ML20125A997
| ML20125A997 | |
| Person / Time | |
|---|---|
| Site: | Oyster Creek |
| Issue date: | 11/30/1992 |
| From: | Marisa Herrera, Mehta H, Ranganath S GENERAL ELECTRIC CO. |
| To: | |
| Shared Package | |
| ML20125A983 | List: |
| References | |
| NUDOCS 9212090165 | |
| Download: ML20125A997 (66) | |
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OC N_ UClBar Erierg ' l 9212090165 921202 PDR ADOCK 05000219 .' JP PDR w
i NNX k RFV. 2 i )-; ) AN ASME SECTION VIII-EVALUATION-
, OF THE OYSTER CREEK DRYWELL FOR WITHOUT SAND CASE
) PART 2 STABILITY ANALYSIS
-(Revision 2)
November 1992 3-prepared for ! ) GPU. Nuclear' Corporation. L Parsippany, New Jersey- ) 4 1
- prepared by.- ;
)- GE Nuclear Energy'
' San: Jose, California
) J1
bbhXhkfREV.2 D AN ASME SECTION VIII EVALUATION 3 OF THE OYSTER CREEK DRYWELL FOR WITHOUT SAND CASE PART 2 g STABILITY ANALYSIS (Revision 2) O Prepared by: M.L.Herrera, Senior Engineer 3 Materials Monitoring & Structural Analysis Services g Verified by: + H. S. Mehta, Principal Engineer Materials Monitoring & Structural Analysis Services Approved by:
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l - l S. Ranganath, Manager Materials Monitoring & Structural Analysis Services O o i
) k b X hk REV 2 TABLE OF CONTENTS )
69a
- 1. INTRODUCTION 11 1.1 General 1-1 1.2 Report Outline 1-1 1.3 References 12
- 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 3-4 3.5 Loads 3-4 3.6 Stress Results 3-8 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 5-1 ii
.- _ _ . _ _ . _ _ _ _ . ~ . _ . _ . _ _
) N 5kX k REV. 2 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-P 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 3-10 Application of L'oads to Match Seismic Stresses - 3-19 )' Refueling Case 3 Application of Loads to Match Seismic Stresses - 3-20 Post-Accident Case ) 4-1 Calculation of Allowable Buckling Stresses - Refueling 4i2 4-2 Calculation of Allowable Buckling Stresses - Post-Accident 4-3 )- 5-1 Buckling Analysis Summary 5 - iii
9 I ) X REV. 2-LIST OF FIGURES 3 Figure Page . No. Title A 3 1-1 Orywell Configuration 1-3 Capacity Reduction factors for Local Buckling of 2-1 2-8 Stiffened and Unstiffened Spherical Shells-D- 2-2 Experimental Data Showing increase in Compressive 2 Buckling Stress Due to Internal Pressure 3 2-3 Design Curve to Account for Increase in Compressive ~2-10 Buckling Stress due to Internal Pressure 2-4 Plasticity Reduction Factors for Inelastic Buckling 2-11 3-1 Oyster. Creek Drywell Geometry 3_-21 32 Oyster Creek-Drywell 3-D Finite Element Model 3-22 3-3-3 Closeup of Lower Drywell. Section of FEM.(Outside View) 3-23 3-4 Closeup of Lower Drywell Section of FEM (Inside View) - 3-24 3-3-5 Flat Plate Buckling Analysis Results for Free Edge 25 Boundary Conditions- ) 3-6 -View of Refined Mesh in the Sandbed Region 26-3-7 Symmetric Boundary Conditions for Stress Analysis.- '3-27 , )' 3-8 -Symmetric and Asymmetric Buckling Modes 28-3-9 Application of loading _to Simulate Seismic-Bending =3-29 9 iv
kbXbfREV.2 , LIST OF FIGURES Figure Page No. Title ,lh D 3-10 Meridional Stresses - Refueling Case 3-30 3-11 Lower Drywell Meridional Stresses - Refueling Case 3-31 D 3-12 Circumferential Stresses - Refueling Case 3-32 3-13 Lower Drywell Circumferential Stresses - Refueling Case 3-33 5 3-14 Meridional Stresses - Post-Accident Case 3-34 3-15 Lower Drywell Meridional Stresses - Post-Accident Case 3-35 D 3-16 Circumferential Stresses - Post-Accident Case 3-36 3-17 Lower Drywell Circumferential Stresses - Post-Accident 3-37 Case 3-18 Sym-Sym Buckling Mode Shape - Refueling Case 3-38
, 3-19 Sym-Asym Buckling Mode Shape - Refueling Case 3-39 3-20 Sym-Sym Buckling Mode Shape - Post-Accident Case 3-40 24 D
D y
1 . bX REV. 2- ) !. 1. INTRODUCTION
- 1.1 General i
To address local wall thinning of the Oyster Creek drywell, GPUN has l prepared a supplementary report to the Code stress report of record ] [1-1] which is divided into two parts. Part includes all of the I Code stress analysis resdits 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 j 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 F removing the sand on the code . aess analysis and buckling evaluation, j respectively. Buckling =of the entire drywell shell is considered in j this analysis with the sandbed region being the area of primary concern.
I 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 intsection 4. Section 5 presents the summary of results and~ conclusions, r 1-1 _M#'T
i 1 kkbfXhkfREV._'2 -3 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- Drywell -
) Part 2 Stability Analysis," GE Report' No. 9-2, DRF#- 00664, November 1990, prepared .for GPUN.
1-4 "An ASME Section VIII ' Evaluation of the .0yster Creek Drywell :-
) Part 1 Stress Analysis," GE Report No. 9-3, DRF# 00664, February 1991, prepared for GPUN. )-
1 i -
) - ) )
1-2
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k.X REV. 2
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Figure 1-1 Drywell Configuration ) 1-3
i
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bX REV. 2
- 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 evaluate = in three steps. In the first step, a theoretical elastic buckling stress, ogg, is determined. This value may be calculated either by classical buckling equations or by finite element analysis. Since the drywell shell geometry is complex, a three dimensional finite -element analysis
, approach is followed using the eigenvalue extraction technique. More J
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, aj, accounts for the difference between classical buckling theory and actual tested buckling stresses for 1 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., aje aj. When the elastic buckling f) stress exceeds the proportional limit of the material, a plasticity reduction factor, nj, is used to account for non-linear material behavior. The inelastic buckling stress for fabricated shells is given by njajaj,, 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 - njojaj,/FS ) 11
_ _ . _ _ _ .... _ . _ _ _ _ . r q
! hX REV. 2 ; )
{ in Reference 21, the safety factor for the Design and Level A & D ; j service conditions is specified as 2.0. A safety factor of 1.67 is t
! cpecified for Level C service conditions (such as the poct a:cident
- condition).
l The determination of appropriate values for capacity and plasticity
- reduction factors is discussed next.
2.2 Determination of Capacity Reduction Factor l 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 i are based on extensive data compiled by W11er it 3). The factors j appropriate for a spherical shell geometry such as that of the drywell , in the sandbed region, are shown_ in Figure 21 (Figura 15121 of ' 1 Referenra 2 1). The tail (flat) and of the curves ara used for , ! onstiffened 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, og is determined to be-0.207. i The preceding value of the capacity reduction factor is very l conservat for two reasons. First, it is based on the assumption l that the 5 ical shell has a uniform thickness equal to the reduced l thickness. JWeVer, the dryWell shell has a 9feater thickness abOV9 the sandbed region which would reinforce the sandbed region. Second, it is assumed that the ci$ e forential stress is zero. The tensile . circumferential - stress has the effect of rounding the shell and 4 reducing the effect of imperfections introduced during the fabrication i and construction phase. A modification of the aj value to account for g the presen9 of tensile circumferential stress is discussed in i
. Subsection 2.3.
The capacity reduction factor values given in Reference 2-1 are i applicable to shells which meet the tolerance requirements of NE-4220 - L i 2 -. --. - - . - _ - . _ . . - - _ - - . -.---- -. . . . . . . ..-
f X REV. 2 of Section !!! [24). Reference 25 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 111, NE vessel, it is justified to use the approach outlined ) in Code Case H 284. 2.3 Modification of Capacity Reduction f actor 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 og must be given in the Design report." The effect of hoop tensile stress on tho buckling strength of cylinders has been extensive 11y documented (2 6 through 211]. Since the methods used in accounting /or 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 (26) presented a comprehensive set of test data, including tnose 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, o .p The capacity reduction factor og can then be modified as follows: ai, mod - og + op ) ) 2-3
) ORF# 00664 INDEX 9 4, REV. 2 The buckling stress in untaxial compression for a cylinder or a sphere of uniform thickness with no internal pressure is given by the ) following:
t Sc- (0.605)(aq)Et/R
=
(0.605)(0.207)Et/R Where, 0.605 is a constant. 0.207 is the capacity reduction factor,aj, and E,t and R are Young's Modulus, wall thickness and radius, respectively. In the presence of a tensile stress such as that
) produced by an internal pressure, the buckling stress is given as follows:
Sc mod = (0.605)(aj + ap)Et/R
=
(0.605)(0.207+a)Et/R p
=
((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)*, where ,p is the internal pressure. Miller (212] gives the following equation that -fits the graphical relationship between X and AC shown in Figure 2-3:
h
'AC = ap /0.605 = 1.25/(5+1/X)
The pre:eding approach pertains to cylinders.- Along the similar
) linas, Miller -(2-13] hai developed an approach for spheres as described next.
4 The non dimensional parameter X. is essentially (op/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)8 24
X k REV. 2 When the tensile stress magnitude, 5, is known, the equivalent internal pressure can be calculated using the expression: p- 2tS/R ; Based on a review of spherical shell buckling data (214, 215), ) Miller (213) proposed the following equation for AC: at:(sphere)- 1.06/(3.24 + 1/X) ) The modified capacity reduction factor, ai, mod, for the drywell geometry was obtained as follows: ' 8 1, mod = 0.707 + AC(sphere)/0.605 ) 2.4 Determination of Plasticity Reduction Factor When the elastic buckling strus exceeds the proportional limit of the ) material, a plasticity reduction factor, nj, is used to account for the non linear material behavior. The inelastic buckling stress for fabricated shells is given by njagoj,. 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 ajoj,/oyand ay is the material yield strength. Figure 2-4 shows the relationship in graphical form. ) nj - 1.0 if a s 0.55
- (0.45/A) + 0.18 if 0.55 < A s 1.6
= 1.31/(1+1.15A) if 1.6 < A s 6.25
- 1/4 if A > 6.25
) 2.5 References 21 ASME Boiler and Pressure ' Vessel Code Case N 284, " Metal 1 Containment -Shell Buckling Design Methods, Section Ill, Division . 1, Class MC", Approved August 25, 1980. I 2-5
k fX k REV. 2 l 22 Letter (1985) from C.D. Miller to P. Rajut
Subject:
Recommended ' Revisions to ASME Code Case N 284. ) 23 Miller, C.D., " Commentary on the Metal Containment Shell Buckling i Design Methods of the ASME Boiler and Pressure Vessel Code," I December 1979. ) 24 ASME Boiler & Pressure Vessel Code, Section !!!, Nuclear Power Plant Components. 25 " Justification for Use of Section !!!, Subsection NE, Guidance in Evaluating the Oyster Creek Drywell," Appendix A to letter dated December 21, 1990 from H.S. Mehta cf GE to S.C. Tummine111 of GPUN. ) 26 Harris, L.A., et al, "The Stability of Thin Walled Unstiffened Circular Cylinders Under Axial Compression Including the Effects of Internal Pressure," Journal of the Aeronautical Sciences, Vol. ; ) 24, No. 8 (August 1957), pp. 587-596. 27 Lo, H.,' Crate, H., and Schwartz, E.B., " Buckling of Thin-Walled Cylinder Under Axial Cnmpression and Internal Pressure," NACA TN ) 2021, January 1950. 2-8 Fung, Y.C., and Sechler, E.E., " Buckling of Thin-Walled Circular Cylinders Under Axial Compression and Internal _ Pressure," Journal ) of the Aeronautical Sciences, Vol. 24, No. 5, pp. 351-356, May 1957. -* 2-9 Baker, E.H., et al., "Shell Analysis. Manual," NASA, CR 912 (April ) 1968), 2-10 Bushnell, D., ~ " Computerized Buckling Analysis of Shells," Kluwer Academic Publishers,-1989 (Chapter 5). ) 2-11 Johnson, B.G., " Guide to Stability. Design ' Criteria- for Metal Structures," Third Edition (1976), John Wiley & Sons. 'i 2-6 - Larv+t*PPmF
) k)S!XNkfREV.2 j 2-12 Hiller, 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 Hethods 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
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I bX REV. 2
- 3. FINITE ELEMENT MODEllNG AND ANALYSIS 3.1 Finite Element Buckling Analysis Methodology This evaluation of the Oyster Creek Drywell buckling capability uses I
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 I structural stiffness matrix, (K), the stress stiffness matrix, [5], and the applied stresses, o ap, are developed and saved from this static analysis. A buckling pass is then run to solve for the eigenvalue or load factor, A, for which elastic buckling is predicted using the equation: ) ( (K) + A [S) ) (u) = 0 where: A is the eigenvalue or load factor. I {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. ihe critical stress, o cr, at a certain location of the structure is thus calculated as: ) 8cr " A C 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
kX REV. 2 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 incrcase 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 tup 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 drywall 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 ctrved surfa.es. ) 3-2
1 ) k)fX REY. 2 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 sttiteners in the cylindrical region of the upper drywell. The section properties of these stiffeners are susnarized in ) Table 3 2. The mesh size in the sandbed region of the model was refined for the purpose of buckling evaluation. The mesh refinement was conducted as ) follows. Buckling analyses of flat plate finite element models with different mesh sizes were conducted and the calculated load factors were compared with the available theoretical values. The analyses considered both the fixed and free edge boundary conditions. The ) results of these analyses showed that with a 3"x3" mesh, the finite element predicted load factors were within a few percent of the theoretical values. Figure 3-5 shows the results of one of the flat plate analyses. Based on these analyses, it was concluded that an ) appropriate mesh size is achieved when the element size in the sandbed region is = 3"x3". Figure 3-6 shows the view of the refined mesh. As discussed in Subsection 3.6, the refined mesh was important for the ) buckling analysis but had ,little effect on the stress magnitudes in the sandbed region. 3.3 Drywell Materials ) The drywell shell is fabricated from SA-212, Grade 8 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. ) 3-3
I bX REV. 2 3.4 Boundary Conditions ) There are two sets of boundary conditions, one for the stress analysis and the other for the buckling analysis. The stress oaalysis boundary conditions are discussed first. ) 3.4.1 Boundary Conditions for Stress Analysis Symmetric boundary conditions are defined for both edges of the 36' drywell model for the static stress analysis as shown in Figure 3 7. ) 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. ) 3.4.2 Boundary Conditions for Bucklin0 Analysis Three sets of boundary conditions are used at the edges of the pie slice model: symmetric at the both edges (sym-sym), symmetric at one ) edge and asymmetric at the other edge (sym-asym), and asymmetric at the both edges (asym-asym). This is required to capture all possible buckling mode shapes that the model is able to predict. Figure 3 8 graphically illustrates the various boundary conditions. With the ) symmetric boundary conditions, the nodes at the edges can displace radially but the rotation is not allowed. In the asymmetric boundary conditions, the nodes at the edges are allowed to rotate but the radial displacement is not allowed. The load factors were determined ) for each of the three sets of boundary conditions and the one with the smallest value was used for the Code margin evaluation. 3.5 Loads ) The loads are applied to the drywell finite element model in the manner which most accurately represents the actual loads anticipated ) .4
) b 5kX k REV. 2 on the drywell. Details on the application of loads are discussed in the following paragraphs. ,
I
)
1 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 l 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 norcal operati6i, 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 combinatinn 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
bX REV. 2 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/ftt 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 19 The same method is used to apply the additional masses to the model for the post-accident case as summarized in Table 3 7. ) 3.5.3 Pressure Loads The 2 psi external pressure load for the refueling case is applied to
)
the external faces of all of the drywell and vent shell elements. The compressive axial stress at the transition from vent shell to beam elements 'is simulated by applying equivalent axial forces to the nodes of the shell elements. Considering the post-accident case, the drywell is assumed to - be flooded to elevation 74'-6" (894 inches). Using a water density of
?
3-6
MxW nov. 2 62.3lb/ft3 (0.0361 lb/in 3), the pressure gradient versus elevation is calculated as shown in Table 3 8. The hydrostatic pressure at the ) bottom of the sandbed region is calculated to be 28.3 psi. According to the elevation of the element centerline, the appropriate pressures are applied to the inside surface of the shell elements. ) 3.5.4 Seismic Loads Seismic stresses have been calculated for the Oyster Creek Drywell in Part 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 3-0 drywell model by applying downward forces at four elevations of the model (A: 23'-7",B: 37'-3",C: 50' 11" and D: 88'-9") as shown on figure 3-9. 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. 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 3-9. 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 )
3-7
) h h!X k REV. 2 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. The mesh refinement produced less than 1% change in the calculated stress magnitudes from
) those obtained with the previous mesh in which the elements in the sandbed region were approx. 12"x12". The stresses reported in these Subsections are based on the refined mesh model. ) 3.6.1 Refueling Condition Stress Results The resulting stress distributions for the refueling condition are shown in Figures 3-10 through 3-13. The red colors represent the most ) tensile stresses and the blue colors, the most compressive. Figures 310 and 3-11 show the meridional stresses for the entire drywell and lower drywell. The circumferential stresses for the same areas are shown on Figures 3-12 and 313. The resulting average meridional ) stress at the mid-elevation of the sandbed region was found to be; l
c -7588 psi Rm The circumferential stress' averaged from the bottom to the _ top of the sandbed region ist { Rc 4510 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 3-14 through 3-17. _The red colors represent the most tensile stresses and the blue colors, the most compressive. Figures 3-14 and 3-15 show the ) 3-8
b .X k REV. 2 meridional stresses for the entire drywell and lower drywell. The circumferential stresses for the same areas are shown on Figures 316
)
and 3-17. The resulting average meridional stress at mid elevation of the sandbed region was found to be; o Am -12000 psi
) P The circumferential stress averaged from the bottom to the top of the sandbed region-is; }
pac = +20210 psi 3.7 Theoretical Elastic Buckling Stress Results
)
After the completion of ~ stress runs for the Refueling and Post-Accident load combinations, the eigeavalue buckling runs are raade 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 The first buckling analysis was conducted using the sym sym boundary
- - conditions. The lowest- (i.e.,- first) -load factor - for this case was found to be 6.14 with the critical buckling occurring in the sandbed region. The critical buckling mode shape is shown in Figure 318.
The red color indicates sections of the shell which displace radially outward and the blue, those areas which displace inward. The first six 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 8.89. Therefore, buckling is- not a concern outside of the sandbed region.
.l The lowest- load factors for the sym asym and asym-asym boundary conditions were determined _to be 6.23 and 7.22, respectively. Figure
- i. 3-19 shows the buckling mode shape with sym asym' boundary conditions.
- It is clear from- these load _ factor values- that the sym sym boundary i 3-9 i
) N bkX k REV. 2 condition load far. tor of 6.14 is the lowest one. Multiplying the load factor of 6.14 by the average meridional stress from section 3.6.1, ) the theoretical elastic buckling stress is found to bet Rie = 6.14 x (7588 psi) = 46590 psi ) 3.7.2 Post Accident Condition Buckling Results Considering the post accident case with symmetric boundary conditions, ) the load factor was calculated as 4.085. The sym asym boundary conditions gave a load factor of 4.206 for the first mode. Based on the refueling condition buckling analyses, it was concluded that the load factor for the asym asym condition will be higher than both the sym sym and sym asym load factors. Thus, the sym-sym boundary ) conditions gave the lowest load factor and thus are controlling. The critical mode shape for the sym sym boundary conditions is shown in figure 3 20. As expected, this mode shape is associated with the sandbed region. ) Multiplying the load factor of 4.085 by the applied stress from section 3.6.2 results in a theoretical elastic buckling stress of ) Pale 4.085 x (12000 psi) = 49020 psi 3.8 References ) 31 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-130?-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. 91, DRF # 00664, November 1990, prepared for-GPUN. y
-3 10
) k)SfXNkfREV2 Table 31 ) Oyster Creek Drywell Shell Thicknesses Section Thickness (in.1 ) 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
kX REV. 2 Table 3 2 Cylinder Stiffener Locations and Section Properties Elevation Height Width Area Bendina Inertia fin4 ) fin) g_ lin)- 11n81. Horizontal Vertical 966.3 0.75 6.0 4.5 13.5 0.211 1019.8 0.75 6.0 4.5 13.5 0.211 1064.5 0.50 6.0 3.0 9.0 0.063 1113.0(I) 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 Procerty Value Young's Modulus 29.6x106 p39 Yleid' Strength 38000 psi Poisson's Ratio 0.3 Density 0.283 lb/in 3 3-12
h fX REV. 2 Table 3 4 ) Oyster Creek Drywell Load Combinations ) CASE I - INITIAL TEST CONDITION Deadweight + Design Pressure (62 psi) + Seismic (2 x DBE) CASE 11 - FINAL TEST CONDITION ) Deadweight + Design Pressure (35 psi) + Seismic (2 x DBE) CASE 111 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 0 175'F or 35 psi 0 281'T) + Seismic (2 x DBE) ) CASE VI - POST ACCIDENT CONDITION Deadweight + Water Load 0 74'6" + Seismic (2 x DBE) ) - ) J ) 3-13
) kbfXNkREV.2 Table 3 5 ) Adjusted Weight Densities of Shell to Account for Compressible Material Weight ) Adjusted Shell Weight Density ThicknesLLind (1b/in 33 i ) : 1 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
) X N REV. 2 ) T8ble 3 6 Oyster Creek Orywell Addition 81 Weights - Refueling Condition 6 OfA0 Pthtft, ul5C. TOTAL 5 F00f LCAD Pit (CAO Pit t0A0 Ptt ) (LEVAf!ON WilGHT Vi!GHT LCA05 LCAD RANGt 36 Clo. f 0F N00t$ Or tutt n00g 8Atr n00g (feet) (Ibf) (1bf) (Ibf) (Ibf) LOA 0 (Ibf) (LEMitf5 APPLICAfl04 (Ibf) 15.56 50000
........ ........ ........ ........ ........ ........ ........... ......... itf
..(....l ..
50000 16 164100 168100 20 11200 11200
" 15 20 !!9300 !!930 6
) tti 556000 116 110 38tt 1911 556000
" ti 25f $56000 55600 8 161 169 6950 3475 26 11100 11100 30 64100 51500 !!5600 30.25 105000 100000 20$000
" 26 30 331700 33170 8 !?9 187 4146 .2073 31 16500 16500
) 32 150 750 33 15450 15450 34 28050 20050 35 1500 !$00
" 31 35 62250 6225 8 188 196 778 389 36 1550 1550
) 40 41000 43350 84350
" 36 40 85900 8590 8 197 205 1074 537 50f 1102000 1102000
" 45 50# 1102000 !!0200 4 418 426 13775 6886 54 7850 7850
" 51 55 7850 185 8 434 444 98 49 56 56400 24000 80400
) 60 95200 700 20000 !!$900
" $6 60 196300 19830 65 $2000 8 454 442 2454 !!!?
20000 72000
" 61 65 72000 7200 8 472 440 900 450 70 5750 5750
" 6870 5750 575 8 508 516 72 36
) 73_ 8850 8850
" 11 75 8850 885 4 526 534 '111 $$
82.11 !!S50 11650
" 81 85 21650 till 8 553 561 271 135 87 1000 1000 90 15000 15000
" 86 90
} 93.15 10100 20700 18000 1600 8 571 579 200 100 94.75d 888000 698000 95.75 20100 20100
" $196 738800 73880 8 589 597 9235 4618 TOTAll: !!84150 388200 862000 3434350 3434350 343435
) f - LOA 0 TO St APPLit0 IN VERf! CAL Ott[CTION ONLY. 6 MISCELLANEOUS LOA 05 INCLUDC $98000 LS WAftt WtlGHT AT 94.75 FT. ELEVATION - 100000 L8 (OVIPMENT 000t WE!6HT AT 30.25 FT. ELEVATION AND WELD PAD LIVE LOA 05 or 24000. 20000 AND 20000 AT $8. 60 ANO 65 FT. ELEVAT10N5 t!FWGT.WK1 ). 3 15
.g )
X N REV. 2
) Table 3 7 Oylter Creek Drywell Additional Weight $ . Polt. Accident Condition 6
) OtAD Pthtit. Milt. TOTAL 5 F00f LOAD Pt8 LOAD Pt8 LCAD Plt (LEVAfl0N WilGHT VtlGHT LCA05 LOAD RANGt 36 OtG. f 0F h00t$ 0F FULL h00t HALr host I (feet) (1bf) (1bf) (1bf) (1bf) LCA0 (1bf) (Lt>tht$ APPLICAf!DN (1bf) (ILf) j 15.56 50000 50000 16 168100 160100 l 20 lit 00 lit 00
) " 15 20 229300 !!930 $ 116 119 38tt 1911
!!# $56000 556000
" 21 25# $56000 55600 8 161 169 6950 1475 26 11100 11100 30 64100 51500 115600 30.25 105000 105000
" 26 30 231700 23170 8 -179 187 2896- 1848
) 31 16500 16500 32 750 750 33 15450 !$450 34 20050 20050 35 1500 1500
" 31 35 62250 6225 8 184 196 778 389
) 36 1550 1550 40 41000 43350 84350
" 36 40 85600 8590 4 197 205 1014 537 50# 1102000 1102000
" 45 50# 110tG0 110200 8 418 426 13775 6868 54 74$0 7850
** $1 55 1850 785 8 436 444 98 49
) 56 56400 56400 60 95200 100 95900 ;
" $6 60 152300 15230 8 454 462 1904 952 65 52000 $2000
" 61 65 52000 5200 8 471 480 650 325 70 5750 5750
) " 68 70 5750 575 8 508 516 ft 36 73 8850 8850
" 71 75 8850 885 8 526 534 111 - 55 82.17 21650 21650
" 8145 21650 till 8 553 581 271 135 81 1000 1000 90 15000 15000
) " 84 90 16000 1600 8 571 579 200 100 93,75 20700 20700 95.75 20100 20100
" 91 96 4 0800 4080 8 549-597 $10 255 TOTAL $t 2164150 388200 0 !$12350 2572350 257235
)
f . LOAD 70 tt APPLit0 IN VttflCAL DittCilM ONLY.
& ho N!$CELLANE005 LOA 05 FOR THil CON 0! TION.
FL000W61.W1 3-16
)
k)fX REV. 2
) Table 3 8 ,
Hydrostatic Pressures for Post-Accident, Flooded Condition WATER DENSITY: 62.32 lb ft3
) 0.03606 lb in3 FLOODED ELEY: 74.5 ft 894 inches ) ELEMENTS ANGLE ABOVE ABOVE EQUATOR ELEVATION DEPTH PRESSURE N0 DES (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.E0 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 26.7 240 242, 257-259 108 -41.89 166.6 727.4 26.2 142 148 147'151, 243, 256 ) 112 39.43 130.2 713.8 25.7 152 155, 244, 255 116 -36.93 194.6 699.4 25.2 156-159, 245, 254 120 34.40 209.7 684.3 24.7 160 165, 246, 253 124 -31.87 225.2 668.8 24.1 166 173 130 29.33 241.3 652.7 23.5 17418$.247,252 248 251 138 26.80 257.6 636.4 23.0 184 195 148 -24.27 274.4 619.6 22.3 196 207
[ - 161 170
-20.13
-14.38 302.5 342.7 591.5 551.3 21.3 19.9 208 215 216 223
- 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 h~
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 L 463 45.50 746.6 147.4 5.3 494 501 J 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)
Ft000P.WK1
)
3-17
khfXNkfREV.2 Table 3 9 i Heridional Seismic Stresses at Four Sections 20 Shell Meridional Stresses Elevation Model Refueling Post Accident Section (inches) Nade losil fosi) A) Middle of Sandbed 119 32 1258 1288 B) 17.25' Below Equator 323 302 295 585 C) 5.75' Above Equator 489 461 214 616 D) Above Knuckle 1037 1037 216 808
) ) ) - ) ) )
3-18 l
bX REY. 2 Table 3 10 c Application of Loads to Match Seismic Stre55c5 - Refueling Case 2 0 Sil5MIC STRt55tl AT 5tCT!0N (pst) SECTION: 1 2 3 4 2 0 N00t: 32 302 -461 1037 COM8Atl51VE STRt55E5 FRON 2 0 ANALT515 'ELEV 119.3" 322.5" 469.1" 912.3" 0.058" $E!SNIC DEFLECTION: 788.67 155.54 103.46 85.31 HORIZ, PLUS VERTICAL Stl5MIC INERTIA: 469.55 139.44 110.13 130.21 TOTAL SE!5MIC COMPRi n M <t45E5: 1258.2i 294.98 213.59 215.52 3-0 STRES$t5 AT 5(CTION (psi) 30 ................................... INPUT SECTION: 1 2 3 4 LOAD 3 0 Noots: 53 35 170-178 400-408 526 534 SECTION INPUT 3 0 UNIT LOAD DESCRIPTION ELEV: 119.3" 322.5" 449.1" 912.3" A 1000 lbs at nod 6 563 through 549 85.43 37.04 34.94 55.23 . 8 500 lbs at 4276435, 1000 lbs at 428-434 89.98 39.92 36.76 0.00 C 500 lbs at 1976205, 1000 lbs at 198-204 97.64 43.37 0.00 0,00 0 500 lbs at 1616169. 1000 lbs at 142-188 89.85 ~ 0.00 0.00 0.00 DE51REDCOMPRE551VESTRESSES(psi): 1258.22 294.98 213.59 215.52 30 INPUT l LOAD SECTION LOAD TO BC APPLICO TO MATCH 2-0 STRES$t5 RESULTING STRE55t1 AT SECTION (psi) ~ A 3902.2 333.37 148.05 136.34 215.52* B 2101.4 188.87.- 83.89 77.25 0.00 C 1453.6 141.93 63.04 0.00 0.00 0 6811.8 594.05 0.00 0.00 0.00 SUN: 1254.22 294.98 213.59 215.52 4 Sil5UNFL.WK1 a 4 3-19
--<v . ,, . ,,,-,-m . . ~ . ,,y-, ,. r-.r-
- 7. REV, 2 Table 3-11 Application of Loads to Match Seismic Stresses - Po5t-Accident Case 2 0 $E15NIC STRES$ts AT $ECTION (pst)
SECTION: 1 2 3 4 2-0 N00E 32 302 461 1037
) CO*RES$1VE STRESSES FROM 2-0 ANALY$!$ ELEV: 119.3* 322.5* 489.1" 912.3" 0.058" $!!$NIC DEFLECTION: 788.67 155.54 103.46 85.31 HOR!2. PLU$ VERTICAL $El$Mic INERTIA: 499.79 429.39 512.76 723.14 TOTAL SEl$NIC COMPRES$1VE STRESSES: 1268.46 544.93 616.22 808.45 )
3-0 $tRES$t$ AT SECTION (pst) 30 ................................... INPUT $ECTION: 1 2 3 4 LOAD 3 0 N00ES: 53 05 170-178 400 404 526 534
$ECTION INPUT 3 0 UNIT LOAD DESCt!Pfl04 ELEV: 119.3" 322.5* ) 449.1" 912.3" A 100) lbs at nodes 563 through 549 85.43 37.94 34.94 $5.23 8 500 lbs at 4276435, 1000 lbs at 428 434 89.84 39.92 36.76 0,00 C 500 lbs at 197&205, 1000 lbs at 198 204 97.64 43.37 0.00 0.00 0 500 lbs at 161&l69, 1000 lbs at 162 168 - 89.85- 0.00 0.00 0.00 ) DE31tEDCOMPat$$1YESTRESSES(psi): 1284.46 584,93 616.22 808.45 3-0 INPUT 1.0A0 ) ..... . . .. . .. .. .. .. . .. ... .. ... ... ... .. .. ....... ....... ..'.. .
A 1250,51 14637.9 555.36 511.45 808.45 8 2850.2 254.17 113.78 -104.77 0.00 C -1941.7 -109.54 -44.21 0.00 0.00 0 318.8 -28.64 0.00 0.00 0.00
) SLM:' 1284.46 544.93 816.22 808.45
$El5FL.wl
)
)
3-20 _ - _ _ _ _ _ _ _ _ _ _ _ _ _ _ - _ _ _ . - - _ \
I. DRF# 00664 I INDEX 9-4, REV. 2 i t I i l i n 1 i DRYWELL t
, yt! ,
ll llglllll qilllllllllig , . igime , _ l l t TH' .640*c .. A e t;.*a ELEV. BT 5"
. ;; p l.); '
. - a . sg
.a 4THg.640 Q g ;Jgy g;g hMNMM; 5.Ed ;$d
! *DEMMM f $Nk
.g ELEV,5r O' y
ya4.. 4$i ...- 4 ., ,
.g 3,,'y q
k*,Auf S i ., 4 3 s .
~'
I .) '
- . d 31 i
p q ;. NJ i t 'U ,W.
> . 4. ~ ':. , , c2. .. h 4,1.' 'NII :
. c. .
t ) a n - c- ~ /8. .. . > :D wwl ll
. .- c Ty,qqr,'
- c
?
.v -- (t,,4t .~,, f
.#;,4.<
1 5 j".:)V s 1 M7dl #I.edM R Wi & ' 4 THK. 676" I Figure 3-1. Oysw reek Drywell Geometry l 1 l l l l 3-21
AMSYS 4.4 DEC 4 1999 1 15:96: 31 PLCT Mo. 1 FRIF7 ELLMENTS RIAL MUM yu =1 YU =-9.8 DIST=718.750 XT =393.931 ZT =639.498 sa c -9e N ROID HIDDEN Y 0
\ .
OYSTER CREEX DRTurfL AHAI.YSIS - OYCR10 CMO SAND, PCOT-ACO. Figure 3-2. Oyster Creek Drywell 3-D Finite Element Model
. _ _ _ _ _ __________y___._..___m
___m...._____ ___y__ y________y_________ ._________m_________m _ _ _ _ _ _ _ _ ._. ONSYS 4.4 i i DEC 4 1999 13:96:41 PLOT Mo. 2 PREP 7 n rMENTs RIAL NUM XU =1 YU =-9.8 DIST=238.376 ' XT =429.452 T =216. SOS AMC =-?P l - CIMTRo' - *DDEN . l J i i W 4 I M i i 1 i s
- s I I I
, oYSTE3 CR m DRYWELL ANALYSIS - CYCR10 CMO SAND. PCST-ACO.y Figure 3-3. Closeup of Lower Drywell Section of FEM (Outside View) i t
_____.__y___ _ _____ . _ . _ _ . _ _ _ _ _ _ . _.______y__.__________. .._.__.___y _. i .I AMSYS 4.4 . ) DEC 4 1999 1 15:27:32 3 PLCT Mo. PREF 7 NS REAL MUM XU =-1 YU =-9.3 DIST=238.376 5 XT =429.452 CT =216.52s ANC==9e N ROID HIDDEN t I 1 ', sa I I I
'\g 1
l l l 1 , oYSTIR CREEx DRYWII.L ANALYSIS - oYcRio (No saMD, rost-ace. > i 4 Figure 3-4. Closeup of Lower Drywell Section of FEM (Inside View) i i } 4-__ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _. . _ . _ ___
iExdiREV.2 5' X 5' FLAT PLATE ' Simply Supported B.C.'s
3.5 Loading
5000 PSI Compressive Stress (Top) 3.45 - Mode Shape: _ All Exhibit Half Wave 12" X 12" 6" X 6" 3" X 3" - 3.4 - I ' i A i 3.35 - i i B
~
n e S ! ! - u- 3.3 - ! ! i ! ! g ! ! a ! ' _ 3.25 - ! !
' . - _ _ _ 4 lhagrgje.a!, =_3.lR_ _ _ _ _ _, _ _ _
3.2 - i i I
. I.
3.15 - l l 6 i i I 3.1 i l i .. . . i 0 100 200 300 400 No. of Elements Figure 3-5. Flat Plate Buckling Analysis Results for Free Edge Boundary i Conditions
.l l
l 3-25 I [ .l
ANSYS
~ '
4.4A
.'. . - OCT 21 1992
. * ~ - - - -
13:11:52 POSTI ELEMENTS REAL NUM XV -l ZV -l
*DIST-108.004
. *XF -37.271 8 , -
I*YF l*2F
-3.226
=373.733 l ANGZ--90 s_ .
CENTROID HIDDEN
,, . , ,, . f: '. : . ? N. . ,
. rr Pj. ay g ;
. l h' ,.
l 1
~
5
~. ,
I, ' - - l. l . .- Figure 3-6 View of Refined Mesh in the Sandbed Region OYSTER CREEK DRYVELL REFINED MODEL
A NS*lO 4.4 1 DIC 4 1999
- ' 1~:19:37
$2-4rw.
PLCT Mo. PRI?? EI.DDTS 6 L- . -44c TY?I NUM a ? EC S*1MROLS , a << !
- NU =1 vu =-0.3 DIST=713.736
-, XT =093.G01 c 27 =539.498 ANC".=-?Q
]
A g ^ CDTROID HIDDEM
- . D !
T-
- m_
t y - i Z 1 , e { i J 3 , g i i ~% l 4 \ %
- n orsrza cmcxx paruzr.L anar.vsrs - cycato .cno sano, rest-ace.p n .r Figure 3-7 Symetric Boundary Conditions for Stress Analysis
s 0 INx hj REV.- 2
) Center of Drywel Sphere '
\ Planes of
/ s Symmetry
)
/
e b 36 3 s
\
/ \
/ g
\ l \
, l
) \,' Unbuckled Shape
\ Budded Shape
\ V ent- f Radal Displacement 1 -
)
( No Rotation -/ Symmetric Buckling of Drywell - Center of
- Drywel s?s Sphere /\ Planes of f Syir.r-try
/ b 36')x
/ \
~
j \
.- . / \ .
e s .-
/
/'.... - - - - - - -
x ,- m
/
\
) / \ Bucided Shape Vent Rotation
[No Radal Disp. ) j
\
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SYM.CRW Figure 3-8 Symetric and Asymetric Buckling Modes
) 3-28
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- OYSTER CREEK DRWELL ANALYSIS - SYM-SYM, N0 SAND, REFUELING l
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] ] OYSTER CREEK DRYWELL ANA' YSIS - OCRFF" ' F (NO SAND, REFUELING) ,
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). b X k REV. 2
- 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 7.59 ksi. Since the applied compressive stress is also 7.59 ksi, it indicates that the safety factor is equal to the Code required value of 2.0. The calculated allowable value of 7.59 ksi is conservative since the ) knockdown factors were calculated conservatively and a ur,i formly corroded thickness of sandbed is assumed. The allowable compressive stress for the Post-Accident, flooded case is 12.93 ksi as compared to the applied compressive stress of 12.0 ksi. Therefore, for both ) cases, the drywell meets the required ASME Code safety factors. )- ) ) ) 4-1
I ' NhkXYkfREV.-2 Table 4-1 Calculation of Allowable Buckling Stresses - Refueling Case r Parameter Value Theoretical Elastic Instability Stress, aj, (ksi) 46.59 Capacity Reduction Factor, og 0.207 Circumferential Stress, 'c (ksi) 4.51 Equivalent Pressure, p (psi) 15.81 "X" Parameter 0.087 AC 0.072 Modified Capacity Reduction Factor, ai, mod 0.326 Elastic Buckling Stress, 8, - 1, mod ie-(ksi) 15,18 Proportional Limit Ratio, A - ,/ 'y 0.40 Plasticity Reduction Factor, nj 1.00 Inelastic Buckling Stress, 'j - nt ,'(ksi) 15.18 Code Factor of Safety, FS 2.0 Allowable Compressive Stress, all " 1/FS (ksi) 7.59 Applied Compressive Meridional Stress, 8,-(ksi) _ 7.59
) ) - )
4-2
( DRF# 00664 INDEX 9 4, REV. 2 Table 4-2 i Calculation of Allowable Buckling Stresses - Post-Accident Crse I- Parameter __. L11gg Theoretical Elastic Instability Stress, age (ksi) 49.020 Capacity Reduction Factor, aj 0.207
)
Circumferential Stress, oc (ksi) 20.21: ' Equivalent Pressure, p (psi) 70.84 "X" Parameter 0.39 AC 0.183 L Modified Capacity Reduction Factor, ai, mod 0.509 Elastic B'rkling Stress, o, at, mod 91e (ksi) 24.94 Proportional Limit Ratio, A = o ,/ oy 0.656 Plasticity Reduction Factor, nj -0.866-
)
Inelastic Buckling Stress, og = ng o, (ksi) 21.59-Code factor of Safety, FS. 1.67 Allowable Compressive Stress,a c ll " ci/FS (ksi) 12.93 Applied Compressive Meridional Stress, o, (ksi) 12.0
) ) )'
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_ _ _ _ _ _ _ - _ - _ _ _ - _ \
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- 5.
SUMMARY
AND CONCLUSIONS 5 The results of_ this buckling analysis for _ the refueling and post; accident load combinations are summarized in Table. 5-1. . - The app 1_ied and allowable compressive meridional stresses shown in Table 51 are for the sandbed region which is the most limiting region in terms o' '
)
buckling. This analysis demonstrates that the Oyster Creek drywell has adequate margin against buckling with no sand support for an assumed sandoed shell thickness of 0,736 inch. This thickness is the-95% confidence projected thickness for the 14R outage. Therefore, for
)
both cases,.the drywell meets the required ASME Code safety factors,_ t
) -
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J 5-1
kXh REV. 2: Table 5-1 Buckling Analysis-Summary load Combination Refuelina Post-Accident
)- Service Condition -Design Level C-Factor of Safety Applied 2.00 1.67 )
Applied Compressive Meridional Stress (ksi)- 7.59 12.0 Allowable Compressive Meridional Stress (ksi) 7.59 12.93 Actual Buckling Safety Factor 2.00 -1.80 .
) .
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5-2
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