ML20087E778
| ML20087E778 | |
| Person / Time | |
|---|---|
| Site: | Oyster Creek |
| Issue date: | 02/28/1991 |
| From: | Frederickson C, Mehta H, Ranganath S GENERAL ELECTRIC CO. |
| To: | |
| Shared Package | |
| ML20087E760 | List: |
| References | |
| DRF-00664-01, DRF-00664-R01, DRF-664-1, DRF-664-R1, NUDOCS 9201220137 | |
| Download: ML20087E778 (62) | |
Text
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,l k kX k REVJ j 2 AN ASME SECTION VIII EVALUATION OF THE OYSTER CREEK DRYWELL FOR WITHOUT SAND CASE .
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AN ASME SECTION VIII EVALUATION 0F THE OYSTER CREEK DRYWELL FOR WITHOUT SAND CASE PART 2 STABILITY ANALYSIS 7 s 1 ; Prepared by:Md- [ ~17 C.D. Frederickson, Senior Engineer-Materials Monitoring & Structural Analysis Services Reviewed by: W -
-H. S. Mehta, Principal Engineer
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Materials Monitoring & Structural Analysis Services - Approved by: -% 4 cM,.^ ^>h I S.Ra:ganath[ Manager Materials. Monitoring & Structural Analysis Services f i
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TABLE OF CONTENTS j j em l 1.. INTRODUCTION 11
-1.1 General- 11 ;
1.2 - Report Outline 1-1 ;
- 1.3 l ._ References - 12 t 2.- BUCKLING ANALYSIS METHODOLOGY- 21
-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 25 2.5 References- 2-5
- 3. FINITE ELEMENT MODELING AND' ANALYSIS 31
- 3 .1 Finite _ Element Buckling Analysis _ Methodology . 31 3.2 Finite Element Model 13 2:
3.3 -Drywell Materials -3 3
.3.4 Boundary Conditions. ~3
- 3.5 Loads 3-4 3.6 Stress Results 3-7 2 3.7- Theoretical Elastic Buckling _ Stress Results 3 3.8 References '3 -
- 4. ALLOWABLE BUCKLING STRESS EVALUATION 4-1 -
'5.
SUMMARY
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f hfX REV. O LIST OF TABLES Table Page No. Title No. 3-1 Oyster Creek Crywell Shell Thicknesses 3 11 32 _ Cylinder Stiffener Locations and Section Properties 3-12 33 -Material Properties for SA 212 Grade B Steel 3-12 3-4 Oyster Creek Drywell Load Combinations 3 13 35 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 h;ights - Post-Accident 3-16 3-8 Hydrostatic Pressures-for Post-Accident, Flooded Cond, 3-17 18 3-9 Meridional Seismic Stresses at Four Sections 3-10 Application of Loads 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 4-2 4-2 Calculation of Allowable Buckling Stresses - Post Accident 4 3 S-1 Buckling Analysis Summary 5-2 l iv
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LIST OF FIGURES Figure Page ! No. Title _Hp2 11- Drywell Configuration 13 . 2-1 Capacity Reduction Factors'for Local Buckling _of 27 ! Stiffened and Unstiffened Spherical Shells 2-2 Experimental Data Showing Increase in Compressive 28 Buckling Stress Due to Internal Pressure 2-3 Design Curve to Account (9r increase in Compressive 2-9 Buckling Stress due to Internal Pressure , 2-4 Plasticity Reduction Factnrs for inelastic Buckling 2-10 3 - 1. Oyster Creek Drywell Geometry 3 21 3-2. Oyster Creek Drywell 3-0 Finite Element Model 3 2e 3-3' Closeup of Lower Drywell Section of FEM (Outside View)- 3-23 3-4 Closeup of Lowar Orywell Section of FEM (Inside View) 3-24 35! Boundary Conditions of Finite Element Model 3-25
-3 6 Application of Leading to Simulate Seismic Bending 3-26 3-7 Meridional Stresses - Refueling Case 3 27 3-8 Lower Drywell Meridional Stresses - Refueling Case 3-28 V
._ _ ~. .- . - . . - . . . bX h REV, O LIST OF FIGURCS Figure Page 4
No. Title 1% , 3-9 Circumferential Stresses - Refueling Case 3-29 3-10 t.ower 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 5 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 L I vi
DR # 00664 IN)EX 9 4, REV. 0
- 1. INTRODUC110N 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. " art 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 suppler.entary report for the degraded drywell is for the present configuration !with sand support in the lower sphere). One option w: -s 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 thrsil,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. l 1 l l l 1-1
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-. k4lbEX Nkf REV. 0 1.3 . References 1
11 " Structural Design of the Pressure Suppression Containment Vessels," by Chica90 Bridge & Iron Co., Contract # 9 0971, 1965. , 12 "An ASME Sectica Vill Evaluation of the Oyster Creek Drywell - Part 1. Stress Analysis," GE Report No. 9-1, DRi# 00664, November 1990, prepared for GPUN. 13 "An ASME Section Vi!! Evaluation of the nyster Creek Drywell - Part 2 Stability Analysis," GE Repo-t No. 9-2, ORF# 00664, November 1990, prepared for GPUN. 1-4 "An ASME Section Vill Evaluation of the Oyster Creek Drywell - Part 1 Stress Analysis," GE Report No. 9 3, DRF# 00664, February 1991, prepared for GPUN. d 1-2
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- 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 (21 and 22).
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, oje, is determined. This value may be calculated either by classical buckling equations or by finite element analysis. Since the drywell shell
- geometry is conplex, a three dimensiont' finite element analvsis approach is followed using the eigenvalue extraction technique. More aetails on the eigenvalue determin6 tion are given in Section 3.
In the second step, the theoretical elastic buckling s trer.s is modified by the appropriate capacity and plasticity reduction factors. The capacity reduction factor, og, accounts for the difference between classical buckling theory and actual tested buckling stresses for fabricatert shells. This difference is due to imperfections inherent in febricated shells, not accounted for in classical buckling thory, which can canse significant reductions in the critical buckling stress. Thus, the elsstic buckling stress for fabricated shells is given by the product of the theoretical elastic buckling stress and d e capacity reduction factor, i.e., oje ag. When the elastic buckling s'ress exceeds the proportional limit of the material, a plasticity reduction factor, ng, is used to account for non linear material behavior. The inelastic buckling stress for fabricated shells is given by ngajoj,. In the final step, the allowable comoressive stress is obtained by dividing the buckling stress calculated in the second step by the safety factor, FS: A110wabic ?ompressive Stress - ngojoj ,/FS 21
4 hh!X REY. 0 in Referencs 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 Lovel C service conditions (such as the post accident condition). The' determination of appropriate values for capacity and plasticity riduction factors is discussed next. t.1 Determination of Capacity Reduction Factor s'e capacity reduction factor, at, 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 (Figuro 15121 of Reference 21). The tail (flat) end of the curves are used for unstiffened shells. The curve marked 'Untaxial 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, aj 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 uriiform thickness equal to the reduced th ; ness. 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 ir.troduced during the fabrication and construction phase. A modification of the og value to account for the presence of tensile circumferential stress is discussed in Subsection 2.3. The capacity reduction factor values given in Reference 21 are applicable to shells which meet the tolerance requirements of NE-4220 22
ftbEX N AEV. I ! of Section !!! [24). Reference 25 compares the tolerance requirements of NE 4220 to the requirements to which the Oyster Creek drywell shell was f abricated. The comparison shows that the Oyster Cresk drywell shell was erected to the tolerance requirements of NE 4220. Therefore, although the Oyster Creek drywell is not a Section !!!, NE vessel, it is justified to use the approach outlined in Code Case N 284. 2.3 Modification of tapacity Reduction factor for Hoop Stress lhe orthogonal tensne stress has the effect of rounding fabricated shells and reducing the effect of imperfections on the buckling strength. The Code Case N 284 (21 and 2 2) notes in the last paragraph of Article 1500 that, "The influenct of internal pressure on a shell structure may reduce the initial imperfections and therefore t higher vahes of capacity ' eduction factors may be acceptable. - Justification for higher valhos of og must be gi'ven in the Design report." The effect of hoop tensile stress on the buckling strength of cylinders has been extensive 11y documented [2 6 through 211). Since the methods used in accounting fc: 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 a1 [2 6) presented a comprehensive set of test data, including those from R:ferences 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 h 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 + ap 23
bEX REV. 1 The buckling stress in untaxial compression for a cylinder or a sphere of uniform . thickness with no internal pressure is given by the following: Se- (0.605)(og)Et/R (0.605)(0.207) Et/R Where. 0.605 is a constant. 0.207 is the capacity reduction factor.og, 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)(o3 + op)Et/R (0.605)(0.207+o)Et/R p ((0.605)(0.207) + AC) Et/R Where AC is o p /0.605 and is given for cylindrical geometries in tha 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. Hiller (212) gives the following equation that fits the graphical relationship between X end AC shown in Figure 2 3: AC = op /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 rion-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
9 kEXhkfREY.1 When the tensile stress magnitude, S, is known, the equivalent internal pressure can be calculated using the exprassion: p= 2tS/R Based on a review of spherical shell buckling data (214, 2 15), Miller (2 13] proposed the following equation for AC: AC(sphere) = 1.06/(3.24 2,'Y. The modified capacity reduction factor, g. ,3 , f- the drywell geometry was obtained as follows: 0 1, mod = 0.207 + AC(sphere)/0.605 2.4 Determination of Plasticity Reduction Factor
- When the elastic buckling stress exceeds the proportional limit of the material, a plastietty reduction factor, ng, is used to account for the non linear material behavior. The inelastic buckling stress for fabricated shells is given by ngagaj,. 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,/ay and ay is the material yield strength. Figure 2 4 shows the relationship in graphical form.
nj - 1.0 if A 5 0.55
- (0.45/4) + 0.18 i f 0. 55 < A 5 1. 6
= 1.31/(1+1.154) if 1.6 < A s 6.25
= 1/A if A > 6.25 2.5 References 21 ASME Boiler and Pressure Vessel Code Case N 284, " Metal Containment Shell Buckling Design Methods, Section !!!, Division 1, Class MC", Approved August 25, 1980.
25
N EX k REV. 1 22 Letter (1985) from C.D. Miller to P. Raju;
Subject:
Recommended Revisions to ASME Code Case N 284. 23 Miller, C.D., " Commentary on the Metal Containment Shell Buckling Design Methods of the ASME Boiler and Pressure Vessel Code." December 1979, 24 ASME Boiler & Pressure vessel Code, Section 111, Nuclear Power Plant Components. 25 "JustifiestiQh f2r Use of SGClion Ill, Subsection NE, Guidance in Evaluating the Oyster Creek Orywell," Appendix A to letter dated December 21, 1990 from H.S. Mehta of GE to S.C. Tumminelli 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. 567 596. 2-7 Lo, H., Ci' ate, H., and Schwartz, E.B., " Buckling of Thin Walled Cylinder Under Axial Compression and Internal Pressure," NACA TN 2021, January 1950. 28 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. 351356, May 1957. 29 Baker, E.H., et al., "Shell Analysis Manual," NASA, CR.912 (April 1968). 2-10 Bushnell, O., " Computerized Buckling Analysis of Shells," Kluwer Academic Publishers, 1989 (Chapter 5), t-2 11 Johnson, B.G., " Guide to Stability Design Criteria for Metal Structures," Third Edition (1976), John Wiley & Sons, t L 26
klXggjpty,! 0 l 212 Miller. C.D., " Effects of Internal Pressure on Axial Compression Strength of Cylinders " CBI Technical Report No. 022891, February 1991. 213 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., heoretical and D.perimental Duckling Loads of Imperfect Sp . .a1 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 i Tension," A!AA Journal, Vol. 1. No. 10, October 1963, pp. 2316 2319. ; 4 r b 9 4, 27
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-o .r 2 E t Figure 2 2 Experimental Data Showing increase in Compressive Buckliag Stress Due to Internal Pressure (Reference 2 6) 29
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- 3. FINITE ELEMENT H0DELING AND ANALYSIS 3.1 Finite Element Buckling Analysis Methodology This evaluation of the Oyster Creek Drywell buckling capability uses thb Finite Element Analysis (FEA) program ANSYS (Reference 3 1]. lhe 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), i and the applied stresses, c 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:
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( (K) + A (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 elasti,c 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, a cr, at a certain location of the structure is thus calculated as: act - A cap 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
.- N 5lx k 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 31, 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 elu. tic quadrilateral shell (STL 33) element because it is better suited for modeling curved
' surfaces.
3-2
NNX$kREV.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 SilF4 beam elements are then connected to this centerline node to model the axial and bending stiffness of the vent and header. Spring (ST!F14) elements are used to model the verttcal header supports inside the torus. ANSYS SilF4 heam elements are also used to model the stiffeners in the cylindi x . region of the upper drywell. The section properties of these stiffeners are summarized in Tabic 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 a m 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. 33
k)b[XhIfRIV.0 1 3.5 Loads The loads are applied to the drywell finite element model fn the manner which most accurately represents the actual loads anticipated on the drywell. Details on the application of loads are discussed in the following paragraphs. 3.5.1 Load Combinations All load combinations to be considered on the drywell are summarized on Table 3 4. The most limiting load combinations in terms of possible buckling are those which cause the most compressive stresses in the - sandbed region. Many of the design basis load combinations include high internal pressures which would create tensile stresses in the shell and help prevent buckling. The most severe design load combination identified for the buckling analysis of the drywell is the
- refueling condition (Case IV). This load combination consists of the following loads: ;
Dead weight of vessel, penetrations, compressible material, equipment supports and welding pads. Live-loads of welding pads and equipment door Weight of refueling water External Pressure of 2 psig Seismic inertia and deflection loads for unflooded condition ,. The normal' operation condition with seismic is very similar to this condition, however, it will be less severe due to the absence of the refueling water and equipment door weight. The most severe load combination for the emergency condition is for the post-accident (Case VI) load combination including: 34
X REV. 0 Dead weight of vessel, penetrations, compressiple 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 loatis 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/ft8 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 lable 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
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t bX REV. 0 3.5.3 Pressure Loads
'The 2 psi external pressure load for the refueling case is applied to the external faces of all of the drywell and vent shell elements. The compressive axial stress at the transition from vent shell to beam eTements is simulated by applying equivalent axial forces to the nodes of the shell elements.
Considering the post accident case, the drywell is assumed to be flooded to elevation 74' 6" (894 inches). Using a water density of 62.3 lb/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 olevation 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 raport, Reference 3 3. Meridional stresses are imposed on- the drywe 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 0: 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 - . - - . . . . . . . - - , = _ = . - -. - . - - - - . - _ . - - . . . - . _ - , .. - _.
NU[X k REV. 0 To find the correct loads to match the seismic strest.es, 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 appl:ed 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 , strest at the mid elevation of the sandbed region was found to be; ORm " 7580 psi 37
=_. - - - . .-_----.a.~ - - - - _ . -. . . -- = , . :- .-_
bbXhkREV.O The circumferential stress averaged from the bottom to the top of the sandbed region is; c' 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 311 . 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 ci rcumf erential stresses for the same areas are shown on figures 313 l and 3-14. The resulting average meridional stress at mid elevation of the sandbed region was found to be; OPAm = 11960 psi The circumferential stress averaged from the bottom to the top of the sandbed region is; OPAc = +20080 psi
+
- y. j ;
n 38
b x N 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 l described in Section 3.1. This 'nalysis determines the theoretical el'astic buckling loads and buck 1 M9 mede shapes. 3.7.1 Refueling Condition Burt W Wsults As shown on Figure 3-15, it % ;:w610le for the drywell to buckle in two different modes, in the e.Ne of symmetric buckling shown on Figure - 315, each edge of thu M' dryvell model experiences radial
. displacement with no rotation. This mode is simulated by applying symmetric boundary conditiont to thu 3 D model the same as used for the stress run. Using these boundary conditions for the refueling
- case, the critical load facter 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 arn restrained from expanding radially. This case is considered by applying asymmetric boundary conditions at the edges of the 3 D_model. With the two pass approach used by ANSYS, it is possible' to study asymetric 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. 39
-p-~- .e --+r a- , ----we v- e --,T"T--
. i tfX Rf' 3 i
i 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 fouild to bet ORie 7.67 x (7580 psi) 58,100 psi ; 3.7.2 Post Accident Condition Buckling Results Considering the post accident case with symmetric boundary conditions, the load factor was calculated as 5.18. Multiplying this load f actor - by the applied stress from section 3.6.2 results in a theoretical elastic buckling stress of oPAle 5.18 x (11960 psi) = 61,950 psi The critical mode shape for this condition is shown in figure 318. Again, the critical buckling mode is in the sandbed region. 3.8 References-3 l_. 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. 32 GPUN Specification SP 1302 53 044 Technical Specification for i 4 19,'inry Cuatainment Analysis - Oyster Creek Nuclear Generating ' A Station; Rev. 2,' October 1990, 33 "An ASME Section Vill Evaluation of the Oyster Creek Drywell - Part 1 Stress Analysis," GE Report No. 9 1, ORF # 00664, November 1990, prepared for GPUN. 3 10
- kblXhIfREY.0 l
Table 3 1 ) { Oyster Creek Orywell Shell lhicknesses Stetion 1hicknt}s (in.) Sandbed Region 0.736
- Lower Sphere 1.154 ;
Mid Sphere 0.770 Upper Sphere 0.722 Knuckle 2.562$ Cylinder 0.640 Reinforcement Below Flange 1.250 Reinforcement Above Flange 1.500 Elliptical Head I.1875 Ventline Reinforcement 2.875 Gussets 0.075 Vent Jet Deflector 2.500 Ventline Connection 2.500 Upper Ventline 0.4375 Lower Ventline 0.250
- 95% confiL9nce projected thickness to 14R.
3 11 __._-._..,___ _ ,.~.--- _ _ _ _._. _ _ _.-_._. _ _ _, _ .---,__.. -_. .
kkbEXhkfREV.O t i Table 3 2 Cylinder Stiffener Locations and Section Properties Elevation Height Width Area gendina Inertia fin 4) fin) _{inl_ fin) fint) Horizontal 'ertical L 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(l) 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 8 Steel Material Procerty __ Value Young's Hodulus 29.6x106 p3g Yield Strength 38000 psi Poisson's Ratio 0.3 Density 0.283lb/in 3 3 12
UEXYkfREV,O Table 3 4 Oyster Creek Orywell load Combinations CASE 1 INITIAL TEST CONDITION Deadweight + Design Prv.sure (62 psi) + Seismic (2 x DBE) CASE !! FINAL TEST C0HD1110N Deadweight + Design Pressure (35 psi) + Seit.mic (2 x DBE) CASE 111 NORMAL OPERATING CONDITION Deadweight + Pressure (2 psi external) + Seismic (2 x DBE) CASE IV REFUELING CONDIT103 , Deadweight + Pressure (2 psi external) + Water Load + Seismic (2 x DBE)
' CASE V - ACCIDENT CONDITION Deadweight + Pressure (62 psi 0 175'F or 35 osi 9 281*F) +
Seismic (2 x DBE) CASE VI - POST ACC10ENT CONDITION Deadweight + Water Load 0 74'6" + Seismic (2 x DBE) 3-13
bbfXNkfREV.0 Table 3 5 Adjusted Weight Densities of Shell to Account for Compressible Material Weight t Adjusted Shell Weight Density Thickness fin.) lib /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 8 e t 3-14
l i
. X REV. O f i
Table 3 6 i Oylter Creek Orywell Additional Weighti . Refuelinn Condition 6 j DEAD Pihtit. Nilt. TOTAL $ F00f LOAD P!I LOAD Ptt LOAD Ptt ! t(tVAfl0N WilGHT VI!GHT LOAD $ LCAD rah 4E 36 Ot0. # Of h00tl 0F FULL h00t MALF h00t 7
-{ feet) (1bf) (1bf) (1bf) (Ibf) LOAD (Ibf) (Linthfl APPLICAf!DN (1bf) (tbf) {
......... ........ ........ ........ ........ ........ .+...... ........ ........... ......... .........
16.$6 50000 $0000 ; il 168100 160100 to 11200 11100
" ll*!D !!9300 !!930 4 116.!!9 36t! 1911-tif $56000 556000- .
" 21.t 5# - $16000 65600 8 161 169 6950 3475 l 26 11100 11100 30 64100 $1600 11$600 36.tl - 10$000 100000 20$000
" 16 30 331700 33170 6 179 187 4146 2073 31 16500 16500 ;
32 760 150 33 15450 16450
- 34. 200$0 260$0 .
35 1500 1500
" 31 35 6ftl0 6ttl 8 168 196 778 389
.36 1550 1850 40 41000 a3350 84350
" 36 40 86900 4590 8 197.t05 1014 $37 50f 1102000 1101000
" 45 50f 1101000- 110100 B' 418 426 13775 f.668 i 54 1850 1850
" $155 7850' 185 8 436 444 98 49
$6 56400 24000 80400
-60 95200 700 20000 ' 115800 -
" $4 60 ~196300 19630 8 464 462 tela lit?
-65 '$2000 20000 72000
" 61 65 ?!000 7200 8 47t.480 900 450 10 6750 5750
" 66 70- $760 575 9 608 516 ?! 36 13 '6450 6450
" 11 75 - M50 68$ 8 5tl.134- til 55 22.17 !!6$0 fl650
" 84 05 !!650 !!65 6 -5$3 561 171 135 87 1000 1000 90 -- !$000 l$000 .
" 66 90 16000 1600 8 $71 579 200 100 93.7$ 20700. 10700 94.75# . 664000 668000 95.75 ' 20100 20100
$49 697 9235 4618
" 91 94 736400 73840 $
TOTAll: tif4150 364200' 46t000 3434350 3434350 343435 f . LOAD 10 St APPLit0 IN VERTICAL O!AECTION Outf. ; 6 . NilCELLAht0US LOAD $ thCLUDE $98000 L8 WAf tR WlGHT Af 94,75 Ft. (LEVAfloh 100000 L8 (OU!PMtNT 000R Wil6HT At 30.tl Ft. ELEVAfl0N AND WELD PAD LIVt LOAD $ OF 34000, 20000 AND 20000 AT $6. 60 AND ll Ft. ELEVAflDNS REFW6T.WK1 3 15 __ L._
4 kt X REV, 0 [ D itble 3 7 j Oylter Creek Drywell Additions) Weicht5 . Polt. Accident Condition 5 OtA0 Pthttt. ul5C. TOTAL 5 FOOT LOAD Pit LCAD P(4 t0A0 P(R ' (L(VA110N vtlGHT WEIGHT LOA 05 LOAD RANG ( 36 0(6. f Of h00($ OF FWL h00( HALF N00! (ibf) (Ibf) (1bf) (1bf) LOAD (1bf) (L(M(Nil APPLICAfl0N (1bf) (1bf) (feet) 15.56 50000 50000 ' is 164100 168100 to lit 00 11200
!!9300 ft930 6 116 119 362t 1911
" 15.t0 .
556000
= 1.'f $$6000 -
556000 55600 8 161 160 6950 3475 l
" 21 25d 26 11100 11100 ^
l 30 64100 51500 115600
-30,tl- 10$000 10$000 131700 23170 8 179 107 703 184G 3
" 46 30 .
31 165M 16500 l 32 750 750 33' 15450 15460 34 200$0 200$0 , al 1500 1500 6tt$0 litl 8 IM.194 778 309 "3135 , 34 1550 1550 l 40 41000 43350 H350 15000 8500 8 197 205 3074 537
" 36 40 SW 110!04 1102000
" 45.$0# 1102000 110200 8 alt.at6 13775 6664 54 7850 1850 7650 705 8 436 444 94 49
" 51 35 54 56400 54400 E3 95!00 700 95600 152300 15!30 8 454 462 1904 952
" 56 60 65 $2000 $2000 Et000 5200 8 472 400 650 - 3t$ !
" 41 65 !
70 5750 5760
- 5750 575 t 500 516 72 36
" 64 70 73 M50 M50
" 11 75- 6450 tel 8 526 534 111 $$
82.17 !!He 21650
!?! 135 21650 8 553 561
" 04 05 till 47 1000 10 8 90 15000 150M
" 84 90 16000 1600 8- 571 579 200 100 -
93.75 20700 -20700 95.75 20100- 20100 40000 4000 8 589 507 510 255 -
" 91 M -
l- ........ ........ ........ ........ ........ ........ 0 !$72350 !$7230 257235 TOTAL 5 - -2164150 364200-f . LOAD TO BC APPLit0 It VERT!tAL CIRECTION ONLY. ,
& . NO MilCtLLAht0U5 LOA 05 FOR THil CON 0lil0N, '
FL000Wei,W1 3 16
- -, . . - - . . . . _ . v , - - . . ~ . --...-,- -..- - .
7.. bEX REV. 0 - Table 3 8 , L>^ fHydrostatic' Pressures for Post Accident, Flooded Condition
-WATER DENSITY: 62.32 lb/ft3 0.03606'b/in3 ,
FLOODED ELEV:-- 74.5 ft 894 inches . ANGLF ELEMECS . AB0Vu , ABOVE EQUATOR ELEVATION.- DEPTH ' PRESSURE - NODES (degrees)) -(inch) -(inch) (psi) ELEMENTS ,
- 27. 53.32 -110.2 783.8 28.3 1 40: -51.97 116.2 '77/.8 28.1 13 24 53? ' -50.621 122.4 771.6 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 66 '55 57-.
92 ~-46.201 143.9' 750.1 27.1 5R-54 138-141-,58 142-147,'240 242,60257 259 102 -44.35 153.4 740.6 26.7-108 -41'.89 : 166.6 727.4 26.2 148I 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, 254 120 434.40l -209.7- 684 3 24.7 -160-165,, 246,.253 5124 -31.87 225.2- f68.8 24.1 166-173,'247
- 1301 -29.33:
1 241.3 652.7 23.5 174-183,24$252 251 138 -26.80 257.6' 636.4 23.0- 184-195
'143 '-24.27- 274.4- -619.6 2.2 . 3 - 196-207 161. 1-20.13. 302.5- 591.5- 121.3 20C-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-t" 427 '25.50- 627.8 266.2 -9.6 462-4691 436~ 30.50- 660.2: 233.8- z8.4 470-477
;445 35.50 690.9: 203.1 7.3 478-485 454" ,40.50 719.8 -)?4.2 -6.3 486-493 463 45.50, 746.6 M/.4 5.3 494-501 472. 50.30 771.1 '122.9 '4.4 502-509
_481 54.86' 790.5- 103.5 3.7 510-517'- 3.2 518-525
'490 .-- 805.6 88.4 499 - 1820.7 73.3 2.6 526-533 508. -- 835.7 58.3 2.1 -534-541 517 -- 850.8 43.2 1.5 542-549-526 - R8'i . 3 . 8.7 0.3 550-557
.. . 187.3 706.7 25.5 340-399 (Ventline)
Ft000P.WK1 3-17 a x .~ . - - _
w k)lbEX h k REV.- 0 l 1 Table 3-9 Heridional Seismic Stresses at Four Sections 2-D Shell Meridional Stresses Elevation Model Refueling Post-Accident __Section (inches) Nadt (osi) (osi) 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
.hkX REV O Table 3 10 Application'of loads to Match Seismic Stressel Refueling Case 2 0 5tl5MIC STRt55t5 AT 5ttil0N (psi) 5tCT10Ni 1 2 . 4 2 0 N00C: 32 302 461 1037 ComPRES$1VE STRE55tl FROM 2 0 ANALT115 ELEV !!9.3" 322.5" 469.1" 912.3" 0.054" Stl5MIC DEFLECTION: 788.67 155.54 103.46 85.31 HORl!. PLU$ VERTICAL SE15MIC INERTIA: 469.55 139.44 110.13 130.21 .
TOTAL SE!5MIC COMPet$51VE STRESSE5: 1258.!! 294.94 213.59 215.52 3 0 STRES$ts AT 5tCT10N (pst) - 3.o . INPUT 5ttil04: 1 2 3- _ 4 , LCAD 3 0 N00t5: 53 65 170 178 400-408 526 534 f stCTION - thPUT 3 0 UNIT LOAD tt$CRIPTION ELEV: 119.3" 322.5" 449.1" 912.3" A 1000 lbs at nodes 563 through 549 85.43 -31.94 34.94 55.23
-500 1bs at 4276435, 1000 lbs at 420-434 89.84 39.92 36.76 0.00 8
500 lbs at 1976205, 1000 lbs at 194-204 97.64 43.37 0.00 0.00 C-500 the at 1616169, 2000 lbs at 162 164 49.85 ' O.00 0.00 0.00 0 OtllREDCOMPAC551VESTRES$t$(pet): !!54.22 294.94 213.59 215.52 30 INPUT LOAD RESULT!>6 STRES5ts AT SECTION (pst) SECTION LOAD TO St APPLIED TO MATCH 2-0 STRt35C5 A- 3902.2 333.37 148.05 -136.34 215.52' 2101.4 .184.87 83.89 -77.25 0.00 8 1853.8 141.93 63.04 0.00 0.00 C 6811.6 594.05 0.00 0.00 0.00 0- ....... ....... ....... .......- O: !!54.22 294.94 213.59 215.52
- SE15UNFL.wt 3-19
. . ~
ORr* 00664 INDEX 9 4, REV. 0 >
?
Table 3-11 Application of Load 5_to Match Seismic Stre55e5 . Post. Accident Case 2 0 SEl5NIC STRESSES AT SECTION (psi) SECTION: 1 2 -3 4 2 0 N00E: 32 302 461 1037 COMPRES$1VE STRE55E5 FROM 2 0 ANALYSil ELEV1 119.3" - 322.5" 489.1" 912.3" 0,054" SE!5NIC DEFLECT 104:- 188.67 155.54 103.46 85.31 Hot!! PLUS VERTICAL SEI5MIC INERTIA: 499.79 429.39- ~ 2.78 123.14 TOTAL SE!5Mit COMPRESSIVE STRESSES: 1288.48 $44.93 616.22 808,45 3-0 $fRE55E5 AT SECil04 (psl) , 3.o: . INPUT- !'U10N: 1 *2 3 4 LCAD 3-0 N00Est 53-65 170-178 400-408 526 534 , SECTION- -INPUT 3-0 UNIT LOAO DE5CRIPT!DN ELEV: 119.3" 322.5" 489.1" 312.3" A- 1000 lbs at nodes 563 through Set 85.43 37.94 34.54 55.23 0 500 lbs at 4276435. 1000 lbs at 428-434 89.88 39.92 36.76 0.00 C 500 lbs at 197&205.1000 lbe at 198 204 - 97.44 - 43.37 0.00 0,00 0 . 500 the et 1616168,1000 lbe at 162154 89.85 , 0.00 -0,00 0.00 DESIREDCOMPAE551VESTRESSES(pet): 1284.46 544.93 616.22 80s.45
~
3-0
+
INPUT LOAO - , SECTION LOAO TO K APPLIED TO MATCH 2-0 STRE55ES RE5ULTING STRESSES AT SECT 10N (psi) A- 14637.9 1250.51 555.34 511.45 808.45 8- 113.78 404.77 0.00 2850.2 256.17 1 C 1941.7 -149.54 -84.21 0.00 0.00 0 3w.8 -28.84 0.00 0.00 0.00 SUM: 1288.46 544.93 418.22 608.45' SE15FL.WK1 I 3-20
ORf8 00664
- INDEX 9-4, REV. O r
4 DRYWELL l
- TH ;6.40*c. - g.G -
4;W9.. M.o ELEV. 87 5* l yy;h,.sk>t Q%N kTHg%gp gg ggj i EWWM7 mW SM l ad8diriM
. f MML .
;, . - s . n ,
; ), .. a 1 .
r/c i -t \ l i. 4*;ff" 1b ' ' s A . i,. . )
.a...l4SL. a .!..
,.: . . . . ,.+t i. r .
I i" ELEV. 51' 0*
' ~
\
'Q 1x .
' ;;. ,41 s >
- t v
\ \
..e. .. ,.
c, . : y.
.,Ji i -
.' .v?
' * * ' ' t.
-?;;'. ; ' . ? ';l4' y.
..$s ' ,\
'q
.L...La61 i % m.s . a , . .-.
- - - j ,. - 7,g*= 4
- 6 ** m- m
{
'~
'.- >A 3
.n -
*m %
,,~ ,,
+' y q
*>. 0 6 y
g 'Mf?R It. THK, .676* Figure 3-1. Oyster Creek Drywell Geometry l 3-21
. . . t t i I i l ( 1 ANSYS 4.4 l 1 ' DEC 4 1999 ' 15:06:81 j 1 PLOT MO. 1 t
- PREF 7 ELEMENTS h REAL MUM I xe =1 YU =-9.8 DIST=718.736 l XT =383.931 ,
2T =639.498 i nMC2=-Te CENTROID HIDDEM l i i i
+
i t I y i i i i Y n i t N ! l t l i i i f l l
- i t
. . i OYSTER CREEX DRYWELL ANALYSIS - OYCR10 (MO SAND. POST-ACC. ) ;
.i i i > i 4 Ficure 3-2. Ovster Creek Orywall 3-0 Finite Flamant Morial i
a 4 ANSYS 4.4 1 DEC 4 1999 i i' 13:96:41 PLOT MO. 2 ' PREF 7 ELD 4ENTS REAL MERM Xtf =1 YU =-0.3 DIST=288.376 XF =429.432
! 2F =216.329 anc2=-se CENTROID HIDDEN O
's
'N
\ {
oYSTEN CREEN DRYWELL ANALYSIS - OYCR10 (HC SAND, POST-ACC. )
. Figure 3-3. Closeup of Lower Drywell Section of FEM (Cutside view)
i t ANSYS '4.4 1 DEC 4 1999 15:97:SZ PLOT MO. 3 FREF7 ELIMENTS REAL NUM t f' XU =-1 YU =-9.8 DIST=288.376 XT =429.432 ' '2T =216.529 AMC*1.=9 e , l CENTROID HIDDEN !
- 6
! L t i \ I 1 ! Y 4 ! f I i , s l i
\x I l
OYSTER CREEN DRYWELL ANALYSIS - CYCR10 ( p SAMD, TOST-ACC. ) i J 4 Floure 3-4 Clnseem of Inwar Orvwell Section of FFM fingine View)
AMSYS 4.4 1 DEC 4 1999 ' 15:19:37 PLOT 290. E c PREP 7 ELINENTS
- TYPE Mt!M EC SVMDOLS M4 =1 YU =-e.3 i l DIST=718.786 i XT =383.831-l 2F =639.439 CEN b f HIDDEM t
Y X i 1 l f f i 4
% \ %
OYSTER CREEN DRYWELL ANALYSIS - OtfCR10 4NO SAHD, FOST-ACC.
%N I
?
I t i Fioure 3-5. Boundary Conditions of Finite Element Model j
,. "5 ij:'**
# 53 . . . .
*agdjg 488Ra mde,umg den emgJ 2 a g
6 a 8
.i Ie Z
- s
.0 f:
.w _
a J e 0 N
- ?
i LA 3 26
AMSYS 4.4 t MOU 16 1999 , 99:53:34 PLOT MO. 2 PosTL STRESS l
, STEF=1 1TER=1 SY ( AUC) ,
it!A'fs DMX =e.221773 SMH =-8174 SMX =695 . 947 e e XU 21 m YU =-e.S s es
- es# DIST=719.786 '
peeksed e se n ews WT =3e3.931 m =639.438 s e e r**ee 9e CENTMID HIDDEM
-8174
-7199 re reeee..
tHE
-3247
%.?.Y TYiY(e e . , . .,r. j 7 Y o w m v. ;
6**>>>>M ' 2 , O VeVr'rY.
\yr VeW<',
U W '<>;I* >l NElrie ' e eees.. 1 - ! OYSTER GEER DRYWELL AMAINSIS - OYCRIS (MO SAND, RETUELINCA Figure 3-7. Meridional Stresses - Refueling Case
l 5
- . ,co Rn ... a n g 3,3 EVEEv=
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e
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Q t a 0 0 t 3-28
p I l ! ir;!i ;ll;t ! -
- i' ,illlt iiitI!;) Ii! ! :? ii t -!
M E 3 D 1S 7 618 D 9 S 7 833 I 4.9 E ) 17 79 49 R C 244 3. 4. H73"48I654 13 .T U 255 93 D54 918gI32M147 765 3OS A S6: O I246 Y13 M1==11(M 9 ===6d1===== 736
.37 R S 3TTTR DM T T NU:OsEE DEXNX S N AO9LoTTXILMMM N9PPSISMEDSS UUITgCQ XYDX I
W I L E U F E R D N A S _O N 1 C 3 1 R C - t Y e s O a C
- o n
S i l I e S u f Y e
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- 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 41 and 42 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 Pefueling case and 1.67 for the Post-Accident cese te 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. l l l. 4-1
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Table 41 . Calculation of Allowable Buckling Stresses - Refueling Case l Parameter Value Theoretical Elastic Instability Stress, aj, (ksi) 58.10 - Capacity Reduction Factor, aj 0.207-Circumferential Stress, oc (ksi) 4.49 , LEquivalent Pressure, p (psi) 15.74 .]
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-AC 0.0716 .
Modified Capacity Reduction Factor, ai, mod _ 0.325 Elastic Buckling Stress. o, - ai~, mod 81e (ksi)- 18.88 . Proportione 1.imit Ratio, A = a,/ay .
- 0.497L i P1asticity. Reduction Factor, nj 'l.00 ;
Inel_astic~lnckling Stress, og - nga, (ksi). 18.88 Factor of Sifety, FS 2,0 ; 9 Allowable Comvressive Stress, 0,11_- aj/FS (ksi) . 9.44 [
- Applied- Compres.ive Meridional Stress, a, (ksi) - 7.58 - -
- Margin - - ((a,13 / r ,) . - 1] x 'a X 24.5%
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t. ORF# 00664 ; INDEX 9 4, REV. 1 ' Table 4 2 Calculation of Allowable Buckling Strosses - Post Accident Case Parameter Value Theoretical Elastic Instability Stress, ogg (ksi) 61.95 Capacity Reduction Factor, aj 0.207 Circumferential Stress, oc (ksi) 20.08 Equivalent Pressure, p (psi) 70.38 "X" Parameter 0.387
^AC 0.182 Modified Capacity Reduction Factor, ai, mod 0.508 Elastic Buckling Stress, a, - ai, mod Oie (ksi) 31.47 -
Proportional Limit Ratio, A = 0,/oy O.828 Plasticity Reduction Factor, nj 0,724 Inelastic Buckling Stress, oj - nja, (ksi) 22.78 Factor of Safety, FS 1.67 Allowable Compressive Stress,a a ll
- 81 /FS (ksi) 13.64 Applied Compressive Meridional Stress, om (ksi) 11.96 Margin - ((ca ll/8 )m - 1] x 100% 14%
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- 5.
SUMMARY
AND CONCLUSIONS The results -of this buckling analysis for the refueling and post-accident loat: combinatios s are sunimarized in Table 51. The app 1ted and allowable compressive meridional stresses shown in Table 51 are fdr the sandbed region which is the most limiting region in terms of buckling. This analysis demonstrates that the Oyster Creek drywell has adequate margir. against buckling with no sand support for an as 'ned sandbed shell thickness of 0.736 inch. This thickness is the C confidence projected thickness for the 14R outage. O e JI 5-1
f bEX h k REV. I 1 Table 5-1 Buckling Analysis Summary load Combination Refuelino Epst Accident 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% l l 5-2}}