ML20087E767
| ML20087E767 | |
| Person / Time | |
|---|---|
| Site: | Oyster Creek |
| Issue date: | 11/30/1990 |
| From: | Frederickson C, Mehta H, Ranganath S GENERAL ELECTRIC CO. |
| To: | |
| Shared Package | |
| ML20087E760 | List: |
| References | |
| DRF-00664, DRF-00664-R01, DRF-664, DRF-664-R1, NUDOCS 9201220133 | |
| Download: ML20087E767 (62) | |
Text
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- 1 DRi# 00664 IN)EX 9 2, REV.1 AN ASME SECTION Vill EVALUATION OF THE OYSTER CREEK DRYWELL PART 2 STABILITY ANALYSIS November 1990 prepared for 4
GPU Nuclear Corporation Parsippany, New Jersey prepared by GE Nuclear Energy San Jose, California I
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AN ASME SECTION Vill EVALUATION 0F THE OYSTER CREEK DRYWELL PART 2 STABILITY ANALYSIS 7 ~
Prepared by: d A [w -
C.D. Frederickson, Senior Engineer Materials Monitoring &
Structural Ana' lysis Services ,
Reviewed by: k b '
H. S. Mehta, Principal Engineer Materials Monitoring &
Structural Analysis Services Approved by:
S. Ranganath, Manager Materials ilonitoring &
Structural Analysis Services i
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. TABLE OF CONTENTS Peut 1.AINTRODUCT10N 1l ,
1.1 Genetal 11 1.2 _ Report Outline 1-1 l'. 3 References 'l 1-
- 2. BUCXLING ANALYSIS METHODOLOGY 2 - l '- ..
2.1 . Basic Approach 2-1 ;
2.2 Determination of Capacity Reduction' Factor 2 2.3' Modification of Capacity Reduction Factor for 2-3 "
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.- l
- 3.2. Finite Element Mode 10 3-2 3.3- ' Drywell Materials 3-3 3.'4 Boundary Conditions' 3-3 ,
3.5-:Loadsi 3-4 ,
3.6 Stress Results 37-3.7l Theoretical _ Elastic Buckling-Stress lResults 3 A- 3.8 Reforencesx 3-10
- 4. ALLOWABLE BUCKLING. STRESS EVALUATION 4-1 5;
SUMMARY
AND CONCLUSIONS 5-1 i
- . .. - 111 p; .
tDEXhffREV.0 LIST OF 1ABLE3 Page Table Title No, No. .
3-1 Oyster Creek Drywell Shell Thickness 3-11 3-2 Cylinder Stiffener Locations and Section Properties 3 12 3-3 Material Properties for FBX-2128 Steel 3-12 3-4 Oyster Creek Orywell Load Combinations 3-13 3-5 Adjusted Weight Gensities of Shell to Account for 3-14
- _ Compressible Material Weight ,
3-6 Oyster Creek Orywell Additional Weights - Refueling 3 15 -
3-7 Oyster Creek Drywell Additional Weights - Post Accident 3-16 3-8 Hydrostatic Pressures for Post-Accident, Flooded Case 3-17 3-9 Meridional Seismic Stresses at Four Sections 3-18 ,
10 Application of Loads to' Match Seismic Stresses - 3 19 Refueling Case 3-11 Application of Loads to Match Seismic Stresses - 3 20 Post Accident Case 41 . Calculation of Allowable Buckling Stresses - Refueling 4-2 4-2 Calculation of Allowable Buckling Stresses - Post-Accident 4-3 51 Buckling Analysis Summary 52 iv
kbX REV. 1 LIST OF FIGURES Figure Page No. Title No.
1-1 Drywell Configuration 1-2 2-1 Capacity Reduction Factors for Local Buckling of 2-8 Stiffened and Unstiffened Spherical Shells 22 Experimental Data Showing Increase in Compressive 29 Buckling Stress Due to Internal Pressure (Reference 2-6) 2-3 Design Curve to Account for Increase in Compressive 2-10 Buckling Stress Due to Internal Pressure (Reference 2-11) ,
2-4 Plasticity Reduction Factors for Inelastic Buckling 2-11 3-1 Oyster' Creek Drywell Geometry 3-21 3 Oyster Creek Orywell 3-D Finite Element Model 3-22 3-3 Closeup of Lower Dry'vell Section of FEM (Outside View) 3-23 3-4 Closeup of Lower Drywell Section of FEM (Inside View) 3-24 3-5 Boundary Conditions of Finite Element Model '
3 25 3-6 Application nf Loading to Simulate Seismic Bending- 3 3-7 Meridional Stresses - Refueling Case 3-27 3-8 Lower Drywell Meridional Stresses - Refueling Case 3-28 l
3-9 Circumferential Stresses - Refueling Case 3-29 I
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LIST OF FIGURES ,
figure Page No. Title ,jn 3-10 Lower Drywell Circumferential Stresses Refueling Case 3 30 3-11 Meridional Stresses - Post-Accident Case 3 31 3 12 Lower _Drywell Meridional Stresses Post-Accident Case 3 32 3-13 Circumferential Stresses - Post-Accident Case 3 33
-3 14 Lower Drywell Circumferential Stresses - Post-Accident 3-34 Case ,
3 15 Symmetric and Anti-Symmetric Buckling Modes 3 35 3-_16 Synmetric Buckling Mode Lhape - Refueling Case 3-36 3-17 Anti Symmetric Buckling Mode Shape Refueling Case 3-37 3-18 Buckling Mode Shape - Post'-Accident Case 3-38 vi
DRF# 00666 INDEX 9 2, REV. 0
- 1. INTRODUCTION 1.1 General To address local wall thinning of the Oyster Creek drywell, GPUN has planned to prepare a supplementary report to the Code stress report of record [1 1). - For convenience, the supplementary report is divided into two parts. Part 1 of the supplementary report (1-2] includes all .
of the Code stress analysis results other than the buckling capability for the drywell shell. This report addresses the buckling capability of the drywell shell shown in Figure 1-1 and constitutes the second part of the supplementary report. Buckling of the entire drywell shall is considered in this analysis with the sandbed region being the area of primary concern.
1.2 Report Outline ,
Secti-~ S of this report cutlines the methodology used in the buckling '
capabliity evalua. tion. 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 corhinations are covered in section 4.-
- Section 5 presents the summary of results and conclusions.
1.3 References 1-1 " Structural Design of the Pressure Suppression Containment Vessels," by Chicago Bridge & Iron Co., Contract # 9 0971, 1965.
1-2 "An ASME Section Vill Evaluation of the Oyster Creek Orywell," GE Report No. 9-1, DRF# 00664, November 1990, prepared for GPUN.
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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 [ References 21, 2-2]. Following the procedure of this Code Case, the allowable compressive stress is evaluated in three steps.
In the first step, a theoretical elastic buckling stress, oje, is determined. This value may be calculated either by classical buckling equations or by finite element analysis. Since the drywell shell geometry is : complex, a three dimensional finite element analysis approach is followed using the eigenvalue extraction technique. More details on the eigenvalue determination are given in Section 3.
In the second step, the theoretical elastic- buckling stress is modified by the appropriate capacity and plasticity reduction factors.
The capacity reduction factor, ej, accounts for the difference between classical -buckling theory and actual tested buckling stresses for fabricated shells. This difference is due to imoerfections 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., aj,oj. When the elastic buckling stress exceeds the proportional limit of the material, a plasticity reduction factor, gj, is used to account for non-linear material behavior. The inelastic buckling stress for fabricated shells is given by njajoje.
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 2-1
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Ini Reference 2-1, the safety factor for the Design and Level A & B service conditions :is specified as 2.0. A safety factor of 1.67 is
- specified for Level- C- service conditions (such as the post-accident j
- floooed _ condition).
- The Determination - of appropriate values for- capacity and plasticity reduction factors is discussed next.
2.2 Determination of_ Capacity Reduction Factor :
The capacity reduction factor, aj, is used to account for reductions in actual buckli_ng strength due to the existence -of _ geometric .
imperfections. The_ capacity reduction factors given in Reference 2-1 are based on extensive data compiled by Miller (2-3). The factors appropriate-for a spherical shell geometry _such as that of the drywell in; the - sandbed region, are shown in Figure 2-1_(Figure 1512-1 of
- Reference 2-1). The . tail (flat) end of. the curves are used for unstiffened , ells. 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, a4 is determined to be 0.207. ,
- The preceding .value of the capacity reduction factor .is very conservative for two reasons. First, it is based on the assumption that the spherical _shell. has a uniform thickness equal _to the reduced thickness. However, the drywell shell has a greater thickness above the - sandbed region which would reinforce the sandbed region. Second, itiis assumed' that the circumferential stress is zero. The tensile circumferential stress has the effect; of rounding. the shell and reducing the effect of imperfections introduced'during.the fabrication
-and construction phase. A modification 'of'the aj value to account for the : presence of ' tensile circumferential stress is discussed in Subsection 2.3.
The capacity reduction factor values given in Reference 2-1 are l
applicable to shells which meet the tolerance requirements of NE-4220 L
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of Section -!!! (24]. Appendix A of- Reference 2 5 compares the tolerance requirements of NE-4220~ to the requirements to which the ,
'0yster 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 Ill, NE vessel, it is justified to use the approach
~
outlined in Code' Case N 284.
2.3 Modification of Capacity Reduction Factor for Hoop Stress The orthogonal tensile stress has the effect of rounding fabricated shells and reducing the effect of imperfections on the buckling strength. The Code Case N 284 (2-1 and 2-2) notes in the- last paragraph of Article 1500 that, "The influence of internal pressure on a shell structure may reduce the initial imperfections and therefore higher- values of capacity reduction factors may be acceptable. -
Justificat' ion for - higher values of aj must be given in the Design report." ;
The effect of hoop tensile stress on the buckling strength of cylinders has been extensive 11y documented [2 through 2-11). Since the methods used in accounting for the effect of tensile hoop stress for the' cylinders and spheres- are similar, the test data and the methods- for the cylinders. are first~ reviewed.- Harris, et al (2-6]
presented a comprehensive set of test data, including those' from-
- References 2 7 and 2-8, which clearly showed ~that in_ternal 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 t'n buckling ~ stress as a function of nondimensional pressure.
This increase -in buckling c.pacity is _ accounted for by defining. a separate reduction factor, op. The capacity reduction factor og can
'then be modified.as follows:
p 01, mod " og + ap 2-3
ItNEXNhfREV.1 I
The buckling stress in nniaxial compression for a cylinder or a sphere of uniform . thickness with no internal pressure is given by the following:
Se- (0.605)(ag)Et/R (0.605)(0.207)'Et/R Where, 0.605 is a constant 0.207 is the capacity reduction factor.ag, 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)(og + op)Et/R (0.605)(0.207 + op )Et/R
[(0.605)(0.207) + AC) Et/R .
Where AC is pa /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 [2-12) gives .the following equation that fits the graphical relationship between X and AC shown in Figure 2-3:
AC - ap/0.605 - 1.25/(5+1/X)_
The -preceding approach pertains to cylinders. Along the similar lines, Miller [2-13] has developed an approach for spheres as described next.
The non-dimensional parameter X is essentially (og/E)(R/t). Since in the case of a sphere, the hoop stress is one-half of that in the cylinder, the parameter X is redefined for spheres as follows:
X(sphere) - (p/8E)(2R/t)2 2-4
EX k REV. 1 When the tensile stress magnitude, S, is known, the equivalent internal pressure can be calculated using the expression:
p= 2tS/R Based on a review of spherical shell buckling data (2-14, 2 15),
Miller (2-13] proposed the following equation for AC:
AC(sphere) = . 00/(3.24 + 1/X)
The modified capacity reduction factor, oi, mod, for the drywell geometry was obtained as follows:
ai, 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 plasticity reduction factor, nj, is used to account for the non-linear material behavior. The inelastic buckling stress for fabricated shells is given t.y ngajoje. 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,/oy and oy is the material yield strength. Figure 2-4 shows the relationship in graphical form.
nj = 1.0 if A s 0.55
- (0.45/4) + 0.18 if 0.55 < A 51.6
- 1.31/(1+1.15A) if 1.6 < A s 6.25
= 1/4 if A > 6.25 2.5 References 2-1 ASME Boiler and Pressure Vessel Code Case N 284, " Metal Containment Shell Buckling Design Methods,Section III, Division 1, Class MC", Approved August 25, 1980.
2-5
kib$XNhfREV.1 2-2 Letter (1985) from C.D. Miller to P. Raju;
Subject:
Recommended Revisions to ASME Code Case N 284.
23 Miller, C.D., " Commentary on the Metal Containment Shell Buckling Design Methods of the ASME Boiler and Pressure Vessel Code," ,
December 1979.
24 ASME Boiler & Pressure Vessel Code, Section !!!, Nuclear Power Plant Components.
25 " Justification for Use of Section III, Subsection f4E, Guidance in Evaluating the Oyster Creek Drywell," Appendix A to letter dated December 21, 1990 from H.S. Mehta of GE to S.C, Tumminelli of GPUN.
2-6 Harris, L.A., et al, "The Stability of Thin-Walled Unstiffened '
Circular Cylinders Under Axial Compression Including the Effects of Internal Pressure," Journal of the Aeronautical Sciences, Vol.
24, No. 8 (August 1957), pp. 587 596.
2-7 Lo, H., Crate, H., and Schwartz, E.B., " Buckling of Thin Walled Cylinder Under Axial Compression and Internal Pressure," NACA TN 2021, January 1950, 2-8_ Fung, Y.C., and Sechler, E.E., " Buckling of Thin Walled Circular Cylinders Under Axial Compression and Internal Pressure," Journal of the Aeronautical Sciences, Vol. 24, No. 5,- pp. 351-356, May 1957.
2-9 Baker, E.H., et al., "Shell Analysis Manual," NASA, CR-912 (April 196B).
2-10-Bushnell, D., " Computerized Buckling Analysis of Shells," Kluwer Academic Publishers,1989 (Chapter 5),
j 2-11 Johnson, B.G., " Guide to Stability Design Criteria for Metal Structures," Third Edition (1976), John Wiley & Sons.
2-6 l
kibEXNhfREV.i 2-12 Miller, C.D., " Effects of ernal Pressure on Axial Compression :
Stren9th of Cylinders " CB (.;anical Report No. 022691, February i 1991- l 213 Miller, C.D., 'Evaluetion of Stability Analysis Methods ilted for
. the Oyster Creek Orywell," CBI Technical Report Prepared for GPU Nuclear Corporation, September 1991.
2 14 Odland, J., " Theoretical and Experimental Buckling Loads of Imperfect Spherical Shell Segments," Journal of Ship Research, Val. 25. No.3, September 1981, pp. 201 218. i 2 1$ Yao, J.C., " Buckling of a Truncated Hemisphere Under Axial
'fension," AIAA Journal, Vol. 1, No. 10, October 1963,-pp. )
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- 3. FINITE ELEMENT MODELING AND ANALYSIS 3.1 finite Element Buckling Analysis Methodology
' This evaluation of the Oyster Creek Drywell buckling capability uses ,
the Finite Element Analysis (FEA) program ANSYS (Reference 31). The l
ANSYS program uses a two step eigenvalue formulation procedure to !
perform linear elastic bucklir.g analysis. The first step is a static analysis of the structure with all anticipated loads applied. The '
structural stiffness matrix, [K), the stress stiffness matrix, [S),
and the applied stresses, a,p, are developed and saved from this
. static analysis. A buckling pass is then run to solve for the eigenvalue or load factor, for which elastic buckling is predicted using the equation: ,
( [K) + 1 (S) ) (u) = 0 where: A is the eigenvalue or load factor. ,
(u) is the eigenvector representing the buckled shape of the structure.
This load factor is a multiplier for the applied stress state at which the onset of elastic buckling will theoretically occur. All applied loads (pressures, forces, gravity, etc...) are scaled equally, For ,
example.- a load factor of 4 would indicate that the structure would buckle for a load condition four times that defined in the stress pass. The critical stress, a ct, at a certain location of the structure is thus calculated as:
O cr
- A #ap Inis theoretical elastic buckling stress is then modified by the capacity hnd plasticity reduction f actors to determine the predicted I buckling stress of the fabricated structure as discussed in Section 2.
This stress is further reduced by a factor of safety to determine the a110wable comprpssive stress.
31 l
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.3. 2 finite Element Model ,
I The Oyster Creek drywell has been previou!!y analyzed using a simplified axisymmetric model to evaluate the buckling capability in the sar.cbed region [ Reference 32). 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. t The geometry of the Oyster Creek drywell is shown in figure 31.
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 i significant structural interaction between the drywell and torus.
4 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 0.700 inch. This is less than the 95'4 confidence projected thickness for outage 14R. Figure 3 4 shows the view from the inside of the drywell with the gussets and the vent jet deflector. ,
The drywell and vent shell is modeled using the 3 dimensional plastic quadrilateral' shell (STif43) element. Although this element has-plastic capabilities, this analysis is conducted using only elastic behavior. This element type was chosen over the elastic quadrilateral shell (ST!f63) element because it is better suited for modeling curved i surfaces. ,
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At a distance of 76 inches from the drywell shell, the vent is l sirnplified 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 STlF4 beam elements are then connected to this ;
centerline node to model the axial and bending stiffness of the vent ;
and header. Spring ($T!f14) elements are used to model the vertical f header supports inside the torus. ANSYS ST!f4 beam elements are also used to model the stiffeners in the cylindrical region of the upper drywell. The section properties of these stiffeners are summarized in Table 3 2.
The sandbed region at the base of the drywell was designed to provide a smooth transition in reduce thermal and mechanical discontinuities.
The sand provides lateral support to the drywell sphere in this region. The-foundation stiffness for the sandbed is considered to be ,
366 psi /in per Reference 2.4.10 of Reference 3 2. AftSYS SilF14 spring elements are extended radially outward from each node of the shell in ,
the sandbed region to model the sand support as shown in figure 3 3.
The stiffness for each of these sand spring elements is calculated by multiplying the foundation stiffness of the sand by the contributory area of each node in the sandbed region.
3.3 Drywell Materials The drywell shell is fabricated from SA 2128 fBX steel, 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 i to the perforations.
3.4 Boundary Conditions
- Symmetric boundary conditions are defined for both edges of the 36*
1 drywell model for the static stress analysis as shown on figure 3 5.
-This allows the nodes at this boundary to expand radially outward frc:n 33
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th X N REV. 0 the drywell centerline and vertically, but not in the circumferential direction. Rotations are also fixed in two directions to prevent the ,
boundary from rotating out of the plane of symmetry. Nodes at the bottom edge of the dr/well are fixed in all directions to simulate the fixity of the shell within the contrate foundation. Nodes at the ends .
of the sand spring elernents and the header support spring elements are also fixed.
3.5 Loads The loads are applied to the drywell finite element model in the manaer which most accurately represents the actual loads anticipated on the drywell. Details on_the application of loads are discussed in the following raragraphs.
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 stretses in the sandbed t egion. 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 ahalysis 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 c*ondition with seismic is very similar to this condition, however, it will be less severe due to the absence of the refueling water and equipment door weight.
3-4 i
NNX REV. 0 The most ' severe load combination for the emergency condition is for the post accident (tase VI) load combination including:
Dead weight of vescel, penetrations, compressible mat'erial 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 luads for flooded condition The application of these loads is described in more detail in the following sections.
3.5.2 Gravity Loads
' he gravity loads include dead weight loads of the drywell shell,*
T weight of the compressible material and penetrations and live loads.
The drywell shall 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 acceteration. The compressible material weight of 10 lb/ft! is added by adjusting the weight density of the shell to also include the
-compressible material. The adjusted weight densities for the varions shell thicknesses are summarized on Table 3 5. The compressible material is assumed to cover the entire drywell shell (not including the vent) up to the elevation of the flange.
~
The additional dead weights, penetration weights and live loads are applied as additional nodal masses to the model. As shown on Table
-3 6 for the _ refueline 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 35
If X REV. 0 method is used to apply the additional masses to the model for the post accident, flooded case as summarized in Table 30. r 3.5.3 Pressure Loads
- The 2 psi external pressure load for the refueling case is applied to the external faces of all of the drywell and vent shell elements. The compressive axial stress at the transition from vent shell to beam elements is simulated by applying equivalent axial forces to the nodes :
of the shell elements. .
Considering the post accident, flooded case, the drywell is assumed to .
be flooded to elevation 74' 6" (894 inches). Using a water density of 62.3 lb/ft3 (0.0361 lb/in ) the pressure gradient versus elevation is 3
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 prestures are applied to the inside surface of the shell elemants.
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 en the drywell during a seismic event due to, a 0.058" deflection of the reactor building and due to horizontal and vertical inertial ioads on the drywell.
The meridional stresses due to a seismic event are imposed on the 3 D drywell model by applying downward forces at four elevations of the model (A: 23'-7",B: 37' 3",C: 50' 11" and 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 i
the load distribution in the lower drywell while also providing a .
l raasonably accurate stress distribution in the upper drywell.
i 36
X REY. 0 To flad the correct loads to match the seismic stresses, the total r 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 ad the corresponding meridional stresses for the refueling (unflooded) and post accident (flooded) seismic cases are summarized in Table 3 9.
Unit loads are then applied to the 3 0 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 ccrrect 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 310 and 3 11 for the refueling and post accident cases, respectively.
3.6 Stress Results ,
The resulting stresses for the two inad combinations desvaed 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 i 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 en figures 3-9 and 3 10. The resulting average meridional stress at the mid elevation of the sandbed region was found to be; 1
arm = 7097 psi 9
1 37
lbfx REV. 0 The circumferentici stress averaged frnm the bottom to the top of the sandbed region is; orc 277 psi 3.6.2 Post Accident Condition Stress Results
'The application of all of the loads described for the post accident condition results in the stress distributions shown in Figures 311 through 3 14. The red colors represent the most tensile stresses anit the blue colors, the most compressive. Figures 's ll and 3-12 show the meridional stresses for the entire drywell and lower drywell. The circumferential stresses for the same areas are shown on Figures 313 and 3 14. The resulting average meridional ; tress et mid elevation of the sandbed region was found to be; oPAm " '9693 psi The circumferential stress averaged from the botton to the top of the sandbed region is; O pac = +4049 psi 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 described in Section 3.1. This analysis determines the theoretical elastic buckling loads and buckling mode shapes.
3.7.1 Refueling Condition Buckling Results As shown on Figure 3-15, it is possible for the drywell to buckle in two different modes. In the case of symmetric buckling shown on '
Figure 315, each edge of the 36' drywell model experiences radial displacement with no rotation. This mode is simulated by applying symmetry boundary conditions to the 3 D model the same as used for the
. stress run. Using these boundary conditions for the refueling case, the critical load factor was found to be 14.32 with the critical ,
buckling occurring in the sandbed region. The critical buckling mode shape is shown in Figure 3-16 for applied symmetry boundary conditions. The red color indicates sections of the shell which displace radially outward and the blue, those areas which displace inward.
The first four buckling modes were solved for in this eigenvalue buckling analysis with no buckling modes found outside the sandbed region for a load factor as high as 16.32. Therefore, buckling is not a concern outside of the sandbed region.
It is also possible for the drywell to buckle in the anti symmetric ,
manner shown in Figure 315. For this mode, the edges of the 3 D ,
model are allowed to rotate but are restrained from expanding radially. This case is considered by applying anti symmetric boundary l
conditions at the edges of the 3 D model. With the.two pass approach I used by ANSYS, it is possible to study anti-symmetric buckling of the l drywell when the stresses are found based on' symmetry tioundary conditions. The resulting load factor found using anti symmetric b'oundary conditions is 16.81. The mode shape for this case' is shown on Figure 3 17.
3-9 ,
hb[X REV. 0 Because the load factor is lower for symmetry boundary conditions with the same applied stress, the symmetric buckling condition is more limiting. Multiplying the load factor of 14.32 by the average meridional stress from section 3.6.1, the theoretical elastic buckling stress is found to bei O Rie = 14.32 x (7097 psi) = 101,650 psi 3.7.2 Post Accident Condition Buckling Results Considering the post accident case with symmetry boundary conditions, the load factor was calculated as 9.91. Multiplying this load f actor by the applied stress from section 3.6.2 results in a theoretical elastic buckling stress of oPAie = 9.91 x (9693 psi) = 96,060 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 3-1 DeSalvo, G.J., Ph.0, and Gorman, R.W., "ANSYS Engineering Analysis System User's Manual, Revision 4.4," f,wanson Analysis Systems, Inc., May 1, 1989.
32 GPUN Specification SP 1302 53 044, Technical Specification for Primary Containment Analysis - Oyster Creek Nuclear Generating Station; Rev. 2, October 1990.
l 33 "An ASME Section Vill Evt.'uation of the Oyster Creek Orywell -
Part 1 Stress Analysis," GE Report No. 9-1, ORf # 00664, November 1990, prepared for GPUN.
i l
3-10
i bEXh!hfREV.0 i
Table 3 1 Oyster Creek Orywell Shell lhicknesses Section Thickness (in.)
Sandbed Region 0.700 Lower Sphere 1.154 Mid Sphere 0.770 Upper Sphere 0.722 ,
Knuckle 2.5625 '
Cylinder 0.640 Reinforcement Below Flange 1.250 Reinforcement Above flange 1.500 Elliptical Head 1.1875 -
Ventline Reinforcenent 2.875 Gussets 0.875 Vent Jet Deflector 2.500 ,
Ventilne Connection 2.500
-Upper Ventline 0.4375 Lower Ventline 0.250 s
1 I 3 11
Ifd!XhIhfREV.O Table 3 2 Cylinder Stiffener Locations and Section Properties 4
' Elevation Height Width Area Eendina Inertia (in )
(in) (in) fin) (int) . Horizontal Vertical 966.3 0.75 6.0 4.5 13.5 0.211 1019.8 0.75 6.0 4.5 13.5 0.711 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 a 2 beam sections, one 2.75x7" and one 1.0x7.375" I'
Table 3 3 Material Properties for FBX 212B Steel Material Property Value_
Young's Modulus 29.6x106 pgg .
) eld Strength 38000 pst Poisson's Ratio 0.3 Density 0.283 lb/in 3 3 12
i XNhfREV.0 Table 3 4 Oyster Creek Drywell Load Combinations CASE I INITIAL TEST CONDITION Deadweight +DesignPressure(62 psi)+ Seismic (2xDBE)
CASE 11 FINAL TEST CONDITION Deadweight + Design Pressure (35 psi) + Seismic (2 x DBE)
CASE Ill NORMAL OPERATING CONDITION Deadweight + Pressure (2 psi external) + Seismic (2 x DBE)
CASE !V REFUELING CONDITION Deadweight + Pressure (2 psi external) + Water Load + ,
Seismic (2 x DBE)
CASE V ACCIDENT CONDITION Deadweight + Pressure (62 psi 9 175'T or 35 psi 9 281*f) +
-Scismic (2 x DBE)
CASE VI POST ACCIDENT CONDITION Deadweight + Water Load 9 74'6" + Seismic (2 x DBE) 3 13
kUfxh!IfREv.0 Table 3 5 Adjusted Weight Densities of Shell to Account for Compressible Material Weight Adjusted Shell Weight Density Thickness (in.) (1b/in3) 1.154 0.343 0.770 0.373 0.722 0.379 2.563 0.310 0.640 0.392 ,
1.250 0.339 ,
1 a
3-14
ORf: 00664 INDCX 9 2, REV 0 Table 3 6 Oylter Creek Drywell Additional Weight 5 . Refueling Condition d
6 10fAL $ FOOT LOAD PER (CAD F(8 LOA 0 Pit DEAD P(Wit 8. MISC.
LOA 0 Ban 4[ 36 0(6 f 0F W30(5 0F rutt 400( MALF n00(
(LtVAT10N VtlGHT Vil6HT LOA 05 (it,f)
(1bf) tiththT3 APPLICAfl0N (1bf)
(1bf) (1bf) (ibf) (Ibf) LOAD (feet) 15,56 50000 50000 16 It>8100 168100 20 11100 11100
!!9300 !!930 6 116 119 alt? 1911
" 15 10 tri 556000 556000
$$600 8 101 169 6950 3475
$16000
" 21 25#
t$ 11100 11100 30 64100 51500 115600 205000 30.tl 105000 100000 33170 8 179 167 4146 1073 331700
" 26 30 '
31 16500 16500 32 750 750 33 15450 15450 '
34 18050 18050 35 1500 1500 8 188 196 778 389 6ftSO 6tti
" 31 35 36 1550 1550 '
40 41000 43350 44350 8 197 105 1074 537 85900 8590
" 36 40
$0i 1102000 1101000 8 416 416 13775 E868
!!02000 110200
" 45 507 54 7850 7850 8 436 444 98 49 7850 765
" 51 55 56 56400 24000 80400 95200 700 20000 115900 60 196300 19630 8 454 462 <2454 !!!7
" 5f 50 65 $2000 20000 72000 8 472 480 900 450 72000 7200
" 61 65 70 5750 5750 8 508 516 ft 36 5750 575
" 66 70 73 8850 8850 8 St6 534 111 55 8850 885
" 71 75 82.17 !!650 !!650 8 553 561 271 135 21650 2165
" 8135 87 1000 1000 90 15000 15000 8 $71+579 200 100 16000 1600
" 86 90 93.75 20700 20700 94,758 698000 698000 95.75 10100 20100 8 589 597 9235 4618 738800 73880
" 91 9 A 862000 3434350 3434350 343435 107ALS: !!84150 388200
- . LDAD 10 BE APPLito IN VCRTICAL CIRECil0N ONLY.
& . MISCELLAht0US LOA 05 INCLUDE 698000 L8 VAlt8 VilGHT AT 94.75 FT. (LivAf!0N 100000 L8 (QUIP'1[NT 000R WilGHT *AT 30.25 FT. ELEVATION AND VELD PA0 LIVE LOADS OF 24000. 20000 AhD 20000 At 56. 60 AND 65 FT [LEVAi!0h5 8tFVGT.WK1 3-15
b x N REV. O Table 3 7 Oyster Creek Drywell Additional Weights . Post. Accident Condition 6
10fAL 5 FOOT LO4 Pt8 10A0 Ptt LDAD PtR OfA0 PtNt?8. MISC.
RAut 36 OtG. # OF h00tl 0F FULL n00t MALF heet (LlyAT104 VilGHT VtlGHT LCA05 LOAD (1bf) (1bf) (LLMth15 APFLICAT104 (1bf) (Itf)
(feet) (1bf) (Itif) (1bf) LOAD 50000 50000 15.56
!$ 168100 158100 20 11!00 !!!00
!!9300 !!930 6 116 119 38:t lill
. 15 20 555000 tri 556000 55600 8 161 169 6950 3475 556000
.. fl.15s 26 11100 11100 64100 $1500 115600
+
30 30.25 105000 105000 8 179 187 1896 1448 231700 23170
. 26 30 31 16500 16500 750 750 32 33 15450 15450 34 28050 26050 ,
35 1500 1500 6tti 8 . 188 196 778 389 62250
. 31 35 '
36 1550 1550 40 41000 43350 84350 8 197 205 1074 537 85900 8590
. 3g.40 f,0f 1102000 !!02000 6888
!!0200 8 418 426 13775 1102000
.e 45 50f 54 7850 7850 436 444 98 49 7850 785 8
.* 51 55 56 56400 56400 60 95!00 700 $5900
- 8 454 46! 1904 952 15t300 15230
. 56 60
?5 5!000 52000 47t*480 650 3t5 5!000 5!00 8
. 61 65 70 5750 5750 8 508 516 ft 36 5750 575
. 66 70 73 6850 8850 8 5!6 534 111 55 8850 885
.. yg.y$
82.17 21650 !!650 IJ5
!!65 8 553 561 271
!!650
. 31 85 87 1000 1000 90 15000 15000 8 571 579 200 100 16000 1600
" 56 90 93.75 20700 20700 95.75 20100 20100 4080 8 589 597 510 !$$
- 9. gg.96 40800
........ ........ ........ ........ ........ -..***. 257235 TOT ALS: !!64150 388200 0 2572350 2572350 t . LCAD 10 St APPLit0 th VERf! CAL DIRECT 10N Ohlf.
6 NO M15CILLAkt005 LOADS FOR TH!$ 00holfl0N.
9 FLOOD 41.VK!
3-16 i
i ' " ^ - ~ - = ~ - - - - n . .. ___, __ ,
khbbXhkREV.0 l
l Table 3 8 i
Hydrostatic Pressures for Post. Accident, flooded Condition l WATER DENSITY: 62.32 lb ft3 0.03606 lb in3 FLOODED ELEY: 74.5 ft .
894 inches ANGLE ELEMENTS ABOVE ELEVATION DEPTH PRESSURE AD0VE EQUATOR NODES (degrees) (inch) (inch) (psi) ELEMENTS 53.32 110.2 783.8 28.3 1 12 27 40 51.97 116.2 777.8 28.1 13 24 50.62 122.4 771.6 27.8 25 36 53 66 49.27 128.8 765.2 27.6 37 48 47.50 137.3 756.7 27.3 49 51, 61 66 ,55 57 79 52 54, 138 141 .58160 92 46.20 143.9 750.1 27.1 153.4 740.6 26.7 240 242, 257 259 102 44.35 142 148 147,151, 243, 256 108 41.89 166.6 727.4 26.2 39.43 180.2 713.8 25.7 152 155, 244, 255 112 156 159, 245, 254 116 36.93 194.6 699.4 25.2 34.40 209.7 684.3 24.7 160 165, 246, 253 120 31.87 225.2 668.8 24.1 166 173 124 652.7 23.5 174185,24724$.252 251 130 29.33 241.3 26.80 257.6 636.4 23.0 184 195 138
-24.27 274.4 619.6 22.3 196 207 148 208 215 161 20.13 302.5 591.5 21.3 14.38 342.7 551.3 19.9 216 223 170 224 231 179. 8.63 384.0 510.0 18.4
-2.88 425.9 468.1 16.9 232 239 189 430 437 197 2.88 468.1 425.9 15.4 8.63 510.0 384.0 13.8 438 445 400 446 453 409 14.38 551.3 342.7 12.4 20.13 591.5 302.5 10.9 454 461 418 462 469 427 2E.50 '27.8 266.2 9.6 30."# f30.2 233.8 8.4 470 477 436 478 485 445 35,40 t t0.9 203.1 7.3 40.5L '49.8 174.2 6.3 486 493 454 494 501 463 45.50 /46.6 147.4 5.3 50.50 771.1 122.9 4.4 502 509 472 510 517 481 54.86 790.5 103.5 3.7 805.6 80.4 3.2 518 525 490 .
526 533 499 . 820.7 73.3 2.6 835.7 58.3 2.1 534 541 508 .
542 549 517 . 850 9 43.2 1.6
. 80', . 3 8.7 0.3 550 557 526
. . . 13i 3~ 706,7 25.5 340 399 (Ventline)
FLOOOP.WK1 l
3 17
i IbbfxNkfREV.O Table 3 9 !
Meridional Seismic Stresses at four Sections .
20 Shell Meridional Stresses Elevation Model Refueling Post Accident Section (inches) H,ogg Iosi)_ (osi)_
119 32 1258 1288 A) Middle of Sardbed B) 17.25' Below Equator 323 302 295 58$ ,
C) 5.75' Above Equator 489 461 214 ' 616 1037 1037 216' 808
. 0) Above Knuckle s 4
4 9
i t
4 4
3 18
i b X N ACV. 0 :
4 Table 3 10 Application of Leads to Hatch Seilmic Strellel . Refueling Cale j F
' !0lil$NjCSTiltll($At$(Cfl0N(Fli) f lEC110Ns I t 3 4 l 2 0 h00ti St H! ell 101 l (LtV lit.3* St!.l* 449.l* tit.3* -
ConPRtillVt littlitl Ftom 2 0 A>&Lilll 180.lf Ill.64 103. 4 $$.31 l 0 MI* Stl5NIC OtFLLCtl0N 11023 130.!! i 469.66 !)f.44 Mott!. PLU$ vttilCA6 lt!!NIC thttilAt 4...... ........ .......
i 1858.t! 2H 94 163.59 !!l.lt I
TOTAL ltilNIC CDNPitillVI STRLlltli t
i lap $14tilfiAt$ttt!0N(pst)- , j e
................................... t 30 SICfl04: 1 t 3 4 r l> pvt I 3 0 h00tli $3 65 l'*0 178 400 404 118 514 LOAD Mt . l* 449.l* 912.3" (LtVi 119.3* ;
intCt10N l> PVT. 3 0 UNif LDAb Otttt.lPfl04 -
ll.43 31.94 34.H ll.!) l A 1000 lb4 61 nodes 6 0 thfevg5 lu ;
89.64 39.lt H.76 0.00 8 500 lbs et 4176435, 1000 lbs at 424 434 0.00 l ti,H 43.37 0.00 C 6001H et itiuM.1000 lbs et lH t04 St.86 0.00 0.00 0.00 0 ' $00 lbs et 1816169. 1000 the et ilt.l u ....... ....... -....... .......
Ot118t0ComPetllivtlittl5tl(tst): litt.it !94.96 !!).59 til.it .
30 i l>PUT LOAD httutil>6 lfRLilts At $tC110N (pst) lttfl04 LCAD TO DE APPLl!D 10 NATCH 2 0 littlltl
- A 3902.3 3H.37 144.0$ lH.34 til.it IH.87 83.09 77.25 0.00 >
e ~ !!01.4 -
141.93 .63.04 0.00 0.00 I' 1453.6 C
0.00 9.00 0.00
-0 L Hil.$ lH . 05 -
Stet;- !!$4.!! 294.94 til.ll til.lt !
?
e f
StilWL.W! ,
4 L
t, b
i 3 19
. , _ . . . . _ . ,.L 1. , , , ...J.a_,__.... ,._.-._.,,,.m...,_.,,_ _ _. _... . ,... .,_. .. - .... . ..,_ . _ ,,,,_,,_ _ . ,. 4 ,. w _ f
i
_.0Rr, 00666 INXX 9 7. REV. 0 Table 3 11 g i
f Application of Loads to Hatch Seilmic Strellel Polt. Accident Cale fii e
iF
!.0 tillNIC littllti At stCT10N (pst) ;
StC110N: I ! 3 4 .
t 0 h00tt 32 30t all 101/ . I Ctmeratst!Vt $1 stilts ItDM t 0 ANAlfill (Livt 119.3* 3tt.l* 449.l* 912.3* l p
t
- 0.0H StinM!C Mrticil0N:
f64.6! 156.54 103.46 66.31 499.79 4tt.39 lit.76 ?t3,14 NORit. PLV$ vttfitAL lillMIC INttilA: ....... ....... ....... ....... ,
........te............................... '
T0fAL ltllMit COMPRtlllVE littlltle itH.46 $44.93 Gil.!! 004.45 3 0 littlltl AT lici!04 (psi) ,
- ................................... t 3.p . r littl0N 1
- t 3 4 thPVT f 3 0 #00tle 63 65 170 178 400 408 $t6 134
, 1040 att.5" 409.1" $12.3* ,
input 3 0 UNif LOAD OllttlPfl0N ELtV: 110.3" SICfl0N. ....... ....... ....... .......
-....... ......................................... I
$$.43 31.44 34.94 _ ll.23 A 1000 the et nodes $43 throwth $69- 36.76 0.00 ;
-49.48 30.92 I $00 lbs et 4t?6435.10001he et 4!8 434 43.37 0.00 0.00 i 91.64 C 600 lbe et 19fkt06.10001he et 1H 204 - 0,00 0.00 0.00 .
500 lbs et 1616169,1000 lbs et 162.lAt 48.6%
0- ....... ....... ....... ....... _j
. Miltt0_COMPetlllVE STRtlill (pst): !!48.46 - H4.93 616.t! tot 45 ,
-30
' thPUT
. LOAD .atlutilNslittlltsAtlittl0N(pst) !
littl04_ LOAD 10 DC APPLit0 TO MAtcM t.0 littlltl ... ............................... .!
A 14637.9 !!$0 51 $$$.36 $11.45 808.45 8
, 28$0.2 fH.lf !!3.78 '104.T! 0.00 169.58 64.!! 0.00 . 0.00 (
C-1941.7 '
.it.64 0.00- 0.00 0.00 0
318.6 SUM: litt.46 $44.93 616.72 808.46 l
lillIL.W1 ,
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iYhdki ., h gllllll"T?";g41 l
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yisu,;ri;:o %?:ryef.Q R THl Qy ;{; Q . ,jggg pggy j lifd3di! 'J.Th. !N5 aTjdidas&l f ENk ,
ELEV. 51' 0"
' 7-Me% $4Cf4MV '
Hy pg E THK. ,676' Figure 3 1. Oyster Creek Drywell Geometry G
3 21
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OYSTER CREEX DRYMELL ANALYSIS - OYCR10 CSAND, REFUELIP9G3 i
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ibEXhkREV.0
- 4. ALLOWABLE BUCKLING ST .35 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 factor: are first calculated as described in sections 2.2 and 2.3. After reducing the theoretical instability stress by this reduction factor, the plasticity reduction factor is calculated and applied. The resulting inelastic buckling stresses are then divided by the factor of safety of 2.0 for the Refueling case and 1.67 for the Post-Accident case to obtain the final allowable compressive stresses.
The allowable compressive stress for the Refueling case is 10.44 ksi.
Since the applied compressive stress is 7.10 ksi, there is a 47?; ,
margin. 1he allowable compressive stress for the Post Accident, flooded cast is 14.34 ksi. This results in a margin of 485; for the '
applied compressive stress of 9.69 ksi.
~
41
1 EXhhfREV.1 Table 41 Calculation of Allowable Buckling Stresses - Refuelinn Case Parameter Eglug Theoretical Elastic Instability Stress, aj, (ksi) 101.65-Capacity Reduction Factor, oj 0.207 Circumferential Stress, oc (ksi) -0.28 Equivalent Pressure, p (psi) 0.000
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"X" Parameter 0.000 AC 0.000 Modified Capacity f< eduction Factor, ai, mod 0.207 Elastic Buckling Stress, o, - ai, mod C ie (ksi) 21.04 .
Proportional Limit Ratio, A = a,/oy . 0.554 Plasticity Reduction Factor, nj 0.993 Inelastic Buckling Stress, aj - nja, (ksi) 20.89 Factor of Safety, FS 2.0 Allowable Compressive Stress,a a ll
- 31 /FS (ksi) 10.44 Applied Compressive Meridional Stress, a, (ksi) 7.10 Margin = [(aa ll/8 m) - 1] x 100% 47%
4 4-2
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X REV. - 1 {
Table 4 2 +
Calculation of Allowable Buckling Stresses Post Accident Case Parameter Value
- Theoretical Elastic Instability Stress, aj, (ksi) 96.06 ,
Capacity Reduction Factor, at 0.207 Circumferential Stress, og (ksi) 4.05 Equivalent Pressure, p. (psi) 13.50 ,
"X" Parameter 0.082 AC- 0.069 Modified Capacity Reduction Factor, ai, mod 0.32 Elastic- Buckling Stress, o, - ai, mod 8te (ksi) 30.74 ' -
Proportional Limit Ratio, A = a,/ay -
0.809' P1asticity Reduction Factor, nj.
0.736 Inelastic -Buckling Stress, og - nja, (ksi) 22.62
~ Factor of Safety, FS 1.67 Allowable Compressive Stress,a a ll?81 /FS (ksi) 13,55
,' AppPed-Compressive Meridional Stress, a, (ksi) 9.69 Margin'=[(cll/8 )m - 1] x:100% 39.7% :
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I I 5.
SUMMARY
AND CONCt.USIONS The results of this buckling analysis for the refueling and pos't-accident load combinations are summarized on Table 51. The applied and allowable compressive meridional stresses shown in. Table 51 are for the sandbed region which is the most limiting region in terms of buckliiig. This analysis demonstrates that the Oyster Creek drywell has adequate margin against buckling for an assumed sandbed shell thickness of 0.700 inch. This thickness is less than the 95%
confidence projected thickness of 0.736 inches for the 14R outage.
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5-1
kkbEXh!hfREV.1 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) 7.10 9.69 r
Allowable Compressive Meridional Stress (k'si) 10.44 13.55 ,
Buckling Margin 47% 40%
5-2 ,