ML20065Q389
| ML20065Q389 | |
| Person / Time | |
|---|---|
| Site: | Oyster Creek |
| Issue date: | 11/30/1990 |
| From: | Frederickson C, Mehta H, Ranganath S GENERAL ELECTRIC CO. |
| To: | |
| Shared Package | |
| ML20065M096 | List: |
| References | |
| NUDOCS 9012170175 | |
| Download: ML20065Q389 (58) | |
Text
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)
AN ASME SECTION Vill EVALVATION
, OF THE OYSTER CREEK ORYWELL PART 2 STABillTY ANALYSIS November 1990 -
prepared for GPU Nuclear Corporation Parsippany, New Jersey p
prepared by GE Nuclear Energy
, San Jose, California D 9012170175 901205 PDR ADOCK 05000219 P PDR
O bEX 0 h REV O O
AN ASME SECTION Vill EVALVAT10t1 0F THE OYSTER CREEK DRYWELL 0
PART 2 STABillTY ANALYSIS O
Prepared by: d A [aMb 7 C.D. Frederickson, Senior Engineer o
Materials Monitoring &
Structural Analysis Services Reviewed by:_ L'b '
H. S. Mehta, Principal Engineer Materials Monitoring &
g Structural Analysis Services Approved by: 7 S. Ranganath, Manager O
Materials Monitoring &
Structural Analysis Services O
O O-i
i kikfxhhREV.O i
- O TABLE OF CONTENTS l
t i U i
- O 1. INTRODUCTION 11 l.1 General 11 1.2 Report Outline 1-1
'O 1.3 References 11
- 2. BUCKLING t.NALYSIS METHODOLOGY 2-1
'O 2.] Basic Approach 21 2.2 Determination of Capacity Reduction factor 22 2.3 Modification of Capacity Reduction f actor for 23 .
Hoop Stress O 2.4 Determination of Plasticity Reduction f actor 24 2.5 References 24
- 3. FINITE ELEMENT MODELING AND ANALYSIS 31 0
3.1 Finite Element Buckling Analysis Methodology 31 3.2 Finite Element Model 3-2 3.3 Drywell Materials 33 3.4 Boundary Conditions 3-3 0
3.5 Loads 34 3.6 Stress Results 3-7 3.7 Theoretical Elastic Buckling Stress Results 39 3.8 References 3-10 0
- 4. ALLOWABLE BUCKLING STRESS EVALVATION 41
- 5.
SUMMARY
AND CONCLUSIONS S-1
'O O
iii
D kt X REV. 0 9 LIST OF TABLES Table Page ,
No, Title !L 9
3-1 Oyster Creek Drywell Shell Thickness 3 11 32 Cylinder Stiffener Locations and Section Properties 3-12 3
3-3 Material Properties for fBX-212B Steel 3-12 3-4 Oyster Creek Drywell Load Combinations 3 13 J
3-5 Adjusted Weight Densities of Shell to Account for 3-14 Compressible Material Weight O 3-6 Oyster Creek Drywell Additional Weights Refueling 3-15 37 Oyster Creek Drywell Additional Weights - Post-Accident 3-16
- O 38 Hydrostatic Pressu es for Post Accident, Flooded Case 3-17 3-9 Meridional Seis nic Stresses at four Sections 3-18 0 3-10 Application of Loads to Match Seismic Stresses - 3 19 Refueling Case 3 11 Application of Loads to Match Seismic Stresses - 3-20 0 Post Accident Case 41 Calculation of Allowable Buckling Stresses - Refueling 4-2 D 4-2 Calculation of Allowable Buckling Stresses - Post- Accident 4-3 51 Buckling Analysis Summary 5-2 O
iv
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) LIST OF FIGURES Figure Page No, Title fio .
11 Drywell Configuration .-2 2-1 Capacity Reduction factors for Local Buckling of 25 p Stiffened and Unstiffened Spherical Shells 2-2 Increase in Buckling Stress due to Internal Pressure 26
> 23 Plasticity Reduction factors for inelastit Buckling 2-7 3-1 Oyster Creek Drywell Geometry 3-21 p 32 Oyster Creek Drywell 3 D finite Element Model 3-22 3-3 Closeup of Lower Drywell Section of FEM (Outside View) 3-23
> 3-4 Closeup of Lower Drywell Section of FEM (Inside View) 3-24 3-5 Boundary Conditions of finite Element Model 3-25 5 3-6 Application of Loading to Simulate Seismic Bending 3-26 3-7 Meridional Stresses - Refueling Case 3 27 p 3-8 Lower Drywell Meridional Stresses - Refueling Case 3-28 3-9 Circumferential Stresses Refueling Case 3-29 p 3 10 Lower Drywell Circumferential Stresses - Refueling Case 3-30 3-11 Meridional Stresses - Post Accident Case 3-31 D !
v
l bbfXYhfREV.O LIST Of flGURES O
figure Page No. ,_
Title No.
O 3-12 Lower Drywell Meridional Stresses Post-Accident Case 3-30 3 13 Circumferential Stresses - Post Accident Case 3 33 0
3-14 Lower Drywell Circumferential Stresses Post Accident 3-34 Case 3 15 Symmetric and Anti-Symmetric Buckling Modes 3 3S o
3 16 Symmetric Beckling Mode Shape Refueling Case 3 36 3-17 Anti-Symmetr9c Buckling Mode Shape - Refueling Case 3-37 0
3 ~18 Buckling Mode Shape - Post Accident Case 3 38 O
O O
O vi I
'O ktbEX REV. 0
- 1. INTRODUCTION
- O 1.1 General To address local wall thinning of the Oyster Creek drywell, GPUN has
- o 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
- O for the drywell shell. This report addresses the buckling capability of the drywell shell shown in Figure 1-1 and constitutes the second part of the supplementary report. Buckling of the entire drywell shell is considered in this analysis with the sandbed region being the !
'O area of primary concern.
1.2 Report Outline
- O 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.
"O Section 5 presents the summary of results and conclusions.
1,3 References O
l1 " Structural Design of the Pressure Suppression Containment Vessels," by Chicago Bridge & Iron Co., Contract # 9 0971, 1965.
- O l-2 "An ASME Section Vill Evaluation of the Oyster Creek Drywell," GE Report No. 9-1, DRF# 00664, November 1990, prepared for GPUN.
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- 2. BUCKLING ANALYSIS METHODOLOGY D
2.1 Basic Approach The basic approach used in the buckling evaluation follows the D methodology outlined in the ASME Code Case N-284 [ Reference 2 1). Following the procedure of this Code Case, the allowable compressive stress is evaluated in three steps, a S in the first step, a theoretical clastic 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 9 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 clastic buckling s t 4'e s s is 4 modified by the appropriate capacity and plasticity reduction factors. The capacity reduction factor, oj, accounts for the difference between classical buckling theory and actual tested buckling stresses for fabricated shells. This difference is due to imperfections inherent d 9 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 f abricated shells is given by the product of the theoretical clastic buckling stress and G the capacity reduction factor, i.e., oie 01 When the elastic buckling stress < xceeds the proportional limit of the material, a plasticity reduction factor, nj, is used to account for non linear material behavior. The inelastic buckling stress for fabricated shells is O given by njojoie' In the final step, the allowable compressive stress is obtained by dividing the buckling stress calculated in the second step by the S safety factor, FS: Allowable Compressive Stress njojojp/f 5 9 2-1 l
1 iO i b x k REV. O ! In Reference 21, the safety factor for the Design and Level A & B
'O service conditions is specified as 2.0. A safety f actor of 1.67 is specified for Level C service conditions (such as the post-accident ,
flooded condition).
'O The Determination of appropriate values for capacity and plasticity reduction factors is discussM next.
2.2 Determination of Capacity Reduction Factor
- O lhe capacity reduction factor, og, 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 O are based on extensive data compiled by Miller (2-2]. The factors appropriate for a spherical shell geometry such as that of the drywell in the sandbed region, are shown in figure 21 (Figure 15121 of Reference 21). The tail (flat) end of the curves are used for
-O unstiffened shells. The curve marked ' Uniaxial compression' is applicable since the stress state in the sandbed region is compressive in the meridional direction but tensile in the circumferential direction. From this curve, og is determined to be 0.207.
- O 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
- 0 thickness. However, the drywell shell has a greater thickness above the sandbed region which would reinforce the sandbed region. Second, it is assumed that the circumferential stress is zero, lhe tensile circumferential stress has the effect of rounding the shell and O reducing the effect of imperfections introduced 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.
O The capacity reduction factor values given in Reference 21 are applicable to shells which meet the tolerance requirements of NE-4220 0 2-2
O kt X k REV. O of Section 111 (2 3). Appendix A of Reference 2-4 compares the O tolerance requirements of NE-4220 to the requirements to which the Oyster Creek drywell shell was fabricated. The comparison shows '. hat the Oyster Creek drywell shell was erected to the tolerance requirements of NE-4220. Therefore, although the Oyster Creek drywell O is not a Section 111, 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 O The orthogonal tensile. Ltress has the effect of rounding fabricated shells and reducing tFe effect of imperfections on the buckling strength. The method Gescribed in Reference 2 5 was used to quantify O this effect. The buckling stress in uniaxial compression for a sphere of uniform thickness is given by the following: Se = (0.605)(0.207) Et/R O Where, 0.605 is a constant, 0.207 is the capacity reduction factor and E,t and R are Young's Modulus, wall thickness and radius, respectively. Reference 2 5 states that in the presence of a tensile O stress such as that produced by an internal pressure, p, the modified buckling stress is as follows: Sc , mod = [(0.605)(0.207) + AC) Et/R O Where AC is given in graphical form in Figure 2 2. As can be seen in Figure 2-2, AC is a function of the parameter X-(p/4E)(2R/t)2 When the tensile stress magnitude, S, is known,- the equivalent internal O pressure can be calculated using the expression: p- 2tS/R O The AC term is then incorporated in the capacity reduction factor itself by defining a mocified capacity reduction f actor, oi, mod; oi, mod = 0.207 + AC/0.605 O 1 2-3
!O bX REV. 0 2.4 Determination of Plasticity Reduction factor
- 0-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
- O fabricated shells is given by njojoje. Reference 2-6 gives the mathematical expressions shown below [ Article -1611 (a)] to calculate the p'asticity reduction factor for the meridional direction elastic buckling stress. A is equal to ojojp/oy and o y is the material yield O strergth, figure 2-3 shows the relationship in graphical form.
nj = 1.0 if'A 5 0.55
- (0.45/6) + 0.18 if 0.55 < A s 2.6 O - 1.31/(1+1.15A) if 1.6 < a .; 6.25
= 1/A if 6 > 6.25 2.5 References
- O 2-1 ASME Boiler and Pressure Vessel Code Case N-284, " Metal Containment Shell Buckling Design Methods, Section 111, Division 1, Class HC", Approved August 25, 1980.
O 2-2 Miller, C.D., " Commentary on the Metal Containment Shell Buckling Design Methods of the ASME Boiler and Pressure Vessel Code," December 1979.
- O 2-3 ASME Boiler & Pressure Vessel Code, Section 111, Nuclear Power Plant Components.
'O 2-4 "An ASME Section Vill Evaluation of the Oyster Creek Drywell," GE Report No. 9 1, DRF# 00664, November 1990, prepared for GPUN. 2-5 Johnson, B.G., " Guide to Stability Design Criteria for Metal .O Structures," Third Edition (1976), John Wiley & Sons. 2-6 Letter (1985) from C.D. Miller to P. Raju;
Subject:
Recommended Revisions to ASME Code Case N 284. !O j 2-4
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- 3. FINITE ELEMENT MODEllNG AND ANALYSIS
- O 3.1 Finite Element Buckling Analysis Methodology This evaluation of the Oyster Creek Drywell buckling capability uses 0 the Finite Element Analysis (FEA) program ANSYS [ Reference 31). The ANSYS program uses a two step eigenvalue formulation procedure to perform linear elastic buckling analysis. The first step is a static analysis of the structure with all anticipated loads applied. The O structural stiffness matrix, (K), the stress stiffness matrix, (S),
and the applied stresses, cap, 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 O using the equation: ( (K) + A (S) ) (u) = 0 o where: 1 is the eigenvalue or load factor. (u) is the eigenvector representing the buckled shape of the structure.
.O 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 O 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:
l i o cr
- A Cap O
l This theoretical clastic buckling stress is then modified by the l capacity and plasticity reduction factors to determine the predicted !O 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. O-3-1
Nb[x 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 0.700 inch. This is less than the 95% confidence projected thickness for outage 14R. Figure 3 4 shows the view from the inside of the drywell with the gussets and the vent jet deflector.
The drywell and vent shell is modeled using the 3 dimensional plastic quadrilateral shell (STIF43) element. Although this element has plastic capabilities, this analysis is conducted using only clastic
> behavior. This element type was chosen over the elastic quadrilateral shell (STlF63) element because it is better suited for modeling curved surfaces.
32
kt X W REV. O At a dista?ce 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 centerlin of the vent radially outward to each of the shell nodes of the vent. ANSYS STIF4 beam elements are then connected to this a centerline nede to model the axial and bending stiffness of the vent and header. Spring (STIF14) elements are used to model the vertical neader supports inside the torus. ANSYS STIF4 beam elements are also used to model the stiffeners in the cylindrical region of the upper e 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 . 1 p a smooth transition to reduce thermal and mechanical discontinuities. The sand provides lateral support to the drywell sphere in this region. The foundation stiffnass for the sandbed is considered to be 366 psi /in per Reference 2.4.10 of Reference 3-2. ANSYS STlf14 spring g 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 D area of each node in the sandbed region. 3.3 Drywell Materials 3 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 D. properties were modified to account for the reduction in stiffness due to the perforations. 3.4 Boundary Conditions D 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 9 3-3
kibX REV. O the drywell centerline and vertically, but not in the circumferential > direction. Rotations are also fixed in two directions to prevent the boundary from rotating out of the plane of symmetry. Nodes at the bottom edge of the drywell are fixed in all directions to simulate the fixity of the shell within the concrete foundation. Nodes at the ends p of the sand spring elen,ents and the header support spring elements are also fixed. 3.5 Loads D The loads are applied to the drywell finite element model in the manner which most accurately represents the actual loads anticipated on the drywell. Details on the application of loads are discussed in p the following paragraphs. , 3.5.1 Load Combinations y 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 p include high internal pressures which would create tensile stresses in the shell and help prevent buckling. The most severe design load combination ider,tified for the buckling analysis of the drywell is the refueling condition (Case IV). This load combination consists of the y following loads: Dead weight of vessel, penetrations, compressible material, equipment supports and welding pads. p 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. D 3-4
y' kX REV.-0 i r The.most severe load combination for the emergency condition is for h the post; accident (Case VI)-load combination including: Dead weight of vessel, penetrations, compressible material and equipment supports b Live load of personnel lock Hydrostatic Pressure of Water for Drywell flooded to 74'-6" External Pressure of 2 psig Seismic inertia and deflection loads for flooded condition .; ): The application of these loads is described in more detail in the following sections, a > 3.5.2 Gravity Loads; r The gravity 1 1oads include dead weight loads of the drywell shell, weight of the-compressible material and penetrations and live loadt. . 1 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 h acceleration. The compressible material weight .of 10 lb/f t? is added by.. adjusting the weight - density of -the shell : to also - include the j compressible material. The adjusted weight densities for the~various - shell--thicknesses are - summarized on- Table ' 3-5. The compressible' - h 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 - )f 3 6- for the refueling case,?t eh to at l 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 These applied masses automatically impose gravity loads
.on Table 3-6.
on the drywell model- with the defined acceleration of 19 The same )1 3-5
O DR1 00664 If0EX 9-2, REV. O method is used to apply the additional masses to the model for the O post-accident, flooded case as summarized in Table 3-7. 3.5.3 Pressure Loads O 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 O 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 O calculated as shown in Table 3-8. The hydrostatic pressure at the bottom of the sandbed region is calculated to be 28.3 psi. According to the elevation of the element centerline, the appropriate pressures O are applied t the inside surf ace of the shell elements. 3.5.4 Seismic Loads O Seismic stresses have been calculated for the Oyster Creek Drywell in Part 1 of this report, Reference 3-3. Meridional stresses are imposed on the drywell during a seismic ever t due to a 0.058" deflection of the reactor building and due to horizontal and vertical inertial loads O on the drywell. The meridional stresses due to a seismic event aie impnsed on the 3-0 drywell model by applying downward forces at four elevations of the O 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 O 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.
O I 3-6
) ifX REV. O To find the correct loads to match the seismic stresses, the total ) seismic stress (due to reactor building deflection and horizontal and vertical inertia) are obtained from Reference 3 3 at the four sections of interest. The four sections and the corresponding meridional stresses for the refueling (unflooded) and post-accident (flooded) )- seismic cases are summarized in Table 3-9. Unit loads are then applied to the 3-D model in separate load sf eps at each elevatic, 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 unknowr.s, 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 c;. ors represent the most D-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 . Tne circumferential stresses for the same areas are
-shown on Figures 3-9 and 3-10. The resul ting average meridional D-stress at the mid elevation of the sandbed region was found to be;
~
oRm - -7097 psi ) D 3-7
O ~ NDfX REV. O The circumferential stress averaged from the bottom to the top of the O sandbed region is; O Rc = 277 psi O 3.6.2 Post Accident Condition Stress Results The application of all of the loads described for the post-accident condition results in the stress distributions shown in Figures 3-11 O through 3-14. The red colors represent the most tensile stresses and-the blue colors, the most compressive. Figures 3 11 and 3-12 show the meridional stresses for the entire drywell and lower drywell . The circumferential stresses for the same areas are shown on figures 3-13 0_ and 3-14. The resulting average meridional stress at mid elevation of the sandbed region was found to be; oPAm
-9693 psi O
The circumferential stress averaged from the bottom to the top of the sandbed region is; O pac = +4049 psi O O O O O 3-8
O bX REV. O 3.7 Theoretical Elastic Buckling Stress Results O 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 O elastic buckling loads and buckling mode shapes. 3.7.1 Refueling Condition Buckling Results O As shown on Figure 3-15, it is possible for the drywell to buck 7 in two different modes. In the case of symmetric buckling shown on Figure 3-15, each edge of the 36' drywell model experiences radial displacement with no rotation. This mode is simulated by applying O 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 O 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.
.O 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 O a concern outside of the sandbed region.
It is also possible for the drywell to buckle in the anti-symmetric manner shown in Figure 3-15. For this mode, the edges of the 3-D O model are allowed te rotate but are restrained from expanding radially. This case is considered by applying anti symmetric boundary conditions at the edges of the 3-D model. With the two pass approach used by ANSYS, it is possible to study anti-symmetric buckling of the 10 drywell when the stresses are found based on symmetry boundary conditions. The resulting load factor found using anti-symmetric boundary conditions is 16.81. The mode shape for this case is shown on Figure 3 17. O 3-9
) fbEX 0 h REV. O Because the load factor is lower for symmetry boundary conditions D with the same applied stress, the symmetric buckling condition is more limiting. Multiplying the load factor of 14.32 by the average meridional stress from section 3.6.1, the theoretical elastic buckling stress is found to be; D oRie = 14.32 x (7097 psi) = 101,650 psi 3.7.2 Post-Accident Condition Buckling Results 5 Considering the post-accident case with symmetry boundary conditions, the load factor was calculated as 9.91. Multiplying this load factor by the applied stress from section 3.6.2 results in a theoretical D elastic buckling stress of U PAie - 9.91 x (9693 psi) = 96,060 psi D The critical mode shape for this condition is shown in Figure 3-18. Again, the critical buckling mode is in the sandbed region. 3.8 References D 3-1 DeSalvo, G.J., Ph.D, and P,cman, R.W., "ANSYS Engineering Analys.s System User's Manual, Re,'sion 4.4," Swanson Analysis Systems, Inc., May 1, 1989. D 3-2 GPUN Specification SP-1302-53-044, Technical Specification for Primary Containment Analysis - Oyster Creek Nuclear Generating Station; Rev. 2, October 1990. 9 3-3 "An ASME Section Vill Evaluation of the Oyster Creek Drywell - Part 1 Stress Analysis," GE Report No. 9-1, DRF # 00664, November 1990, prepared for GPUN. D D 3-10
O khbfX h h REV. O O Table 31 Oyster Creek Drywell Shell Thicknesses e Section Thickness (in.) Sandbed Region 0.700 Lower Sphere 1.154 g- Mid Sphere 0.770 Upper Sphere 0.722 Knuckle 2.5625 Cylinder 0.640 g Reinforcement Be' low Flange 1.250 Reinforcement Above Flange 1.500
. Elliptical Head 1.1875 Ventline Reinforcement 2.875 g Gussets 0.875 Vent Jet Deflector 2.500 Ventline Connection 2.500 Upper Ventline 0.4375 g Lower Vantline 0,250 0
O e 3-11
.O_ 0 gtNEX N REV. 0 O Table 3 2 Cylindar Sti.'fener Locations and Section Prop'rties o Elevation Height Width Area Bendina iner'ia (in4) l (in) (in) (in) (in2) Horizontal _ Drtical I I
966.3 0.75 6.0 4.5 la 5 0.211 j o ! 1019.8 0.75 6.0 4.5 13.5 0.?ll j 1064.5 0.50 6.0 3.0 9.0 0.063 .O 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 O (1) - This stiffener is made up of a 2 beam sections, one 2.75x7" and one 1.0x7.375" O Table 3 3 O Material Properties for FBX-2128 Steel I-l- Maurial Property Value t ,0 Young's Modulus 29.6x106 psi l l Yield Strength 38000 psi l [ Poisson's Ratio 0.3 O Density 0.283 lb/in 3 l lO l 3-12 l
'()- kkbkx h h REV. 0 () . Table 3-4 Oyster Creek Drywell Load Combinations O CASE I - INITIAL TEST CONDITION Deadweight + Design Pressure (62 psi) + Seismic (2 x DBE) () CASE II - FINAL TEST CONDITION Deadweight + Design Pressure (35 psi) + Seismic (2 x DBE) CASE 111 - NORMAL OPERATING CONDITION () Deadweight + Pressure (2 psi external) + Seismic (2 x DBE) CASE IV - REFUELING CONDITION Deadweight + Pressure (2 psi external) + Water Load + () Seismic (2 x DBE) CASE V - ACCIDENT CONDITION Deadweight + Pressure (62 psi 0 175'F or 35 psi @ 281*F) +
- C) Seismic (2 x DBE)
CASE VI - POST ACCIDENT CONDITION Deadweight + Water Load 0 74'6" + Seismic (2 x DBE) O O O O 3-13
n hbkxhhfREV.O Table 3-5 O Adjusted Weight Densities of Shell to Account for Compressible Material Weight O Adjusted Shell Weight Density Thickness (in.) (ib/in 3) O 1.154 0.343 0.770 0.373 0.722 0.379 0 2.563 0.310 0.640 0.392 1.250 0.339 O O O O O O 3-14
C-RF* 00664 kt4DEX9-2,REV.O n* Table 3 6 Oyster Creek Drywell Additional Weights . Refueling Condition 5 O DEAD PENETR. M150. TOTAL 5 FOOT LOAD PER LOA 0 PER LOAD PER LOAD RANGE 36 OIG. # OF N00E5 0F FULL N00E HALF N00E ELEVATION VEIGHT VElGHT LCA05 (lbf) (lbf) ELEMENT 5 APPLICATION (1bf) (1bf) (feet) (ibf) (1bf) (ibf) LOAD 15.56 50000 50000 16 165100 168100 0 20 1 200 11200 229300 22930 6 116 119 3822 1911
" 15 20 22d 556000 556000 556000 55600 8 161-169 6950 3475 ** 21 25#
26 11100 11100 30 64100 51500 115600 30.25 105000 100000 205000 4146 2073 O o rs.3o 3337eo 3337o e 379 187 31 16500 16500 32 750 750 33 15450 15450 34 28050 2B050 35 1500 1500 ' 6225 8 188-196 778 389 62250 Q " 31 35 36 1550 1550 40 41000 43350 84350 8590 8 197 205 1074 537 85900
** 36 40 50# 1102000 1102000 110200 8 416-426 13775 6888
" 45 50# 1102000 54 7850 7850 785 8 436 444 Bs 49 7850 0 ** 51 55 56 56400 24000 80400 60 95200 700 20000 115900 19630 8 454 462 2454 1227
" 56-60 196300 65 52000 20000 72000 7200 8 472-4B0 900 450
" 61 65 72000 70 5750 5750 g' " 66 70 5750 575 8 508 516 72 36 73 8850 8850 885 8 526 534 111 $5 8850
" 71-75 82.17 21550 21650 2165 8 553 561 271 135
" 81-85 21650 87 1000 1000 15000 15000 0 " 86 90 90 16000 1600 8 571-579 200 100 93.75 20700 20700 94.75# 698000 698000 95.75 20100 20100 73B80 8 589-597 9235 4618
" 91 96 738800 O T01At5 2184150 388200 e62000 3434350 3434350 343435 f - LOAD TO BE APPLIE0 IN VERTICAL DIRECTION ONLY.
& . MIST [LLANIOUS LOADS INCLUCE 698000 LB WATER VEIGHT AT 94.75 FT. ELEVATION 100000 LB EQUIPMENT 000R WElGHT AT 30.25 FT. ELEVATION AND WILO PA0 LIVE j
LOA 05 Or 24000, 20000 AND 20000 AT 56, 60 AND 65 FT. ELEVATIONS RtrwcT.vti O 3-15
) DRr* 00664 IN)EX 9 2, REV. O D Table 3-7 Oyster Creek Drywell Additional Weights - Post. Accident- Condition TOTAL 5 FOOT LOAD PER LOAD PER LOAD PIR DEAD PENETR. MISO. g LDA05 LOAD RANGE 36 DEG. # OF N00($ OF FULL NODE HALF N00E ELEVATION VEIGHT VIIGHT (1bf) (1bf) ELEMENTS APPLitATION (lbf) (lbf) (feet) (1bf) (Ibf) (1bf) LOAD 15.56 50000 50000 16 168100 168100 20 11200 11200 22930 6 116-119 3822 1911 229300 D " 15-20 22f 556000 556000 55600 8 161 169 6950 3475 556000
" 21 25#
26 11100 11100 30 64100 51500 115600 30.25 105000 105000 23170 8 179-187 2896 1448 231700
" 26 30 E 31 16500 16500 32 750 750 33 15450 15450 34 28050 28050 35 1500 1500 6225 8 188 196 778 389 62250
" 31 35 36 1550 1550 g 84350 40 41000 43350 8590 8 197 205 1974 537
** 36 40 85900
;** 4:02000 1102000 8 418-426 13775 6688 1102000 110200
" 45-50f 54 7850 7850 785 8 436-444 98 49
** 51 55 7850 56 56400 56400 D
60 95200 700 95900 8 454 462 1904 952 152300 15230
" 56-60 65 52000 52000 5200 8 472 480 650 325 52000
** 61 65 70 5750 5750 575 8 508 516 72 36 5750
" 66-70 I 73 8850 8850 55 8850 885 8 526 534 111
" 71 75 82.17 21650 21650 2165 8 553 561 271 135
" 81 85 21650 87 1000 1000 90 15000 15000 1600 8 571-579 200 100
" 86-90 16000 g 20700 93.75 20700 95.75 20100 20100 4080 8 589 597 510 255
" 91 96 40800 TOTALS: 2184150 388200 0 2572350 2572350 257235 D f - LOAD 10 BE APPLIED IN VIRTICAL DIRECTION ONLY.
& . NO MISOCLLANIOUS LOADS FOR THl5 CONDITION, FLOOOVGT.VK1-D 3-16
O INX REV. O O Table 3-8 Hydrostatic Pressures for Post-Accident, Flooded Condition O WATER DENSITY: 62.32 lb/ft3 0.03606 lb/in3 FLOODED ELEV: 74.5 ft . 894 inches O ANGLE ELEMENTS AB0VE ABOVE EQUATOR ELEVATION DEPTH PRESSURE ] . b.. !$$$$. .b.SI f._ .b."! .. ..b!!.... ........... b b ........... 27 -53.32 110.2 783.8 28.3 1-12 40 -51.97 116.2 777.8 28.1 13-24 O 771.6 27.8 25-36 53 -50.62 122.4 66 -49.27 128.8 765.2 27.6 37-48 79 -47,50 137.3 756.7 27.3 49-51, 61-66 ,55-57 92 -46.20 143.9 750.1 27.1 52-54, 138-141 ,58-60 102 -44.35 153.4 740.6 26.7 240-242, 257-259 108 -41.89 166.6 727.4 26.2 142-147'151, 148- 243, 256 O- 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 -34.40 209.7 684.3 24.7 160-165, 246, 253 124 -31.87 225.2 668.8 24,1 166-173, 247, 252 130 -29.33 241.3 652.7 23.5 174-183, 248-251 138 -26.80 257.6 636.4 23.0 184-195 148 -24.27 274.4 619.6 22.3 196-207 O 161 -20.13 302.5 591.5 21.3 208-215 170 -14.38 342.7 551.3 19.9 216-223 179 -8.63 384.0 510.0 18.4 224-231 188 -2.88 425.9 468.1 16.9 232-239 197 2.88 468.1 425.9 15.4 430-437 400 8.63 510.0 384.0 13.8 438-445 e 409 14.38 551.3 342.7 12.4 446-453 418 20.13 591.5 302.5 10.9 454-461 427 25.50 627.8 266.2 9.6 462-469 436 30.50 660.2 233.8 8.4 470-477 445 35.50 690.9 203.1 7.3 478-485 454 40.50 719.8 174.2 6.3 486-493 463 45.50 746.6 147.4 5.3 494-501 8 472 50.50 771.1 122.9 4.4 502-509 481 54.86 790.5 103.5 3.7 510-517 490 - 805.6 88.4 3.2 518-525 499 - 820.7 73.3 2.6 526-533 508 - 835.7 58.3 2.1 534-541 517 - 850.8 43.2 1.6 542-549
, 526 -
885.3 8.7 0.3 550-557 187.3 706.7 25.5 340-399 (Ventline) FLOODP.WK1 9 3-17
O ktbkx b REV O O- Table 3 9 Meridional Seismic Stresses at Four Sections O 2-0 Shell Meridional Stresses , Elevation Model Refueling Post Accident Section (inches) Node (osi) (osi)
'O A) Middle of Sandbed 119 32 1258 1288 B) 17.25' 8elow Equator 323 302 295 585 O v C) 5.75' Above Equator 489 461 214 616 D) Above Knuckle 1037 1037 216 808 O'
O O-O O' O 3-18
O DRF4 00664 INDEX 9-2, REV. 0 0 Table 3-10 Application of Loads to Match Seismic Stre5se5 - Refueling Case O 2 D SEl$MIC STRESSES AT $ECTION (psi) 2 3 4 SECTION: 1 2 D N00E: 32 302 461 1037 Q COMPRES$1VE STRESSES FROM 2-0 ANALYSIS ELEV: 119.3" 322.5" 489.1" 912.3" 788.67 155.54 103.46 85.31 0.058" SEl5MIC DEFLECTION: HORIZ. PLUS VERTICAL $EISHIC INERTIA: 469.55 139.44 110.13 130.21 TOTAL SE15Mit COMPRES$1VE STRESSES: 1258.22 294.98 213.59 215.52 0 3 D STRESSES AT SECTION (psi) 3.c 4 SECTION: 1 2 3
]NPui t AD 3 0 NO Es: 53 65 17 178 400 408 526-534 0 ELEV: 119.3" 322.5" 489.1" 912.3"
$ECTION INPUT 3-D UNIT LOAD DESCRIPTION 1000 lbs at nodes 563 through 569 85.43 37.94 34.94 55.23 A
500 lbs at 427&435, 1000 lbs at 428-434 89.88 39.92 36.76 0.00 B 500 lbs at 197L205, 1000 lbs at 198-204 97.64 43.37 0.00 0.00 C 500 lbs at 161L169, 1000 lbs at 162-168 89.85 0.00 0.00 0.00 0 g DESIREDCOMPRESSIVESTRESSES(psi): 1258.22 294.98 213.59 215.52 3-D INPUT LOAD RESULTING STRESSES AT SECTION (psi) O stCTION LOAD 10 BE APPLIED TO MATCH 2-0 STRESSES A 3902.2 333.37 148.05 136.34 215.52 2101.4 188.87 83.89 77.25 0.00 8 1453.6 141.93 63.04 0.00 0.00 C 6611.6 594.05 0.00 0.00 0.00 0
--~~~~' -'--' ~~~~~~~ ~~~-~~~
0 SUM: 1258.22 294.98 213.59 215.52 SEISUNFL.VK1 O O 3-19
. _ . . _ _ _ _ - - _ _ _ _ - _ _ . _ . _ ~_ _ _ _ _ _ __
t
- OL,.
kt kX REV. 0
.O Table 311 Application of Loads to Match Seismic Stresses - Post. Accident Ca$e
- Os -
l 2DSEl$MitSTRES$ESATSECTION(psi) 1 2 3 4 SECTION:
- O. 32 302 461 1037 2 0 N00E
119.3" 322.5" 489.1" 912.3" COMPRE$$1VE STRESSES FROM 2 0 ANALY11$ ELEV: 788.67 155.54 103.46 85.31 D.058* SEISHIC DEFLECT 10NL MORIZ. PLU$ VERTICAL SEl$MIC INERTIA: 499.79 ~ 429.39 512.76 723.14 TOTAt $tl5 HIC COMPRES$1VE STRESSES: 1288.46 584.93 616.22 808.45 r LO' : 3DSTRES$ESATSECT10N(ps'i)' 30 3 4 j
$ECTION: 1 2 INPUT-
'O LOAD; 3-D NODES: 53 65 170 178 400 408- 526 534- . ;. ELEV: 119.3* 322.5" 489.1" 912.3"-
.SECTION INPUT 3 D UNIT LOAD DESCRIPTION-1000 lbs at nodes $63 through 569= 85.43 37.94 34.94 -55.23 A
~500 lbs at 427&435, 1000 lbs at 428-434 89.88 33.92 36.76 0.00 8
97.64 43.37 0.00- '0.00 C 500 lbs-at 197&205. 1000 lbs at 198 204 500 lbs at-1616169, 1000 lbs at 162 168 89.85 0.00= 0.00- -0.00 LO- D DESIRED COMPRE$$1VE STRES$ES (pst): 1288.46: 584,93 616.22 808.45
-j 1 i' lNPUT.
c O- toad .- RESULTINGSTRESSESATSECTION(psi)' SECil0N . LOAD TO BE APPLIED TO MATCH 2 D STRES$ts A -14637.9 -1250.C: 555.36 511.45 -808.45 2850.2- 256.17: 113.s3 104.77 0.00
- 8
-1941.7 189.58 84.21~ 0.00 0.00-C-
318.8- 28.s4 -0.00 0.00 0.00-D .O: ....... ....... ....... ....... SUM: 1288.46 '584.93 616.22 808.45 s
$E11FL.VK1 LO
.O-3 20
O I O DRYWELL : (G($% #s;;; R. TH-ELEV. 87' 5" y.640$:!O s < - ihy gusyM5WN O R TH' j .640";fi}g f(,;; @g; g #0MG %$$R KW gWNhk$ fMk@kk
/ hh $$w ~
llk,Ydith$? U~%MfQ l? Mhijs }% R.THK 676" Figure 3 1. Oyster Creek Drywell Geometry 9 9 3-21
o o o o o o o o o o o l gSys 4,4 l 1 y i d99=1 PR 7%DhTs RF.AL Ngg M =1 W _ pjSTE7fff7.s1 2r :ssa . D D HIppg Y O n , DD DRyggg A LYSZS - CRLQ gg*Nb. REyggg Figure 3.g~ Oyster Creek 0FW ell 3-0 Finite Element Model
! O o o o o o o o o o o 4 AMSYS 4.4 1 MOU 13 1999 l' 14:55:17 PLOT Mo. 2 ! ! PREP 7 ELDtENTS REAL NUM XU =1 ! I YU =-9.9 DIST=288.375 ' XT =429.452 , 2F =216.528 s AMC2=-99 i' CENTROID HIDDEN t t Y 0 i i r c r s. . - s , s , 3. t i i
\
s ! N I I I s OYSTER CREEE DRYWELL ANALYSIS - OYCR1Q CSAND, REFUELING) , t [ Figure 3-3. Closeup of Lower Drywell Section of FEM (Outside View)
O O O e O O O O O O O nHsYs 4.4 1 NOU 13 1999 14:56:23 PLOT MO. 3 PREP 7 ELEMENTS REAL MUM XU =-1 YU =-9.9 DIST=288.376 XT =429.452 2F =216.529 ANC2=9H CENTROID HIDDEM Y 5 ccc.ss sis:
'\
\
N
\
l l OYSTER CREEX DRYWELL ANALYSIS - OYCR1Q (SAND. REFUELING) Figure 3-4. Closeup of Lower Drywell Section of FEM (Inside View)
=
O O O O O O O O O O O ANSYS 4.4 1 OCT 15 1999 99:31:26 - sc hk b "kN MEhTS TYPE MUM 9 BC SYMBOLS
% - B.8 P ~. DIST=718.786 d XF =393.931 g 2F =639.498 AMC2=-99 g CENTROID HIDDEM M
G y
<e w
M
*a w '
l t 3 % Ns $8 OYSTER CREEN DRYMELL ANALYSIS - OTCRI A 'CS AND, UNIT LOAD CA SE) { Figure 3-5. Boundary Conditions of Finite Element Model
O O O O O O O O O O O ANSYS 4.4 1 OCT 15 1999 99:32:26 PLOT MO. 2 PREP 7 ELDIENTS TYPE MUM BC SYMBOLS XU =1 YU =-9.8 DIST=718.786 XT =393.931 2F =639.498 AMC1,=-99 CENTROID HIDDEM y u t 1 . OYSTER CREER DRYWELL ANALYSIS - OYCR1A CSAND. UNIT LOAD CA SE) Figure 3-6. Application of Loading to Simulate Seismic Bending
O O -O O O O O O- O O O ANSYS 4.4 1 NOU 14 1999 98:12:59 PLOT MO. 1 POST 1 STRESS
's STEP =1 ITER =1 SY CAUC)
NIDDLE ELEM CS DMX =9.299378 SMM =-9993 SMX =442.997 l rm
....sesse XU =1 r- YU =-9.8
""' DIST=719.796 M XT =393.931
;.'.;ri: & g633 4's M
CENTROID HIDDEM
-9993
-7964 26
.e c c e e ren -
49 e r r r c rees -3311 e r y yycre. -1434 y e e e e v e re. 442.997 ' m r e d eve >w ' I 4 > > F)>> r O 'r 'r'r'rY.
). & H kwv.
- i. e e ;- e '; ', * '
rI fr- E r > F F
.....r...
A 5 s ! OYSTER CREER DRYWELL ANALYSIS - OYCR1Q (SAND, REFUELIHG) Figure 3-7. Meridional Stresses - Refueling Case
_ _ _ . . . . . , , . . . . .g Vs 1 A h959 p * ,s % ,g i a s *UC) e - T ~ ~ ,. . . - s' hk yc
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g f ? OYSTER CREEX DRYWELL ANALYSIS - OYCR14 CSAND, "EFUELINC) . 59Ure 3.g, lower OD'Wey y Merinionay Stressg, - Reryg,'"9 Casg
O O O # O O O O O O O ANSYS 4.4 1 NOU 14 1999 98:13:14 PLOT MO. 2 POST 1 STRESS
', STEP =1 ITER =1 SY (AUC)
MIDDLE ELEM CS DMX =9.299378 SMN =-3363 SJX =2873 XU =1 YU =-9.8 DIST=718.796 XF =393.931 g g 6 g.498 CENTROID HIDDEF E -3363 E -267G
-1977 errecers -591.439 181.478 l 14e7 y __
2873
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u A s.a. s u.t s s OYSTER CREEN DRYWELL AHALYSIS - OYCR1Q CSAMD, REFilELIHG) Figure 3-9. Circumferential Stresses - Refueling Case t - _ _ _ _ _ - _ _ - . _ - _ - - - _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ - _ _ _ _ - _ - - _ _ _ _ _ _ _ _ _ _
//
O O O O O O O O O O O t ANSYS 4.4 I NOU 14 1999 98:13:19 PLOT Ho. 4 POSTL STRESS STEF=1 ITER =1 SY <AUC)
. . ..: . . . . , , HID yty ,DLE, c
', DMX =9.19678
) .l , } l l _} } l SMN =-3363 g u..s. .. .
SMX =2873 s: e. e , .
.s , .>
n e >. , >
'e' s es a MU T1 J 7 -3~ ,
vu =-9.s l b 7;.i_s 2m sf i. - . i. . .i. A. . g. . . .' :.T!.j*p DIST=2se.376 xr =42v.452
;. a--= =-
a fd
--T-jj hkC25 hb'
_s CINTROID HIDDEN E ~3363 2679 Y CC.i?
- 4! Q 1487 hh1 439 191.478 2873 8 .
A 2 t i OYSTER CREEN DRYWELL ANALYSIS - OYCR1Q CSAND, REFUELING) Figure 3-10. Lower Drywell Circumferential Stresses - Refueling Case
o e e e o e e o o o o ANSYS 4.4-1 NOU 15 1999 9?:36:33 f PLOT No. 2 POST 1 STRESS STEP =1
- gN ITER =1 l SV <nUG)
MIDDLE ELEM CS DMX =9.48297 SMN =-12329 SMX =2719 XU =1
- stb 7 5 796 etettwe XF =393.931 e
g g g.498 s is s isin CENTROID HIDDEN L
-12329
-19659 k- -
- 8258 V:::::%
rnen:&
-3 m
-639.211 y 2719 t'
i fiL Of hE % ' :
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e=mg J . O.-kHGG%,1 r :-
- 1+ >H y s+. 22l OYSTER CREEX DRYWELL ANALYSIS - --
^^
-'HD. POST-ACCID.)
Figure 3-11. Meridional Stresses - Post-Accident Case
5 b A H TM to W N tp A H D4 99 TS bnN wSS N
.m a HM OTn ENn Tm O WM3 e*** OW SS H @ .H p WNH *tD S @ S A N S N e n H nM a M *Hb SSNHewdH210M@b O en .. HHv USIN Hinens iiiiiin N.ib 41111 11 Il il ll11Il1111I N
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6 v C - - -- v v v ANSYS 4.4 I NOU 15 1999 G?:36:17 PLOT P90. 1 POST 1 STRESS
',g STEF=1 ITER =1 SY (AUC)
MIDDLE j ELEM CS DMX =9. 48G!S7 8 SMM =-4594 StX =12763
$ XU =1 as YU =-S.8 mi DIST=719.786 mI XF =393.931
*I Q =6 g.438 CINTROID HIDDEM M Nk 6 EH8 sm49 e?e6 137G3 F + + ,"r'r~r Y, f 4 4 h b b l- M I
1 OYSTER CREEX DRYWELL ANALYSIS - ^ ^^ ^" ND , PV .CID.) Figure 3-13. Circumferential Stresses - Post-Accident Case
j O O O O O o o o o o o , 1 i ! i i I
,' t AMSYS 4.4 l ! 1 NOU 13 1999 l 9?:37:38 i PLOT MO. 3 PosTL STRESS ! ,' STEP =1 ITER =1 l
l 4 e--,--,-.. . --. SY (AUC) l l MIDDLE ELEM CS 3 DMX =9.311487 i f SMM =-4594 SMX =12763 i XU =1 , ! YU =-C.8 I ) D I ST=290 . 3*/ 6 ! XF =429.452 ! 2 "g =2g.328 CENTROID HIDDEM
-4594 E
" -2666 l
5138 3e49 1 8996 i 12763 Y i i f I i c e r nu l ( t _ OYSTER CREEX DRYWELL ANALYSIS - OYCR10 CSAND, POST-ACCID.) l l
' Figure 3-14 Lower Drywell Circumferential Stresses - Post-Accident Case !
O O Center of Drywell ,\ Sphere /\ Symmetry Planes of
/y;M z'
,v- \;
O ,/- 36 -\,
/ \
/ \
/ \ ,
, j \ .s
'1 j Unbuckled Shape O
.. l. /
....g
/
N
\ Buckled Shape
'N Vent f Radial Displacement i O I No Rotation /
Symmetric Buckling of Drywell O Center of s Drywell f\ Sphere Planes ' r3 /} g ey
,/' 36 \
/ \
/ \
/
\
O -' '
' \ ~ ,
4/ - Unbuckled Shape j % ...-.... . . . . . 3
/ \ Buckled Shape O 'N Vent f Rotation )
(No Radial Disp. / Anti-symmetric Buckling of Drywell O SYM.DRW Figure 3 15. Symetric and Anti-Symetric Buckling Modes O 3-35
O O O O O O O O O O O
! AMsYs 4.4
- :1 wou 14 isse 98:23:22 PLOT MO. 1 POST 1 STRESS STEF=1 i TTER=1 1 FACT =14.322 CLOMAL
, DMX =9.963618 i SMH =-9.962299 SMX =9.942W65 4
XU =1 i YU =-o.S 1 DIST=192.351 XF =327.422 i L gzIg 235 1 M -e.se:2es
-9,991731
! M -9.991254 l 5"dTU53' e.634E-m3 s 9.991131 ! Y 9.963993 o. l l I l
\' \ '\ \ s OYSTER CREEK DRYWELL ANALYSIS - OYCR1R CSAMD. REFUELIMC)
Figure 3-16. Syenetric Buckling Mode Shape - Refueling Case
g 1 1,%e i RI=r:
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fi eps y t \
*vsta Ebe yMYHuz **LVsz, , *Vcmq (34rr3 D Q'Hc>
Figur, ,_ "tI-S n c gUcky ? Mode Sg, ' ' Refugg '*9 Ca3, J
1 _ _ .o ____ o o o o o o o o o o I i i 1 1 ! ANSYS 4.4 ] 1 NOU 16 1999 . Os:15:23 i PLOT NO. 1 1 POST 1 STRESS STEP =1 ITER =1 ; FACT =9.911 ! 4 UX i D CLOBAL ! DMX =9.983995 l SMN =-9.992392 SMX =M.9H2399 l XU =1 i YU =-9.s i DIST=196.859
- XF =327.422
. SKc:23243 12s
-9.992392 M -9.991760
- -9.991233
' L?d31"if e.9m3E-e3 9.99144 Y 9.992599 l 5 i, N OYSTER CREEX DRYWELL ANALYSIS - OYCR1U CSAND, POST-ACCID.) Figure 3-18. Buckling Mode Shape - Post Accident Case
) ktk X REY. 0
- 4. ALLOWABLE BUCKLif4G STRESS EVALVAT10t4 h
Applying the methodology described in Section 2 for the modification of the theoretical elastic buckling stress, the allowable compressive stresses are now calculated. Tables 4-1 and 42 summarize the p calculation of the allowable buckling stresses for the Refueling and Post Accident conditions, respectively. The modified capacity reduction f actors are first calculated as described in sections 2.2 and 2.3. After reducing the theoretical instability stress by this e 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. D The allowable compressive stress for the Refueling case is 10.44 ksi. Since the applied compressive stress is 7.10 ksi, there is a 47% margin. The allowable compressive stress for the Post Accident, % e flooded case is 14.34 ksi. This results in a margin of 48% for the applied compressive stress of 9.69 ksi. D D D 3
. I 4-1
l b Nbxh RIV. 0 D lable 4 1 Calculation of Allowable Buckling Stresses - Refueling Case D Parameter Valur Theoretical Elastic Instability Stress, ogg (ksi) 101.65 D Capacity Reduction Factor, og 0.207 Circumferential Stress, oc (ksi) 0.28 Equivalent Pressure, p (psi) 0.000
*X* Parameter 0.000 D AC 0.000 Modified Capacity Reduction factor, oi, mod 0.207 Elastic Buckling Stress, og - oi, mod Die (ksi) 21.04 Proportional Limit Ratio, A - og/oy 0.554 D Plasticity Reduction factor, nj 0.993 Inelastic Buckling Stress, oj ngoe (ksi) 20.89 Factor of Safety, FS 2.0 Allowable Compressive Stress,a c ll " Di /IS (ksi) 10.44 D Applied Compressive Meridional Stress, og (ksi) 7.10 Margin - [(og/cm) - 1) x 100% 47%
D D D D 4-2
g.. ) ktI]fx k REV. O Table 4-2 Calculation of Allowable Buckling Strasses Post Accident Case
)
Parameter Value Theoretical Elastic Instability Stress, oje (ksi) 96.06
) Capacity Reduction factor, og 0.207 Circumferential Stress, oc (ksi) 4.05 Equivalent Pressure, p (psi) 13.50 "X" Parameter 0.164 AC 0.114 ?
Modified Capacity Reduction Factor, oi, mod 0.396 Elastic Buckling Stress, og ai, mod O ie (ksi) 38.04 Proportional Limit Ratio, A = og/oy 1.001 p Plasticity Reduction factor, nj 0.630 Inelastic Buckling Stress, og = njoe (ksi) 23.95 factor of Safety, FS 1.67 Allowable Compressive Stress,a c ll
- Di /FS (ksi) 14.34
$ Applied Compressive Meridional Stress, om (ksi) 9.69 Margin - [(og/c m) - 1) x 100% 48%
D D D D 4-3
kt fx h h REV. 0
- 5.
SUMMARY
AND CONCLUSIONS 3 The results of this buckling analysis for the refueling and post-accident load combinations are summarized on Table 51. The applied and allowable compressive meridional streT.s shown in Table 5-1 are
- for the sandbed region which is the most 1'miting region in terms of buckling. This analysis demonstrates that ,he 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%
D confidence projected thickness of 0.736 inches for the 14R outage. D D D D D D D 5-1
O kt kX hh REV. O O Table 51 Buckling Analysis Summary O toad Combination Refuelina Post-Accident O Service Condition Design Level C factor of Safety Applied 2.00 1.67 O Applied Compressive Meridional Stress (ksi) 7.10 9.69 Allowable Compressive Meridional Stress (ksi) 10.44 14.34 O Buckling Margin 47% 48% O O O O O 52
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