .215.52 3-0 STRESSES AT SECTION (psi) 3-D INPUT SECTl ON: - 1

• 3 4 LOAD 3-0 ND0CIES: 53-65 170-178 400-408 526-534 SECTION INPUT 3-0 UNIT LOU DESCRIPTION ELEV: 119.3" 322.S- 489.1" 912.3" I

A 1000 lbs at nodn 563 through 569 85.43 37.94 34.94 55.23 500 lbs at 427435. 1000 lbs at 428-434 89.88 39.92 . 36.76 . 0.00 C 500 lbs at 1971205. 1000 lbs at 198-204 17.64 43.37 0.00 0.00 0 500 lbs at 1611169. 1000 lbs at 162-168 89.85 . 0.00 0.00 0.00 2____52 DESIRED COMPRESSIVE STRESSES (psi): 1258.22 294.98 213.59 215.52 3-D INPUT LOAO SECTION LOAD TO BE APPLIED TO MATCH 2-0 STRESSES RESULTING STRESSES AT SECTION (psi)

A 3902.2 333.37 148.05 136.34 215.52 2101.4 168.87 83.89 77.25 0.00 C 1453.6 141.93 U3.04 0.00 0.00 .

0 6611.6 594.05 0.00 0.00

• 0.00 213____

SUM: 1258.22 294.98 213.59 215.52 SEISUNFIL.WI 3-19

?NDEX -2, REV. 0 Table 3-11 Application of Loads to Match Seismic Stresses - Post-Accident Case 2-0 SEISMIC STRESSES AT SECTION (psi)

SECTION: 1 2 3. 4 .

2- NWOOt: 32 302 461 1037 .

COMPRESSIVE STRESSES FROM 2-D ANALYSIS ELEV: 119.3" 322.S5 4t9.1" 912.3" O.0S8- SEISMIC DEFLECTION: 788.67 155.54 103.46 85.31 HORIZ. PLUS VERTICAL SEISMIC INERTIA: 499.79 429.39 512.76 723.14 Ii--- -T-

- -------- 584.__

TOTAL SEISMIC COMPRESSIVE STRESSES: 1288.46 St84.93 516.22 808.45 3-D STRESSES AT SECTION (psi) 3-0 INPUT SECTION: 1 2 3 4 LOAD 3-0 NODES: 53-65 17O-178 400-408 526-534 SECTION INPUT 3-D WilT LOAD OESCRIPTION ELEV: 119.3" 322.5" 489.1" 912.3-A 1000 lbs at nodes 563 through 569 8S.43 37.94 34.94 55.23 500 lbs at 427U435. 1000 lbs at 428-434 89.68 39.92 36.76 0.00 C 500 lbs at 197120.S 1000 lbs St 198-204 97.94 43.37 0.00 0.00 0 SOO lbs at 1I6&169. 1000 lbs at 162-168 69.85 0.00 0.00 0.00 DESIRE.COMPRESSIVE STRESSES (psi): 128.46 584.93 81i.22 808.45 3-D INPUT LOAD SECTION LOAD TO BE APPLIED TO MATCH 2-0 STRESSES RESULTING STRESSES AT SECTION (psi)

A 14637.9 1250.51 555.36 511.45 808.45 2850.2 256.17 113.78 104.77 0.00 .

c -1941.7 -189.58 -U4.21 0.00 0.00 D -318.8 -28.64 0.00 0.00 0.00 5____3 SWU: 1288.46 St4.93 616.22 608.45 SElSFL .W1 3-20

-- w m-

DRYWELL i ELEV.51I0" 66 .676*

N-iTHK.

Figure 3-1. Oyster Creek Drywell Geometry 3-21

- - -M

• MSYS 4.4 1 NOU 1.3 19" 14:53:53 REAL "lUN WU =1 XJ=363 . at 2? =639.49 ANGZ---90 CENROID HZDV3N w

• w OYSTER CRVX DRYWLL ANALYSIS - OYCQ (S4AD. RZIIELING)

Fmniro I-7 fnvetpr rria-nt n#ls.,wii l.fn Ffnit+ Flawartt Mndal

l I N8 13 199 1 4:33:17 WU =1

~ yST-2A,% 7 6 XT =4290.453 aNG7z -_

CENTROID HIxDENM riWm

. . . . s . % I s . I I OYSTER CRZEK DRYMELL ANALYSIS - OFCRJA CSAND. REFfILING)

FS VIe Flgure 3-3. Closeup of Lower Drywell Section of FEN (Outside View)

• NSYS 4.4 NOU 13 1996 14: 6::23 1.

REAL "UK1 W(U =.-1 EYS E7 376 xT =4aa,45a 2r =21_.!28 AR4x2=98 HIDOKH CD4?ROID O .TE .RVRIELAMVI VRQC~I.RFEIG

.1c n w '.A

• l asPlln of I Pwer novel.l -+t.4"n f as ITm"4Ci4 VWOO

ANSYS 4.4 1 OCT 15 1996 89:31:26 TVPZ MMN BC SYNES

=--0. a DIST=718 .786

)4F =362.031 2Jr =639.490 29w0tB HI DDUN w.

1%3 U'

L AA LSI__ U .

OYVSTER CREEI DRY MZLY m sis - Sand *e UNIT LOD CA BE)

-. - - I 0. , . .- P _.

a

hNsYS 4.4 1 OCT 15 199S 6I: 32: 26 XC S1YOL DR5_7 8 786W

,c .39a.31 41 Zlr =3.631

=639.498 g AO H I DDA to "3

at OYSTER CREE2X DRYWELL ANALYSIS - @YCRI+/-A CSAHD, UNIT LOAD IN)- '

I Finnsr J -1 Annlicatinn nf tnadinn ton gigflathp Cqigmic RndinO

ANSYS NOU 4.4 14 1990 es8:la:59 SY (AUrC)

DmX =6. 2n"79 SMX =442.367 X(U =1 WU =-8.8 DZST=7+/-8o786 XT =33.831 X at So-76' 74

-3311

-1434 w 442.967 OYSTER CR-X DR-M-LL A.-SIS - O'CRIQ ',, -. R.WUILIN()

7 , -Moidin trpcat - Apflin rat

Nov 14 1993 08:14:37

. STEM-_

ITUI~

SY COUG31 DM( =9.im670 SvX =4_ .. 87 XCU =1

. _ . .- 834 Disimas.376 XT =42g8.452 E

CINTROIP HIDDSN

-s.,

-79'4

-3311

-1434 442.907 IOYSTER CRtEV DRVMELL AIMLYSIS -OYCR+/-Q CSAHD* REFUELING)

F. ntirfo I-A Inwor nflvwoil Hrsidinnxl Str.Potp - RofIPlirn C3ase

m ANSYS 4. 4 NOV 14 1996 18:13:14 SX (PG)ml DMX =6.266378 SM. -3363 SEX =2873 XU =1 WU =-G..

DIST=710 .786 Xir =383.931 Ale - '

E IJ1l0D HIDDIR

.- 2670 i

16.4719 Ile N  : w2973 148?

210727 OYSTER CREEK DRYMELL AIALYSIS -OYCRI.Q (SANDl REFULING)

, F"ntra 1.4 rIr-rfprrntjAi

• ,e . qtf*tqo . Rafoollnn or.x

I ANZkS 4.4 NOU 14 1999 68:15:10 STKP=1 SX (<UG)

SX =2873 XU =1 2 _N YU =-' .'*376DU

=-~fR33 DIST=268 XT =420.452 CENILROID HIDDEN

• -- 336a 1467

.~ 2873 CP OW 33T CM= DRVWKLL AM~LYSIS - OV~CR1Q CAD XULfG figure 3-10.

F. .Lower Orywell Circumferential Stresses - Refueling Case

" ANSYS 1401J5 1"9S 4.4 83:6: 35 STEII~

ITER=l DMX =9. 4w?

8MH a-12320 SME =271U XU =

WU' =--- a DIS?=718 .786 XJr =33.631 CENTOIDHIDDEN

-. 2329

__-18653

• 3'93

-6369all

. 2713 OYSTER C.EX DRVMZLL ANALVSIS .0 w SOST-ACCID.)

,, ' . Fn11 1-1i Mpridinnal Strpgto. - Pnqt-Arrid.'nt ramp

111111 ANSYS n_4_4 4.4 I

89:37 : 36

'I i713=1 tTR=l SY c SUG) mIX CM4 =9.311487

=-12396 SNX =2710 DST-288 .376 Xr =429.452 WIOIP HIxDAEN A4o-3976

-630.211 L4a 2716 L:a NMa I

OYSTER CoEn VRYNELL OtYMLYSIS - OVCRlU CSA". POST-ACCID.)

FlPiteroe,.1.1 Inwer Drywell No-idinAnal qtravvt - Pnet..Lp4flnt rag^

ANSYS 4.4 1OU LS 199t 89:56:17

.STE=1 ITXR=I spiX IAUG30 DMX =0*4=07

&M* =-4594 SMX =12763 Xu =1

lX==YU =-@..8 DIST=7180786 XI =3S3.831 498 CypTROtv H xDDU E -4594

.-2666 5849 12763 i,

. YTE .RE ALVI-RIIL ?0 OT-CI

.*~7~ . d .. ._.f. ~~I~

A.

I ANSYS 4.4 Mau +/-5 9ss9

., - _ -sX C&UG)0 DW~=6.3+/-1487 CM# z-4594 SMX =13763 XU =1 DI ST=28 . 376

__ 2*l528 CEWROID HIDxDEN

.- 4594

__ =-269X 12763 OYSTER CflEZX DRYVELL ANALYSIS - OYC!U CSARD OST-ACCKD.)

. Ffatorp -14. tIner nrvwii c+af0 00.c - ft. ---

_ . -.. 1.. --. ..--- -- IIII-. ~l~ .

1i 1.

.II

I i

iI

. I Unbuckled Shape .i I I

.I Budded Shape i 0

Vent (Radial Displacement No Rotation ), i Symmetric Buckling of Drywell Unbuckled Shape Budded Shape Vent Rotation -

No Radial Disp. /

Anti-symmetric Buckling of Drywell .

SYM.DRW Figure 3-15. Symuetric and Anti-Symmietric Buckling Modes 3-35

-M ANSVO 4.4 STE?--l KTKR~

.ACT.14.322 4 .~HO DI St1" IVXGLO3*l. .3 6:2-327 22 DNX =.. 6_s XV?=-..

DI ST 92 035 XI =327.422

." =-s 689208-

• .W

.a' E-9.99+/-254

-9.9L731

=.a 09.91131

- 11 ~~- -9. m22..- -

OYSTER CREWC DRYMELL AHILYSIS - OCCRIE CSAM D8 RIUKLIHG)

Finigr 1-15 Svmmetric Buckl int NrMmo hano - of,..lin" a2:

• HsY 4.4 Nov n 1996 69:26:n3

.. STlP=

KTK=1 }

FmCT=16 .812

. ._=ll DMX =6.663214 CNN. =- .2937 SMt =G. 0214

.XU =1 YU =-G.s DIS?=192 .3s1 Xr =327.422 E-6.992164

. _l GlX136

-°. 662837

-6.661492

a. B+/-1197 68 6e0.6214 OYSTER CAtE( DRYINLL AIMLYSIS - OuCRt" CSAMD, FfUKLIKG) 4 4

, F4nn,,o 1.17 -qvmmotvr e Peu14vny MAngl4 C.nao - DftOl4flfir

aNSYS NOV 4.4 1.6 1990 U6:1S:29 STE?=1 TBCTS. 911 DIX =0.6995

." =-; .602302 SNX =9.M5 XU =1 Yuq =.

=4.

XF =327.422 128

-.. .02362

- -0. iJ76f WE . 0.X0144 0.062589 OYSTER CREBX DRYJELL AdMLYSIS - _YCRUV CLANV PIDOST-ACCKD.)

. ~~~ ^.. TO- '0v1 Alnn 14nrl I*--- near* A-SA6 eP...

RF#F 00664 YNDEX 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 47x 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

DEX 9-P,REV. 1

• Table 4-1  :

Calculation of Allowable Buckling Stresses - Refueling Case Parameter Value Theoretical Elastic Instability Stress, aie*(ksi) .101.65 Capacity Reduction Factor, a. 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, O; mod 0.207 Elastic Buckling Stress, ae - *i,mod °ie (ksi) 21.04 r.

Proportional Limit Ratio, A - °elay 0.554 Plasticity Reduction Factor, Ai 0.993 Inelastic Buckling Stress, al - qje (ks1) 20.89 Factor of Safety, FS 2.0 Allowable Compressive Stress, aall. m 7j/FS (ksi) 10.44 Applied Compressive Meridional Stress, am (ksi) 7.10 Margin - C(aallIam) - 1] x 100% 47%

'I 4-2

.NUA T,2 REV. 1 Table 4-2 Calculation of Allowable Buckling Stresses - Post-Accident Case Parameter Yalu Theoretical Elastic Instability Stress, ale (ksi) 96.06 Capacity Reduction Factor, ae - 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, ai mod 0.32 Elastic Buckling Stress, ae 'i,mod aie (ksl) 30.74 Proportional Limit Ratio, A - oelay 0.809 Plasticity Reduction Factor, -i 0.736 Inelastic Buckling Stress, a° - ?7iae (ksi) 22.62 Factor of Safety, FS 1.67 Allowable Compressive Stress, all - oa/FS (ksi) 13.55 Applied Compressive Meridional Stress, am (ksi) 9.69 Margin - [(aall/om) - 1] x 100% 39.7%

4-3

~RF# 00664

-NDEX 9-2, REV. 0

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 I

mwwmw

DRSf 900664 NEX.9-2, REV. 1 Table 5-1 Buckling Analysis Summary Load Combination

Refueltn Post-Accident Service Condition Design Level C Factor of Safety Applied 2.00 1.67 Applied Compressive Meridional Stress, (ksi) 7.10 9.69 Allowable Compressive Meridional Stress (ksi) 10.44 13.55 Buckling Margin 47% 40%

5-2

IREV. 0 AN ASNE 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

. 4X'. 1 vmmmmmmmmmmmmmm

-NOEXT-2! REV. 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

VN EXO-66REV. 1 TABLE OF CONTENTS 1.. INTRODUCTION 1-1 1.1 General 1-1.

1.2 Report Outline 1-1.

1.3 References 1-2

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

. . 1 3.5 Loads *3-4 3.6 Stress Results .3-7 3.7 Theoretical Elastic Buckling Stress Results 3-9 3.8 References 3-1X1

4. ALLOWABLE BUCKLING STRESS EVALUATION 4-1 S.

SUMMARY

AND CONCLUSIONS 5-1 iii

RF# 08664 YNEX 9-4, REV. 0

.LIST OF TABLES Table Page No. Title No.

3-1 Oyster Creek Drywell Shell.Thicknesses 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 - 3-19 Refueling Case 3-11 Application of Loads to Match Seismic Stresses - 3-20 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

RF#000664 TNDEX 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 2-8 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-IC 3-1 Oyster Creek Drywell Geometry 3-21 3-2 Oyster Creek Drywell 3-D Finite Element Model 3-22' 3-3 Closeup of Lower Drywell Section of FEM (Outside View) 3-4 Closeup of Lower Drywell Section of FEM (Inside View) 3-24 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 3-2; 3-8 Lower Drywell Meridional Stresses - Refueling Case .3-2B v

WMNWN

ORF# 00664

• NOEX 9-4, REV. 0 LIST OF FIGURES Figure Page No. Title No.

3-9 Circumferential Stresses - Refueling Case 3-29 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 Asymmetric Buckling Modes - 3-35 3-16 Symmetric Buckling Mode Shape - Refueling Case 3-36 3-17 Asymmetric Buckling Mode Shape - Refueling Case 3-37 3-18 Buckling Mode Shape - Post-Accident Case 3-38 vi

RF# 00664 INOEX 9-4, REV. 0 I. 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 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.

1-1

RF# 00664 NDEX 9-4, REV. 0 1.3 References 1-1 "Structural Design of the Pressure Suppression Containment.

Vessels," by Chicago Bridge & Iron Co.,Contract # 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 Dryweli -

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

6EX 8-2, REV. 0

• b.

• \ o9 Prcos A - -

Poems' )^

4..

. %X II 6)~~ I-( '.I R

I 4

.bi.1 -

Iuxc - Sos\};11.

o10

&awa

.~ ~~ 0r'A'I.

k%.'%.v%

(I E1w. 15' 77.  :

- W- its. ,. 0 0 ELTAV (aI %-t, i

rbskI. , . . k

_ Sw,*ic, SkUt(V)

Figure 1-1 Drywel 1 Configuration 1-3

%61X l-2,REV. O

-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 [2-1 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, lie, 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 elgenvalue extraction technique. More details on the elgenvalue 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, al, 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., aleri. When the elastic buckling stress exceeds the proportional limit of the material, a plasticity reduction factor, Ad, is used to account for non-linear material.

behavior. The inelastic buckling stress for fabricated shells is given by 1#ficrie 7

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 - 1ilioie/FS 2-1

?WXYT REV. t' 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, aj, 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, al 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 et, 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

NE REV. I 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 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, ap. The capacity reduction factor a, can then be modified as follows:

ai,mod 1 + p .

2-3

NWXEA I'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)(aj)Et/R

- (0.605)(0.207) Et/R.

Where, 0.605 is a constant, 0.207 is the capacity reduction factor, 1,'

and E,t and R are Young's Modulus, wall thickness and radius, respectively. In the presence of a tensile stress such -as thai:

produced by an internal pressure, the buckling stress is given as:

follows:

Scmod - (0.605)(aj +:ap)Et/R

- (0.605)(0.207 + a )Et/R

- [(0.605)(0.207) +.AC] Et/R Where AC is ap/0. 605 *and is given for cylindrical geometries in thi -

graphical form in Figure 2-3. As can be seen in Figure 2-3, AC -isa 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 (ao/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 now

.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, 'imod' for the drywell geometry was obtained as follows:,

`i,mod 207 + AC(sphere)/0.6O5 2.4 Determination of Plasticity Reduction Factor When the elastic buckling stress exceeds the.proportional limit of the material, a plasticity reduction factor, A;, is used to account for the non-linear material behavior. The inelastic buckling stress for fabricated shells is given by qjeirie. 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 aaiie/ly and ay is the'material yield strength. Figure 2-4 shows the relationship in graphical form.

7 - 1.0 ifAA 0.55

- (0.45/A) + 0.18 if 0.55 < A S 1.6

- 1.31/(1+I.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

- EX0964, 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 irn 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 Unstiffenecd 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 TH 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 1

WREV. 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 Axial.

Tension," AIAA Journal, Vol. 1, No. 10, October 1963, pp.

2316-2319.

2-7

TNUEX T-T,REV. I as as It CA 0.2 0.0 0

Figure 2-1 Capacity Reduction Factors for Local Buckling of Stiffened and Unstiffened Spherical Shells 2-8

N EX 8-, REV. 1 I0 a

SUOHY (NAA) a

--- Iit-O FUGSaSECHLER at_

A l

-- 'I lo _LOC1E.

,,. CRATE SCHfWARTZ.-=_

• . S NAA1 1- !. 1 1.0a IBESTFICUV 6

VdC 4 TI E ?

.10

.01 O (V)2 Figure 2-2 Experimental Data Showing Increase in Compressive Buckling Stress Due to Internal Pressure (Reference,2-6) 2-9 I

?N'9EX09T,4 REV.. 1 I0 I

a 6

4 I I I a

1.0 S

6 S - -- l T

. AC 4 2

0.10

. . U .. U =

a 6

4 a

0.0o la. -0 V a a .1I0a -

0.01 6JO a

... r o Figure 2-3 Design Curve to Account for Increase in Compressive Buckling Stress Due to Internal Pressure (Reference12-11) 2-10 I

VNE 9-4 REV. I

.. I .

L.a 4 1 9 9 P 141

.4A.

0mv SS,- -

a.1.

6 .S U,,+S tA a _ #n -a 2I ao e.* 4AA -4 7. _6

& 's/T. c Figure 2-4 Plasticity Reduction Factors for Inelastic Buckling 2-11

.1,

RF# 064

?NDEX 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, X, for which elastic buckling is predicted using the equation:

([K] + X [S] ) (u) 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, Gcr, at a certain location of the structure is thus calculated as:

dcr S Xap 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

5NREXO864, 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

RV"qNEX9 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.

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

_ I

R1F# 00664

.NDEX 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 zof 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 for the post-accident (Case VI) load combination including:

3-4

?RE%08664 N EX -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 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 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 10.

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 1g. 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 I

• DRF#X00664 INOEX 9-4, REV. 0 3.5.3 Pressure Loads The 2 psi external pressureiload 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/ft3 (0.0361 lb/in3), 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: 50W-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

TNDEX @-4, REV. 0 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; oRm - -7580 psi 3-7 M

E REV.0 The circumferential stress averaged from the bottom to the top of the sandbed region is; ORc - 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;

°PAm ' -11960 psi The circumferential stress averaged from the bottom to the top of the sandbed region is; 0PAc - +20080 psi 3-8

REV. 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 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 I

RF# 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; URie - 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 OPAie - 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 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.

3-10

YRF# 00664 NOEX 9-4, REV. 0 Table 3-1 Oyster Creek Drywell Shell Thicknesses Section Thickness (in.}

Sandbed Region 0.736 -

Lower Sphere 1.154 Mid Sphere 0.770 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

• 95% confidence projected thickness to 14R.

3-11

DRF#X00664 INDEX 9-4, REV. 0 Table 3-2 Cylinder Stiffener Locations and Secltion Properties Elevation Height Width Area Bending Inertia (in4)

(in) (in) (in) (in2l Horizontal Vertic.al 966.3 0.75 6.0 4.5 13.5 0.211 1019.8 0.75 6.0 4.5 13.05 - 0.1 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 2 beam sections, one 2.75x7" and one 1.0x7.375" Table 3-3 Material Properties for SA-212 Grade B Steel Material Propertv Value Young's Modulus 29.6x106 psi Yield Strength 38000 psi Poisson's Ratio 0.3 Density 0.283 lb/in 3 3-12

RF# 00664 INDEX 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 @ 175'F or 35 psi e 281"F) +

Seismic (2 x DBE)

CASE VI - POST ACCIDENT CONDITION Deadweight + Water Load e 74'6" + Seismic (2 x DBE)

I 3-13

IRF#X 064.

NDEX 9- , REV. 0 Table 3-5 Adjusted Weight Densities of Shell to Account for Compressible Material Weight Adjusted Shell Weight Density Thickness (in.) (lb/in 3) 1.154 0.343 0.770 0.373 0.722 0.379 2.563 0.310 0.640 0.392 1.250 0.339 3-14

RF#006 YNOEX 906, REV. 0 Table 3-6 Oyster Creek Drywell Additional Weights - Refueling Condition OEA1 PENE TR. mISC. TOTAL 5 FOOT LOAD PER LOAD PER LOAD PER ELEVATION WE16NHT WE]1GHT LOADS LOAD RANGE 36 DE6. E OF NOOES OF FULL NODE HALF NOOE (fast) (lbf)I (lt bf) (lbf) (lbf) LOAD (lbf) ELEMENTS APPLICATION (Ibf) * ()11f) 15.56 sooaI0 50000 16 161 1100 168100 20 12 1200 11200

^ 15-20 229300 22930 6 115-119 3822 :1911 220 5560cI0 556000

• 21-25# 556000 55600 Sa 161-169 6950 .3475 26 12 1100 11100 30 64100 51lS00 115600 30.25 105000 100000 205000

• 26-30 331700 33170 8 179-187 4146  :!073 31 1500 16500 32 750 750 21 33 1450 13450 34 1!1050 280SO 35 1500 1500 21 62250 i6225

• - 31-35 a 1U8-196 778 389 36 lSSO 1550 40 41000 431350 84350

• 38-40 85900 8590 * . 197-205 1074

• 537 SO# 1102000 1102000

'- 45-500 1102000 110200 ,8 418-426 13775 488 8

54 P8so 7850

• - S1-SS5 7850 785 8 436-444 98 . 49.

56 56400 24000 80400 60 95200 700 20000 115900

• 4 56-60 196300 19630 8 454-4t2 2454 1227 65 52000 20000 72000

'4 61-65 72000 7200 8 472-480 900 450 70 7SO 5750

• - 66-70 5750 575 8 S08-_SI 72 38 73 1850 8850

• - 71-75 USO 6 526-534 ill '55.

J2.17 216S0 216S0

'481-85 21650 2165 8 . 553-561 271. 135 87 1000 1000 90 --1! 5000 15000

• 4 86-90 16000. 1600 a 571-573 200: - 100 93.75 20700 20700 94.750 69J000 698000 95.75 20100 20100

• 91-96 738800 73880 8 589-S97 9235 4618 2184150_ _ 3 TOTALS: 2194150 3U1 5200 862000 34343SO 3434350 . 343435

• - LOAD TO BE APPLIED IN VERTICAL DIRECTION ONLY.

& - MISCELLANEOUS LOADS INCLUOE 898000 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 REFiGT.W1 .

3-15

TNDEX P-T4 REV. 0 Table 3-7 Oyster Creek Drywell Additional Weiqhts - Post-Accident Condition.

DtAD PENETR. MISC. TOTAL 5 FOOT LOAD PER LOAD PER LOAD PER ELEVATION VEI6HT WEIGHT LOADS LOAD RANGE 36 DES. * # OF NODES OF FULL NMOE HALF NODE (feet) 50bf) (lbf) (Ibf) LOAD (Tbf) ELEMENTS APPLICATION (lbf) (lbf) 15.56 50000 50000 16 168100 168100 20 11200 11200 15-20 229300 22930 6 116-119 3822 1911 220 556000 556000 0* 21-251 556000 5500 a 161-169 6950 3475 26 11100 11100 30 64100 51500 115600 30.25 105000 105000 25-30 231700 23170 8 179-187 2896 1448 31 16500 16500 32 750 750 33 15450 154S5 34 28050 28050 35 1500 1500 31-35 62250 6225 8 lU8-196 778 389 38 3SS0 1550 40 41000 43350 64350 0* 36-40 85900 8590 a 197-205 1074 537 500 1102000 1102000

• 45-509 1102000 110200 8 418-426 13775 6888 54 7850 7850 98 - 49 51-55 76SO 785 8 436-444 56 56400 564W

• 60 95200 700 95900

" 56-60 152300 15230 8 454-462 1904 952 65 52000 52000 0 61-65 52000 5200 8 472-480 650 325 70 5750 5750 5750 8 508.-Si 72 36 U-70 575 73 8850 U50 71-75 "50 8S a 526-534 111 55 82.17 215SW 21650 8 553-561 271 135 81-85 21650 2165 87 1000 1000 90 . 15000 15000 e6-90 16000 1600 a 571-579 200 100 93.75 20700 20700 95.75 20100 20100 91-96 40800 4080 .B 589-597 510 255 TOTALS: 2164150 398200 0 2S72350 2572350 257235 J - LOAD TO SE APPLIED IN VERTICAL DIRECTION ONLf.

& - NO MISCELLANEOUS LOAD FOR THIS CONDITION.

FL00N6T.J1 3-16

gE 64, REV. (3 Table 3-8 Hydrostatic Pressures for Post-Accident, Flooded Condition WATER DENSITY: 62.32 lb/ft3 0.03606 lb/1n3 F1LOODED ELEV: 74.5 ft 894 inches ANGLE ELEME1TS ABOVE ABOVE EQUATOR ELEVATION DEPTH PRESSURE NODES (degrees) (inch) (inch) (psi) ELEMENTS

,'7 -53.32 110.2 783.8 28.3 1-12 410 -51.97 116.2 777.8 28.1 13-24

53 -50.62 122.4 771.6 27.8 25-36.-

I56 -49.27 128.8 765.2 27.6 37-48

t9 -47.50 137.3 756.7 27.3 49-51, 61-66 ,55-57 92 -46.20 143.9 750.1 27.1 5 -54, 138-141 ,58-456 11)2 -44.35 153.4 740.6 26.7 42-147, 240-242, 257-259 118 -41.89 166.6 727.4 26.2 148-151, 243, 256 112 -39.43 180.2 713.8 -25.7 152-155, 244, 255 116 -36.93 194.6 699.4 25.2 156-159, 245, 254 1,'0 -34.40 209.7 684.3 24.7 160-165, 246, 253 1;4 -31.87 225.2 668.8 24.1 166-173 247 252.

130 -29.33 241.3 652.7 23.5 -174-183, 24i-251 i:ia -26.80 257.6 636.4 23.0

• 184-195 148 -24.27 274.4 619.6 22.3 196-207 1(51 -20.13 302.5 591.5 21.3 208-215

-14.38 342.7 551.3 19.9 216-223 17T9 -8.63 384:0 510.0 18.4 224-231 1138 -2.88 425.9 468.1 16.9 232-239 197 2.88 468.1 425.9 15.4 430-437 400 8.63 510.0 384.0 13.8 438-445 409 14.38 551.3 342.7 12.4 446-453.

418 20.13 591.5 302.5 10.9 454-461 427 25.50 627.8 266.2 9.6 462-469 4:36 30.50 660.2 233.8 8.4 470-477 445 35.50 690.9 203.1 7.3 478-485 454 40.50 719.8 174.2 6.3 486-493 4,63 45.50. 746.6 147.4 5.3 494-501 472 50.50 771.1 122.9 4.4 502-509 4:B1 54.86 790.5 103.5 3.7 510-517 490 805.6 88.4 3.2 518-525 499 820.7 73.3 2.6 526-533 5D8 835.7 58.3 2.1 534-541 517 850.8 43.2 . 1.6 542-549 526 885.3 8.7 0.3 550-557 187.3 706.7 25.5 340-399 (Ventline)

FtOODP.WKI 3-17

9RF# R0664 NOEX 9-4, REV. 0 Table 3-9 Meridional Seismic Stresses at Four Sections 2-0 Shell Meridional Stresses Elevation Model Refueling Post-Accident on Se cti-i (inches) Node (DSi 1-119 32 1258 1288 A) Middle of Sandbed Below Equator 323 302 295 585 B) 17.25 489 461 214 616 C) 5.75* Abo ve Equator 1037. 1037 216 808 D) Above Knu rcek 3-18

- REX"94 REV. 0 Table 3-10 Application of Loads to Match Seismic Stresses - Refueling Case 2-0 SEISMIC STRESSES AT SECTION (pti)

SECTION: 1 2 3 4 2-0 NOOE: 32 302 461 1037 COMPRESSIVE STRESSES FROM 2-0 ANALYSIS ELEV: 119.3" .322.3S 489.1" 912.3" 0.058" SEISNIC DEFLECTIDN: 788.67 155.54 103.46 85.31 HOR12. PLUS VERTICAL SEISMIC INERTIA: 469.55 139.6U 110.13 130.21 TOTAL SEISMIC COMPRESSIVE STRESSES: 1258.22 294.98 213.59 215.52 3-0 STRESSES AT SECTION (psi) 3-D INPUT SECTION: 1 2 3 4 LOAD 3-0 NODES: 53-8S 170-178 400-406 S26-531 SECTION INPUT 3-0 UNIT LOAD DESCRIPTION ELEV: 119.3" 322.5 - 489.1" 912.3" A 1000 lbs at nodes 503 through 569 85.43 37.94 34.94 55.23 SO lbs at 427243S. 1000 lbs at 428-434 . 89.88 39.92 36.76 0.00 C S00 lbs at 1971205, 1000 lbs at 198-204 97.64 -43.37 0.00 0.00 0 500 lbs at 161U169. 1000 lbs at 162-168 89.85* 0.00 0.00 0.00 9.___8_

DESIRED COMPRESSIVE STRESSES (psi): 1258.22 294.93 213.59 215.52 3-D INPUT LOAD SECTION LOAD TO BE APPLIED TO MATCH 2-0 STRESSES RESULTING STRESSES AT SECTION (psi)

A 3902.2 333.37 148.05 136.34 215.52 2101.4 188.87 83.69 77.25 0.0c C 14S3.6 141.93 63.04 0.00 594.05 0.00 0.00 O.Wc 2.OC5

'D 6811.6 SUIN: 1258.22 294.98 213.59 215.52 SEISUNFL.W.1 3-19 I

OREFV 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 (psi)

SECTION: 1 2 3 4 2- NOODE: 32 302 461 1037 COMPRESSIVE STRESSES FROM 2-D ANALYSIS ELEY: 119.3" 322.5" 489.1" 912.3" 0.058" SEISMIC DEFLECTION: 788.67 155.54 103.46 85.31 HORIZ. PLUS VERTICAL SEISMIC INERTIA: 499.79 429.39 512.76 723.14 TOTAL SEISMIC COMPRESSIVE STRESSES: 1288.46 584.93 616.22 608.45 3-0 STRESSES AT SECTION (psi) 3-D INPUT SECTION: I .2 3 4 LOAD 3-0 NOOES: S3-6S 170-178 400-408 526-534 SECTION INPUT 3-0 UNIT LOAD DESCRIPTION ELEV: 119.3" 322 .5 489.1" 912.3" A 1000 lbs at nodes 563 through 569 65.43 37.94 34.94 55.23 500 lbs at 427143S. 1000 lbs at 428-434 89.88 39.92 36.76

• 0.00 C 500 lbs at 1972205. 1000 lbs at 198-204 97.64 43.37 0.00 0.00 0 500 lbs at 161&169. 1000 lbs at 162-168 89.65- S0.00 0.00 0.00

__16__2 DESIRED COMPRESSIVE STRESSES (psi): 1288.46 .584.93 -616.t2- 808.45 3-D INPUT LOAD SECTION LOAD TO BE APPLIED TO hATCH 2-0 STRESSES RESULtIN6 STRESSES AT SECTION (psi)

A 14637.9 1250.51 SS5.36 511.45 808.45 -

2850.2 256.17 113.78 104.77. 0.00 C -1941.7 -189.58 -64.21 0.00 0.00 D -318.8 -28.64 0.00 0.00 0.00 SUN: 1288.46 584.93 616.22 808.45 SEISFL.1W1 3-20

APEX -T, RE-V. 0 DRYWELL ELEV.51'fO ITHK.678' Figure 3-1. Oyster Creek Drywell Geometry 3-21

NSYS 4.4 1 DEC 4 1990 REAL MON uJ =1 SYSAM?.786 XI =383.031 Zr =639.498 afG2=-90 CZTROIOD HIDDEN w

P3 Po.

L C O O NSC OYSTER ConE DRYIM 0MUlSIS - ONTRO CHO UNA POST-ACC*

Fiaure 3-2. Oyster Creek Orvwell 3-0 Finite Element Mndel

AySVS 4.4 1 DEC 4 1998 13: 6: 41 REAL MIN XU =1

=S284.376 XT =420.452 2' =216 .5320 CEt1ROQD HIDDEN

.M.

. up 7X I .1 OYSTER CRERI4 DRVL AWILYSIS -OYCRJO CHO SANDo POST-ACC.

. Figure 3-3. Closeup of Lower Drywell Section of FEM (Outside ViewI

ANSYs 4.4 DEC 4 1990 18:07:X2 REAL HtU MU =-I lhs T=8. 3 76

)( =420.43a 27 =216.520 ANG290 CENSROID, HIDDA w

1%

I I OYSTER CHICX DRYVELL ANALYSIS - OYCRlO CHO SAHD. POST-ACC.

1.

fiaure.3-4. eCnseti of l.ower Orwell Section of FF I(Inside View)

ANtSw 4.4 I DEC 4 1990 15: 19:37 Al TYPE HUM UPS1 DC SYMBOLS DIST=718 .786 Xr =383.031 ZV =639 .498 4otivD H IDDKN w

Y RE

-m -I. . I OYSTER COAEX DRYWALL

. I

-;6 Fiaure 3-5. Boundary Conditions of Finite Element Model

ANSYS 4.4 1 OCT 15 1999 N 69:32:26 Gkl no. 2TS TYPE HUM DC SYNBOLS rU =-0.8a

=U 1

DIST=718

• 786 Xr =303.931 2 =639.498

§=AO0D HI DayM

~1

\ .

• R

.( .R ' - . - :. ', , U L C

.OVSTER COAEX DRYWEL ANALYSIS - OYCRIA (SAHD UNIT LOAD CA CZ)

,Finfora tI.A Annlicatinn of I nadin t tn iomlato la4tmir Randiinn

lI AN-A u l4

.NOU N6 199 q ' .STKP=1 I TKR=1 I . AsN

.aDMX

.ST (AUG

=0.221779

=-9174 SNX =695.647

,XU =1

'EU =-S.8

.~ -. DIST7 1 .7X86 XF =363.g31

_ C tOID, HIDDEN

'-0174

7198 rof f tfe -3247

. nD OlSIM CUEEX DRYWEELL (ANALYSIS - OES Cal SAND. EFf14 .

e 37 s Figure 3-7.: Meridional Stresses - Refueling Case.--

fiNSYS 4.4 1OU 16 1996

. . STEP=Jl IT1+/-

- -,-- - -i -S S _AUG)

• DMX

• • • ~~S SS*I 8M =1B.139473

=-S174

~' SMX =695.947 XU =1 Y~U =-s.s DI S7280 376 XF =429.452

-- 0174

-7108

-3247

-1276 695.947 OYSTER CREf DRYPELLAtWLYSIS -iCRIS CNO SAlD. REYUELI e4mi.-s I A It-or nrvwpll Mor4idinno, qtrpttq - ggfliplinq Iaq<P

hAISYS 4.4 NW4 L6 1996

. . .09:33:39

. 0STEP--

iT=1 SX c OUGII DI.IX =.221779 SH -3547 SAX =6754

.XU YU =1 =-6.s DIST=710.*796

. i . XT =363.631 X~e-S3&. 49s ;

l -

lCEITROW

-3547

-24623HIDDE 2176 445

'754 OYSTER CHEN DRYWELL.ANbLYSIS -OCf+/-S c.o ,ND'. RE.ELI Tioure 3-9. Crvumferential Stresses -Refuelin Case

NSYS 4.4 NOU 16 1996 59: 15L STEP--I ITERI~

DMX =9.139473 SHM =-3547 SMX =6734

,(U =1 YU =-9.a DI ST=280.376 X(7 =422P.452 E-2403 CENT"IID HIDDEN

~-3547 T2176 4465 6754 CD OYSTER CZEEX DRYIELL AMLYSIS OACIS

- CHO SAND. BEFEx Figure 3-10. Lower Drywel Circumferential Stresses Refueling Case I..

• aNS9IS 4.4 1 Irw if i:?iw 16:30:38 STEP=1 I TE=1 DMX =9 . 479734 ratm SM" =-13155 SMX =3894 "09SlM XU =1

,T- YU =-e.8

.,4 DI 8T=718.706 XF =363.931 9;.49a .

CDMTROID HIDDSI E -13155

-11269

-3683

_ 65.136 3894 w

. 41 OYSTER CBEX DR SELL SAM#. POST-ACC4e 6..

Fintirp t -11 ridinnal Stresvws - Pnct-Arridpnt CatR

.NSVS 4.4 N4ov i9 1990 16: 33: 46 STEP+/-l IT ERm1 r "r r- , . SY CAUG)

I. * * * * * * * *I D U..i

• I. *a * * *3.EL X' '  :** . *1 . * * . DMX =0. 479734 4t be .u2 - e.f * -F. * * .a .l SF-la =MH-13+/-53 SMt =3894 3\ fr a

___ -- Sa -i-ap ~YU - -.

X* * ' 4.-t-

' * ~ ~ j Lu DIST=286

_= 1-* .376 3s; t * .- ,X =42;.453

_ .- 3683 105.136 3894 OYSTER C*ERN DRYWELL

-'OYCRO ANALYSIS O SA'D' POST-ACC.

r .sa Figure 3-12. Lower Drywell Meridilonal Stresses - Post-Accident Case

Abjeue A a 1 OU LS9 1990 1i: 39: 42 gLOT NO. 3 WOST STteSa STEt=1 LX ( AU C)

DMX =0.479734 SHM =-5295 SMX =27791 X' =1 YU =-e.s DI 37=710.7e6 XJ =383.931 By 6&.498.

CETKROID HIDDKM

-5295

- -1538 13126

~ 20459 27791 w

OYSTER CoREE DRYMEM AALVSSIS SAND. POST-ACC.

Figure 3-13. Circumferential Stresses - Post-Accident Case - .

ANSYS 4.4 NOU L9 1996 16:32:33 KTXi=1 SX C*UG)

DHX =6.47 9734 GMN -5205 SNX =27791 XU =1 YU =-0.8 DI ST=288.376 XT =426.452 U-sads CENIXOUD HIDDEN 13126

~29459 27791

,. I Fiqure 3-14. Lower Drywell Circumferential Stresses - Post-Accident Case

?961X 9-4, REV. 0 Center of.

Drywel Sphere Planes of Symmetry Usudded Shape Budded Shape:

Vent Racial Displacement

( No Rotation )

Symnletric Buckling of Drywell Urbudded Shape Budded Shape Vent I Rotation \

-kNo Radial Disp . A Asymmetric Budding of Drywel SYM.DRW Figure 3-15. Synetrlc and Asymmetric Buckling Modes 3-35

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.9.'W744 OYSTER CREEX DRYWELL ANALYSIS -

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• NDEX -64REV. 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 41%

margin. The allowable compressive stress for the Post-Accident, flooded case is 13.77 ksi. This results in a margin of 15% for the applied compressive stress of 11.96 ksi.

4-1

N AT06 REV. 1 Table 4-1 Calculation of Allowable Buckling Stresses - Refueling Case Parameter Value Theoretical Elastic Instability Stress, a.ie (ksi) 58.10 Capacity Reduction Factor, a 0.207 Circumferential Stress, ac (ksi) 4.49 Equivalent Pressure, p (psi) -15.74 "X" Parameter 0.0865 AC 0.0716 Modified Capacity Reduction Factor, eimod 0.325 Elastic Buckling Stress, ae - a1i,mod Oie (ksi) 18.88 Proportional Limit Ratio, A a ve/ay 0.497 Plasticity Reduction Factor, i 1.60 Inelastic Buckling Stress, "i niae (ksi) 18.88 Factor of Safety, FS 2.0 Allowable Compressive Stress, 0all - /FS (ksi) 9.44 Applied Compressive Meridional Stress, am (ksi) 7.58 Margin - [(Gall/Cm) - 1] x 100 24.5%

4-2

DRF# 00664 INDEX 9-4, REV. 1 Table 4-2 Calculation of Allowable Buckling Stresses - Post-Accident Case Parameter Theoretical Elastic Instability Stress, oie (ksi) 61.95 Capacity Reduction Factor, a 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 - Oimod aie (ksi) 31.47 Proportional Limit Ratio, A - aey ' 0.828 Plasticity Reduction Factor, Adi 0.724 Inelastic Buckling Stress, ai '7ioe (ksi) 22.78' Factor of Safety, FS 1.67 Allowable Compressive Stress, orall Or/FS (ksi) ' 13.64 Applied Compressive Meridional Stress, Gm (ksi) 11.96:

Margin - t((all/am) - 1] x 100% 14X 4-3

RF# 00664 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

RREX08644, REV. 1 Table 5-1 Buckling Analysis Summary Load Combination Refuelin Post-Accidul, Service Condition Design Level C Factor of Safety Applied 2.00 1.67 Applied Compressive Meridional Stress (ksi) 7.58 11.96 Allowable Compressive Meridional Stress (ksi) 9.44 13.64 Buckling Margin 24.5% 14.0%

5-2