ML061020614

Letter Forwarding, an ASME Section Viii Evaluation of Oyster Greek Drywell for Without Sand Case Part 1 Stress Analysis, Attachment 2 to Letter Dated 04/07/06

ML061020614
Person / Time
Site:	Oyster Creek
Issue date:	03/04/1991
From:	Devine J GPU Nuclear Corp
To:	Document Control Desk, Office of Nuclear Reactor Regulation
References
2130-03-20289, 5000-91-2026, C320-91-2035, TAC MC7624 DRF # 00664, Rev 0
Download: ML061020614 (162)

Text

{{#Wiki_filter:ATTACHMENT 2 (GPU Letter to NRC dated March 4,1991)

_PU Nuclear Corporation Qne UDoer Por Roac _uclea Parsmooany. New Jeesey 37C:5 201-318 7000 TELEX 136 482 Writer's Qirect Dial N-.-e, March 4, 1991 5000-91-2026 C320-91-2035 U. S. Nuclear Regulatory Commission Att: Document Control Dosk Washington, DC 20555 Gentlemen:

Subject:

Oyster Creek Nuclear Generating Station (OCNGS) Docket No. 50-219 License No. DPR-16 Oyster Creek Drywall Containment

References:

(1) GMUN letter dated December S, 1990 - Drywell Structural Reports and Water Intrusion Summary In the Reference 1 letter, GPUW coritted to provide to you the structural design reports supporting drywell sand removal at Oyster Creek. As you know, our investigations indicate utrongly that the presence of sand is a major contributor to the high corrosion rates observed in the sandbed region (elevation 8'-11* to 12'-3') of the drywall; for that reason we consider sand removal to be an important elaemnt in our program to eliminate the corrosion threat to drywall integrity. Attachment I to this letter provides this information in the form of G0 Reports Index No. 9-3 and 9-4, An MSHM Section VIII gvaluation of the Oyster Creek Drywall for Without Sand Came Stress and Stability Analysis." This two (2) part report covers the structural analysis of the Oyster crook drywell with the wand cushion removed and conservatively assume a uniform drywall nandbed region corroded thickness of 0.736 at the end of Cycle 13 operation. The report demonstrates that if the sand cushion were removed the drywell would remain in full compliance with ASHE Code requirements. Further measurements of the drywoll shell thickness are being made during the current 13R outage and will serve as a basis for refined corrosion rate projections. Assuming that thems updated corrosion rate projections are consistent with the Attachment I ana"lysis, PUN plans to proceed with the removal of the drywall sand cushion during Cycle 13 operation as a prudent and positive step in arresting corrosion at Oyster Creek. tZ :C320=35 GPU Nuclear Carporaion is a subsidiary of Generfa Pubiic Utilities Coporstion

C320-1l-20 3 S DryweLl Containment page 2 if you have any questions or comments on this submittal or the overall drywall corroision program, please contact Mr. Michael Laggarte Manager, Corporate Nuclear Licensing at (201) 316-7968. Sincorely, J. C. DeVine, Jr. Vice Preslent, Technical Functions Attachment JCD/R'b/plp ccz Administrator, Region 1 Sonior NRC Resident Inspector Oyuter Crook NRC Project Manager Z:c320;203

DRF # 00664 INDEX No. 9-3 REV. 0 AN ASME SECTION VIII EVALUATION OF OYSTER CREEK DRYWELL FOR WITHOUT SAND CASE PART I STRESS ANALYSIS February 1991 prepared for GPU Nuclear Parsippany, Corporation New Jersey prepared by GE Nuclear Energy San Jose, California

?N'EX N8. 3-3, REV. 0 AN ASME SECTION ViII EVALUATION OF OYSTER CREEK DRYWELL FOR WITHOUT SAND CASE PART I STRESS ANALYSIS Prepared by: C - C.D. Frederickson, Senior Engineer Materials Monitoring & Structural Analysis Services G. W. Contreras, Engineer Materials Monitoring & Structural Analysis Services Reviewed by: H. S. Mehta, Principal Engineer Materials Monitoring A Structural Analysis Services Approved by: S. Ranganath, Manager Materials Monitoring & Structural Analysis Services

TMEX 88. 9-3, REV. 0 TABLE OF CONTENTS

1. INTRODUCTION 1-1

1.1 Background

1-1 1.2 Supplementary Code Stress Analyses 1-2 1.3 Scope of Present Analysis 1-3 1.4 Report Outline 1-3 1.5 References 1-3

2. ANALYSIS BASES 2-1 2.1 Drywell Geometry and Materials 2-1 2.2 ASME Code Allowable Values 2-3 2.2.1 Thickness Reductions From Local Corrosion 2-4 Effects 2.2.2 Allowable Stresses for Post-Accident 2-5 Condition 2.3 Load Magnitudes and Combinations 2-S 2.4 Temperature Gradients 2-6 2.5 References -

2-7

3. DRYWELL FINITE ELEMENT ANALYSIS 3-1 3.1 Description of Finite Element Models 3-1 3.1.1 Axisymmetric Model 3-1 3.1.2 Pie Slice Finite Element Model 3-1 3.2 Load Application on Pie Slice Model 3-2 3.2.1 Gravity Loads 3-3 3.2.2 Pressure Load 3-3 3.2.3 Seismic Loads 3-4 3.3 Stress Results for Various Load Cases and 3-4 Combinations I

?EX 8.6-3, REV. 0 TABLE OF CONTENTS (CONT'D) Page No. 3.4 Temperature Stress Analysis 3-S 3.5 References 3-6

4. SEISMIC LOAD DEFINITION 4-1 4.1 Finite Element Model 4-1 4.2 Dynamic Analysis Methodology and Response Spectra 4-1 4.3 Post-Accident Seismic Analysis 4-2 4.4 Analysis for Relative Support Displacement Effects 4-2 4.5 References 4-3

5. CODE STRESS EVALUATION 5-1 5.1 Code Stress Evaluation of Regions Above the Lower 5-1 Sphere 5.2 Elastic Stress Analysis of Sandbed and Lower 5-2 Sphere 5.2.1 Small Displacement Solution Results 5-2 5.2.2 Large Displacement Solution Results 5-4 5.3 Code Evaluation of the Sandbed and Lower Sphere 5-4 5.3.1 Primary Stress Evaluation 5-4 5.3.2 Extent of Local Primary Membrane Stress 5-5 S.3.3 Primary Plus Secondary Stress Evaluation 5-6

6.

SUMMARY

AND CONCLUSIONS 6-1 APPENDIX A DETAILED RESULTS FOR AXISYMMETRIC MODEL TEMPERATURE STRESS ANALYSIS ii

?N E 8. 9-3, REV. 0 LIST OF TABLES Table Page tno. Title No. 2-1 As-designed and Projected 95% Confidence 2-8 thicknesses used in the Code Stress Evaluation 2-2 Allowable Stresses for Drywell Shell in 2-9 Section VIII Analysis Z-3 Allowable Stresses for Post-Accident Condition 2-10 2-4 Load Combinations specified in the Parsons 2-11 Report (Reference 2-3) 2-5a Dead Weight Loads 2-12 2-5b Penetration Loads 2-13 2-5c Live Loads 2-15 3-1 Load Cases Considered in the Finite Element 3-7 Analysis 3-2 Adjusted Weight Densities of Shell to Account 3-8 for Compressible Material Weight 3-3 Oyster Creek Drywell Additional Weights - 3-9 Refueling Condition 3-4 Oyster Creek Drywell Additional Weights - 3-10 Accident and Post-Accident Condition 3-5 Hydrostatic Pressures for Post-Accident 3-11 Condition iii

HEXd R 3-3, REV. 0 LIST OF TABLES (CONT'D) Table Page No. Titlg No. 3-6 Meridional Seismic Stresses at Four Sections 3-12 3-7 Application of Loads to Match Seismic 3-13 Stresses - Accident Condition 3-8 Application of Loads to Match Seismic 3-14 Stresses - Post-Accident Condition 3-9 Description of Load Combinations in Terms of 3-15 Unit Load Case Sum 5-1a Comparison of Calculated Stresses to Code 5-7 Allowable Values (Nominal Drywall Wall Thicknesses Above Lower Sphere) 5-lb Comparison of Calculated Stresses to Code 5-8 Allowable Values (95% Projected Drywell Wall Thicknesses Above Lower Sphere) 5-2a Comparison of Calculated Primary Stresses to 5-9 Code Allowable Values (Small Displacement; Lower Sphere and Sandbed) S-Zb Comparison of Calculated Primary Stresses to 5-10 Code Allowable Values (Large Displacement; Lower Sphere and Sandbed) 5-3a Comparison of Calculated Primary Plus Secondary 5-11 Stresses to Code Allowable Values (Small Displacement; Lower Sphere and Sandbed) iv

MSEX N8. 9-3, REV. 0 LIST OF TABLES (CONT'D) Table Page No. Title No. 5-3b Comparison of Calculated Primary Plus Secondary 5-12 Stresses to Code Allowable Values (Large - Displacement; Lower Sphere and Sandbed) V

RF #0664 ?NDEX N8. 9-3, REV. 0 LIST OF FIGURES Figure Page No. FIGURE9 No. 1-1 Drywell Configuration 1-5 3-1 Complete Axisymmetric Finite Element 3-16 Model of Drywell 3-2 Sand Bed Region of Drywell Finite Element 3-17 Model 3-3 Knuckle Region of Drywell Finite Element 3-18 Model 3-4 Cylindrical Region of Orywell Finite 3-19 Element Model 3-5 Upper Cylindrical Region of Drywell 3-20 Finite Element Model 3-6 Oyster Creek Drywell Pie Slice Finite 3-21 Element Model 3-7 Inside Closeup View of Lower Orywell 3-22 Section 3-8 Application of Loading to Simulate 3-23 Seismic Stresses 3-9 Below Curb Drywall Model Nodalization 3-24 for Temperature Analysis During Accident Condition vi

TNOEX N. 3-3, REV. 0 LIST OF FIGURES (CONT'D) Figure Page No. FIGURE No. 3-10 Example of Calculated Temperature 3-25 Distribution at Various Elapsed Times 3-11 Meridional Stress Distribution in the 3-26 Sand Bed Region from Temperature Distribution at t-210 Seconds 3-12 Circumferential Stress Distribution 3-27 in the Sand Bed Region from Temperature Distribution at t-210 Seconds 5-1 Circumferential Stresses for Accident 5-13 Condition V-1 in 'With Sand' and 'Without Sand' Cases - Small Displacement 5-2 Plot of Accident Condition V-i Meridional 5-14 Stresses for 'Without Sand' Case - Small Displacement 5-3 Circumferential Membrane Stress 5-15 Distribution Using Small Displacement Option 5-4 Circumferential Membrane Stress Magnitudes 5-16 at Four Meridional Planes in Sandbed Region - Small Displacement 5-5 Beam With Transverse Plus Axial Loading 5-17 5-6 Circumferential Membrane Stress 5-18 Distribution Using Large Displacement Option vii

T X 88. 9-3, REV. o LIST OF FIGURES (CONT'D) Figure Page No. FIGURE No. 5-7 Comparison of Circumferential Membrane 5.19 Stress Magnitudes With Large and Small Displacement Options 5-s Circumferential Membrane Stress Magnitudes 5-20 at Four Meridional Planes in Sandbed Region - Large Displacement viii

NE'X '80.9-3, REV. 0

1. INTRODUCTION

1.1 Background

The Oyster Creek Nuclear Generating Station utilizes a GE BWR Nuclear Steam Supply System and a steel Mark i pressure suppression type containment vessel system. The pressure suppression system consists of a drywell, a pressure suppression chamber (torus) which stores a large volume of water and a connecting vent system between the drywell and the water pool. The drywell, sometimes referred to as the containment vessel or containment structure, houses the reactor vessel, reactor coolant recirculation loops, and other components associated with the reactor system. Figure 1-1 shows the drywell along with the pertinent dimensions. The drywell is a combination of a sphere, cylinder and 2:1 ellipsoidal dome and it resembles an inverted light bulb. The spherical portion of drywell near the base includes a sandbed region that provides an elastic transition zone which is intended to ameliorate abrupt thermal and mechanical discontinuities. The pressure suppression system was designed, analyzed and constructed by Chicago Bridge & Iron Company (CBI). A recent inspection of the steel shell (November 1986) prior to restart from the 11R outage in the sandbed region revealed that some degradation of the shell had taken place during the years since completion of construction. Subsequent inspections also indicated minor thickness degradations in the upper spherical and cylindrical sections of the drywell. A detailed description of the previous analyses pertaining to Oyster Creek drywell is given in Reference 1-1. An ASME Code stress analysis addressing the drywell thickness degradation is documented in Reference 1-2. The analyses in Reference 1-2 are based on the present configuration in the sandbed region, i.e.. it is assumed that the sand is present. One of the option GPUN Is exploring to mitigate further 1-1

TAOEX P8.63-3, REV. 0 corrosion in the sandbed region, is to remove the sand. The purpose of the stress analyses presented in this report is to evaluate the drywell per ASME Section VIII for this modification. 1.2 Supplementary Code Stress Analyses The Code of record for the stress analysis of Oyster Creek drywell is Section VIII, 1962 Edition and Nuclear case Interpretations 1270 N-5, 1274 N-5 and 1272 N-5. The CBI stress report (Reference 1-3) augmented by the recent GE report (Reference 1-2) constitutes the Section VIII Code stress report of record for the drywell. The GE report is a supplementary stress report to 'the CBI stress report and addresses aspects of Code compliance as they relate to the local wall thinning observed in the Oyster Creek drywell. The stress analyses in this report as in the previous GE report [1-2] are guided by GPUN Technical Specification for primary containment analysis [1-4]. Based on the ultrasonic (UT) inspection results, the projected 95% confidence thickness value for the drywell shell in the sandbed region is 0.736 inch. However, in several previous Oyster Creek drywell analyses, as discussed in Reference 1-1, a conservative thickness value of 0.700 inch was used. A shell thickness of 0.700 inch in the sandbed region was used in the stress analyses documented in Reference 1-2. In the first part of the stress analysis report of Reference 1-2, the nominal or as-designed thicknesses were assumed everywhere except in the sand bed region. The thickness in the sand bed region was assumed as 0.700 inch compared to the as-designed thickness of 1.154 inch. Later, the local thinning In areas other than the sand bed region of drywell was addressed. The second part of Reference 1-2 report addressed the buckling evaluation of drywell shell. 1-2

bE& 2'.9-3, REV. 0 1.3 Scope of Present Analysis The stress analyses described in this report address the case when the sand has been removed from the sandbed region (called the 'without sand case'). A companion report [1-5] addresses the buckling evaluation for this case. The finite element models used in the Reference 1-2 analyses were modified for this case by removing the spring elements representing sand stiffness. It will be shown that this change affects only the stresses in the sandbed and adjacent region. The stresses in the other regions of the drywell are essentially unaffected. 1.4 Report Outline Section Z of the report describes the drywell geometry, materials, ASME Code allowables and load combinations used in the evaluation of applied stresses. Also discussed is the temperature gradient definition in the sand bed region under OBA conditions. Section 3 includes the details of drywell finite element analysis. Seismic load analyses are covered in Section 4. Section 5 presents the Code stress evaluation results to meet the Code criteria. Finally,- the summary and conclusions are discussed in Section 6. The Appendix includes calculated stresses from some of the unit load cases. 1.5 References 1-1 Yekta, M4., "OC Drywell Structural Evaluations," GPUN Technical Data Report No. 926, Rev. 1, February 6,1989. 1-3

Rf # 00664 TNDEX NO. 9-3, REV. 0 1-2 a. 'An ASME Section VIII Evaluation of the Oyster Creek Drywell - Part 1 - Stress Analysis," GE Index # 9-1, ORF # 00664 (November 1990).

b. "An ASME Section VIII Evaluation of the Oyster Creek Drywell -

Part 2 - Stability Evaluation," GE Index # 9-2, DRF # 00664 (November 1990). 1-3 "Structural Design of the Pressure Suppression Containment Vessels," by Chicago Bridge & Iron Co..Contract t 9-0971, 1965. 1-4 GPUN Specification SP-1302-53-044, Technical Specification for Primary Containment Analysis - Oyster Creek Nuclear Generating Station; Rev. 2, October 1990. 1-5 "An ASME Section VIII Evaluation of the Oyster Creek Drywell for Without Sand Case - Part 2 - Stability Evaluation," GE Index # 9-4, DRF # 00664 (February 1991). 1-4

Y 0EX 08663 R REV. 0 O W a Ltut& X O.II >Y _54 ELIvt ^ tL,4e 's, ) Figure 1-1 Drywall Configuration 1-5

VNEX N8. 9-3, REV. 0

2. ANALYSIS BASES 2.1 Drywell Geometry and Materials The spherical section has an inside diameter of 70 ft. which intersects the 33 ft. diameter cylindrical portion.

A transition knuckle is provided at the connection of the sphere to the cylinder (Figure 1-1). The drywell is 105'-6" high. The plate thicknesses vary from a maximum of 2.625 in. at the transition between the sphere and the cylinder down to a minimum of 0.640 in. in the cylinder. The head wall thickness is 1.188 in. The head, which is 33 ft. in diameter, is made with a double tongue and groove seal which permits periodic checks for tightness. Ten vent pipes, 6'-6" in diameter, are equally spaced around the circumference to connect the drywell to the vent header inside the pressure suppression chamber. The drywall interior is filled with concrete to elevation 10'-3" to provide a level floor. Concrete curbs follow the contour of the vessel up to elevation 12'-3" with cutouts around the vent lines. On the exterior, the drywell is encapsulated in concrete of varying thickness from the base elevation up to the elevation of the top head. From there, the concrete continues vertically to the level of the top of the spent fuel pool. The base of the drywell is supported on a concrete pedestal conforming to the curvature of the vessel. A structural steel skirt was first installed to provide interim support for the vessel during erection. A portion of the steel skirt was left in place which serves as one of the shear rings that provides horizontal restraint for the drywall during an earthquake. The proximity of the biological shield concrete surface to the steel shell varies with the elevation. The concrete is in full contact with the shell over the bottom of the sphere at its invert elevation 2'-3" 2-1

NDEX ON8. 9-3, REV. 0 up to elevation 8'-1l 1/4". At that point, the concrete is stepped back 15 inches radially to form a pocket which continues up to elevation 12'-3'. That pocket is currently filled with sand which forms a cushion which is intended to smooth the transition of the shell plate from a condition of fully clamped between two concrete masses to a free standing condition. This sand filled pocket is referred to here as the sandbed. In the analyses described in this report it is assumed that the sand has been removed. Up from elevation 12'-3' there is a 3-inch gap between the drywell and the concrete biological shield wall which is filled with foam material that provides insulation but no structural support. An upper lateral seismic restraint, attached to the cylindrical portion of the drywell at elevation 82'-6", allows for thermal, deadweight, and pressure radial deflection, but not for lateral movement due to seismic excitation. All penetrations for piping, instrumentation

lines, vent

ducts, electrical

lines, equipment accesses, and personnel entrance have expansion joints and double seals where applicable.

The materials of construction for the drywell are given in Specification 5-2299-4 [2-11. The drywell shell, i.e., the sphere, cylinder, dome, and transitions, was constructed from SA-212, Grade B High Tensile Strength Carbon-Silicon Steel Plates for Boilers and other Pressure Vessels ordered to SA-300 specification. The following steels were used in the construction of penetrations, reinforcements, and appurtenances: SA-300 Steel Plates for Pressure Vessels for Service at Low Temperatures. SA-333 Seamless and Welded Steel Pipe for Low Temperature Service. SA-350 Forged or Rolled Carbon and Alloy Steel Flanges, Forged Fittings, and Valves and Parts for Low Temperature Service. 2-2

INDE 9869-3, REV. 0 ASTM A-36 Structural Steel. Table 2-1 shows the as-designed thicknesses used in the Code stress evaluation of the drywell shell 11-2]. Also shown in the same Table are the projected 95% confidence thickness values in the locally corroded areas [2-2]. These latter thicknesses are used in the primary stress evaluation presented in Subsection 5.2. 2.2 ASME Code Allowable Values The Oyster Creek drywell vessel was designed, fabricated and erected in accordance with the 1962 Edition of ASME Code, Section VIII and Code Cases 1270N-5, 1271N and 1272N-5. The Code Case 1272 N-5 limits the general membrane stresses to 1.1 times the allowable stress values given in Table UCS-23 of Section VIIl. The combined general membrane, general bending, and local membrane stresses are limited to 1.5 times the general membrane stress allowables. Finally, the Code Case limits the sum of the primary plus secondary stresses to three times the allowable stresses given in Table UCS-23. The allowable stress value given in Table UCS-23 for SA 212, Grade B is 17500 psi. Accordingly, the allowable stress values for various categories of stresses are shown in Table 2-2. The original Code of record and the Code Cases do not provide specific guidance in two areas. The first relates to the size of a region of increased membrane stress due to thickness reductions from local or general corrosion effects, and the second pertains to the allowable stresses for service level C or post-accident conditions. In the first case, guidance was sought from Subsection HE of Section III. The justification for the use of this guidance is provided in a report prepared by Dr. W.E. Cooper of Teledyne [2-5]. In the latter case, the Standard Review Plan document was used as guidance with details discussed in Reference 2-6. The allowable limits obtained are discussed next. 2-3

?gF # OC664 OEX N. 9-3, REV. 0 2.2.1 Thickness Reductions from Local Corrosion Effects Consideration of local corrosion effects can be achieved by application of the requirements for Local Primary Membrane Stresses. A thorough discussion of this is presented in Reference 2-5. The discussion presented here is extracted from that reference. The NE-3213.10 definition of Local Primary Membrane Stress is: Cases arise in which a membrane stress produced by pressure or other mechanical loading and associated with a primary or discontinuity effect produces excessive distortion in the transfer of load to other portions of the structure. Conservatism requires that such a stress be classified as 'local primary membrane stress even though it has some characteristics of a secondary stress. A stress region may be considered local if the distance over which the membrane stress intensity exceeds 1.1 S., does not extend in the meridional direction more than 1.0(Rt), where R is the minimum midsurface radius of curvature and t is the minimum thickness in the region considered. Regions of local primary membrane stress intensity involving axisymmetric membrane stress distributions which exceed 1.1 Smc shall not be closer in the meridional direction than 2.51(Rt), where R is defined as (R1+R2)/2 and t is defined as (tl+t 2)/2, where tj and t2 are the minimum thicknesses at each of the regions considered, and R, and R2 are the minimum midsurface radii of curvature at these regions where the membrane stress intensity exceeds 1.1 Smc. Discrete regions of local membrane stress intensity, such as those resulting from concentrated loads acting on brackets, where the membrane stress intensity exceeds 1.1 Smc shall be spaced so that there is no overlapping of the areas in which the membrane stress intensity exceeds 1.1 SWc. The value of Smc from NE of Section III is equivalent to 1.1 S from Section VIII. 2-4

MOEX '88. -3, REV. 0 There is no Code limit for the extent of the region in which the membrane stress exceeds 1.0 SMc but is less than I.1SMC. This io% variation in the allowable stress was provided because of the "beam on elastic foundation" effects of such local regions, the stress decays as one moves away from the thin region, but overshoots general membrane stress value by a small amount as the effects dampen out with distance. Thus, this provision is not equivalent to a 10%. increase in the allowable stress which can be taken advantage of in the original design. However, given a design which satisfies the general *Code intent, as the Oyster Creek drywell does as originally constructed, it is not a violation of Subsection NE requirements for the membrane stress to be between 1.*Smc and 1.iSmc over significant distances. Based on the preceding discussion, a limit of 1.lSmc will be used in evaluating the general membrane stresses in areas of the drywell where reduced thicknesses are specified. 2.2.2 Allowable Stresses for Post-Accident Conditfon In the post-accident condition, the drywell is flooded to elevation 74'-6". The allowable stress values for this condition are given In Table 3.8.2-1 of Reference 2-4. Table 2-3 shows the allowable stress values used for the post-accident condition. 2.3 Load Magnitudes and Combinations The loads to be considered in the Oyster Creek drywell stress analysis, and the load combinations are specified in Reference 1-4. References 2-1 and 2-3 also contain similar descriptions of the loads and load combinations. Table 2-4 shows these load combinations. The Cases I and II pertain to test loads imposed on the drywell prior to plant startup. These loads are enveloped by the loads specified in Case V - Accident Condition. Therefore, separate calculations were not conducted for Cases I and tI, 2-5

RF # OQ664 TNOEX No. 9-3, REV. 0 A comparison of the load combinations shown in Table 2-4 and those given In Reference 2-4 is covered in Reference 2-6. From that comparison it was concluded that the load combinations in Table 2-4 essentially envelope those described in Reference Z-4. The dead load, live load and other equipment loads used in the stress calculations were obtained from an earlier study by C8l (Reference No. 2.4.3 of Reference 1-4], and are shown in Tables 2-5a though 2-5c. In the dead weight loading, the weight of the compressible material attached to the drywell was separately added. This weight was taken as 10 lbs. per sq. ft. of drywell surface [Reference No. 2.4.2 of Reference 1-4]. The additional weight on the cylindrical portion of the drywell during the refueling was obtained from Reference No. 2.4.3 in Reference 1-4 as 561 lbs/inch of drywell cylindrical region circumference. The stresses from seismic loads were separately calculated as described in Section 4. 2.4 Temperature Gradients The drywell shell is essentially at a uniform temperature during all of the operating conditions except the accident condition. During the accident condition it is assumed that the drywell shell except the region below the curb (i.e., the sand bed region) is at the same temperature as that of the environment inside the drywell, An analysis of the meridional temperature distribution in the sand bed region during the accident condition was reported in Reference 1-4. The meridional temperature results in Reference 1-4 are given as a function of elapsed time from the start of the accident condition to 4500 seconds. These temperature distributions are used in Section 3 to calculate the stresses. 2-6

NEX, 98'. 9-3, REV. 0 2.5 References 2-1 Technical Specification S-2299-4; Design, Furnishing, Erection and Testing of the Reactor Orywell. and Suppression Chamber Containment Vessels (1964). 2-2 "Forcasted Orywell Thicknesses to 14R," letter dated October 5, 1990 from S.C. Tumminelli of GPUN to H.S. Mehta of GE, dated. 2-3 'Primary Containment Design Report," prepared by The Ralph M. Parsons Company, FSAR Amendment 15. 2-4 Nuclear Regulatory Commission Standard Review Plan, Section 3.8.2, Steel Containment, Rev. 1, July 1981. 2-5 "Justification for use of Section III, Subsection NE, Guidance in Evaluating the Oyster Creek Drywell," Appendix A to letter dated December 21, 1990 from H.S. Mehta of GE to S.C. Tumminelli of GPUN. 2-6 "Comparison of FDSAR and SRP Load Combinations," Appendix D to letter dated December 21, 1990 from H.S. Mehta of GE to S.C. Tumminelli of GPUN. 2-7

N;EI R8. 3-3. REV. 0 TABLE 2-1 As-designed and Projected 95% Confidence thicknesses used in the Code Stress Evaluation As-designed Thicknesses L nu DrYwelIRecon Cylindrical Region Knuckle Upper Spherical Region Middle Spherical Region Lower Spherical Region Except Sand Bed Area Sand Bed Region 0.640 2.625 0.722 0.770 1.154 1.154 Projected 95% 14R Thicknesses 0.619* 2.625 0.677 0.723 1.154 0.736

• no on-going corrosion 2-8

%R # 08664 IDEX NO. 5-3, REV. 0 TABLE 2-2 Allowable Stresses for Drywell Shell in Section VIII Analysis (Except Post-Accident Condition) Primary Stresses General membrane General membrane plus bending 19300 psi 29000 psi Primary olus Secondary Stresses Surface stresses including thermal effects 3x17500 or 52500 psi NOTE: The general membrane stress allowable value of 19300 psi is equal to 1.1x17500, where 17500 psi is the allowable stress value for the drywell material in Table UCS-23 of Section VIII. 2-9

Y O00664 XNO. 9-3, REV. 0 TABLE Z-3 Allowable Stresses for Post-Accident Condition Primary Stresses General Membrane General Membrane plus Bending 38000 psi 1.5x General membrane or 57000 psi Se&9ndarv Stresses Primary plus Secondary 70000 psi NOTE: The above allowable stresses are based Standard Review Plan, Section 3.8.2., Steel Containment 2-10

9NEX' 8l-3, REV. 0 Table 2-4 Load Combinations specified in the Parsons Report (Reference 2-3) CASE I - INITIAL TEST CONDITION Deadweight + Design Pressure (62 psi) CASE II - FINAL TEST CONDITION Deadweight + Design Pressure (35 psi) + Seismic (2 x OBE) + Seismic (2 x DBE) CASE III - NORMAL OPERATING CONDITION Deadweight + Pressure (2 psi external) + Seismic (2 x DBE) CASE IV - REFUELING CONDITION Deadweight + Pressure (2 psi external) + e 118'-3" + Seismic (2 x DOBE) Water load at water seal CASE V - ACCIDENT CONDITION Deadweight + Pressure (62 psi & 175 F or 35 psi & 281 F) + Seismic (2 x DOBE) CASE VI - POST ACCIDENT CONDITION Deadweight + Water Load 9 74' 6W + Seismic (2 x DBE) Notes: (1) The loads shown above predominate. Reference 2-3 contains all of the loads. (2) DBE is the design basis earthquake. 2-11

?OEx' 8N863, REV. 0 TABLE 2.5a Dead Weight Loads Item Elevation (ft.) Weight in lbs Upper Header 60.00 36000 Lower Header 40.00 41000 Upper Weld Pads 65.00 40000 Middle Weld Pads 60.00 40000 Lower Weld Pads 56.00 48000 Top Flange 95.75 20100 Bottom Flange 93.75 20700 Stabilizers 82.17 21650 Upper Beam Seats 50.00 1102000 Lower Beam Seats 22.00 556000 12 Ft Diam. EQ DOOR 30.2S 48000 Personnel Lock 30.00 64100 Vents 15.56 S0000 13 Ft Diam EQ DOOR 30.25 57000 Upper Weld Pads 65.00 12000 Middle Weld Pads 60.00 19200 Lower Weld Pads 56.00 8400 2-12

Ndt N8'.9-3, REV. 0 TABLE 2-Sb Penetration Loads Penetration ID Elevation (ft.) Weight in l§s x - 54A 87.00 1000 x 5 5 A Thru H 16.00 150000 x - 6 16.00 6000 x - 7A Thru D 30.00 45600 x - 8 26.00 2450 x - 9A, 9B 34.00 22600 x-10, 11 26.00 8650 x 12, 45 31.00 16500 x - 13A, 138 33.00 15450 x - 14,15,39B 70.00 5750 x - 43, 44 54.00 7850 x - 16A,B 73.00 8850 x 17 90.00 2750 x 18, 19 20.00 900 x - 20,21,22 40.00 850 x - 23,24,34A,B 20.00 6000 x-25 90.00 3750 x - 27 90.00 1000 x - 28A-G 34.00 5450 x - 30AB. 32A 16.00 3700 x - 31AB, 53 16.00 3750 x - 26 20.00 3900 x - 3SA Thru G 16.00 900 2-13

RF# 00664 MNEX NO. 9-3, REV. 0 TABLE 2-Sb (Cont'd) Penetration Loads Penetration I Elevation (ft.) Weiaht in lbs x - 36 60.00 700 x - 37 A Thru D 40.00 8100 x - 38A Thru D 40.00 8100 x - 42 20.00 400 x - 39A 30.00 850 x - 40 AB, 46A 30.00 2400 x - 46B, 52 30.00 1650 x - 49, 50 35.00 1500 x - 51 32.00 750 x - 100AB, 1048 40.00 2500 x - 105A,D+107A 40.00 2500 x - 100C,D,G+104 40.00 4150 x - 1058,C+1068 40.00 2550 x - lOOE, 103A,10 40.00 2500 x - 102B 40.00 850 x - lOlA-F 40.00 5100 x - 104BD 40.00 1650 x - 548 90.00 1000 x - 55 A+B 90.00 2000 x - 102A,104A,10 40.00 2650 x - 100F,103B 40.00 1850 x - 29A,B,47,48 90.00 4000 x - 32B,33A,33B 16.00 3750 x - 40C0 36.00 1550 x - 41 90.00 500 2-14

NEX N8. 9-3, REV. 0 TABLE 2-Sc Live Loads Item Elevation (ft.) Weight in lbs Upper Header 60.00 4200 Lower Header 40.00 7150 Upper Weld Pads 65.00 20000 Middle Weld Pads 60.00 20000 Lower Weld Pads 56.00 24000 Equip Door 30.25 100000 Personnel Lock 30.00 15000 2-15

NEX NO. l-3, REV. 0

3. DRYWELL FINITE ELEMENT ANALYSIS 3.1 Description of Finite Element Models The drywell was modelled for finite element analysis using the ANSYS computer program [3.1).

Two finite element models, an axisymmetric model and a 36' pie slice model, were used in the stress analysis. Both of these models are essentially the same as those used in the stress analyses [1-2] except that the elements representing sand stiffness were eliminated. The axisymmetric model was used in determining the stresses for the seismic and the thermal gradient load cases. The pie slice model was used for dead weight and pressure load cases and to evaluate the stresses for load combinations. The pie slice model includes the effect of vent pipes and the reinforcing ring on the stress state in the sandbed and adjacent region. 3.1.1 Axisymmetric Model The axisymmetric model is shown in Figures 3-1 through 3-5, where Figure 3-1 is an overview, and Figures 3-2, 3-3, 3-4, and 3-5 show the sand bed, knuckle, cylindrical, and upper most cylindrical regions, respectively. The geometry as described in Subsection 2.1, along with References 3-2 and 3L3, was used in generating this model. The model was developed using axisymmetric solid elements (STIF 25), with the lower most portion being fixed in all directions. This element has asymmetric load capability which was required for the seismic evaluation. Seismic evaluations are discussed in Section 4. 3.1.2 Pie Slice Finite Element Model Taking advantage of symmetry of the drywell with 10 ventlines, a 36' section was modeled. Figure 3-6 shows the 36' pie slice finite element model of the drywell. This model includes the drywell shell 3-1

N N8. 3-3, REV. 0 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 drywall and torus. The various colors in Figure 3-6 represent the different shell thicknesses of the drywall and ventline. Figure 3-7 shows the view from the inside of the drywell with the gussets and the vent jet deflector. The drywell and vent shell are modeled using the 3-dimensional plastic quadrilateral shell (STIF43) element. At a distance of 76 inches from the drywell shell, the ventline modeling was simplified by 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 ventline. ANSYS STIF4 beam elements are then connected to this centerline node to model the axial and bending stiffness of the ventline and header. Spring (STIF14) elements are used to model the vertical header supports inside the torus. ANSYS STIF4 beam elements are also used to model the stiffeners in the cylindrical region of the drywall. Symmetric boundary conditions are defined for both edges of the 36' drywell segment. This allows the nodes at this boundary to move radially outward from the drywell centerline and vertically, but not in the circumferential direction. Rotations are also fixed in twc 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 concretec foundation; 3.2 Load Application on Pie Slice Model The loads are applied to the drywell finite element model in the manner which most accurately represents the actual loads anticipated on the drywell. Details on the application of loads are discussed in the following paragraphs. 3-2

r?F R# 00664 NDEX NO. 9-3, REV. 0 3.z.1 Gravity Loads The gravity loads include dead weight loads of the drywell shell, weight of the compressible material and penetrations and live loads. The drywell shell loads are imposed on the model by defining the weight density of the shell material and applying a vertical acceleration of 1.0 9 to simulate gravity. The ANSYS program automatically distributes the loads consistent with the mass and acceleration. The compressible material weight of 10 lb/ft2 is added by adjusting the weight density of the shell to also include the compressible material. The adjusted weight densities for the various shell thicknesses are summarized in Table 3-2. The additional dead weights, penetration weights and live loads are applied as additional nodal masses to the model. As shown on Table 3-3 for the refueling condition 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 in Table 3-3. These applied masses automatically impose gravity loads on the drywell model with the defined acceleration of

19.

The same method is used to apply the additional masses to the model for the accident and the post-accident conditions as summarized in Table 3-4. 3.2.2 Pressure Load The appropriate pressure load is applied to the internal/external faces of all of the drywall and vent shell elements. The 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. In the post-accident condition, the drywell is assumed to be flooded to elevation 74'-6" (894 inches). Using a water density of 62.3 lb/ft3 (0.0361 lb/in3), the pressure gradient versus elevation is calculated as shown in Table 3-5. The hydrostatic pressure at the 3-3

dEX g. 9-3, REV. 0 bottom of the sandbed region is calculated to be 28.3 psi. According to the elevation of the element centerline, the appropriate pressures are applied to the inside surface of the shell elements. 3.2.3 Seismic Loads Seismic inertia and displacement stresses were first calculated using the axisymmetric model. The seismic meridional stresses determined from the axisymmetric model were then imposed on the pie slice model by applying downward forces at four elevations of the model (A: 23'-7",B: 37'-3",C: 50'-1IV and 0: 88'-9") as shown on Figure 3-8. Using this method, the meridional stresses calculated from the axisymmetric model are duplicated at four sections of the pie slice model including 1) the mid-elevation of the sandbed region, 2) 17.250 below the equator, 3) 5.75' above the equator and 4) just above the knuckle region. These four sections were chosen to most accurately represent the loading in the lower drywell while also providing a reasonably accurate stress distribution in the upper drywell. Table 3-6 shows the meridional stress magnitudes at the four sections. Unit loads are then applied to the pie slice model in separate load steps at each elevation shown in Figure 3-8. The resulting stresses at the four sections of interest are then averaged for each of the applied unit loads. By solving four equations with four unknowns, the correct loads are determined to match the stresses shown in Table 3-6 at the four sections. The calculation for the correct loads are showr in Tables 3-7 and 3-8 for the accident and post-accident conditions, respectively. 3.3 Stress Results for Various Load Cases and Combinations Only the two orthogonal stress components meridional and circumferential - are significant at the maximum stress locations in the drywell shell. A review of the component stresses indicated that. the calculated shear stress magnitudes are insignificant compared to the values for the total meridional and circumferential stresses. Therefore, the orthogonal stress magnitudes and the principal stress; 3-4

RNEd 98. l-3, REV. 0 magnitudes were essentially the same. Also, the maximum stress was equivalent to the stress intensity at the locations evaluated. The stresses for the seismic inertia, seismic displacement and temperature load cases (see Table 3.1) were calculated using the axisymmetric model. The details of the temperature stress analysis is described in the next Subsection and the procedures used in the calculation of the seismic stresses are covered in Section 4. The calculated values of the membrane and membrane plus bending stresses for temperature case are tabulated in Appendix A. The seismic stresses were incorporated in the pie slice model to determine the overall stress resultants for the accident and post-accident load combinations. The temperature stresses determined from the axisymmetric model were separately added to the accident condition stresses obtained from the pie slice model. The multipliers applied to the various unit load cases (Table 3-1) to obtain total stresses for a particular load combination are shown in Table 3-9. The resulting stresses for these load combinations are discussed and compared with the Code allowables in Section S. 3.4 Temperature Stress Analysis The thermal response in the sand bed region to a DBA LOCA has been analyzed by GPU in Reference 1-4. Figure 3-9 shows the meridional nodes below the drywell floor, for which the calculated temperatures as a function of elapsed time are reported in Reference 1-4. An example of the calculated temperatures is shown in Figure 3-10. From a review of the temperature distributions, two intermediate time steps were identified as possibly yielding the most severe thermal stresses. At 60 seconds, the largest temperature gradient occurs over a two inch meridional length. At 210 seconds, the maximum temperature is achieved. In addition, a third time step, 690 seconds, was evaluated to verify that a more deeply penetrating temperature condition would not result in higher stresses than the first two cases. 3-S

9N'EX' '. l-3, REV. 0 The predominant stresses for each of these cases occurred near the top of the sand bed region (near the 0.736" to 1.154" transition) and were in the circumferential and meridional directions. It was found that the thermal stresses at 210 seconds yielded the more severe stress condition. Figures 3-11 and 3-12 show the meridional and circumferential stress distributions in the sand bed region. 3.5 References 3-1 Gabriel J. DeSalvo, Ph.D. and John A. Swanson, Ph.0, "ANSYS Engineering Analysis System User's Manual," Revision 4.1, Swanson Analysis System, Inc. Houston, PA, March 1, 1983. 3-2 CB&I Orwg. 9-0971 sheet number 4, Rev. 1, "Drywell - Field Weld Joint" 3-3 CB&I Drwg. 9-0971 sheet number 7,.Rev. 5, "Drywell - Cylindrical Shell & Top Head" 3-6

sNOEX 88. 9-3, REV. 0 TABLE 3-1 Load Cases Considered in the Finite Element Analysis Case No. Loading 1 Pressure 2 Gravity-i (Accident Condition) 3 Gravity-2 (Refueling) 4* Unflooded Seismic 5 Flooded Seismic 6 Flooded Hydrostatic Pressure 7 Seismic Relative Support Displacement 8 Temperature Gradient Ouring DBA Load Cases Analyzed by Axisymmetric Finite Element Model 3-7

UNSEX 88."9-3, REV. 0 TABLE 3-2 Adjusted Weight Densities of Shell to Account for Compressible Material Weight Shell _Thigtnessfinj) Adjusted Weight Density (lb/in3) 1.154 0.770 0.722 2.563 0.640 1.250 0.343 0.373 0.379 0.310 0.392 0.339 3-8

NbEX N8. 3, REV. 0 TABLE 3-3 Additional Weights - Oyster Creek Drywell Refueling Condition ELEVATION fat) 15.56 16 20 Z2Z

• -21-254 Z6 30 30.25 o 26-30 31 32 33 34 35

" 31.35 36 40 35-40 45-50f 54 51-55 55 60

• 56-60 65
  • 61-655 70 6C-70 73
• 71-75 12.17 1-85 87 90

'- 66-90 93.75 94.751 95.75 A 91-96 TOTAI.S: DEAD (Ibf)

• 50000 PENETR.

WEIGHT Olbf) 168100 11200 556000 64100 105000 41000 1102000 56400 95200 S2000 11100 51500 16500 750 15450 26050 1500 1550 43350 7850 mISC. TOTAL LOADS LOAD t1bf) (1bl) 500O0 168100 11200 556000 11100 115600 100000 205000 16500 7SO 15450 Z8oSO 1500 1550 64350 1102000 7850 24000 50400 20000 115900 20000 72000 5750 ISO 21650 1000 15000 20700 8000 695000 20100 5 FOOT RANGE LOAD 229300 556000 LOAD PER 36 DEG. (Obf) 22930 55600 6 a 116-119 161-169 3822 1911 6950 3475

• OF NODES OF ELEAENTS APPLICATION LOAD PER FULL NODE LOAD PER HALF NODE (lbf)

._..4._.. 331100 33170 S2250 5225 a 179-167 a IU-19O 4146 2073 778 369 700 85900 1102000 78SO 196300 7tO0 5750 1850 21650 8590 110200 715 19630 72O0 S75 853 21H5 5750 e6so B 197-205

a.

418-426 8 436-444 8 454-462 8 472-460 a 508-516 a 526-534 8 $53-581 a 571-579 a 589-597 1074 13775 98 2454 900 72 11I 271 200 537 6888 49 1227 450 38 55 135 100 21550 00.0 15000 16000 1600 20700 20100 86--3--5 218J150 3J$200 J62W00 3434350 73UO0 343450 73880 343435 9235 4618 LOAM TO IE APPLIED IN VERTICAL DIRECTION ONLY. - miSCELLANEOUS LOA INCLUDE 696000 LS WATER VEI6T AT 94.75 FT. ELEVATION 100000 LI EQIJPININT OM WE1IHT AT 30.25 FT. ELEVATION ED0 WELD PAD LIVE LOADS OF 24000. 20000 ANO 20000 AT S5. 60 AN0 65 FT. ELEVATIONS RECFWGT.1K 3-9

NEX, 8.69-3, REV. 0 TABLE 3-4 Oyster Creek Drywell Additional Weights Accident and Post-Accident Condition R. MISC. TOTAL. S FOOT LOAD PER HT LOADS LOAD tAog 36 ODG. f OF ) (1bf) (Ibf) LOAD (lbf) ELEMENTS A ELEVAT ION (feet) 15.56 156 20 15-20 Z2, 26 30 30.25 Z6-30 31 32 33 34 35 31-36 36 40 40 50s 45-501 54 0* 51-55 56 60 58.60 6s 61.65 70 86710 73 937.75 12.,1? "' sl..aS 81 90 86*40 93,75 95.75

• 91.96 TOTALS:

DEAD WEIGHT VE IVG) lbr) 530000 PENE5T VEIGI lbf 16611 It21 00 00 SS6000 64100 105000 41000 1 o02000 56400 9S200 52000 ll0oo 51500 16500 7SO 15450 B00SO 1500 1550 43350 7850 700 50000 11100 11200 556000 11100 115600 105000 16500 750 t5450 28050 1500 1550 8.4350 1202000 7850 58400 95900 52000 5750 U50 21550 1000 15000 20700 20100 229300 556000 ZZ930 55600 8 a NODES OF hDPLICATION 116-119 l61-169 LOAD PER FULL moot (lbf) LOlA PER HVLF WOVE (Thf I 1 1..... 3122 1311 69S0 3475 231700 23170 52250 5225 6 179-151 158-196 2896 1448 77t 389 SS900 1102000 M65o 152300 52000 5750 6O 21M10 $590 110200 'SI a 8 a 5750 "50 lS230 S200 575 us 216S a a a a a 197-20S 416-425 436-444 454-462 472-410 508-516 526-534 553-561 1074 13775 90 1904 650 72 111 271 200 537 6888 49 952 325 36 55 135 100 21550 1000 15000 20700 20100 16000 1000 4060 257.23S a 5s1-Si9 8 58591 40800 2184150 380200 0 2512350 2572350 P - LOAD T0 BE APPLIED IN VERTICAL DIRCCTION ULYT. L - NO MISCELLANEOUS LOADS FOR THIS CONDITION. 510 255 FLOOW T.W1 3-10

9NEX N8. 9-3, REV. 0 TABLE 3-5 Hydrostatic Pressures for Post-Accident Condition WATER DENSITY: FLOODED ELEV: 62.32 lb/ft3 0.03606 Ib/in3 74.5 ft 894 inches ELEMENTS ABOVE NODES 27 40 53 66 79 92 102 108 112 116 120 124 130 138 148 161 170 179 188 197 400 409 418 427 436 44$ 454 463 472 481 490 499 508 517 526 ANGLE ABOVE EQUATOR (degrees) 53.32 -51.97 -50.62 -49.27 -47.50 -46.20 -44.35 -41.89 .39.43 -36.93 -34.40 -31.87 -29.33 -26.80 -24.27 -20.13 -14.38 -8.63 -2.88 2.88 8.63 14.38 20.13 25.50 30.50 35.50 40.50 45.50 50.50 54.86 ELEVATION (inch) 110.2 116.2 122.4 128.8 137.3 143.9 153.4 166.6 180.2 194.6 209.7 225.2 241.3 257.6 274.4 302.5 342.7 384.0 425.9 468.1 510.0 551.3 591.5 627.8 660.2 690.9 719.8 746.6 771.1 190.5 805.6 820.7 835.7 850.8 885.3 DEPTH (inch) T 783.8 777.8 771.6 765.2 756.7 750.1 740.6 727.4 713.8 699.4 684.3 668.8 652.7 636.4 619.6 591.5 551.3 510.0 468.1 425.9 384.0 342.7 302.5 266.2 233.8 203.1 174.2 147.4 122.9 103.5 88.4 73.3 58.3 43.2 8.7 PRESSURE (psi) 28.3 28.1 27.8 27.6 27.3 27.1 26.7 26.2 25.7 25.2 24.7 24.1 23.5 23.0 22.3 21.3 19.9

18.4 16.9 15.4 13.8 12.4 10.9 9.6 8.4 7.3 6.3 5.3 4.4 3.7 3.2 2.6 2.1 1.6 0.3 ELEMENTS 1-12 13-24 25-36 37-48 49-SI, 61-66,55-57 52-54, 138-141.58-60 142-141, 240-242, 257-259 148-1SI, 243, 2S6 152-155, 244, 255 156-159, 245, 254 160-165, 246, 253 166-173, 247, 252 174-183, 248-251 184-195 196-207 208-215 216-223 224-231 232-239 430-437 438-445 446-453 454-461 462-469 410-477 478-485 486-493 494-501 502-509 510-517 518-525 526-533 534-541 542-S49 550-557 187.3 706.7 25.5 340-399 (Ventline)

FLOODP.WK1 3-11

NEEX 98'.33, REV. 0 TABLE 3-6 Meridional Seismic Stresses at Four Sections Elevation Section (inches) A) Middle of Sandbed 119 B) 17.25 Below Equator 323 C) 5.75 Above Equator 489 D) Above Knuckle 1037 2-D Shell Model 32 302 461 1037 Meridional Stresses Accident Post-Accident 1DS1J (ps;i) 1258 1288 295 585 214 616 216 808 3-12

9dE &8'.4-3, REV. 0 TABLE 3-7 Application of Loads to Match Seismic Stresses - Accident Condition 2-0 SEISMIC STRESSES AT SECTIO% (pill SECTION; 2-0 XOOD: ELEV: COMPRESSIVE STRESSES FNM 2-D tANAYSIS 0.058S SECSMIC DEFLECTION: HORIZ. PLUS VERTICAL SEISMIC INERTIA: TOTAL SEISSIC COMPRESSIVE STRtSSES: 1 32 119.3" 7U.61 465.55 2. 302 155.54 13g." 3 461 £03,"s 110.13 4 1031 9IZ.31 85.31 130.21 125_.22 23..... 12SU.ZZ 294.90 213.S9 215.52 3-0 INPUT LOAD SECTION A I C 0 SECTION: 3-0 0W0S: ELEY: INPUT 3-0 UNIT LOAD DESCRIPTION 3-0 INPUT LOAD SICT;ON A a C 1000 lbs St fiO*S 583 thPOUgh 559 500 lbs at 4Z7U43S. 1000 lb. it 428-4U 500 lbs at 1971209. 1000 lba at 196204 500 lbs at 1611169. 1:00 lbs *t 152-16 OESISEo COSWRESSIVE STIESSES (psil: LAD TO St APPLIED TO MATCH 2-0 STRESSES 3902.2 2101.4 1453.6 6611.1 3-0 STlESSES AT SECTION (psi) l 2 3 4 53-65 170-178 400-408 525-534 11S.3' MAt.O 489.1" 912.3-85.43 7.54 34.H 55.23 89.60 39.92 36.75 0.00 97.54 43.37 0,00 0.00 89.85 0.00 0.00 0.00 . v 1258.22 294.98 213.59 215.32 RESULTIN6 STRESSES AT SECTION (psi) 333,37 141.05 136.34 215352 1u.87 63.19 77.25 0.00 141.93 63.04 0.00 0.00 594.05 0.00 0.00 0,00 __58....... --9. _ -._9 ..1... 125812t 294.98 213.59 215.52 5UL: SEI5UNFL.l1 3-13

9NE 8. 3-3, REV. 0 TABLE 3-8 Application of Loads to Match Seismic Stresses Post-Accident Condition SECTION: 2-0 NOCE: ELEY: COMPRESSIVE STRESSES FROM 2-0 ANALYSIS 0.058" SEISMIC DEFLECTION: HCRIt. PLUS VERTICAL SEJIMIC INERTIA: T..OT.. IC TOTAL SEISMIC CtX8PRESSIVE STRESSES: 3-0 INPUT LOAO SECTION A C a SECTION: 3-D NONES: tEEV: INPUT 3-0 UNIT LOAD DESCRIPTION 2-D SEISMIC STRESSES AT SECTION (w)t) 1 2 3 4 3Z 302 461 1037 119.3" 322.5 483.1" 912.3-718.67 ISS.54 103.46 85.31 499.79 429.39 512.76 723.14 ,,w__....... 1Z85.46 584.93 616.22 SO.4! 3-0 STRESSES AT SECTION (psi) 1 2 3 4 53-S 170-170 400-406 526-56:4 119.3' 322.5-403.1 912.3' ..__.. j.... 85.43 37.94 34.94 55.2,; 819.8 39.92 38.76 O.OC 91.64 43.17 0.00 0.011 99.85. 0.00 0.00 0.0t 1288.46 584.93 616.2Z 8a8.4!; RESULTING STRESSES AT SECTION (pgt) 1250.51 555.36 511.45 808.45 256.17 113.78 104.77 0.02 -159.58 -84.21 0.00 0.00 -2B.64 0:00 0.00 0.00 1288.48 584.91 516.22 $08.45 3-0 INPUT SECTION A C 0 1000 lbs at lodos 563 tOrough 589 500 lbs at 427143S. 1000 lbs at 428-434 500 lbs at 1971205. 1000 lbh 4t 19t8204 500 lb* at 1611U69, 1000 ibs St 152-1l8 DESIRED COMPRESSIVE STRESSES (psi): LOAD TO BE APPLIED TO PATCH t-O STRESSES 14637.9 2850.2 -1941.7 -318.8 SIW: SEISfL.WK1 3-14

ME1 98'8. 9-3, REV. 0 TABLE 3-9 Description of Load Combinations in Terms of Unit Load Case Sum Load Comb. I a Nc Load wrmal 0 Conditio Refuel in Accident Accident I Combination CaseWW Constituent Load Cases perating III - (Case I)xO.03226 + Case 2 +/- n(3) Case 4 t Case 7 g Condition IV - (Case 1)xO.03226 + Case 3 +/- Case 4 +/- Case 7 Condition - l V-1 + Case 1 + Case 2 t Case 4 +/- Case 7 + Case 8 Condition - 2 V-2 + (Case 1)x0.565 + Case 2 + Case 4 +/- Case 7 + Case 8 ident Condition VI + Case 2 +/- Case 5 + Case 6 i Case 7 (1) For load'combination definition see Reference 2-3. (2) For unit load case description see Table 3-1. (3) Normal Operation also includes live load due to personnel lock. (4) Load Combination Case Numbers are based on Table 2-44 Post-Acc Notes: 3-15

AKSYS 1 0/1 5/90 2.8923 PREP? ELEtENTS AUTO SCALING DIST-653 XF-210 YF-700 EDGE wI* Im OYSTER CREEK DRYYELL - FE I Figure 3^1 Complete Finite Element Model of Drywell

Am5y5 121 4'9e 14.0624 PREP7 ELEMENTS XNPX=2e00 YMAX-1 75 T06-C =: AIUTO SCALIttG ZU 1 DIST-39. 9 XF-283 YF-140 ZF--.01 13 w SAMD BED REGION - H0 SAND I Figure 3-2 Sand Bed Region of Drywall Finite Elemnt Model

AllsYs 1B/t5/30 3.2315 PREP? ELEMENTS ^ XfI;X-2000 ~Ynr-780 B YMPBXJ970 AUTO SCALING DIST-49. 4 XF u227 YF 825 co3 Figure 3-3 Knuckle Region of Drywell Finite Element Model

ANSYS 1O'15'90 3.0986 PREP7 ELEMENTS XHAX=2000 Ynri-960 YtMAXI 145 MIIUMz 1 AUTO SCALING ZU-1 DIST-101 XF-208 Yr-i 052 EDGE (ha b" Figure 3-4 Cylindrical Region of Drywell Finite Element Model

ANS5YS 10/15/YB 3.3621 PREP7 ELEMENTS XMRX=2B08 YMI1-1 864 YMAX-1 1 t1 MNUl-I AUTO SCALING ZVj-DISTu-e7.9 XF-203 YF-1 090 w OYSTER CREEK Di I FE MODEL Figure 3-5 Upper Cylindrical Region of Drywell Finite Element Model

AtSVS

4. 4A JA0 4 1991 13:a:

22 r7 LNIJ kliWL HUM XV =1 YU =-B.8 ZF =639.498 AHCZ--H9 I amriTox1 HipDm4 w -An4 r LM G - OYSSX~ CXEMX ID"WE AMLY3116 - OYCJUP1 CHtO SAND, ACCIDIENT}l Figure 3-6 Oyster Creek Drywell Pie Slice Finite Element Model

I tmes Jl ia Jo" 4 1991 13:14: 2S Part? J NTS XV =-I VU =-0.8 Zr -216. 528 ANC7=98 C3MTBID HIDDEN Wa N I I OYSTER CREEK DRYMELL ANALYSIS - O*CIIH (H_ SAlD. ACCIDENT> Figure 3-7 Inside Closeup View of Lower Drywell Section

(ib ta L-lC OYStER COCAK DRYWALL A#ZLYKXU OVC"1 Cat Seag. ACCIDEN!) AtiSS

4. 4 JAN 4 1991 P? ISLIENTS FORC WU

=U .Xl =39.031 Zr =2 39.49B RC7r1-90 RICM&OI HIDDEN Figure 3-8 Application of Loading to Simulate Seismic Stresses

9;6EI 0.93, REV. 0 I- -le t fI DW Uoner I I I I I I I~ I

LI I

how 00 k Tm_ Figure 3-9 Below Curb Orywell Model Analysis During Accident Nodalization for Temperature Condition 3-24

N'E~X' R8`9-3, REV. 0 Thu ($3uo5¶ 01S ¶1 120 150 ¶90 210 340 27 30 gist TVWIM11 IN UGMS 9 0 24.7 2". Z1.5 231.3 251.8 2S1. al.? 211.7 0.2 UA.6 230.7 Z.4. 237.3

23. 1 24.,1 2O'.9 241.

2.5 MtO.A 213-4 19.4 W3. 2U-4 228.9 Ud. I 1.3 4.,S lo, Ts.t o0s 210.3 411.3 al?.2 219. 4t U13 7 173 1.9 !91 197,2 202. 304 ¶ 20o a1t.& 1.25

• 3.7 167.9 173 181.4 to, 19.& 1s9.9 201.1 5

10.9 115.3 '65.9 11.9 180.2 11t.1 184.2 192.5 M. 1 119.7 t 163.3 I70 t7O. 179.g 163.6 2 30.1

34.3

'4.. t53.3 W0.4 166.2 171 171 2.23 !21.8 121.7 6 14.3 1it.1 117.6 141,6 166.9 2.5 l 116.5 2.3 136.1 143.5 S49.4 154.7 159.2 13 0 lC9.3 112.5 1 t29.3 13.2

143, 17,4 152 3

1047 107.1 15.! 122.9 129.6 135.5 14.7 14s.3 3.2t 101.1 103.& 110.4 l7.4 123.7 129.4 134.5 139 3.5 94.23 1O.2

• 1.4 112.6 115.1 123.9 128.5 133.3 3.7 96.03 97.4 103 10.5 113.9 li 123.7 12 4

904.31 9."

00.1 1o0 109.9 I11.6 119.1 123.3

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AMrsYs 1' 4/91 14.9424 PDST 1 STEP=t ITER-1 0 STRESS PLOT AUTO SCALtIHG DIST-8.78 XF-292 YF'1 44 MXt11 B1 MN l -9857 of 500B 1 5003 OYSTER CREEK DRYhELL - THERMAL T210e S Figure 3-11 Meridional Stress Distribution in the Sand Bed Region From Temperature Distribution at t-210 Seconds

AISY5 l/ 4/91 14.9778 Posri STEP=1 ITER-1 0 5TRESS PLOT AUTO 5CALING ZV-1 DIST-B. 79 XF-292 YF-144 /MX-1 3475 0 G160e OYSTER CREEK DRYUELL THERMAL T210 8 5 t Figure 3-12 Circumferential Stress Distribution in the Sand Bed Region From Temperature DistrIbution at t-210 Seconds

NEI 9.36.-3, REV. 0

4. SEISMIC LOAD DEFINITION This section briefly describes the general methodology followed in the seismic evaluation.of the drywell. A detailed report on the seismic analysis methodology and the results is included in Reference 4-1.

4.1 Finite Element Model The axisymmetric finite element model was used in the seismic analysis. All of the concentrated loads listed in Tables 2-5a and 2-5b were included in both the flooded and unflooded seismic analyses. Since the lower and upper beams connect to the drywell through pads, the beam weights do not act during the horizontal earthquake excitation. Therefore, the beam weights are active only in the vertical direction. In addition, the live loads listed in Table 2-5c were included in the unflooded seismic analysis. The drywell is constrained at the "reactor building/drywell/star truss" interface at elevation 82'-6" and at its base. The upper constraint was implemented in the finite element analysis by restraining the middle node in the horizontal direction at this elevation. The base constraint is as before, i.e., all nodes fixed. 4.2 Dynamic Analysis Methodology and Response Spectra The seismic input motion spectra were provided by GPUN in Reference 1-4. The seismic motion spectra were for two locations: at the mat foundation and at the upper constraint. Since the ANSYS program can only accept one Input spectrum, the input spectra at the two elevations were enveloped. The response spectrum dynamic analyses were first conducted for frequencies up to the ZPA frequencies of the input motion spectra. The response contributions due to the truncated higher frequency modes 4-1

'?NEXd 8.9-3, REV. 0 were calculated by static analyses in which the total model mass is subjected to support accelerations. These were taken as ZPA accelerations for each of the orthogonal spatial directions. All colinear modal response contributions were combined by the Double Sum Method and the spatial contributions by the SRSS method. The response contributions due to the truncated higher frequency modes were combined with the response totals due to the lower frequency modes included in the analysis by the SRSS method. The resulting total colinear inertia responses were combined with the corresponding responses due to relative support motion by the absolute sum method. These stresses were then combined with the stresses from other loads (e.g., pressure, thermal, etc.) for the Code evaluation. 4.3 Post-Accident Seismic Analysis In the post-accident condition, the drywell is flooded to elevation 74'-6". The weight of the water was lumped at several elevations along the meridian of the drywell. Based on previous experience, the fluid-structure interaction effects were assumed as negligible and the hydrodynamic mass of water was assumed as 80% of the total mass of water which would fill an empty drywell. This exclusion of 20%. mass reasonably accounts for the volume of RPV, shield wall and pedestal. 4.4 Analysis for Relative Support Displacement Effects The drywell is fixed at its base and is laterally constrained by the reactor building'at elevation 82'-6". During seismic excitation, the reactor building would experience relative displacement between the drywell constraint elevation and the basemat. Since the reactor building is much stiffer and much more massive than the drywell, it will take the drywell for a 'ride' during relative support displacement. Therefore, the stresses in the drywell due to relative support displacement were determined and added to those from the seismic inertia loads. 4-2

TPF # 00664 ?DEX NO. 9-3, REV. 0 The horizontal relative displacement of the drywell upper support with respect to the drywell at the basemat was specified as 0.058 inch for ZxDBE condition [1-41. The stresses from this relative displacement were obtained by applying a horizontal displacement of 0.058 inch at the upper support elevation. 4.5 References 4-1 "Seismic Analysis Details," Appendix B of letter dated December 21, 1990 from H.S. Mehta of GE to S.C. Tumminelli of GPUN. 4-3

DRF i 00664 INDEX N8. 9-3, REV. 0

5. CODE STRESS EVALUATION Sections 3 and 4 describe the analyses for shell stresses for the various unit load cases and the limiting load combinations V and VI.

The stress analysis for the 'with sand case' in Reference 1-2a has shown that the accident condition, load combination V-}, and the post-accident condition, load combination VI, represent the limiting load combinations for the Code stress evaluation. This was also determined to be the case for the 'without sand' configuration considered in this report. The removal of sand from the sandbed region affects the stresses only in the sandbed and the adjacent lower spherical region. Therefore, the Code stress evaluation of these regions is described separately from the other regions of the drywell. 5.1 Code Stress Evaluation of Regions Above the Lower Sphere Figure 5-1 shows a plot of the accident condition membrane circumferential stresses for the 'with' and 'without' sand cases as a function of meridional distance. Stresses in both the sandbed and the other drywell regions are included in Figure 5-1. It is seen that in the other regions the stress magnitudes for the two cases are essentially identical. From the preceding it is clear that the stresses in the other regions (i.e., other than the sandbed and the adjacent lower spherical region) are unaffected by removing the sand. Nevertheless, for completeness, the calculated stress magnitudes for these regions from Reference 1-2a are repeated in Tables 5-la and 5-lb. The stress magnitudes shown in Tables 5-la and 5-lb are computed using elastic small displacement analysis. As discussed in Subsection 5.2, the stresses in the sandbed and lower sphere regions were also evaluated using elastic large displacement analysis. A comparison of the component stresses from the small and large displacement solutions for the drywell regions above the lower sphere showed insignificant differences. 5-1

MMEX N. 9-3, REV. 0 In order to evaluate the impact on the penetration analyses, a comparison of the radial and meridional displacements at the equator plane of the sphere (elevation 37'-3") for the with and without sand cases was performed. The comparison showed that the radial displacements in the two cases were essentially identical but the meridional or vertical displacements differed by = 0.04Z inch for load combination V-1. This difference was judged to be small compared to the calculated vertical thermal displacement of = 0.5 inch for the accident condition load combination V-2. 5.2 Elastic Stress Analysis of Sandbed and Lower Sphere 5.2.1 Small Oisplacement Solution Results The maximum stresses are along the meridional boundary of the model (i.e., the plane of symmetry between the vents), so the stresses along this boundary will be considered first. Figure 5-2 shows the plot of meridional membrane stress magnitudes for the accident condition V-I as a function of meridional distance from the bottom of the sandbed. A comparison of the membrane stress magnitudes in Figures 5-1 'without sand' case and Figure 5-2 shows that the circumferential stress is higher than the meridional stress in both the sandbed region and the lower spherical region. This is expected since the absence of sand springs allows more radial displacement of the drywell shell under dead weight and internal pressure. Figure 5-3 shows a plot of the membrane circumferential stress distribution. The maximum value of the circumferential membrane stress is - 23.0 ksi.

Further, this stress exceeds 1.1 Smc (21.2 ksi) for a meridional distance of - 26 inches (see Figure 5-1).

The Code (NE-3213.10) states that cases arise in which a membrane stress produced by pressure or other mechanical loading and associated with a primary or discontinuity effect produces excessive distortion in the transfer of load to other portions of the structure. Such a membrane stress is conservatively classified by the Code as local primary membrane stress. The Code limits the magnitude of this stress to 1.5 Smc (29.0 ksi). A stressed region may be considered local if 5-2

oRF #0664 INDEX N. 9-3, REV. 0 the distance over which the membrane stress intensity exceeds 1.1 SMc does not extend in the meridional direction more than 1.0/(Rt). With R-420 in. and t*0.736 inch in the sandbed region, 1.0/(Rt) is equal to 17.6 inches. Thus, the maximum value of the circumferential membrane stress (23.0 ksi) meets the Code stress limit (29.0 ksi) but its meridional extent over 1.1 Smc is greater than 1.0/(Rt). The meridional extent of 26 in. occurs only at the plane of symmetry between the vent lines. The extent is less at other meridional planes. Figure 5-4 shows the meridional extent of circumferential membrane stress above 1.1 Smc at four meridional planes. Using a weighted average over the circumference of the model, the meridional extent was calculated as 14 inches. This average value is less than 1..O/(Rt) and, thus, meets the meridional extent criterion given in NE-3213.10. The objective of the Code in limiting the meridional extent and magnitude of the local primary membrane stress is to preclude excessive distortion in the transfer of load to other portions of the structure, since such distortion could invalidate the elastic analysis. The small displacement results showed that the maximum radial displacement in the sandbed region was 0.28 inch for the accident condition V-1. This is less than half the modeled thickness of the drywell in that region and, therefore, Is judged not to be excessive. The small displacement analysis conducted previously is conservative because the stiffening effect of the tensile in-plane stresses is not considered. This effect would tend to reduce the local radial deflection (thus, also the local circumferential stress) of the drywell shell in the sandbed region. For example, consider the case of a beam subjected to both transverse and tensile axial loads as shown in Figure 5-5. A small displacement analysis of this configuration considers the bending moments based on the transverse load only. The bending stresses and deflections of the beam are overpredicted based on these bending moments. In a real structure, tensile axial loads in combination with the deflections of the beam 5-3

RF # 00664 NOEX NO. 9-3, REV. 0 produced by transverse loads creates an opposing bending moment. As a result the overall bending moment is reduced, leading to smaller bending deflections and stresses. This stiffening effect can be included only by conducting a large displacement analysis. 5.2.2 Large Displacement Solution Results Based on the preceding discussion, a large displacement analysis was conducted using the same pie slice model and the accident condition V-I loads. A large displacement analysis can be conducted using the ANSYS code by activating the KAY(6) key. When this option is chosen, the ANSYS program first calculates displacements of the structure based on a small displacement analysis. The geometry of the structure is then updated based on the calculated displacements. The loads are again applied to the structure and the displacements are recalculated. The geometry of the structure is continually updated and the displacements are recalculated until the maximum displacement change between successive iterations is reduced below the selected convergence criteria. A convergence criteria of 0.01 inch was chosen for this analysis. In this manner, the ANSYS code accurately accounts for the stiffening of the structure due to in-plane tensile stresses. Figure 5-6 shows the distribution of membrane circumferential stress. Figure 5-7 shows a plot of membrane circumferential stress as a function of meridional distance when the large displacement option in ANSYS was used. For comparison, the stress results from the small displacement solution (Figure 5-1) are also shown in Figure 5-7. It is seen that the maximum value from the large displacement solution is = 21.5 ksi (compared to a 23 ksi in the small displacement analysis), and it exceeds 1.1 Smc (21.2 ksi) over a maximum distance of only 11 inches at the meridional plane between the vent lines. This is clearly less than the 1.0 /(Rt) distance of 17.6 in. Figure 5-8 shows the circumferential membrane stress magnitudes at four different meridional planes based on large displacement solution. Using a weighted average over the circumference of the model, the meridional extent was calculated as a 2 in. 5-4

RF # 00664 INDEX NO. 9-3, REV. 0 5.3 Code Evaluation of the Sandbed and Lower Sphere 5.3.1 Primary Stress Evaluation Tables 5-2a and 5-2b show the maximum values of primary stresses for the accident condition load combination V-1, and the Code allowable values for the small and large displacement solutions, respectively. In the primary membrane stress category, the calculated stress intensities for the sandbed region are based on the average values. The peak value of the circumferential membrane stress in the sandbed region was compared with the local primary membrane stress limits. As expected, a comparison of Tables 5-2a and 5-2b shows that the calculated stress magnitudes using the large displacement option are in general slightly lower than those obtained using the small displacement option. The differences in the stresses are larger in the sandbed region where the radial displacements are larger. The calculated primary stress magnitudes in the sandbed region and lower sphere meet the Code stress limits. 5.3.2 Extent of Local Primary Membrane Stress Paragraph NE-3213.10 of the Code states that a stresses region may be considered local if the distance over which the membrane stress intensity exceeds 1.1 S. does not extend in the meridional direction more than 1.01(Rt), which is - 17.6 Inches. When the small displacement solution is used (5.2.1), the membrane circumferential stress magnitude in the sandbed region exceeds 1.1 Smc over a meridional distance of - 26 inches at the plane of symmetry between the vent lines. However, this distance was found to be 14 Inches using a weighted average considering other meridionals. Furthermore, this distance of 26 inches at the plane of symmetry between the vent lines was reduced to - 11 inches when the large displacement solution was used in which the stiffness matrix is updated based on the deformed shape. Therefore, it is concluded that 5-5

UNSEX' 283-3, REV. 0 the circumferential stress in the sandbed region meets the meridional extent criterion of the Code Paragraph NE-3213.10. 5.3.3 Primary Plus Secondary Stress Evaluation Only two load cases result in significant secondary stresses in the shell. The first is the temperature gradient (accident condition V-1) which produces secondary stresses in the sandbed and lower sphere. The second is the post-accident condition which produces discontinuity bending moments in the shell at the bottom of the sandbed. The post-accident load combination case VI controls. Tables 5-3a and 5-3b show the calculated values of primary plus secondary stresses and a comparison with the allowable values for small and large displacement solutions, respectively. All of the calculated primary plus secondary stress values are within the Code allowable values. 5-6

RF 3 00864 ?NOEX NO. 9-3, REV. 0 TABLE 5-la Comparison of Calculated Stresses to ( Nominal Drywell Wall Thicknesses Limiting Load Combination - Code Allowable Values Above Lower Sphere) V-1 Drywell Region Stress Categ. Calc. Stress Magnitude, Max. (psi) Allowable Stress (psi) Cylinder (t*0.640 in.) Prim. Memb. 19200 19300 Prim. Hemb. + Sending 20280 29000 Knuckle (t-2.625 in.) Prim. Memb. 18430 19300 Prim. Memb. + Bending 20620 29000 Upper Sphere (ts0.722 in.) Prim. Memb. 19090 19300 Prim. Memb. + Bending 26350 29000 Middle Sphere (t-0.770 in.) Prim. Memb. 18460 19300 Prim. Memb. + Bonding 23110 29000 5.7

N gEX 8 9-3, REV. o TABLE 5-1b Comparison of Calculated Stresses to Code Allowable Values ( 95% Projected Drywell Wall Thicknesses Above Lower Sphere) Limiting Load Combination - V-1 Drywell Region Stress Categ. Cale. Stress Magnitude, Max. (psi) All owabl e Stress (psi) Cylinder (t-0.619 in.) Upper Sphere (tS0.677 in.) Middle Sphere (t-0.723 in.) Prim. Memb. Prim. Memb. + Bending Prim. Memb. Prim. Memb. + Bending Prim. Memb. Prim. Memb. + Bending 19850 21200 20970 29000 20360 21200 28100 29000 19660 21200 24610 29000 5-8

'NaEX 28.g9-3, REV. 0 TABLE 5-2a Comparison of Calculated Primary Stresses to Code Allowable Values ( Small Displacement; Lower Sphere and Sandbed ) Limiting Load Combination - V-1 Orywell Region Stress Categ. Calc. Stress Magnitude, Max. (psi) Allowable Stress (psi) Lower Sphere (t-1.154 in.) Prim. Hemb. Local Prim. Memb. Prim. Memb. + Bending 13800 17690 17800 21200 29000 29000 Sandbed (t-0.736 in.) Prim. Memb. Local Prim. Memb. Prim. Memb. + Bending 17430 22970 24950 21200 29000 29000 5.9

UNEX N8.99-3, REV. 0 TABLE 5-2b Comparison of Calculated Primary Stresses to Code Allowable Values ( Large Displacement; Lower Sphere and Sandbed ) Limiting Load Combination - V-1 Drywell Region Stress Categ. Calc. Stress Magnitude, Max. (psi) Allowable Stress (psi) Lower Sphere (t-1.154 in.) Prim. Memb. Local Prim. Memb. Prim. Memb. + Bending 13940 17530 17640 21200 29000 29000 Sandbed (t-0.736 in.) Prim. Memb. Local Prim. Memb. Prim. Memb. + Bending 16540 21540 23130 21200 29000 29000 5-10

YNuEX N8. l-3, REV. 0 TABLE 5-3a Comparison of Calculated Primary Plus Secondary Stresses to Code Allowable Values ( Small Displacement -Lower Sphere and Sandbed ) Drywell Region Stress Categ. Calc. Stress Magnitude, Max. (psi) Allowable Stress (psi) Lower Sphere (t-1.154 in.) Sandbed Region (t-0.736 in.) Prim. + Sec. 29020 (Acc. Load Cond. V-1) Prim. + Sec. 30280 (Post-Acc. Load Cond. VI) Prim. + Sec. 38420 (Acc. Load Cond. V-I) Prim. + Sec. 67020 (Post-Acc. Load Cond. VI) 52500 70000 52500 70000 .5-11

?AF # 00664 NDEX NO. 9-3, REV. 0 TABLE 5-3b Comparison of Calculated Primary Plus Secondary Stresses to Code Allowable Values ( Large Displacement -Lower Sphere and Sandbed ) Drywell Region Stress Categ. Calc. Stress Magnitude, Max. (psi) Al l owabl e Stress (psi) Lower Sphere (t-1.154 in.) Sandbed Region (te0.736 in.) Prim. + Sec. 28860 (Acc. Load Cond. V-1) Prim. + Sec. 30280 (Post-Acc. Load Cond. VI) Prim. + See. 36600 (Acc. Load Cond. V-1) Prim. + Sec. 67020 (Post-Acc. Load Cond. VI) 52500 70000 52500 70000 5-12

(SCurnoIJDferl-l'.ul Ie ItIbrccici.ei .(.oidi I. I I 'lli Smaoll D~ispslucemot Accl d(;mie~l ("ondilions V I i-In .X I-LnnaR C 0 a, U-to E 1I us v 30 281 26 24 22 20 18 16 14 12 10 8 6 4 2 0 0

10) 20(0 Meridional Oistunce fhoin S3 IBo(turui (in) 31)0 40)

P-0C C>O o C Ch %as' 4, ACCDNr4A.DRW Figure 5-1 Circumferential Stresses for Accident Condition V-1 in 'With Sand' and 'Without Sand' Cases - Small Displacement

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(.t u, -V M La V) 0 C, 0a, 30 28 26 24 22 20 18 16 14 12 10 8 6 4 2 I-zro ca-n M 0', 0% Lop IC <3 0 100 200 0(1r Meridional Disloiice Irom SO tHoUusn (in) Figure 5-2 Plot of Accident Condition V-I tHeridional Stresses for 'Without Sand' Case - Small Displacement OYCR 1 M2.DRW

ANSYS 4.4A J*N 17 1991 L3: 54:4? STEP-1I 3X I( UG) K IDLEY Om -29544 SM -0.115997 SCN =32979 XU =1 Y~U =-G.6 wdST=22*

• 96 air =4B.S9 aVr

=-21.071 1fflcz~-!is-1 33 CINOID HI DDEN S9.1159 1021" 12761 = 13313 I (en VSZw C X iC Figure 5-3 Circumferential Membrane Stress Olstribution Using Small Displacement Option.

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I N. 9-3, REV. 0 I i -0 re -_-

e.

-eorv: m:! .eory-I, 3 _ in I. Figure 5-5 Beam With Transverse Plus Axial Loading 5-17

A?5V3 4.4A JaM AS ibli 9: 48: 32. PLOT MO, 2 EgRiSS ITERJ.3 5X CAUG) WIDDLI .CLIS Cs SM4X =21344

• XE =40.755

]t? =-24.621 tar =250.228 A t4=-99 CZNYROID WIDu31 .952 =7184 co: 9577 5167 L19139 OVSTER CAXZX MOWSflL Af4WLWSIS - OVC~tZL SMA. ACCIDINT>l Figure 5-6 Circumferential Membrane Stress Distribution Using Larqe Displacement Dotion.

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0EX# N6 9-3, REV. 0

5.

SUMMARY

AND CONCLUSIONS This report is a supplementary report to the Code stress report (Reference 1-2) of record and addresses aspects of Code compliance as they relate to the local wall thinning observed and the removal of sand from the sandbed region in the Oyster Creek drywell. The loads and load combinations used in the analysis were based on the previous drywell stress analyses and the GPUN technical specification (Reference 1-4). In developing the allowable stress limits guidance was taken from Subsection NE of Section III, ASME Code where the Code of record, Section VIII and Code Case 1272N-5, is not explicit. The stress analysis first considered a model in which everywhere as-designed thicknesses were used except in the sandbed region where the thickness was assumed as 0.736 inch. This served as a basis for evaluating the stresses for the 95% confidence projected thicknesses to 14R. The highest stresses were determined to be from the Case V-1 and VI load combinations in all the different regions of the drywell. It was shown that the primary and secondary stresses are within the allowable limits for both conditions (as-designed thicknesses and 95% projected 14R thicknesses). At the plane of symmetry between the vent lines, the meridional extent of the circumferential membrane stress above I~smc, was in excess of I.0/(Rt). However, using a weighted average considering other meridional planes, this distance was less than 1.O/(Rt). Furthermore, a large displacement solution indicated the extent at the symmetry plane to be also less than 1.0/(Rt). This clearly satisfied the Code criterion for the extent of local primary membrane stress. It is concluded that the Oyster Creek drywell shell will continue to meet the Code of record requirements at least up to 14R with the sand removed from the sandbed region. The analysis for buckling capability of the drywell shell without sand is contained in a companion GE report (Reference 1-5). 6-1

9NE Rd 9-3, REV. 0 APPENDIX A DETAILED RESULTS FOR AXISYMMETRIC MODEL TEMPERATURE STRESS ANALYSIS A-l

N"EX 8.-3, REV. 0 This appendix presents a summary of the finite element analysis results for the temperature stress case (Load Case No. 8 in Table 3-1). The stresses reported in these tables are the nodal stresses. Since there are three nodes across the thickness of the drywell shell (e.g., see Figure 3-3), the stress at the center node is essentially a membrane stress. The difference between the stress at an inner or the outer node and the middle node is indicative of the bending stress at that section. In each of the stress tables, the second and third columns from the left show the radial and vertical coordinates of the center nodes. Four stress components (three normal stresses and one shear stress) are listed for each of the inner, middle and the outer nodes. Table Z-1 shows the wall thicknesses in the various regions of the drywell. To help assess the maximum stress levels, the range of node numbers associated with each wall thickness are given below: Drywell Region Node Number Range Sandbed Region 1 through 96 Lower Spherical Region 100 through 237 except Sandbed Area Middle Spherical Region 241 through 603 Upper Spherical Region 604 through 876 Knuckle 880 through 942 Cylindrical Region 946 through 1449 A-2

ter CreA

a. Data for Thensal Stress at 210 seconds - lb Sand Outside Nodes Radial Terldtonra Hoop Middle Nodes Radial herldional Hop ode X

Y Theta { Imbt) OIrchl (degrtesJ Node Si ST lot ) WPsU) St SXY (pisi 1 ps1 Node Sx ST

fpsi, (psi)

St Sxr (psi I (ps, I Inside Nodes Radial Herldlonal Ioop Node SX SY St (psI) (Psi) fps)) SXY (pi ) 2 241.08 106.93 5 246.51 186.10 5 2SO.28 109.25 It 251.81 110.4* 14 253.45 t1I.66 17 255.03 112." t0 2SS.61 114.06 21 2s8.61 115.25 21 2S9.14 li3.5 29 251.30 111.13 32 262.5s 111.91 35 254.39 12D.21 3S 265.53 121.44 41 267.4? 122.72 44 269.00 123.99 41 210.S2 125.26 50 212.03 125.54 53 273.S4 1227.3 56 2ts.05 129.13 S5 ?16.54 130.43 U 218.04 131.74 45 211.s4 133.01 6a 2t1.03 134.41 it 282.52 13S.75 14 284.01 137.11 11 I2S.48 135.47 to 288.53 139.63 U 26.42 141.21 66 28.88 142.52 59 211.33 143." n 292.11 145.31 9S 294.21 146.77 is 214.65 147.04 101 295.06 141.31 104 296.51 148.12 go? 291.92 155.14 36.00 36.21 36.54 35.81 37.86 3T.35 37.52 31.09 38.15 36.43 38.10 38.98 39.25 39.S2 39.7n 40.06 40.33 40.60 40.11 41.14 41.4t 41.5 41.16

• 2.23 42.50

42. i 43.0$

43.33 43.60 43.87 44.15 44.42 44.49 44.56 44.63 4S.10 1 22Z.09 1034.01 380.65 4 -62.03 86s.03 211.SG 7 10.75 072.26 231.3s 10 -4.22 143.82 133.30 13 1.19 801.16 23.75 19 -0.58 451.11 -106.11 II -6.61 28U.40 -257.49 2t -1.40 88.-6 -42i.52 2S

• I.98

-1S2.71 -603.70 20 -2.75 -431.10 -155.23 31 -3.42 -100.06 -393.71 34 -4.35 -1162.93 -t123.91 31 -4.33 -1652.56 -138l.14 40 -5.92 -2292.89 -156S.9S 43 -5.Z0 -2602.51 -2115.52 46 -1.31 -341a.S9 *-sZ1.os 49 -5.85 -4206.9t .1665.1 52 -8.24 -4900.25 -1U26.11 5S -6.16 -5163.3J -1657.09 56 -3.10 -6531.01 -139t.1t 61 -3.20 -1230.80 -93S.65 64 -6.13 -718.12 -211.20 6G 3.67 -6200.13 U43.46 70 -2.54 -8313.58 1328.22 13 15.J1 -5015.75 3341.44 is 3.35 -7240.83 5179.95 39 33.01 -5780.66 7401.65 62 12.42 -3599.93 "913.31 8S SC.15 -421.51 12302.4? 68 52.43 3152.20 12454.SS 91 -231.59 8521.2a 6323.4S -9.14 -8.96 -8.94

• S.02

-9.43 -10.20 -13.37 -12.97 .14.95 -61.46 -20.13 -23.31 -26.24 -29.41 -32.09 -34.SS -35.67 -31.14 -34.03 -30.95 -23.08 -14.20 2.54 19.01 46.55 13.93 120.94 1S3.85 221.60 245.39 30.60 2 -6.29 5 3.34 a 0.65 13 1.10 14 0.81 I? 0.64 20 0.46 23 0.t 26 -0.04 29 -0.46 32 -0.11 35 -1.39 35 -1.14 41 -2.69 44 -3.01 47 -4.41 SO -4.49 S3 -6.63 56 -5.51 S -9.34 62 -1.04 tS -13.-0 Go -6.13 II -16.73 74 -2.78 17 -21.43 30 7.18 03 -27.11 86 23.22 eS -16.56 92 -20.50 95 66.02 96 -131.17 101 -391.9S 104 21.21 101 17.57 -8.65 -3.20 -S.90 -8.65 -12.63

• 6.6S

-8.80 -4.6s -7.00 -9.26 -106.88 -1.36 -10.11 -152.05 -8.02 -1.5S -211.83 -9.01 -13.55 -372.31 -10.S7 -16.11 -479.58 -12.56 -11.39 -558.S -15.06 -23.16 -693.60 -18.06 -21.66 -781.92 -21.52 -32.38 -863.45 -25.39 -37.71 -910.88 -29.54 -42.79 -918.51 -33.32 -45.36 -814.40 -35.8Z -52.48 -163.67 -41.8s -S1.05 -511.08 -44.97 -53.39 -275." -46.96 -60.51 130..0 -41.29 -56.11 115.68 -45.39 -53.50 1374.72 -40.55 -39.18 22S4.64 -31.93 -31.34 3322.91 -18.67 -1.40 4561.1Z 0.07 13.13 6052.16 25.40 6.73 1701.71 57.91 87.33 9530.91 "9.0s 1W2.5S 11429.45 145.61 M65." 12886.15 206.92 394.74 11835.16 266.24 355.34 4256.37 305.94 -12.39 -11315.26 -261.26 56t.36 -1811.81 -127.53 914.14 -104".76 -116.11 165.6i -9013.26 105.53 37.93 -1569.90 55.3? 3 -238.43 -1053.51 -331.81 -4.21 6 68.72 -1004.2 -291.76 -4.42 9 -11.1l -091.83

• 331.27

-4.16 12 6.59 -764.10 -347.65 -5.42 is 0.81 -631.41 -388.54 -6.38 to 2.00 -484.13 -435.98 -7.76 21 1.63 -316.12 -481.59 6S 24 1.81 -119.42 -535.60 -82.89 27 1.31 114.41 -513.30 -S.16 30 1.18 393.SS -591.68 -16.81 33 1.13 126.71 -551.34 -23.13 36 1.30 1120.45 -Z32.04 -27.94 39 1.83 1561.18 -431.99 -33.34 42 0.03 21)1.51 -269.35 -38.90 45 -0.52 2113.00 -30.40 -44.51 5

• 2.42 3380.30 297.18

-50.23 51 -3.22 4106.06 128.07 -55.56 54 -6.44 4812.80 1272.65 -59.2S 5s -7.28 5659.14 1944.14 -62.32 60 -12.29 6428.10 2743.54 -61.83 63 -12.28 7146.60 3694.30 -60.29 64 -21.46 1746.21 4154.93 -52.12 69 -19.12 3tll.78 6014.92 -42.49 72 -33.16 8306.43 7349.11 -21.43 15 -24.45 8091.66 5779.96 -0.36 15 -481.22 7350.IZ 16230.17 39.13 St -25.36 6026.09 11671.S1 15.06 64 -67.75 336S.39 1291t.21 142.37 87 -10.61 669.81 lpi3?j 191.64 90 -64.t8 -3247.12 11144.07 281.6 93 91.71 -1947.40 2150.06 301.36 96 28.OO -9155.70 -13916.42 -382.15 99 200.69 -6526.92 -13138.3? -554.42 102 -253.59 -7181.11 -12462.12 -293.51 OS 28.76 -6180.1 -10663.11 109.91 IOa i., i2 -34Gu.;! ia5.32 94 703.92 11900.93 -3125.28 1694.61 91 1072.89 80o6.70 -621.51 2643.90 108 31.3$ 577?.4J -9376.01 106.71 103 28.09 6670.21 -417.07 103. 63 106 -S.4S 7802.93 .5629.43 32.0Z PA I PD1T11l.WK0 23-0ct -90

star Cresk Raw Data for Tharil Stress.t 210 secontd - No Sand Outside NOd& Radial Irtdiomal Maop Sl ST S2 (pt) (p1sO (poll 3idile NO&s Radial erldional Hoop cY M_.A Co eV Radial Inside Ad&* Radal Nrldional Hoop NId X I theta NCId lncl (mtob) 1degr"s) (pit z ahT H l sr Sz sx5 (psi) 4,ai1 PIII (Pull (psij (Pal) (psliI tP1) IN0 29.33 ISI.56 4S.37 113 300.74 15."

• 45.55 Ill 302.23 154.42 45.9Z l1l 303.5t 5S.81 45.39 122 304.91 1ST.32 46.47 12s 306.21

£56.t? 44.14 12t 30W1.S 160.23 41.14 131 309.01 161.70 47.21 134 312.3S 1I5.31 4.96 131 315.65 19.06 40.64 1t4 311.91 132.61 4".31 143 32U.1? 3.6 41.29 146 32S.25 160.40 50.16 149 326.40 104.25 5.34 152 331.41 19".4 52.01 35S 334.51 12.01 52G.6 156 331.49 136.03 53.35 161 340.00 1l2.S 53.94 164 342.46 202.8 S4.S2 161 344.93 206.36 SS.0 110 34.34 209.85 5S.68 173 31.71 213.36 56.25 116 352.05 215.90 56.U3 I1t 354.35 210.44 S7.41 102 350.6? 224.03 S1.99 18S 3S6.1S, 227.5 S$.S7 166 $61.04 231.21 S9.14 I1N 363.20 234.14 59.12 194 365.3? 231.61 60.30 191 361.41 242.31 40.8e to0 36A.4S 246.03 S1.4S 203 311.46 249.11 62.03 206 373.43 253.52 62.51 209 315.38 257.30 53.19 212 3?7.26 261.09 13.77 215 379.11 264.91 54.34 19 -26.01 6227.4S -4115.90 I£.4? 112 2.S4 3 -2157.41 -l.71 IIS -11.n7 0342.41 -119S.31 -33.27 1ji a." 0 S.45 -719.30 -51.41 121 -4.62 1619.44 -47.92 -60.21 124 6.32 1112.11t S3.22 -60.36 12 -26.25 6581.51 917,44 -71.28 130

• 0.14 5741.13 12tS.U4

-62.07 133 23.07 4430.63 1137.09 -83.20 134 30.11 3020.16 1740.12 -16.92 13t 21.20 1551.50 1528.91 -49.24 142 36.22 966.99 1213.11 -33.25 14S 10.49 3S1.91 162.73 -20.01 140 6.3S -34.13 W8.AY -10.24 15£ 3.11 -245.04 341.57 -3.43 IS4 0.74 -332.09 170.21 0.16 1S? -0.20 -339.11 49.71 2.65 I U -0.54 -301.07 -14.99 3.S6 161 -0.97 -2S1.44 -S2.87 3.74 564 -0.14 -*0r.45 -70.$I 3.MY 16" -0.91 -141.60 -14.54 2.99 li2 -0.11 -103.09 -68.61 2.40 75 -0.63 -64.5s -59.99 1.81 13 .0.46 -35.54 -46.29 1.28 111 -0.33 -14.29 -36.S4 0.63 1I4 -0.21 -0.01 -25.87 0.46 la -0.13 8.6M -16.90 0.2? 190 -0.06 13.01 -9.82 0.54 193 -0.01 14.49 -4.S6 -0.01 135 0.02 11.95 -0.9 -0.13 19 0.03 12.25 1.34 -0.15 202 0.04 10.08 2.60 -0.15 205 0.04 7.76 3.13 -0.14 206 .04 S..5 3.15 -0.1? 212 0.03 3.81 2.97 -0.09 214 0.0? 2.35 2.44 -0.07 110 -13.52 56.61 -6216.54 33.43 113 13.59 1.6 -5003.4? 6.92 116 -3.5S -2.16 -3923.80 -14.30 S19 6.37 -26.78 -2916.20 -30.559 12? 0.31 -34.03 -2157.SS -42.72 12S 6.23 -43.81 -1458.33 -51.20 128 -5.6s -47.12 -511.11 -56.7? 131 32.25 49.25 -311.53 -59.80 134 32.17 -44.S6 462.63 -55.87 13? IS.15 -39.0 6SS5.0S -SO.73 140 11.03 -30.15 SS6.5I -39.76 143 S.40 -22.34 905.30 -23.61 141 1.43 -14.16 756.51 -16.28 141 -0.3S -3.94 517.73 -10.92 152 -1.51 -4.54 40S.00 -5.11 £15 -2.17 -1.11 257.42 -1.31 156 -1.82 0.14 142.9) 1.00 161 -1.06 0.39 11.33 2.04 164 -1.11 1.31 20.11 2.41 Is? -0.01 1-44 -11.90 2.53 170 -0.53 1.35 -31.52 2.29 113 -0.42 1.13 -3.15 1.93 116 -0.27 0.51 -40.15 1.51 371 -0.14 0.67 -37.05 1.13 1i2 -0.06 0.4* -31.60 0.19 18s 0.00 0.31 -25.24 0.50 In 0.04 0.18 -1.95 0.25 1II 0.06 0.09 -13.30 0.12 194 0.06 0.03 -5.59 0.01 191 0.06 -0.01 -4.9D -0.06 200 0.05 -0.03 -2.38 -0.09 203 0.04 -0.04 -0.32 -0.13 206 0.03 -0.04 0.85 -0.30 209 0.02 -0.04 1.46 -0.09 21? 0.02 -0.03 1.13 -0.07 21S5 0.41 -0.02 1.72 -0.06 lil 114 117 120 l23 326 129 132 13S 138 £43 144 147 M~o 153 151 158 162 16S 168 all 114 Ill too 163 186 t92 195 196 201 204 20? 210 213 2lS 1.32 -U219.11 -83S2.30 52.51 27.62 -8414.St -1216.38 35.53 7.71 -6401.38 -4111.t5 1.78 4.-6. -8361.44 -5160.S1 -6. 4 8.31 -7151.16 -4230.01 -2t.39 6.39 -1265.16 -3462.47 -31.40 8.96 -6)14.93 -2132.34 -39.J2 26.44 -5058.0 -2005.5S -35.38 44.14 -4536.62 -616.54 -32.85 3.01 -3109.95 -43.12 -33.40 1.19 -£920.23 402.13 -25.56 495S l1lS. 596.0t -23.60 -6. U -383.06 629.62 -17.34 -J.01 17.20 SS.64 -11.0 -1.31 237.50 462. 1 1 -6,36 -5.35 330.00 34S.14 -3.49 -3.47 341.23 231.12 -0.19 -3.69 310.29 1s5.12 3.43 -1.40 265113 94.61S l.S -0.7 206.13 41.13 1.49 -0.43 352.3 13.7z 3.54 -0.11 IDS.1? -1.31 1.41 0.06 61.00 -20.21 1.22 0.17 V3.04 -25.75 0.96 0.21 15.29 -26.64 0.13 0.22 0.63 -24.61 0.52 0.20 -6.32 -22.01 0.33 0.11 -12.14 -16.72 0.39 0.14 -14.49 -12.62 0.09 0.o0 -14.02 -8.86 5.01 0.01 -12.39 -5.12 -0.03 0.05 -30.20 -3.26 -0.06 0.03 -r.13 -3.41 -0.56 0.01 -5.14 -0.20 -0.06 0.00 -3.39 0.51 -0.06 0.00 -2.41 0.9% -0.05 Pege 2 NQ8S1210.ylK1 23 Oct-90

ter Creek Raw Oaks for Therwal Stress at 210 seconds - No SAnd Outside Nodes 11Idle Nodes Inside Jodes Radial PerIdional Hoop Radial lerldlonal

Dop, Radial Meridlonal Ibop ald x

Y thet& Node Sl SY St SXY Node SX SY St S Node SX Sy St Sx (inch) limb) (degrees) (pat) fpsOl (psi) (psi) (psi) (pFil (psi) (psi) (psi) (Pol (psI) (psi) 216 380.93 26.74 64.92 217 0.02 221 382.71 272.59 65.54 220 0.01 224 313.49 214.32 65.16 213 0.01 221 384.21 216.04 66.02 226 0.00 230 36S.03 211.71 66.21 229 0.00 233 us.1i 219.st 66.53 232 0.90 236 315.54 281.2S 66.13 US 0.00 23 381.1s 262.00 66.20 238 -0.01 242 M8.1 2ez.74 61.00 241 -0.02 24$ 361.40 283.16 6.IS 244 0.00 246 381.32 Z".t 61.30 24? 0.00 251 388.24 28S.SJ 61.4S 250 0.00 2S4 366.61 216.80 67.60 253 0.00 251 321.06s 267.6 61.1s 255 0.00 260 388.50. 28.84 61.10 259 0.00 263 3".91 US." Ws.05 22 0.00 266 390.32 290.8 68.20 265 0.00 269 390.13 291.90 68.3S 26J 0.00 212 39J.13 222.53 W.50 211 0.00 21S 39t.23 295.31 1.93 214 0.00 216 393.40 298.62 65.36 21 0.00 26l 3S.50 301.77 69.79 rao 0.00 264 395.5 304.14 70.22 233 0.00 2Jt 396.54 301.1 10.15 286 0.00 290 391.1 31O.59 73.06 289 0.00 2n3 393.48 313.68 11.53 292 0.00 291 3".67 316.61 71.94 29S 0.00 219 400.64 313.68 72.37 298 0.00 302 401.58 322.U5 72.30 301 0.00 30S 402.51 325.11 13.23 304 0.00 306 403.41 326.73 13.66 301 0.00 311 404.78 331.;5 14.09 310 0.00 314 405.13 334.80 74.52 3t3 0.00 317 405.31 331.84 74.S 316 0.00 320 406.71 340.15 75.30 319 0.00 323 407.S6 343.95 15.81 322 0.00 1.27 0.59 0.2S 0.07 -0.01 -0.18 -0.22 -0.35 -0.52 -0.60 -0.31 -0.12 -0.62 -0.11 -0.60 -0.5S -0.56 -0.53 -O.49 -0.42 -0.33 -0.24 -e. 17 -0.11 -0.06 -0.03 0.00 0.01 0.02 0.03 0.03 0.02 0.02 0.02 0.01 0.01 1.9S -0.05 218 0.00 1.41 -0.03 221 0.00 1.27 -0.02 224 0.00 1.08 -0.02 221 0.00 0.91 -0.01 230 0.00 0.1S -0.01 233 0.00 016l 0.02 236 0.01 O.S 0.06 219 0.01 0.41 0.03 242 -0.01 0.32 0.00 24S 0.00 0.25 0.00 246 0.00 0.19 0.00 251 0.00 0.14 0.00 254 0.00 0.09 0.00 257 0.00 0.05 0.00 260 0.00 0.01 0.00 263 0.00 -0.02 0.00 266 0.00 -0.0s 0.00 269 0.00 -0.07 0.00 272 0.00 -0.13 0.01 215 0.00 -0.13 0.00 218 0.00 -0.13 0.00 281 0.00 -0.12 0.00 ZU4 6.00 -0.10 0.00 ?S? 0.00 -0.06 0.00 290 0.00 -0.01 0.00 293 0.00 -O.OS 0.00 296 0.00 -0.03 0.00 2n 0.00 0.02 0.00 302 0.00 -0.01 0.00 305 0.00 0.00 0.00 30S 0.00 0.00 0.00 311 0.00 0.w0 0.00 314 0.00 0.01 0.00 311 0.00 0.01 0.00 320 0.00 0.01 0.00 323 0.00 -0.0o -0.01 -0.03 -0.Ot 0.00 0.00 -0.02 -6.01 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 1.55 -0.04 219 -0.01 1.30 -0.01 222 0.00 1.11 -0.02 225 -0.01 1.04 -0.02 228 0.00 0.91 -0.0! 231 0.00 0.3 -0.01 234 0.00 0."6 -0.01 237 0.01 0.51 -0.03 240 0.00 0.56 -0.02 243 -0.01 0.49 0.00 246 0.00 0.43 0.00 249 0.00 0.37 0.00 252 0.00 0.31 0.00 255 0.00 0.21 0.00 258 0.00 0.22 0.00 261 0.00 0.18 0.00 264 0.00 0.14 0.00 267 4.00 0.11 0.00 2J0 0.00

0. 0 0.00 213 0.00 0.03 0.00 275 0.00

-0.03 0.00 29 0.00 -0.06 0.00 282 0.00 -0.01 0.00 285 0.00 -0.01 0.00 28 0.00 -0.01 0.00 291 0.00 -0.06 0.00 294 0.00 -0. 05 0.00 297 0.00 -0.04 0.00 300 0.00 -0.03 0.00 303 0.00 -0.02 0.00 306 0.00 -0.01 0.00 3091 0.00 -0.01 0.00 312 0.00 0.00 0.00 315 0.00 0.00 0.00 31 0.00 0.00 0.00

32) 0.00 0.00 0.00 324 0.00

-1.31 -0.61 -0.27 -0.06 0.05 0.15 0.25 0.31 0.50 0.S9 0.61 0.62 0.62 0.61 0.60 0.58 0.S6 0.53 0.49 0.42 0.33 0.25 0.17 0.11 0.01 0.03 0.00 -0.01 -0.02 -0.03 -0.03 -0.12 -0.02 -0.02 -0.01 -0.01 1.15 -0.04 1.10 -Q.02 1.01 -0.02 1.00 -0.02 0.91 -0.01 0.32 -0.01 0.12 -0.02 0.11 -0.03 0.70 -0.02 0.66 0.00 0.50 0.00 0.5s 0.00 0.49 0.00 0.44 0.w 0.39 0.00 0.35 0.00 0.30 0.00 0.26 0.00 0.z2 0.00 0.13 0.00 0.06 0.00 0.01 0.00 -0.02 0.00 -0.04 0.00 -o.0s 0.00 -0.05 0.00 -0.0 0.00 -0.04 0.00 -0.03 0.00 -0.03 0.00 -0.02 0.00 -0.01 O.0 -0.01 0.00 -0.01 O.0 0.0 0.00 0.00 0.00 Page 3 3WDST?)0.VKI 23-0ct-90

ter Creek Raw Data for ThermAl Stress at 210 seconds - o Snd Oustside Nodes MIiddle Nodes Jnside x03es T adial Ikrldioall Hoop Radial Nerdilonal Heop Radial hetidionl hiop bd d r Te todo Si sY 5z sXv Nd si SY 52K SY St SXY jinch) (inch) (degrees) (psi) (pall fps)) fpsi) (psi) (psi) fpsi) (psi) (psi) (psi) (psi) (poll 326 406.32 347.01 76.24 325 o.00 329 409.06 350.068 6.6? 326 0.00 332 409.77 3S3.15 17.10 331 0.00 335 410.40 356.23 77.53 334 0.00 330 411.14 359.31 17.96 331 0.00 341 419.18 362.40 18.35 340 o.00 344 4)2.41 365.49 7S.02 343 0.00 341 413.01 3U.59 79.J5 34 0.00 3S0 413.58 311.69 79.66 349 0.00 353 414.14 314.0 84.11 3S2 0.00 3S5 414.67 371.91 60.54 355 O.OD 359 415.31 381.02 80.91 35t 0.00 362 415.66 384.14 61.40 341 0.00 365 419.12 381.25 89.83 354 0.00 3U 411.SS 390.33 82.n6 356 0.00 311 416.91 393.51 82.88 3S0 0.00 3J4 417.36 396.64 83.12 373 0.00 317 411.12 311.8 3.5S 311 0.00 360 418.07 402.91 63.98 319 0.00 383 418.39 406.OS 84.41 352 0.00 385 416.68 409.19 s4.8l US 0.00 38 418.95 412.33 8S.21 388 0.00 392 419.20 415.48 3S.10 391 0.00 3S5 419.43 418.63 86.13 194 4.00 3" 411.63 421.6S 86.S6 3II 0.00 401 419.51 424.93 4." 400 0.00 404 413.90 128.088 7.42 401 0.00 407 420.09 431.23 117.5 406 0.00 410 420.20 434.38 66.28 409 0.00 413 410.28 437.S4 66.11 412 0.00 416 420.34 440.69 89.14 415 0.00 413 46.31 4W3.8S 89.57 418 0.00 42? 420.39 447.00 90.00 421 0.00 42S 420.37 4S5.15 90.43 424 o.00 428 420.14 453.31 SO.8 421 0.00 431 420.23 456.46 St.29 430 0.00 0.01 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 o.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.Oj 0.00 326 0.00 0.0o Q.00 329 0.00 0.00 0.00 332 0.00 0.00 0.00 335 0.00 0.00 0.00 338 0.00 0.00 0.00 341 0.00 0.00 0.00 344 0.00 0.00 0.00 347 0.w0 0.o0 0.00 350 0.00 0.00 0.00 3U3 0.00 0.00 0.00 l56 0.00 0.00 0.00 359 0.00 0.0o 0.00 362 0.00 0.00 0.o0 365 0.00 0.00 0.00 366 0.00 0.00 0.00 311 0.00 0.00 o.00 374 0.00 0.00 0.00 377 0.00 0.00 0.00 380 0.00 0.00 0.00 38 0.00 0.00 0.00 3U6 0.00 0.00 0.00 39 0.00 0.00 0.00 392 0.00 0.00 0.00 39S 0.00 0.00 0.00 396 0.00 0.00 0.00 401 0.00 0.00 0.00 404 O.00 0.0o

0.

O 407 0.00 0.00 0.00 410 0.00 0.00 0 00 413 0.00 0.00 0.00 416 0.00 0.00 0.00 419 0.00 0.00 0.00 422 0.00 0.00 0.00 425 0.00 0.00 0.00 428 0.00 0.00 0.00 431 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.0e 0.00 0.00 0.01 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.0e 0.00 0.eo 0.00 0.60 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 o.00 0.00 0.00 327 0.00 0.00 0.00 330 0.00 0.00 0.00 333 0.00 0.00 0.00 336 0.00 0.0 0.00 339 0.00 0.00 0.00 34? O.D. 0.00 0.00 345 0.00 0.00 0.00 348 0.00 0.00 0.00 351 0.00 0.00 0.00 3S4 o.00 0.00 0.00 357 0.00 0.00 0.00 360 0.00 0.00 0.00 363 0.00 0.00 0.00 356 0.00 0.00 0.00 369 0.0o 0.00 0.00 37f 0.00 0.00 0.00 37S 0.00 0.00 0.00 318 0.00 0.00 0.00 381 0.00 0.00 0.00 384 0.00 0.00 0.00 381 D.00 0.00 0.00 390 0.00 0.00 0.00 393 0.00 0.00 0.00 395 0.00 0.00 0.00 399 0.00 0.00 0.00 402 0.00 0.00 0.00 405 0.00 0.00 0.00 406 0.00 0.00 0.00 411 0.00 0.00 0.00 414 0.00 0.00 0.00 417 0.00 0.00 0.00 420 0.00 0.00 0.00 423 0.00 0.00 0.00 426 0.00 0.00 0.00 4?9 0.00 0.00 0.00 432 0.00 -0.01 0.00 0.DO 0.00 0.00 O.DO 0.00 0.00 D.D0 0.0o 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.W0 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.10 O.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.D0 0.0o 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 O.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 D.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 c.00 0.0o 0.00 0.00 0.to ?age 4 NOWt10. 1X 23-Oct-90

Oyster Creek Raw Data for thermal Streu at 210 seconds - So Sand Owtsdde tiodes Wldcile Wades Nneidi Modes Radial Keridional Hoop Radlal Ikrldioalis Hoop Radial 1rtdioo1

1. tip Node I

y Thta Node SI ST Sz SXY Nods SI SI 52 UT Node 5X ST 57 SIr Potch) (inch) Idegrees) (psi) (psil (psi) (psI) (psi) (psil (psi) psi (pit Is (psi) (psi) (psi) 434 420.20 459.62 91.12 433 0.00 431 4*0.0t 492.fl 92.15 431 6.00 440 419.96 465.92 92.56 4U9 0.00 443 419.11 469.01 93.01 442 6.00 446 41l.6U 412.t2 53.44 445 0.06 449 413.43 41S.21 93.51 445

• .00 452 451.20 415.52 94.3 451 0.00 455 415.35 481.65 94.73 4S4 0.00 455 418.48 454.61 "5.I 45S O.oo 461 415.35 4A.S 95.55 460 5.00 464 416.01 411.00 6.or 453 0.00 461 411.12 494.22 96.4 466 0.00 410 411.34 431.36 S6.6 469 0.00 413 4146.31 5.4 61.31 41Z a.00 476 416.S5 503.62 21.74 475 0.00 471 4t6.12 505.14 6.11 416 4.00 482 41S.66 5W.56 18.60 481 0.00 415 415.17 512.91 39.03 454 0.00 488 414.57 516.09 9.46 481 0.00 491 414.14 191.20 99.81 490 0.00 494 413.5 522.3) 100.3?

41S 3 0.00 491 413.91 525.41 100.15 46 0.00 500 4)2.4) 521.51 01.ls 49 0.00 S53 411.5 s3)1.10 101.61 S52 0.00 50S 411.14 s34.69 162.04 sos 0.00 5§ 410.41 531.17 102.41 s6 0.00 Sit 401.17 545.8S 162.90 Sit 0.00 S1S 409.04 S43.92 103.33 III 0.00 518 408.32 546.96 130.6 Sly 0.00 s21 407.s6 55.05 204.519 520 0.00 524

• 06.11 SS3.11 104.6z2 23 0.00 521 405.91 556.16 105.05 375 0.00 530 40S.13 ss9.20 105.46 529 0.00 533 404.28 S.24 10ZS.1 S32 0.00 536 403.41 S5S.2?-

106.34 535 0.00 S39 402.51 M.ZS 105.11 S35 0.00 0.00 0.01 0.00 0.00 0.00 0.00 0.00 6.60 0.00 o." 0.00 0.00 0.00

0. 00 0.00 0.00 0.00 0.00 0.00 10.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 o.se 0.00 0.00 434 0.00 0.00 0.00 431 0.00 0.0a 0.00 440 0.00 0.00 0.00 443 o.00 0.00 0.00 446 0.00 0.00 0.00 449 0.00 p.00 o.oo 452 0.00 0.00 0.00 455 4.00 0.00 0.00 A45 0.00 5.00 0.00 461 8.00 0.00 0.00 464 0.00 0.00 0.00 461 0.00 0.00 0.00 410 0.00 0.00 0.01 413 0.00 0.00 0.00 418 0.00 0.00 0.00 419 0.00 0.00 0.00 4U2 0.00 0.00 0.00 4S8 0.00 0.00 0.00 488 0.00 0.00 0.o0 41 o.00 0.00 0.00 494 0.00 0.00 0.00 411 0.00 0.00 0.00 Soo 0.50 0.00 0.00 503 0.o0 0.00 0.00 5u 0.00 0.00 0.00 509 5.00 0.00 0.00 S12 D.W0 0.00 0.00 55 0.0 0.00 0.00 Ss 0.00 0.00 0.o0 S21 0.00 0.00 0.00 S24 0.00 0.00 0.00 52Y 0.00 0.04 0.00 530 0.00 0.00 0.00 531 0.00 0.00 0.00 536 0.00 0.00 u.00 9

i.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 6.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 5.00 0.60 0.00 0.00 5.00 0.00 0.00 0.00 435 0.00 0.00 0.00 435 0.00 0.00 0.00 441 0.0c 0.00 0.00 444 0.00 0.00 0.e0 441 0.00 0.00 0.00 450 0.00 0.00 0.00 4S3 0.00 0.00 0.00 456 0.00 0.00 0.00 459 0.00 0.00 0.00 462 0.00 0.00 0.00 46S 0.00 0.00 0.00 466 0.00 0.00 0.00 411 0.00 0.00 0.00 424 0.00 0.00 0.00 477 0.00 0.00 0.00 480 0.00 0.00 0.00 4w 3 0.00 0.00 0.00 486 0.00 0.00 0.00 48 0.00 0.00 0.00 49? 0.00 0.00 0.00 495 0.00 0.00 0.00 49e 0.00 0.00 0.00 sol 0.00 0.00 0.00 504 0.00 0.00 0.00 50 0.00 0.00 0.00 S10 0.00 0.00 0.00 513 0.00 0.00 0.00 Sig 0.00 0.00 0.0e Sig 0.00 0.00 0.00 522 0.00 0.00 0.00 525 0.e0 0.00 0.00 525 0.60 0.00 0.00 53t 0.00 0.00 0.00 534 0.00 0.00 0.00 S31 0.00 u.w0 u.w 540 G.Fi 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.06 0.00 5.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 o.0o 0.00 0.00 0.00

a. o 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00a 0.00 O.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.08 0.00 0.00 0.00 0.00 0.00 0.00 0.00

0. 0 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.05 0.00 0.00 O.W 0.Q0 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.01 Page S I"OS1210.vt1 23 0ct-U

Oyster Creek a. Daete for Iherwl Stress at 210 seconds - No Sahd 0u radial PA Mod X 1 Theta Nods SX (inch) (inch) (degrees) p~s]) ttlde Model Middle Nodes ridlonal Hoop Radtal Mlrldional Moop SY St SlY lode SX ST St Sly (psi) Pl (pol pl) (pull (PsI l (psi) (psi I Inside Nodes Radial erldional Hoop Nods 5X S6 St (psi) (pit) (psi) SXT (psi I 542 40163 S7L.3 101.20 541 0.04 S4S 1O1O." 514.32 107.63 SJJ 0s.c S46 39.67 S71.32 106.06 S47 0.00 551 38.568 5"6.3t 106.49 550 0.00 554 397.67 583.31 106.92 Su 0.00 SS7 396.64 568.29 109.35 55 0.00 560 395.58 59.26 109.75 559 0.oo 543 394.S0 S92.23 110.21 562 0.00 566 393.40 5 85.1 110.64 Su 0.00 Sa9 392.2a 59U.13 1t11.0 S 0.00 572 391.13 601.01 111.50 sit 0.00 5M5 3O.13 602.10 111.6S 574 0.00 575 390.32 603.12 111.60 5sl 0.00 S61 319.91 804.14 111.3S S 0.00 564 389.50 605.16 112.10 583 0.00 561 389.06 6B6.16 112.2S u8s 0.00 560 368.61 607.20 112.40 s5o 0.00 S93 316.24 606.21 112.55 S9u 0.0W 595 387.82 609.23 112.70 595 0.00 S9S 387.40 610.24 112.35 S8u 0.00 602 366.91 611.26 113.00 601 0.00 605 3 6.8 £11-.39 113.02 604 0.00 60S 3W6.11 613.20 113.29 WB? 0.00 51t 385.33 615.01 113.56 10i 0.00 614 314.54 616.51 113.83 613 0.00 817 383.J74 15.61 114.09 618 0.00 620 382.93 620.41 £14.36 619 0.00 623 3t2.11 622.20 114.63 622 0.00 626 381.29 623.9 114.90 625 0.00 629 380.45 U25.71 £15.11 62S 0.00 632 319.6t 627.SS 115.44 531 0.00 535 318.16 529.32 11S.70 634 0.00 638 311.91 631.09 115.91 611 0.00 641 311.04 632.61 116.24 640 0.00 644 316.16 634.63 116.51 543 0.00 641 315.25 636.39 l16.7s 646 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.06 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 1.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 542 0.00 0.00 0.00 545 0.00 0.00 0.00 S4 0.00 0.00 0.00 S5S 0.00 0.00 0.04 S51 0.00 0.00 0.00 55 0.00 q.oo 0.00 5"0 0.00 0.00 0.00 Su 0.00 0.00 0.00 566 0.00 0.00 0.00 S59 0.00 0.00 0.00 512 0.00 0.00 0.00 515 0.00 0.00 0.00 578 0.00 0.00 0.00 sol 0.00 0.00 0.00 Sl4 0.00 0.00 0.00 561 6.0o 0.00 0.00 590 0.00 0.00 0.00 553 0.00 0.00 0.00 586 0.00 0.00 o.oo 5ss 0.00 0.00 0.00 602 0.00 0.00 0.00 605 0.00 0.00 0.00 0 0.00 0.00 0.00 all 0.00 0.00 0.00 61M 0.00 0.00 0.00 all 0.00 0.00 0.00 620 0.00 0.00 0.00 623 0.00 0.00 0.00 626 0.00 0.00 0.00 629 0.00 0.00 0.00 632 0.00 0.00 0.00 63S 0.00 0.00 0.00 638 0.00 0.00 0.00 541 0.00 0.00 0.00 644 0.00 0.00 0.00 641 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 5.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 .D00 0.00 0.00 0.00 0.00 5.00 0.00 01.00 0.00 0.00 0.00 0.00 o.oo 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 5.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 541 0.00 546 0.00 S54 0.0o 552 0.00 555 C.0o s55 0.00 561 0.00 564 0.00 567 0.00 S6O 0.00 513 0.00 S15 0.00 579 0.00 562 0.00 5O5 0.00 Su 0.oe 591 0.00 W94 0.00 591 0.00 600 0.00 503 0.00 6" 0.00 609 0.00 612 0.00 G15 0.00 416 0.00 621 0.00 624 0.00 621 0.00 530 0.00 633 0.00 636 0.00 639

0. 00 642 0.00 645 0.00 646 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.OO 0.00 0.00 0.00 0.00 0.00 0.00 0.00 O.00 0.00 0.00 0.00 0.00 0.00

.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 O.O0 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 D.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.DO 0.00 0.00 0.00 0.00 6.00 0.00 0.00 0.00 6.00 0.00 0.00 0.00 0.00 0.00 0.00 D 00 0.00

0. Od 0.00 0.05 0.00 Page 6 N05T210 WK 23-0c t -90

1water Crase Raw DAM for Theail Stress at 210 secans - He Send awNoside s middle Nodes Inside Nogs Radial 1kridional Hoop AdWIl Ikredlaul Hloop Radial Merld11al Omp Made x Y Theta NW14 SI Si 12 5IT Nod 5S St S2T Node sx si %s SKY (Inch) (inch) (dwats) (psi) (pSI) (pall (psi) fpet) IPsi) (psi) (psi I (Pat) psi) (psi) (psi) 50 376.39 U5.14 117.05 U9 0.30 153 313.49

6.

117.31 G52 0.00 656 372-.5 641t.4 I.S8 555 0.00 859 111.6? 443.35 W1.8S gm 0.00 662 370.14 £45.12 118.12 "1 0.00 6s 6 .ea 846.65 I1IS.3S 34 0.00 go8 3U.91 6.58 118.55 es? 6.00 U11 318.30 646.7 116.66 U10 0.00 674 31W." 650.82 119.00 US 0.00 51? 355.48 5-.S) 119.33 U1 0.00 "a 345.30 454.9 119.l 4 673 0.00 583 354.11 J5.06 119.96 ? 5.00 S6U 362.21 455.13 1U0.31 615 0.00 689 351.70 661.19 1201.U3 53 0.00 M9t 360.4? 63J.2S 110.96 69I 0.00 62S 35M.24 665.30 121.29 694 0.00 gs 35W.91 167.34 121.61 a? 0.00 701 355.13 89.317 121.94 700

• 0.00 704 35S.45 511.40 522.96 703 0.00 701 354.11 6n3.42 1Z2.59 l06 0.00 1)0 352.58 U

7S.45 12W.M2 769 D.00 713 351.5 617.43 123.24 Fit 0.00 Ji1 350.26 619.43 1J3.57 71S 5.00 119 348.93 61.4t 123.69 75 0.00 722 341.59 33.40 124.22 I21 0.00 125 345.24 US.S1 124.55 724 1.00 729 344.36 407.34 124.57 i2t 0.00 731 343.50 119.3O 125.20 130 0.00 134 342.12 W91.25 12S.S2 733 0.00 73? 349. 2 393.39 I1S.Js Y3e 0.00 14o 33§.32 69S.13 125.18 731 0.00 143 331.90 691.05 125.50 14* 0.00 746 336.41 606.97 14.83 745 0.00 149 335.03 700.88 121.1s 143 0.00 7SS 332.12 704.66 127.61 154 0.00 0.00 0.00 0.04 0.00 0.00 0.00 0.0 0.00 0.00 0.00 0.00 0.00 0.00 0.0e 0.00 5.00 0.00 0.00 0.00 0.00 0.o0 0.oo 0.00 0.O0 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.0 0.00 0.00 0.00 650 0.00 0.00 0.00 5.00 653 0.00 0.00 0.00 0.00 556 0.00 0.00 0.00 0.00 659 0.00 0.00 0.00 0.00 662 0.05 0.00 0.00 0.00 655 0.00 0.00 Q.00 0.00 me8 0.00 0.00 0.00 0.00 671 0.00 0.00 O.00 0.00 674 0.00 8.00 0.00 6rr 0.00 0.00 0.00 680 o.00 O.00 0.00 683 0.00 0.o0 0.00 6G6 0.00 0.00 0.00 u9 0.00 0.06 0.06 692 0.00 0.00 0.00 69t 0.00 0.00 0.6o G" 0.00 0.00 0.00 701 0.00 o.00 0.00 14 0.00 0.00 0.00 701 0.00 0.00 0.00 710 0.00 0.00 0.00 113 0.00 0,00 0.00 lie o.00 0.00 0.D0 739 0.00 0.00 0.00 722 0.00 5.00 0.00 125 0.00 0.00 0.00 726 0.00 0.00 0.00 731 . 0.00 0.00 0.00 734 O.00 0.00 0.00 137 0.00 0.00 0.00 740 0.00 0.00 o.00 743 0.00 0.00 0.00 745 0.0* 0.00 0.00 749 0.00 0n.oo0 752 0.00 0.00 0.00 5ss 5.00 0.00 0.00 0.00 0.00 0.0o 0.00 0.00 C.o0 0.00 0.00 0.00 0.00 0.00 0.00 5.00 0.00 e.00 0.00 0.00 0.00 0.00 0.00 0.do 0.00 0.00 0.00 0.00 0.00 0.00 0.00 651 0.00 0.00 0.00 654 0.00 0.00 0.00 s51 0.00 0.00 0.00 660 0.00 0.00 0.00 63 0.00 0.00 0.00 666 0.00 0.00 0.00 6E9 0.00 0.00 0.00 672 0.00 0.n0 0.00 615 0.00 0.00 0.00 618 0.00 0.00 0.00 us8 0.00 0.00 0.00 684 0.00 0.00 0.00 681 0.00 0.00 0.00 690 0.00 0.00 0.00 553 0.00 0.00 0.0o 966 0.00 0.00 0.00 699 0.00 0.00 0.00 702 0.00 0.00 0.00 ye5 0.00 0.00 0.00 706 0.00 0.00 0.00 III 0.00 0.00 0.00 714 0.00 0.00 0.00 J11 0.00 0.00 0.00 720 0.00 0.00 0.00 123 0.e0 0.00 0.00 723 0.40 0.00 8.00 12S 0.00 0.08 0.00 732 0.00 0.00 0o.0 135 0.o0 0.00 0.00 738 0.00 0.00 0.00 741 0.00 5.00 0.00 744 0.00 0.00 0.00 747 0.00 o.00 0.00 1so 0.00 0.00 0.00 753 0.0O 0.00 0.00 r56 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 O.00 0.0o 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.60 0.00 0.00 0.00 0.00 o.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.o00 0.0 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 .e00 0.00 0.00 0.W0 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.60 0.00 0.00 0.00 0.00 0.00 0.00 0.0 0.00 0.00 0.00 O.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.ca 5.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 Page 7 P 7S12)O.vKI 23- & t 9C

tster Croek low late for thermal Stress at 210 seconds - No Sand outside Nodes Middle Nodes Inside Nodes Radial 1qertdconul Hoop Redlal Herldional Hoop Radial ertdlf l V"op Nod x I hta Ndsx Sx SL sxr pa ck sit sYz S No sx s sz sOr (nchl (Inch) (digres) W I 1Ps (pi (psi) S) (psi) (psi (ps) (psi) (PSI) (PIt) (pst) (psi) J56 330.85 106.31 126.13 151 0.00 141 329.1I 10."U 128.46 INU 0.00 164 337.51 110.31 128.18 n3 0.00 1U 325.37 112.11 In9.1i n10 0.00 110 324.85 114.02 125.44 1t 0.00 113 323.13 11S.81 129.78 772 0.00 715 321.60 111.10 130.Q1 11

• .00 729 320.03 11S.S3 130.41 11b 0.00 182 310.504 21.34 130.34 761 0.60 785 316.13 123.1S 131.07 184 4.06 160 31S.3S 124.95 131.38 761 0.°'

lot 313.1? 128.14 131.12 1t0 0.00 714 312.11 128.52 132.04 193 0.00 ?1 310.56 1)0.21 132.31 1S6

• .00 5a" 306.24 M12.01 t32.10 199 0.00 603 301.3Z 133.81 133.062 00 0.00 SU 305.68 135.SS 133.35 805 0.00 o08 304.03 731.21 333.68 am 0.50 J12 3W2.38 139.01 134.06 all 0.00 015 300.71 740.1S 134.33 814 0.06 sit 281.03 142.44 134.55 17 0.00 5a1 211.3s 144.13 134.89 820 0.00 26 m9.63 14St. 135.31 U23 O.00 U21 "93.8s J41.5 13S.63 U

J 0.00 830 Z92.23 ?4V.16 131.35 029 5.00 a3 290.51 1508. 135.2 832 0.00 O6M 238.11 MAt.4Y 131.61 035 5.00 an 267.03 154.II 136.94 Sit 0.00 642 265.26 5.JJ 1347.26 U41 0.00 U4S MM.5? 11.35 331.50 044 0.00 845 2MM.51 70U.0 138.33 641 0.00 551 21J.60 162.5 1338.61 8sO 0.00 814 274.61 15S.27 139.21 S3 0.00 6s5 271.53 756.6S 139.15 is 0.00 38 W 68.SS 710.40 140.o1 ass 0.00 663 ?6S.48 112.92 140.34 362 0.00 0.00 0.00 0.00 0.00 0.06 0.00 0." 0.00 0.00 0.00 0.00 0.00 0.00 0.90 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 S."o 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 Js8 O.00 0.00 0.00 ?61 0.00 0.00 0.00 184 0.00 0.00 0.00 is? 0.00 0.00 0.00 110 0.00 0.00 0.00 113 0.00 6.00 0.00 ly6 0.00 0.00 0.00 J19 0.00 0.00 0.00 7J2 0.00 0.00 0.00 38S 0.00 0.00 0.00 781 o.oo 0.00 0.00 1li 0.00 0.00 0.00 714 o.0 0.00 0.00 lo? 0.0 0.00 0.00 am 0.00 0.00 0.00 803 0.00 0.00 0.00 806 0.00 0.00 0.00 Us 0.00 0.00 0.00 s12 0.00 0.06 0.00 61s o.00 0.00 0.00 8l1 o.oo 0.00 0.00 821 0.00 0.00 o.00

• 24 5.00 0.00 0.00 St?

0.00 0.00 0.00 130 0.00 0.00 0.00 033 0.0O 0.00 5.00 835 0.00 0.00 0.00 838 6.04 0.00 0.00 842 0.00 0.00 0.00 34S 0.00 0.00 a.00 848 0.00 0.00 0.00 SS6 0.05 0.00 0.00 154 5.0D 0.00 0.00 3s? 0.00 0.00 0.00 8O 4.00 0.00 0.00 863 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.06 0.00 0.00 o.00 0.00 0.00 0.00 0.0o 0.00 0.00 .00 0.00 0.00 0.00 o.00 0.00 0.00 0.00 0.00 0.0o 0.00 0.00 0.00 0.00 0.00 0.00 159 0.0Do 0.00 0.00 752 0.00 0.00 0.00 165 0.00 0.00 0.00 158 0.00 0.00 0.00 Mn 0.00 0.00 0.00 14 0.00 0.00 0.00 711 0.00 0.00 0.00 160 0.00 0.00 0.00 183 0.C0 0.00 0.00 1S 0.00 0.00 0.00 1n 0.00 0.00 0.00 n1z 0.00 0.00 0.00 1ss 0.00 0.00 0.00 16 0.00 0.00 0.00 $Ii 0.00 0.00 0.00 S 0.00 0.00 0.00 SO 0.00 0.00 0.00 BI 0.00 0.00 0.00 J13 0.00 0.00 0.00 615 0.00 0.00 0.00 p1i 0.00 0.0D 0.00 322 0.00 0.00 0.00 2S5 0.00 0.00 0.00 828 0.00 0.00 0.00 B1t 0.00 0.00 0.00 534 0.00 0.00 0.00 831 0.00 0.00 0.00 340 0.00 0.100 0.00 o 43 0.00 0,00 0.00 846 0.00 0.00 0.00 B" 0.00 0.00 0.00 652 0.00 0.00 0.00 3SS 0.00 0.00 0.00 6as 0.00 0.00 0.00 361 0.00 0.00 0.00 064 0.00 0.00 0.00 0.00 0.00 a.00 O.00 0.00 0.00 0.00 0.00 6.00 0.00) O.00 0.00 0.00 0.o 0.00 0.0 0.00 0.00 0.00 o.00 0.00 0.00 0.00 0.00 6.so a0.00 0.00 0.DO 0.00 0.00 0.00 o.oo 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.80 0.00 0.00 3.06 o.00 0.00 0.00 0.00 0.00 6.00 0.00 0.00 0.00 0.00 0.00 6.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 o0.00 e.0 0.00 0.00 0.00 0.00 0.00 6.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00

0. oa o.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 e.W 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 page a MOST210.W1 23-Oct-l0

vster Creek. RaNw Dte for Iherml Stress at 21O sWcOds -NO llnd Outslde Nodes 3dId1]* Nodes Inside Nodes Radtal herldlonal loop Radtsil erldimnal Hoop Radial MerldIonal Hoop Node X y Theta Nd 5U SY SU SUY Node 5U 5y 5i T Node Sx SI SU S x ( I C) linchl (degrees) fps$) (psi) (pal) (psi (psi) (pfst (psi) Ipsi) (pai) (psti I(p) (psi 666 262.39 715.41 141.38 us 0.00 859 2SS.2S 7J1.88 141.92 au 0.00 612 256.14 785.31 142.46 III 0.00 a85 252.98 782.7r 143.00 334 0.00 618 Z51.48 183.M5 143.26 61F 0.00 683 242.13 784.9 343.51 880 0.00 684 t41.27 186.16 143.91 6U3 0..0 Ut? 744.55 M8M.3 144.43 8 0.00 680 141.68 19.67 144.89 889 0.00 833 239.05 132.19 345.34 892 0.00 S96 236.28 794.63 t415.0 895 0.00 a9 233.41 796.S1 146.26 898 0.00 S92 230.69 711.42 146.1? S0l 0.00 s0 225.fl 681.86 341.52 504 0.00 900 221.39 as5.58 148.32 say 0.00 233 231.22 802.61 349.09 s10 0.00 514 213.41 314.38 149.65 513 0.00 817 209." 013.19 ISO.57 I11 0.00 $20 206.91 S24.21 111.25 919 0.00 123 204.38 829.51 ISI.89 922 0.00 526 202.23 935.08 152.47 925 0.00 925 200.55 640.14 I53.01 925 0.00 932 19.34 646.52 153.48 931 0.00 935 198.31 8S2.30 153.90 531 0.D0 938 1"1.31 858.26 154.25 93t 0.00 141 198.33 158.16 154.2 940 0.00 944 19t.32 60.7S tS4.39 943 0.00 941 398.32 U2.J0 154.s0 941 0.00 950 198.32 U63.71 S54.5S 949 0.0A 95 198.32 354. IS 154. 9152 0.30 95 198.32 865.1s I54.66 9SS 0.00 959 190.32 886.s 154.11 958 0.00 962 191.32 t61.18 JS4.16 961 0.00 965 198.32 868.03 154.82 964 0.00 966 196.32 669.86 154.8? a96 0.00 971 193.32 370.53 154.93 870 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.ao 0.00 0.00 0.00 0.00 0.00 D.00 0.00 0.00 0.00 O.00 0.00 5.00 0.00 0.00 0.00 0.00 0.00 D.00 o.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 066 0.00 0.00 0.00 869 0.00 0.00 0.00 872 0.00 O.00 0.90 8U5 0.00 0.00 0.00 Big 0.00 0.00 0.00o WI 0.00 q.00 0.00 884 0.00 0.00 0.080 as 0.00 0.00 0.00 690 0.00 0.00 0.00 e93 D.00 0.00 0.00 U36 0.00 0.00 0.00 a89 0.00 0.00 0.00 902 0.00 0.00 0.00 90s 0.00 0.00 0.00 906 0.00 0.00 0.00 911 0.o0 0.00 0.00 914 0.00 0.00 0.00 917 0.00 0.00 0.00 920 0.00 0.00 0.00 923 0.00 0.00 0.00 924 0.00 0.00 0.00 3 0.00 0.00 0.00 932 0.00 0.00 0.00 935 0.00 0.00 0.00 938 0.00 0.00 0.00 941 o.0o 1.00 0.00 944 0.00 0.00 0.00 947 .00 0.00 0.00 9so 0.00 0.00 0.00 9S3 5.00 0.00 0.00 956 0.00 0.00 0.00 959 0.00 0.00 0.00 96? 0.00 0.00 0.00 96S 0.00 0.00 0.00 g98 0.00 0.00 0.00 s11 0.W 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 D.00 0.00 0.0o 0.00 a.0a 0.00 0.0* 861 0.00 0.00 8.04 870 0.00 0.e0 0.00 873 0.00 O.00 0.00 376 0.00 G.00 0.00 8?9 o.00 0.00 0.00 ad2 0.00 0.00 0.00 ga5 0.00 0.oa 0.00 one 0.00 0.00 0.00 91 0.00 0.00 5.00 894 0.00 0.00 0.00 391 0.00 0.00 0.00 son 0.00 0.00 0.00 903 0.00 0.00 O.00 906 0.00 0.00 0.00 909 0.00 0.00 0.00 9)2 0.00 0.00 0.00 91S 0.00 0.00 0.00 91a 0.00 0.00 0.00 9?1 0.00 0.00 0.00 924 0.0a 0.00 0.00 92? 0.04 O.00 0.00 930 0.00 0.00 0.00 933 0.00 0.00 0.00 536 0.05 0.00 0.00 939 0.00 0.00 0.00 942 0.00 0.00 0.00 94S 0.00 0.00 0.00 s3u 0.00 0.00 0.00 951 0.00 0.00 0.00 95s4 0.00 0.00 0.00 951 0.00 0.00 0.00 9e 0 0.00 0.00 0.00 963 0.00 0.00 0.00 966 0.00 0.00 0.00 969 0.o0 0.00 0.00 siz Ci.w 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.08 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 o.0o U. UU 0.00 5.00 0.00 0.ou 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.0* 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 5.00 0.08 0.00 0.00 0.00 0.00 0.oO

u. Oa

v. ifc Page 2 gNOS1210.W9 Z3 Oct 99

Dyster Creek RCm Oat& for Therve) Stress at 210 seconds - No Sand Outside 1odes 11MVe Nodes Inside Nodes Radial Herldional Hbop Madial Ykridlonal Woop Radial terrddooal Hoop S C 1 I T tht odeod Si SY SX Sly Noek Si SY St Sly (inch) (inch) (degrees) (psi) (pal1 (ps1) (peI) (pll (ptll lpos) (Ps') (poil (psi)l (poll (pi 914 198.32 s;,.a 154.98 us3 0.o 97? 338.32 813.03 15.04 316 0.00 9gm 198.3? 814.06 SS.09 Wil 0.00 983 198.32 37S.13 IS5.15 96? 0.00 S8U 196.32 016.1 155.20 91$ 0.00 9on 198.32 877.21 1S.25 MS 6.00 992 198.32 ,11.21 ISS.31 91 0.00 99S 196.32 660.5S 155.42 994 0.00 996 t98.32 882.41 55.33 99t 0.0 1001 198.32 8S5.08 155.64 1o0o 0.00 I004 156.32 667.34 155.75 1003 0.00 1007 198.32 181.6£ IS5.85 100I 0.00 1010 126.32 191.86 155.91 S00 0.00 1013 116.312 M94.14 154.G0 t0l1 o.o0 1016 1S.32 08s.41 1S.1 0i15 0.00 tO19 19S.3t O90.60 155.29 1036 0.00 102? 19t.32 900.94 156.40 1021 0.00 1025 198.32 203.20 156.53 1024 0.00 102S 196.32 90S.41 158.61 1027 0.00 1031 110.32 801.13 Is567 30o" 0.00 1034 196.32 910.00 155.51 1033 0.00 1031 1906.2

• 12.21 156.91 1036 0.00 I14o IS5.3?

214.53 157.01 1035 0.00 1043 196.32 916.80 157.11 1042 0.00 l04a 198.32

• 19.06 IS1.21 304S 0.00 1049 195.32 321.33 157.31 1041 0.00 lO2 196.32 323.59 IS7.41 1051 0.00 1SS 1I8.332 2S.66 I5S.SO 10A 0.00 I058 196.32 26.13 157.60 105lt 0.0 1061 198.32 u30.31 151.19 l50 0.eo 3064 355.32 S32.61 157.79 1063 0.00 I06?

189.3t 934.92 1S1.88 3065 0.00 1010 196.32 931.19 157.9? IE9 0.00 1023 196.32 939.45 158.01 1072 0.00 3015 191.32 341.72 150.I6 3o0S 0.00 1019 198.32 943.91 158.2S 10)6 0.00 0.00 0.00 0.00 0.00 0.00 G.00 0.00 0.00 0.00 0.00 o.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 974 0.00 0.00 0.00 977 0.00 0.00 0.00 980 o.o0 9.00 0.00 983 0.00 0.00 0.00 986 0.00 0.00 0.00 989 0.00 0.00 0.00 992 0.00 0.00 0.W0 995 e.00 0.00 0.00 "a 0.00 0.00 0.00 1001 0.00 0.00 0.00 3004 0.00 0.0 0.00 t00? 0.00 0.00 0.00 1010 0.00 0.00 0.00 1013 0.00 0.00 0.00 lolS 0.00 0.00 0.00 1013 6.00 0.03 0.00 1022 0.00 0.00 0.00 1025 0.00 0.00 0.00 l028 0.00 0.00 0.00 1031 0.00 0.00 0.00 1034 0.00 0.00 0.00 1037 0.00 0.00 0.00 1040 0.00 0.00 0.00 1043 0.00 0.00 0.00 1046 0.00 0.00 0.00 1049 0.00 0.00 0.00 1052 0.00 0.00 0.00 11015 0.00 0.00 0.00 0ose 0.00 0.00 0.00 1011 0.00 0.00 0.00 1064 0.00 0.00 0.00 06 0.00 0.00 0.00 100a 0.00 0.00 0.00 1073 0.00 0.00 0.00 1076 5.00 0.00 0.0o 1019 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 8.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 9sz a.oo 0.00 0.00 978 0.00 0.00 0.00 901 0.00 0.00 0.00 964 0.00 0.00 0.00 987 0.00 0.00 0.00 990 0.0e 0.00 0.00 993 0.00 0.00 0.00 g6 0.00 o.00 0.00 999 0.00 0.00 0.00 3002 0.00 0.00 0.00 1005 0.00 0.00 0.00 IO00 0.00 0.00 0.00 1011 0.00 0.00 0.00 1014 0.00 0.00 0.00 1li e0.00 0.00 0.00 1020 0.00 0.0D 0.00 1023 0.00 0.00 0.00 l12s 0.00 0.00 0.00 3029 0.00 0.00 0.00 103? 0.00 0.00 0.00 1035 0.00 0.10 0.00 1036 0.00 0.00 0.00 1041 0.00 0.00 0.00 1046 0.00 0.00 0.00 1041 0.00 0.00 0.00 1050 0.00 0.00 0.00 1053 0.00 0.00 0.00 1055 0.00 0.00 8.00 1059 0.00 O.oW 0.00 106? 0.00 0.00 0.00 3065 0.00 0.00 0.00 1060 0.00 0.00 0.00 1071 0.00 0.00 0.00 1074 0.00 0.00 0.00 10J7 0.00 0.00 0.00 1080 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.0e 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.DO 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00

1. .0 0.00 0.0o 0.00 0.00 0.00

.D00 0.00 0.00 0.00 5.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 8.00 0.00 0.00 0.00 0.00a 0.00 0.00 0.00 0.00 0.00 0.00 0.00 D. 00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 .o00 0.00 0.00 0.00 0.00 0.00 O.00 0.00 0.00 0.00 0.00 0.00 0.05 0.00 0.00 0.00 0.00 0.00 0.00 5.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.o0 0.c0 0.e0 0.00 0.00 0.00 0.so 0.00 Page 10 NGDS12MUI4.k Z3 0ct -90

Oyster Creek Raw Data for Therma Stress at 210 seconds - No Sand Outside Nodes KIMdde 11odes Inside gode5, Radial Nerldloial Hoop Radlial feridlonal fop Radial erldional &Hop Nod A 1 Y Theta Nod SI Sy 52 Sly Node SI ST SX Sly Node sx SY 52 s (Inch) (inch) Idegr"s) (pit) (PSI) (PsI) fpsl (psi) {pslil (pil (psI) IP3t1 (psI) (PsI) (PsI) 1082 198.3t 946.25 150.34 1061 0.00 105 198.32 943.2S 158.41 1084 0.00 10U 198.32 950.2S 158.49 Joe? 0.00 1091 193.32 9WA.2S 156.57 1690 0.00 1094 196.32 954.2S 156.65 1093 0.00 I091 198.3t 56.2S 158.72 IM9 5.0. 1100 198.32 3S7.26 156.16 ON 0.00 Ila] I9".32 ass.16 153.7 11ez G..o 1106 1S6.12 SS9.11 15.A3 2105 0.00 119 198.32 960.06 158.87 1108 0.00 1112 158.32 9S1.0t ISS.90 llt 0.00 1115 198.32 961.97 15.94 1114 0.00 111 19S.32 962.92 258.9? 1117 0.00 1821 198.32 953.87 159.01 1120 0.00 214 19.32 964.32 IS9.04 123 0.00 I12J 198.32 965.71 159.06 1124 0.00 1)30 l9S.32 96S.88 I59.06 1129 0.00 1133 196.32 956.2S 156.10 1132 0.00 1134 138.32 25U.53 1SSI. 113S 0.00 1139 193.32 946.13 159.11 1136 5.00 1142 396.32 961.13 159.15 £141 0.00 1145 196.32 156.13 I.SX11 1144 0.00 1148 l9.32 139.1) 159.22 1141 0.00 1151 138.32 310.73 159.25 1150 0.00 1154 198.32 911.73 159.30 1153 0.00 1151 198.32 172.73 159.33 1lS 0.00 1150 198.32 1)3.13 1S9.37 1159 0.00 1163 11S.32 9314.73 156.40 1162 0.00 1166 198.32 915.73 I59.44 1INS 0.00 1516 198.32 076.73 159.47 116U 0.00 1172 136.32 918.13 159.5S tilt 0.00 1115 136.32 9L8.1) 159.U3 11J4 0.00 I17 198.32 963.33 S9.71 1117 0.00 11l 198.32 965.53 159.t7 1160 0.00 116 198.32 987.73 1S9.86 1183 0.00 ll I9.37 989.94 159.93 INS 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 M.e 0.00 0.00 0.00 0.00 0.00 5.00 0.00 0.00 0,00 0.00 0.00 0.00 0.00 0.0o 0.00 0.00 0.00 5.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 1082 0.00 0.00 0.00 1085 0.00 0.D0 0.00 1068 0.00 0.00 0.00 1091 0.00 0.00 0.08 1094 0.00 0.00 0.00 1091 0.00 0.00 0.00 1100 0.00 0.90 0.00 1103 0.00 0.00 0.00 11t 0.00 0.00 0.00 1109 0.00 0.00 0.00 1112 0.00 0.00 0.00 1115 0.00 0.00 0.00 lIfe 0.00 0.00 0.00 1121 0.o0 0.00 0.00 1124 0.00 0.00 0.00 1127

• .00 0.00 0.00 1130 0.00 0.00 0.00 1133 0.00 0.00 0.00 3136 0.00 0.00 0.00 1139 0.00 0.00 0.00 1142 0.00 0.00 0.00 1lS 0.00 0.00 0.00 1148 0.00 0.00 0.00 1151 0.00 0.00 0.00 1154 0.00 0.00 0.00 IRS?

0.00 0.00 5.00 1160 0.00 0.00 0.00 1163 0.00 0.00 0.0e 1166

• .00 0.00 0.00 1to%

0.00 0.50 0.00 1172 0.00 0.00 5.00 IllS 0.00 0.00 0.00 118) 0.00 0.00 0.00 l181 6.00 0.00 0.00 1164 0.00 0.00 0.00 l1i1 0.00 0.00 0.00 0.00 0.0o 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 1003 0.00 0.00 0.00 1066 0.00 0.00 0.00 1089 0.00 0.00 0.00 109Z 0.00 0.00 0.00 1095 0.00 0.00 0.00 1098 0.00 0.00 0.00 1101 0.00 D.00 0.00 1104 0.00 0.00 0.00 11o1 0.00 0.00 0.0 1110 0.00 0.00 0.o0 1113 0.00 0.00 0.00 1116 0.00 0.00 0.00 1119 0.00 0.00 0.00 1127 0.00 0.00 0.00 112S 0.00 0.C0 0.00 1128 0.00 O.00 0.00 1111 0.00 0.00 0.00 1134 0.00 0.00 0.00 1137 0.00 0.00 5.00 1140 0.00 O.00 0.00 1143 O.W 0.00 0.00 1146 0.00 0.00 0.00 1149 0.00 0.00 0.00 1152 0.00 0.00 0.00 1155 0.00 0.00 0.00 1158 0.oo 0.00 0.00 1161 0.00 0.00 0.00 1164 0.00 0.00 0.00 1167 0.00 0.00 0.00 11M0 0.00 0.00 0.00 1173 0.00 0.00 0.00 1t76 0.00 5.00 0.00 11n 0.00 0.05 0.00 1162 0.00 0.00 0.00 l185 0.00 0.00 0.00 1188 0.e0 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.0C 0.00 0.00 0.00 D.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.0O 0.00 0.0Oa 0.00 0.00 0.00 0.00 O.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00

0. D0 0.W0 0.00 0.00 0.00 0.00 0.00 0.00 0.0o O.w 0.0o 0.00 0.05 0.00 0.00 0.00 0.05 0.o0 0.00 0.00 0.00 0.00 0.00 0.00 0.00 Page I I, N051230.W1K 23-Oct -90

lster Croek Raw Data for Thenmal Stress at 210 sonds - 1b Sand Outside Modes Ile16 Nodes lsIde Nodes ladle) Msrldinal Hoop Radial herldlonal Hop Radlal Nerdlonal HoP hxdb X v thta Node SX Si Si sly Node SK SI St Ski Node SX Sr s2 sZr (Incht 4Incwbl) (dgrees) (pi) (ps I) (ps$ (psi) 4 ) (PSI) 1ps,, (pa' lIPs) (psil tPoll (pal) 1190 198.32 92.14 160.01 1189 0.06 1193 196.32 994.34 160.00 1132 0.00 1198 198.3t 996.54 180.36 fi9s 0.00 199 198.32 398.14 160.23 11t9 0.00 1202 136.32 1000.94 130.30 1201 0.00 120S 191.32 £003.1s 160.37 1204 0.00 1206 19S.32 1005.3S 150.45 1201 0.06 1231 198.32 1007.55 160.52 1210 0.00 1214 196.3Z 1009.15 160.19 If) 0.00 J217 136.32 1010.70 I3.62 1215 0.00 1220 193.32 1011.64 160.6S 1219 0.00 1223 19t.32 1012.61 160.68 1222 0.00 1228 196.32 1013.58 150.11 122S 0.00 1229 191.32 1014.51 150.74 1226 0.00 123? 159.37 1015.47 160.77 1231 0.00 1235 198.32 1015.42 160.61 1234 0.00 1238 193.32 1017.37 130.63 1237 0.00 1241 558. 3 1018.32 55W.08 3240 0.00 1244 198.32 1019.20 160.89 1243 0.00 1241 196.1? 1019.38 160.89 1241 0.o0 1250 198.32 J019.15 160.90 1249 0.00 1253 153.32 1020.13 160.51 1152 0.00 5258 191.32 1020.23 10.9* 1255 0.00 12S 158l.3? 10ZI.23 150.3S 52Ss 0.00 1262 I53.32 1422.23 160.98 1261 0.00 126S 198.32 1023.23 161.01 164 0.00 1268 198.31 1IC4.23 151.04 1277 0.00 1271 138.32 102S.23 161.07 1210 0.00 1211 196.32 1026.23 161.10 1271 0.00 12n7 158.3t l0?1.23 111.13 1278 0.00 1250 198.32 1D28.23 151.55 1219 0.00 1283 193.32 1029.23 165.5 328* 0.00 1286 19.32 1030.23 561.22 3285 0.00 1289 196.32 1012.65 161.2 l2z8 0.00 3292 198.3? 1035.06 161.36 1291 0.00 129S 19.32t 1037.SI 161.44 5294 0.00 0.00 0.00 0.00 0.00 0.00 6.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.06 0.00 0.00 0.00 0.00 G.0G 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 5.00 0.00 0.00 0.00 0.00 0.00 0.00 ]190 0.00 0.00 1193 O.00 0.00 1596 0.00 0.00 1IM9 0.00 0.00 Itar 0.00 0.00 1205 0.00 0.00 1206

• 0o o.00 1251 0.00 0.00 1?14 0.00 0.00 1231 0.00 0.00 1220 0.00 0.00 3223 0.00 0.00 1226 0.00 0.00 1229 0.00 0.00 5232 0.00 0.00 123S 0.00 0.00 1235 0.00 0.00 1241 0.00 0.00 1244 0.00 0.00 1241 0.00 0.00 1250 0.00 0.00 1253 0.00 0.00 1256 0.00 0.00 12S9 0.00 0.00 126?

0.00 0.00 1265 0.00 0.04 1258 0.00 5.00 5211 0.00 0.00 1274 0.00 0.00 127? 0.00 0.00 1290 0.00 0.00 1283 0.00 0.00 1286 0.00 0.00 128I 0.00 0.00 1292 0.00 0.00 U2ss 0.00 O.DO 0.00 0.00 0.00 0.00 0.00 0.D0 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 o.0e 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 .D00 0.00 0.00 0.00 0.00 0.00

• .oO 0.00 0.oo 0.00 0.00 0.00O 0.00 0.00 0.00 0.00 0.00 0.0so 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00e 0.00 0.00 0.00 0.00 9.00 0.00 0.00o 0.00 0.00 0.00 0.00 0.eo 0.00 53.00 list 0.00 0.00 5194 0.00 0.w0 119?

0,00 0.00 1200 0.00 0.00 5203 0.00 0.00 1206 0.00 0.00 1209 0.00 0.00 1212 0 00 0.00 1215 0.00 0.00 1214 0.00 0.00 1221 0.00 0.00 124 0.o0 0.00 122? 0.00 0.00 1230 0.00 0.00 1Z33 0.00 0.00 1236 0.00 0.00 1739 0.00 0.00 1242 0.00 0.o0 124f 0.00 0.00 1248 0.00 0.00 1251 0.00 0.00 1254 0.00 0.00 12St 0.00 0.00 12W 0.00 0.00 1263 0.00 0.00 1268 0.00 0.00 5269 0.00 0.00 1212 0.00 0.00 1WtS 0.00 0.00 1278 0.00 o.a0 1281 0.00 0.00 1284 0.00 0.00 1287 0.00 0.00 1290 0.DO 0.00 1293 0.00 0.00 1296 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.05 0.00 0.00 0.00 O.00 0.00 0.00 0.00 0.00 0.05 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 o.0o 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.o0 0.00 0.00 0.08 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.0o 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 .00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.0w page 12 MoST210.W1 23-oct 90

)yater Crek Raw Date for Thurmli Stress at 210 secondS - No Sand Outside odes Kiddle Nod1 Inside lNdes Radiml heridional Hoop Wadlal Merldional Hoop Radial Merldional Nwop MNd. I r Tbeta Nod* M SY St SI Nodle SX ST Si Sly Node St St sI Isnch) (lnthl (dagreeJi (pal) (psi) (psi) (psf) (psi) (Pal) (psi) (psi) (psi) (psi) (psi) Wpsi) 1296 198.32 1039.94 16t.51 1291 0.00 1301 195.32 1042.36 161.S8 1300 0.00 1334 1S6.32 1044.7 161.6S 1303 6.00 1301 198.32 1041.22 l61.72 13[S 0.0 1310 194.32 1049.S 161.78 1301 0.00 1313 198.32 1052.07 1t1.65 1312 0.00 131i 19.32 10S4.S0 161.12 1315 0.01 1311 196.3t 1055.41 161.95 1318 0.00 1322 198.32 1056.43 161.91 1321 0.00 1325 196.32 165.40 162.00 1324 0.06 1325 196.3t 1056.36 1$2.03 1327

• .00 1331 198.32 106.33 16S.05 1330 0.00 1334 31N.32 1060.29 162.06 1133 5.00 1331 196.32 3061.21 162.11 1336 0.00 1340 198.32 1062.22 12.1 139 0.00 1343 138.32 1063.18 152.16 1342 0.00 1346 191.32 1064.1S 162.11 134S 0.00 1343 198.32 144.2S 352.19 134 0.06 13S2 198.3t 1064.S0 162.19 13Sl 0.00 13SS 196.32 1064.75 162.20 13S4 0.00 1358 198.32 164.6S 162.20 1357 0.00 l311 166)3. 1065.5S 162.23 flu 0.00 1364 1£9.32 I0U.85 362.25 13U) 0.00 13357

1. 9.32 1067.65 362.28 136 0.00 1310 191.32 1O06.1S 162.31 1339 0.00 1313 3S3.32 306.S 112.34 1372 0.00 1316 196.32 1010.ff 162.36 1375 6.a0 1311 I.32 1011.85 16t.39
2. 3S 0.00 13t2 198.32 1012.8S 12.42 1311 0.00 1385 198.32 1013.65 15.44 134 0.00 1388 139.32 1914.5S 162.41 1381 0.00 1391 198.32 1071.07 152.53 1390 0.00 1394 198.32 1019.15 16t.S0 1323 0.0O 1l91 190.32 1031.50 I2.64 139 0.00 1400 198.32 1063.11 162.70 133 0.00 1403 196.32 1065.93 362.16 1402 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.M 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 D.00 0.00

• .00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 o.00 0.00 0.80 0.00 0.00 1296 0.00 0.00 0.00 1301 0.00 0.00 5.00 1304 O.00 0.00 0.00 1301 O.0 0.00 0.00 1310 0.00 0.00 0.00 1313 0.00 0.00 O.00 1316 0.00 0.00 0.00 119 0.00 0.00 0.00 1322 O.00 0.00 0.00 13?5 0.00 0.00 0.00 1326 0.00 o.00 0.00 1331 0.00 0.00 0.00 1334 0.00 0.00 0.00 1331 0.00 0.00 0.00 1340 0.00 0.00 0.00 1343 0.00 0.00 0.00 1346 0.00 0.00 0.00 134 0.06I 0.00 0.00 1352 0.00 0.00 0.00

1. 355 0.00 0.00 0.0w
2. 398 0.00 0.00 0.00 1361 0.00 0.00 0.00 1364 0.00 0.00 0.00 1361 5.00 0.00 0.00 1370 o.00 0.00 0.o0 1373 0.00 0.00 0.00 1316 5.00 0.00 0.00 1371 0.00 0.00 0.00 3332 0.00 0.00 0.00 1365 0.00 0.00 0.00 1388 0.00 0.00 0.00 1391 0.00 0.00 0.00 1311 0.00 0.00 0.00 1391 0.00 0.00 0.00 1400 0.00 0.00 0.00 1403 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 5.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 5.00 0.00 0.00 0.00 0.00 0.00 0.00 329S 0.00 0.00 0.00 130Z 0.00 0.00 0.00 2305 oo00 0.00 0.00 1308 0.00 0.00 0.00 1331 0.00 0.00 0.00 1314 0.00 0.00 0.00 1311 0.00 0.00 5.00 1320 0.00 0.00 0.00 1323 0.00 0.00 0.00 1IZ6 0.00 0.00 0.00 1329 0.00 0.00 0.00 1332 0.00 0.00 0.00 1335 0.00 0.00 0.00 1338 0.00 0.00 0.00 l341 0.00 0.00 0.00 1344 0.00 0.00 0.00 1341 O.00 0.00 0.Co 1350 0.00 0.00 0.00 1353 0.00 0.00 5.00 135 0.00 0.00 0.00 1359 0.00 O.00 0.00 1352 0.00 0.00 0.00 1365 0.0o 0.00 0.00 1358 0.00 0.00 o.00 1311 0.00 0.00 0.00 1314 0.00 0.00 0.00 1377 0.00 0.00 o.00 1330 O.00 0.00 0.00 1393 0.00 0.00 0.00 138s 0.00 0.00 0.00 1339 0.00 0.00 0.00 13192 0.00 0.00 0.00 t395 0.00 0.00 0.00 1390 0.00 0.00 0.00 3401 0.00 0.00 0.00 1404 0.00 0.00 0.w0 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00

0. 00 0.00 0.00 0.00 0.00v 0.00 0.00 5.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 6.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.06 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00D 0.00 0.00 0.00 0.00 0.00G 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 Page 13 oNsizio1.?-

?3-Oct-s0

Oyster Creek law Dat for rheml Stress at 210 seconds - No Sad OutsIde Nods Middle Ibdn Inside Nodes ladl)l Norldouml op Radia1 Morldial lH*p RadIl Meridional Hoop Ibb X T Ihnate Nod SX ST sr Sx r o S 5XKY SIbeh SR sr sz SltY IInh, ) (inch) (aogresel (poll (Pll (pstJ (psi) (pitI (psi) (psi) (psi) IPI (1si) (poll (potl 3406 1".3? 1088.14 152.61 1419 0.00 140 IJ18.32 109.36 1S2.81? 140 0.80 141? 1".32 109f.Sy IUR.§2 1411 0.00 1415 18.3t 309.79 1M2. 9 14)4 0.0 141 198.32 1191.00 113.03 1411 6.00 1401 180.3? 19.00 lU.0 14U0 0.00 1624 1".32 30on.00 163.06 1423 0.66 142 196.32 1100.00 163.11 1425 6.00 1435 19.32 1103.00 163.13 1429 0.06 1431 198.32 1102.00 163.IS 1432 0.00 1434 16.1? tt03.00 163.1 143S 0.00 1435 198.32 1104.ft 183.20 1438 0.00 1442 190.32 110S.00 1U3.23 1441 0.00 1441 18.32 3105.00 163.2S 14U4 0.00 1445 156.32 1101.00 163.2* 14? 0.00 1451 1".41 1106.25 163.29 1450 0.00 14S4 115.63 111S.S0 153.31 14S3 0.00 0.00 0.00 0.00 0.00 0.00 0.00 8.00 0.06 0.00 0.00 5.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 1406 0.00 O.00 0.00 140M 0.00 0.00 0.00 1412 0.00 0.00 0.00 1415 0.00 0.00 5.00 141 0.00 0.00 0.00 1421 0.00 0.00 0.00 146 0.00 5.00 0.00 1421 0.00 0.00 0.w0 1430 0.00 0.19 0.00 1433 0.00 0.00 0.00 1436 0.00 0.00 0.00 141. 0.00 0.00 0.00 1442 0.00 0.00 0.00 144S 0.00 0.0 0.00 1448 0.00 0.00 0.00 34S1 0.00 0.00 0.00 1454 0.00 0.00 0.00 0.00 0.00 0.00 0.00 5.00 0.00 5.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 148? 0.00 0.00 0.00 2410 0.00 0.00 0.00 1413 0.DO 0.05 8.00 1416 0.00 O.00 D.00 1419 0.00 0.00 0.00 14?? 0.00 0.00 0.00 1425 0.00 0.00 0.00 3428 0.00 0.00 0.00 1431 0.01 0.00 0.00 14U 0.00 0.00 0.00 1437 0.00 o.0 0.00 3440 0.00 0.00 0.00 1443 0.O0 0.00 5.00 1446 0.00 0.00 0.00 1449 0.00 0.00 0.00 1452 0.00 0.00 0.00 155 0.00 0.00 o.o0 O.00 0.00 0.00 0.00 0.00 0.00 0.0 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 5.90 0.00 0.00 0.00 O.W 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 G.OO 0.00 0.0 O.o0 D.00 0.00 0.00 0.00 0.00 OW 0.0 0.W0 0.00 0.00 0.00 0.00 0.00 0.0 0.00 D.00 SY 94 88 9J I6W2.69 11000.93 12454.5S 2643.90 101 101 s6 i8 .391.95 974.74 12"6.75 -72J.93 96 96 3S 99 289.00 -9656.70 -13916.42 -564.4

YVAVV4, REV. 0 AN ASME SECTION VIII EVALUATION OF THE OYSTER CREEK DRYWELL FOR WITHOUT SAND CASE PART 2 STABILITY ANALYSIS February 1991 prepared for GPU Nuclear Parsippany, Corporation New Jersey prepared by GE Nuclear Energy San Jose, California

DRF# 00664 INDEX 9-4, RE'. 0 AN ASME SECTION VIII EVALUATION OF THE OYSTER CREEK DRYWELL FOR WITHOUT SAND CASE PART 2 STABILITY ANALYSIS Prepared by: Cf-6. i 7 C.D. Frederickson, Senior Engineer Materials Monitoring & Structural-Analysis Services Reviewed by:_ H. S. Mehta, Principal Engineer Materials Monitoring & Structural Analysis Services Approved by:__ S. Ranganath, Manager Materials Monitoring & Structural Analysis Services i

ThOEXY-4. REV..l TABLE OF CONTENTS Page

1. INTRODUCTION 1-l 11.I General 1-1 1.2 Report Outline 1-1 1.3 References 1-2

2. SUCKLING ANALYSIS METHODOLOGY 2-1 2.1 Basic Approach 2-1 2.2 Determination of Capacity Reduction Factor 2-2 2.3 Modification of Capacity Reduction Factor for 2-3 Hoop Stress 2.4 Determination of Plasticity Reduction Factor 2-4 2.5 References 2-5

3. z.INITE ELEMENT MODELING AND ANALYSIS 3-1 3.1 Finite Element Buckling Analysis Methodology 3-1 3.2 Finite Element Model 3-2 3.3 Orywell Materials 3-3 3.4 Boundary Conditions 3-3 3.5 Loads 3-4 3.6 Stress Results 3-7 3.7 Theoretical Elastic Buckling Stress Results 3-9 3.8 References 3-10

4. ALLOWABLE BUCKLING STRESS EVALUATION 4-1

5.

SUMMARY

AND CONCLUSIONS S-1 iii

QRF" 00664 INDEX 9-4. REV. D3 LIST OF TABLES Table Page No. Title Noo. 3-1 Oyster Creek Drywell Shell Thicknesses 3-11 3-2 Cylinder Stiffener Locations and Section Properties 3-12 3-3 Material Properties for SA-212 Grade B Steel 3-12 3-4 Oyster Creek Drywell Load Combinations 3-13 3-S Adjusted Weight Densities of Shell to Account for 3-14 Compressible Material Weight 3-6 3yster Creek Drywell Additional Weights - Refueling 3-15 3-7 Oyster Creek Orywell Additional Weights - Post-Accident 3-16 3-8 Hydrostatic Pressures for Post-Accident, Flooded Cond. 3-17 3-9 Meridional Seismic Stresses at Four Sections 3-18 3-10 Application of Loads to Match Seismic Stresses - 3-19 Refueling Case 3-11 Application of Loads to Match Seismic Stresses - 3-20 Post-Accident Case 4-1 Calculation of Allowable Buckling Stresses Refueling 4-2 4-2 Calculation of Allowable Buckling Stresses - Post-Accident 4-3 5-1 Buckling Analysis Summary 5 2 iv

QRFz 00694 INDEX 9.4. REV. 0 LIST OF FIGURES Figure Page No. tllo. I-I Drywell Configuration 1-3 2-1 Capacity Reduction Factors for Local Buckling of 2-7 Stiffened and Unstiffened Spherical Shells 2-2 Experimental Data Showing Increase in Compressive 2-8 Buckling Stress Due to Internal Pressure 2-3 Design Curve to Account for Increase in Compressive 2-9 Buckling Stress due to Internal Pressure 2-4 Plasticity Reduction Factors for Inelastic Buckling 2-10 3-1 Oyster Creek Orywell Geometry 3-21 3-Z Oyster Creek Drywell 3-0 Finite Element Model 3-22 3-3 Closeup of Lower Orywell Section of FEM (Outside View). 3-23 3-4 Closeup of Lower Orywell Section of FEM (Inside View) 3-24 3-5 Boundary Conditions of Finite Element Model 3-25 3-6 Application of Loading to Simulate Seismic Bending 3-26 3-7 Meridional Stresses - Refueling Case 3-27 3-8 Lower Drywell Meridional Stresses - Refueling Case 3-28 v

DRF' 00664 INDEX 9-4. REV. 0 LIST OF FIGURES Figure Page No. Title no. 3-9 Circumferential Stresses - Refueling Case 3-23 3-10 Lower Drywell Circumferential Stresses - Refueling Case 3-30 3-11 Meridional Stresses - Post-Accident Case 3-31 3-12 Lower Drywell Meridional Stresses - Post-Accident Case 3-32 3-13 Circumferential Stresses - Post-Accident Case 3-33 3-14 Lower Drywell Circumferential Stresses - Post-Accident 3-34 Case 3-15 Symmetric and Asymmetric Buckling Modes 3-3S 3-16 Symmetric Buckling Mode Shape - Refueling Case 3-36 3-17 Asymmetric Buckling Mode Shape - Refueling Case 3-37 3-18 Buckling Mode Shape - Post-Accident Case 3 8 vi

YRF# onse4 ?NDEX 9-4. REV. 0

1. INTRODUCTION 1.1 General To address local wall thinning of the Oyster Creek drywell. GPUN has prepared a supplementary report to the Code stress report of record

[1-1] which is divided into two parts. Part 1 includes all of the Code stress analysis results other than the buckling capability for the drywell shell [1-2]. Part 2 addresses the buckling capability of the drywell shell shown in Figure 1-1 (1-3]. The supplementary report for the degraded drywell is for the present configuration (with sand support in the lower sphere). One option which is being considered by GPUN to mitigate further corrosion in the sandbed region is to remove the sand. Reference 1-4 and this report evaluate the Influence of removing the sand on the code stress analysis and buckling evaluation, respectively. Buckling of the entire drywell shell is considered in this analysis with the sandbed region being the area of primary concern. 1.2 Report Outline Section 2 of this report outlines the methodology used in the buckling capability evaluation. Finite element modeling, analysis and results are described in section 3. Evaluation of the allowable compressive buckling stresses and comparisons with the calculated compressive stresses for the limiting. load combinations are covered in section 4. Section S presents the summary of results and conclusions. 1-1

DRFd 00664 INDEX 9-4, REV. 0 1.3 References 1-1 "Structural Design of the Pressure Suppression Containment Vessels," by Chicago Bridge & Iron Co.,Contract # 9-0971, 1965. 1-2 "An ASME Section VIII Evaluation of the Oyster Creek Drywell Part 1 Stress Analysis." GE Report No. 9-1, DRF# 00664, November 1990, prepared for GPUN. 1-3 "An ASME Section VIII Evaluation of the Oyster Creek Drywell - Part 2 Stability Analysis," GE Reoort No. 9-2, ORFz 00664. November 1990, prepared for GPUN. 1,4 "An ASME Section VIII Evaluation of the Oyster Creek Drywell - Part 1 Stress Analysis." GE Report to. 9-3, DRF# 00664. February 1991, prepared for GPUN. 1-2

?REE 00664 IN EX 9-4, REV. 0 %L.b -= P.I q' 9 Kn 'tt'A Piws kA POI giT A: 2 as ia,- I bhY F pawrF_ 'j, b-l. t;, 4; Z!A f ~ ¢<i E9 to i Em-F t0s2 ha - t zlotytlLU10*& 3,^~ ) Figure 1-1 Drywell Configuration 1.A

?RF# 00664 NDEX 9-4, REV. 0

2. BUCKLING ANALYSIS METHODOLOGY 2.1 Basic Approach The basic approach used in the buckling evaluation follows the methodology outlined in the ASME Code Case N-284 (2-1 and 2-2].

Following the procedure of this Code Case, the allowable compressive stress is evaluated in three steps. In the first step, a theoretical elastic buckling stress, abie 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 elgenvalue determination are given in Section 3. In the second step, the theoretical elastic buckling stress Is modified by the appropriate capacity and plasticity reduction factors. The capacity reduction factor, ai, accounts for the difference between classical buckling theory and actual tested buckling stresses for fabricated shells. This difference is due to imperfections inherent in fabricated shells, not accounted for in classical buckling theory, which can cause significant reductions in the critical buckling stress. Thus, the elastic buckling stress for fabricated shells is given by the product of the theoretical elastic buckling stress and the capacity reduction factor, i.e., gie1*i When the elastic buckling stress exceeds the proportional limit of the material, a plasticity reduction factor, I1i, is used to account for non-linear material behavior. The inelastic buckling stress for fabricated shells is given by itfjaiie. 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 a Bioasoie/FS 2-1

R ATx 8 REV. 0 In Reference 2-1, the safety factor for the Design and Level A & B service conditions is specified as 2.0. A safety factor of 1.67 is specified -for Level C service conditions (such as the post-accident condition). The determination of appropriate values for capacity and plasticity reduction factors is discussed next. 2.2 Determination of Capacity Reduction Factor The capacity reduction factor, aj, is used to account for reductions in actual buckling strength due to the existence of geometric imperfections. The capacity reduction factors given in Reference 2-1 are based on extensive data compiled by Miller (2-31. The factors appropriate for a spherical shell geometry such as that of the drywell in the sandbed region, are shown in Figure 2-1 (Figure 1512-1 of Reference 2-1). The tail (flat) end of the curves are used for unstiffened shells. The curve marked 'Uniaxial compression' is applicable since the stress state in the sandbed region is compressive in the meridional direction but tensile in the circumferential direction. From this curve, a; is determined to be 0.207. The preceding value of the capacity reduction factor is very conservative for two reasons. First, it is based on the assumption that the spherical shell has a uniform thickness equal to the reduced thickness. However, the drywell shell has a greater thickness above the sandbed region which would reinforce the sandbed region.

Second, it is assumed that the circumferential stress is zero.

The 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 a@ value to account for the presence of tensile circumferential stress is discussed in Subsection 2.3. The capacity reduction factor values given in Reference 2-1 are applicable to shells which meet the tolerance requirements of NE-4220 2-2

DRF#X00664 INEX 9-4, REV. 0 of Section III [2-4]. Reference 2-5 compares the tolerance requirements of NE-4220 to the requirements to which the Oyster Creek drywell shell was fabricated. The comparison shows that the Oyster Creek drywell shell was erected to the tolerance requirements of NE-4220. Therefore, although the Oyster Creek drywell is not a Section III, NE vessel, it is justified to use the approach outlined in Code Case N-284. 2.3 Modification of Capacity Reduction Factor for Hoop Stress The orthogonal tensile stress has the effect of rounding fabricated shells and reducing the effect of imperfections on the buckling strength. The Code Case N-284 [2-1 and 2-2] notes in the last paragraph of Article 1500 that, "The influence of internal pressure on a shell structure may reduce the initial imperfections and therefore higher values of capacity reduction factors may be acceptable. Justification for higher values of a1 must be given in the Oesign report." Haris, et al 2-6] present the most comprehensive set of test data which clearly show that internal pressure in the form of hoop tension, increases the axial buckling stress. Reference 2-6 also contains data from References 2-7 and 2-8. Baker [2-9] and, in a recent book, Bushnell [2-10] also endorse References 2-6 and 2-7 in discussing the effect of internal pressure on the buckling capability of cylindrical shells. Based on References 2-6, Johnson (2-11] recommends a procedure to account for this increase in buckling capability. The data reported in Reference 2-6 are first discussed to show the reasonableness of the procedure recommended by Johnson. Figure 2-2 (Figure 13 from Reference 2-6) shows the experimental data from several sources showing the increase in compressive buckling stress due to internal pressure. The curve recommended by Johnson corresponds to the '90% Probability Curve' in Figure 2-2. As shown later in Tables 4-1 and 4-2, the values of nondimensional pressure parameter (the abscissa in Figure 2-2) in the buckling analysis 2-3

NUETT REV. 0 presented in this report, are between 0.1 and 1.0. In this range, the '90% Probability Curve' in Figure 2-2 is also a lower bound to the experimental data. This clearly shows that the use of Johnson's procedure in this report assures conservative results (i.e., the procedure predicts a smaller increase in the compressive buckling stress than that indicated by the experimental data). The implementation of Johnson's procedure in this evaluation is described next. The buckling stress in uniaxial compression for a sphere of uniform thickness is given by the following: Sc a (0.605)(0.207) Et/R 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.

In the presence of a tensile stress such as that produced by an internal pressure, p, the modified buckling stress is as follows: Sc,mod m ((0.605)(0.z07) + AC] Et/R Where AC is given in 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)z. When the tensile stress magnitude, S, is knowt, the equivalent internal pressure can be calculated using the expression: p a ZtS/R The AC tanm is then incorporated in the capacity reduction factor itself by defining a modified capacity reduction factor, ti mod:

• i,mod. 0.207 + AC/0.605 2.4 Determination of Plasticity Reduction Factor When the elastic buckling stress exceeds the proportional limit of the 2-4

?NWEXY 8-REV. 0 material, a plasticity reduction factor, ni, is used to account for the non-linear material behavior. The inelastic buckling stress for fabricated shells is given by niatiaie. Reference 2-2 gives the mathematical expressions shown below EArticle -1611 (a)] to calculate the plasticity reduction factor for the meridional direction elastic buckling stress. A is equal to taiie/oy and ay is the material yield strength. Figure 2-4 shows the relationship in graphical form. 171 -1.0 if A c O.55 (0.45/A) + 0.18 if 0.55 < A < 1.6 a 1.31/(1+1.15A) if 1.6 c A S 6.25 -1/A if A > 6.25 2.5 References Z-1 ASME Boiler and Pressure Vessel Code Case N-284, "Metal Containment Shell Buckling Design Methods, Section 1II, Division 1, Class MC", Approved August 25, 1980. 2-2 Letter (1985) from C.D. Miller to P. Raju;

Subject:

Recommended Revisions to ASME Code Case N-284. 2-3 Miller, C.D., "Commentary on the Metal Containment Shell Buckling Design Methods~ of the ASME Boiler and Pressure Vessel Code," December 1979. 2-4 ASME Boiler & Pressure Vessel Code, Section III, Nuclear Power Plant Components. 2-5 "Justification for Use of Section MII, Subsection HE, Guidance in Evaluating the Oyster Creek Drywell,' Appendix A to letter dated December 21, 1990 from H.S. Mehta of GE to S.C. Tuminelli of GPUN. 2-5

EX -4, REV. 0 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. S, pp. 351-356, May 1957. 2-9 Baker, E.H., et al., 'Shell Analysis Manual," NASA, CR-912 (April 1968). 2-10 Bushnell, D., 'Computerized Buckling Analysis of Shells," Kluwer Academic Publishers, 1989 (Chapter 5). 2-11 Johnson, B.G., "Guide to Stability Design Criteria for Metal Structures," Third Edition (1976), John Wiley & Sons. 2-6

sX0096-24, REV. ( 032 CA9 0.2 0.0 0 4 a 12 to m X i, l/T 24 23 Figure 2-1 Capacity Reduction Factors for Local Buckling of Stiffened and Unstiffened Spherical Shells 2-7

YNUEX T. REV. 0 10a 6 4 _(NAA) I - it4 L4, CRATE a S1WATTZ-;- .r- _ NAA 2 1.0 a6 4 I -- i ',-t BEST FIT CURVE ' -';- _.u5 l>:-_ 'a-T + ~ - ~- n E t 2 _I,, I I / THEORETICAL OURVE 7I .10 = 6 4 ; 2 .01 .01 t9 f** n^. .9- ,s Se,_, 90 K% PRO BABSILIT Y C UJRV E:. ~ . 1 1' t -,...1 a.t I~ ' -t t t l itI i I I I i* - !..4 i -J 4 6 0 E (t Figure 2-2 Experimental Data Showing Increase in Compressive Buckling Stress Due to Internal Pressure (Reference 2-6) 2-8

?RFD OQ664 NDEX -4, REV. 0 a 6 4 2 1.0 8 6 4 2 0.10 B a 4 a 0.0t 0.01 010 10 rE [ t ] Figure 2-3 Design Curve to Account for Increase in Compressive Buckling Stress Due to Internal Pressure (Reference 2-11) 2-9

R00664REV 1.0 7.o a ' Lli /c Figure 2-4 Plasticity Reduction Factors for Inelastic Buckling 2-10

DRF 00664 INDEX 9-4. REV. 0

3. FINITE ELEMENT MODELING AND ANALYSIS 3.1 Finite Element Buckling Analysis Methodology This evaluation of the Oyster Creek Drywell buckling capability uses the Finite Element Analysis (FEA) program ANSYS [Reference 3-11.

The ANSYS program uses a two step eigenvalue formulation procedure to perform linear elastic buckling analysis. The first step is a static analysis of the structure with all anticipated loads applied. The structural stiffness matrix, [K], the stress stiffness matrix, [S], and the applied stresses. cap' are developed and saved from this static analysis. A buckling pass is then run to solve for the eigenvalue or load factor. A., for which elastic buckling is predicted using the equation: ([K] + X CS] ) (u) = 0 where: X is the eigenvalue or load factor. (u) is the eigenvector representing the buckled shape of the structure. This load factor is a multiplier for the applied stress state at which the onset of elastic buckling will theoretically occur. All applied loads (pressures, forces, gravity, etc...) are scaled equally. For example. a load factor of 4 would indicate that the structure would buckle for a load condition four times that defined in the stress pass. The critical stress,

acr, at a certain location of the structure is thus calculated as:

acr 0 ap This theoretical elastic buckling stress is then -modified by the capacity and plasticity reduction factors to determine the predicted buckling stress of the fabricated structure as discussed in Section 2. This stress is further reduced by a factor of safety to determine the allowable compressive stress. 3-1

• RF:

00664 [NOEX 9-4, 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-Z]. This type of analysis conservatively neglects the vents and reinforcements around the vents which significantly increase the stiffness of the shell near the sandbed region, In order to more accurately determine the buckling capability of the drywell, a three dimensional finite element model is developed. The geometry of the Oyster Creek drywell is shown in Figure 3-1. Taking advantage of symmetry of the drywell with 10 vents, a 36* section is modeled. Figure 3-2 illustrates the finite element model of the drywell. This model includes the drywell shell from the base of the sandbed region to the top of the elliptical head and the vent and vent header. The torus is not included in this model because the bellows provide a very flexible connection which does not allow significant structural interaction between the drywell and torus. Figure 3-3 shows a more detailed view of the lower section of the drywell model. The various colors on Figures 3-2 and 3-3 represent the different shell thicknesses of the drywell and vent. Nominal or as-designed thicknesses, summarized in Table 3-1, are used for the drywell shell for all regions other than the sandbed region. The sandbed region shown in blue in Figure 3-3 is considered to have a thickness of 0.736 inch. This is the 95% confidence projected thickness to outage 14R. Figure 3-4 shows the view from the inside of the drywell with the gussets and the vent jet deflector. The drywell and vent shell are modeled using the 3-dimensional plastic quadrilateral shell (STIF43) element. Although this element has plastic capabilities, this analysis is conducted using only elastic behavior. This element type was chosen over the elastic quadrilateral shell (STIF63) element because it is better suited for modeling curved surfaces. 3-2

RFY 00654 INOEX 9-4, REV. 0 At a distance of 76 inches from the drywell

shell, the vent is simplified using beam elements.

The transition from shell to beam elements is made by extending rigid beam elements from a node along the centerline of the vent radially outward to each of the shell nodes of the vent. ANSYS STIF4 beam elements are then connected to this centerline node to model the axial and bending stiffness of the vent and header. Spring (STIF14) elements are used to model the vertical header supports inside the torus. ANSYS STIF4 beam elements are also used to model the stiffeners in the cylindrical region of the upper drywell. The section properties of these stiffeners are summarized in Table 3-2. 3.3 Orywell Materials The drywell shell is fabricated from SA-212, Grade B high tensile strength carbon-silicon steel plates for boilers and other pressurE vessels ordered to SA-300 specifications. The mechanical properties for this material at room temperature are shown in Table 3-3. These are the properties used in the finite element analysis. For the perforated vent jet deflector, the material properties were modifiec to account for the reduction in stiffness due to the perforations. 3.4 Boundary Conditions Symmetric boundary conditions are defined for both edges of the 36" drywell model for the static stress analysis as shown on Figure 3-5. This allows the nodes at this boundary to expand radially outward frora 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 thie fixity of the shell within the concrete foundation. Nodes at the end of the header support spring elements are also fixed. 3-3

DRF OO4RV INDEX 9-4. REV. 3 3.5 Loads The loads are applied to the drywell finite element model in the manner which most accurately represents the actual loads anticipated on the arywell. Details on the application of loads are discussed in the following paragraphs. 3.5.1 Load Combinations All load combinations to be considered on the drywell are summarized on Table 3-4. The most limiting load combinations in terms of possible buckling are those which cause the most compressive stresses in the sandbed region. Many of the design basis load combinations include high internal pressures which would create tensile stresses in the shell and help prevent buckling. The most severe design load combination identified for the buckling analysis of the drywell is the refueling condition (Case IV). This load combination consists of the following loads: Dead weight of vessel, penetrations. compressible material. equipment supports and welding pads. Live loads of welding pads and equipment door Weight of refueling water External Pressure of 2 psig Seismic inertia and deflection loads for unflooded condition The normal operation condition with seismic is very similar to this condition, however, it will be less severe due to the absence of the refueling water and equipment door weight. The most severe load combination for the emergency condition is for the post-accident (Case VI) load combination including: 3-4

TNOEX 9-4, REV. C' Dead weight of vessel, penetrations. compressible material and equipment supports Live load of personnel lock Hydrostatic Pressure of Water for Drywell Flooded to 74'-A" External Pressure of 2 psig Seismic inertia and deflection loads for flooded condition The application of these loads is described in more detail in the following sections. 3.5.2 Gravity Loads The gravity loads include dead weight loads of the drywell shell, weight of the compressible material and penetrations and live loads. The drywell shell loads are imposed on the model by defining the weight density of the shell material and applying a vertical acceleration of 1.0 g to simulate gravity. The ANSYS program automatically distributes the loads consistent with the mass and acceleration. The compressible material weight of 10 lb/ft2 is added by adjusting the weight density of the shell to also include the compressible material. The adjusted weight densities for the various shell thicknesses are summarized in Table 3-5. The compressible material is assumed to cover the entire drywell shell (not including the vent) up to the elevation of the flange. The additional dead weights, penetration weights and live loads are applied as additional nodal masses to the model. As shown on Table 3-6 for the refueling case, the total additional mass is summed for each 5 foot elevation of the drywell. The total is then divided by 10 for the 36' section assuming that the mass is evenly distributed around the perimeter of the drywell. The resulting mass is then applied uniformly to a set of nodes at the desired elevation as shown on Table 3-6. These applied masses automatically impose gravity loads on the drywell model with the defined acceleration of 1g. The same method is used to apply the additional masses to the model for the post-accident case as summarized in Table 3-7. 3-5

JRFa 00654 INOEX 9-4. REV. 0 3.5.3 Pressure Loads The 2 psi external pressure load for the refueling case is applied to the external faces of all of the drywell and vent shell elements. The compressive axial stress at the transition from vent shell to beam elements is simulated by applying equivalent axial forces to the nodes of the shell elements. Considering the post-accident case, the drywell is assumed to be flooded to elevation 74'-6" (894 inches). Using a water density of 62.3 lb/ft3 (0.0361 lb/in3), the pressure gradient versus elevation is calculated as shown in Table 3-8. The hydrostatic pressure at the bottom of the sandbed region is calculated to be 28.3 psi. According to the elevation of the element centerline. the appropriate pressures are applied to the inside surface of the shell elements. 3.5.4 Seismic Loads Seismic stresses have been calculated for the Oyster Creek Orywell in Part I of this report, Reference 3-3. Meridional stresses are imposed on the drywell during a seismic event due to a 0.058" deflection of the reactor building and due to horizontal and vertical inertial loads on the drywell. The meridional stresses due to a seismic event are imposed on the 3-0 drywell model by applying downward forces at four elevations of the model (A: 23'-7",B: 37'-3",C: 50'-I1" and D: 88'-9") as shown on Figure 3-6. Using this method, the meridional stresses calculated in Reference 3-3 are duplicated at four sections of the drywell including

1) the mid-elevation of the sandbed region, 2) 17.25' below the equator, 3) 5.75' above the equator and 4) Just above the knuckle region.

These four sections were chosen to most accurately represent the load distribution in the lower drywell while also providing a reasonably accurate stress distribution in the upper drywell.

ORF: 00664 INDEX 9*4. REV. 0 To find the correct loaos to match :ne seismic stresses, the total seismic stress (due to reactor buildina deflection and horizontal and vertical inertia) are obtained from Reference 3-3 at the four sections of interest. The four sections and the corresponding ¶eridional stresses for the refueling and post.accident 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 correct loads are determined to iatch the stresses shown in Table 3-9 at the four sections. The calculation for The correct loads are shown on Tables 3-10 and 3-11 for the refueling and post-accident cases, respectively. 3.6 Stress Results The resulting stresses for the two load combinations described in section 3.5 are summarized in this section. 3.6.1 Refueling Condition Stress Results The resulting stress distributions for the refueling condition are shown in Figures 3-7 through 3-10. The red colors represent the most tensile stresses and the blue colors. the most compressive. Figures 3-7 and 3-8 show the meridional stresses for the entire drywell and lower drywell. The circumferential stresses for the same areas are shown on Figures 3-9 and 3-10. The resulting average meridional stress at the mid-elevation of the sanabed region was found to be; aRm a -7580 psi 3-7

DRF X00664 INDEX 9-4, REV. 0 The circumferential stress averaged from the mottom to the top of the sandbed region is; 0Rc 4490 psi 3.6.2 Post-Accident Condition Stress Results The application of all of the loads described for the post-accident condition results in the stress distributions shown in Figures 3-11 through 3-14, The red colors represent the most tensile stresses and the blue colors, the most compressive. Figures 3-11 and 3-12 show the meridional stresses for the entire drywell and lower drywell. The circumferential stresses for the same areas are shown on Figures 3-13 and 3-14. The resulting average meridional stress at mid-elevation of the sandbed region was found to be; "PAm- -11960 psi The circumferential stress averaged from the bottom to the top of the sandbed region is; CPAc a +20080 psi 3-8

ORF= 00664 INDEX 9 4, REV. 3 3.7 Theoretical Elastic Buckling Stress Results After completion of the stress runs for the Refueling and Post-Accident load combinations, the elgenvalue buckling runs are made as described in Section 3.1. This analysis determines the theoretical elastic buckling loads and buckling mode shapes. 3.7.1 Refueling Condition Buckling Results As shown on Figure 3-15, it is possible for the drywell to buckle in two different modes. In the case of symmetric buckling shown on Figure 3-15, each edge of the 36' drywell model experiences radial displacement with no rotation. This mode is simulated by applying symmetric boundary conditions to the 3-D model the same as used for the stress run. Using these boundary conditions for the refueling case, the critical load factor was found to be 7.67 with the critical buckling occurring in the sandbed region. The critical buckling mode shape is shown in Figure 3-16 for symmetric boundary conditions. The red color indicates sections of the shell which displace radially outward and the blue, those areas which displace inward. The first four buckling modes were computed in this eigenvalue buckling analysis with no buckling modes found outside the sandbed region for a load factor as high as 9.94. Therefore, buckling is not a concern outside of the sandbed region. It is also possible for the drywell to buckle in the asymmetric manner shown in Figure 3-15. For this mode, the edges of the 3-0 model are allowed to rotate but are restrained from expanding radially. This case is considered by applying asymmetric boundary conditions at the edges of the 3-0 model. With the two pass approach used by ANSYS, it is possible to study asymmetric buckling of the drywell when the stresses are found based on symmetric boundary conditions. The resulting load factor found using asymmetric boundary conditions is 10.13. The mode shape for this case is shown on Figure 3-17. 3-9

DRF4 00664 INDEX 9.4, REV. 0 Because the load factor is lower for symmetric boundary conditions with the same applied stress. the symmetric buckling condition is more limiting. Multiplying the load factor of 7.67 by the average meridional stress from section 3.6.1, the theoretical elastic buckling stress is found to be; aRie a 7.67 x (7580 psi) = 58,100 psi 3.7.2 Post-Accident Condition Buckling Results Considering the post-accident case with symmetric boundary conditions. the load factor was calculated as 5.18. Multiplying this load factor by the applied stress from section 3.6.2 results in a theoretical elastic buckling stress of GPAie = 5.18 x (11960 psi) - 61,950 psi The critical mode shape for this condition is shown in Figure 3-18. Again. the critical buckling made is in the sandbed region. 3.8 References 3-1 DeSalvo, G.J., Ph.O, and Gorman. R.W., "ANSYS Engineering Analysis System User's Manual, Revision 4.4." Swanson Analysis Systems. Inc., May 1, 1989. 3-2 GPUN Specification SP-1302-53-044. Technical Specification for Primary Containment Analysis - Oyster Creek Nuclear Generating Station; Rev. 2, October 1990. 3-3 "An ASME Section VIII Evaluation of the Oyster Creek Drywell - Part 1 Stress Analysis," GE Report No. 9-1, ORF. 00664, November 1990, prepared for GPUN. 3-10

0RF~ 00664 INDEX 9-4, REV. 0 Table 3-1 Oyster Creek Orywell Shell Thicknesses Section Thickness (in.} Sandbed Region 0.736 Lower Sphere 1.154 Mid Sphere 0.770 Upper Sphere 0.722 Knuckle 2.5625 Cylinder 0.640 Reinforcement Below Flange 1.250 Reinforcement Above Flange 1.500 Elliptical Head 1.1875 Ventline Reinforcement 2.875 Gussets 0.875 Vent Jet Deflector 2.S00 Ventline Connection 2.500 Upper Ventline 0.4375 Lower Ventline 0.250

• 95% confidence projected thickness to 14R.

3-11

DRF# 00664 INDEX 9-4, REV. 0 Cyl inder El evation (in) 966.3 1019.8 1064.5 1113.0(1) 1131.0 (1) Table 3-2 Stiffener Locations and Sect Height Width Area (in) 6in. 4 in2l 0.75 6.0 4.5 0.75 6,0 4 5 0.50 6.0 3.0 2.75 7.0 26.6 1.00 7.38 1.0 12.0 12.0 This stiffener is made up 2.75x7" and one 1.0x7.375" ion Properties Bending Inert-i a (in 41 Horizzontal Vertical 13.5 0.211 13.5 0.211 9.0 0.063 387.5 12.75 144.0 of 2 beam 1.000 sections, one Table 3-3 Material Properties for SA-212 Grade B Steel Material propertv Val ue Young's Modulus 29.6x106 psi Yield Strength 38000 psi Poisson's Ratio 0.3 Density 0.283 lb/in3 U 3-12

?RF#; 00664 ?NDEX 9-4, REV, 0 Table 3-4 Oyster Creek Drywell Load Combinations CASE I - INITIAL TEST CONDITION Deadweight + Design Pressure (62 psi) + Seismic (2 x DBE) CASE II - FINAL TEST CONDITION Deadweight + Design Pressure (35 psi) + Seismic (2 x OBE) CASE III - NORMAL OPERATING CONDITION Deadweight + Pressure (2 psi external) CASE IV - REFUELING CONDITION Deadweight + Pressure (2 psi external) Seismic (2 x OBE) + Seismic (2 x OBE) + Water Load + CASE V - ACCIDENT CONDITION Deadweight + Pressure(62 psi @ 175*F or 35 psi @ 281*F). Seismic (2 x DBE) CASE VI - POST ACCIDENT CONDITION Deadweight + Water Load @ 74'6" + Seismic (2 x DBE) 3-13

DRF4 00554 INDEX 9'.4. REV. 0 Table 3-5 Adjusted Weight Densities of Shell to Account for Compressible Material Weight Shel I Thickness (in.) Adjusted Weight Density ( l b/i n3 ) 1.154 0.770 0.722 2.563 0.640 1.250 0.343 0.373 0.379 0.310 0.392 0.339 3-14

0RF% 00664 INDEX 9-4, REV. O Table 3-6 Oyster Creek Drywell Additional Weights - Refueling Condition ELEVATION (feet) 15.56 ICM 16 20 15-20 25I

• • 21-251 26 30 30.25
• ' 6-30 31 32 33 34 35 31-35 36 40 36-40 50#
• ' 45-50S 54
  • 51-55 56 60 5 St-60 as

'* 61-65 70 o55-70 73 '71-75 82.17 0* 1-85 57 90 86-90 93.75 94.l5d' 95.75 TOTALS DEAD WEIGHT (lbf) 50000 556000 PENETR. WEIGHT (Ibf) 188100 11200 11100 51500 15500 750 15450 26050 1500 misC. TOTAL LOADS LOAO (0bf) (1bf) 50000 18100 11200 556000 l1100 115600 100000 205000 16500 750 15450 28050 1500 5 FOOT RANGE LOAD 239300 556000 LOAD PER 36 DEG. (lbfI G OF ELEEMTS NODES OF APPLICATION 116-119 161-169 LOAD PER FULL NODE (lbf) 3822 5950 4146 22930 55600 6 a LOAD PER HALF NODE (lbf I 1911 3475 2073 54100 105000 331700 33170 a 179-187 a 18-196 62250 6225 1550 778 389 41000 1102000 56400 95200 52000 1550 43350 7850 700 57SO 8s50 64350 1102000 7550 24000 80400 20000 115900 20000 72000 575D U50 21650 1000 15000 20700 698000 698000 20100 85900 1102000 78M0 196300 7ZO0o 5750 8850 21650 8590 110ZOO 75 a S I 19630 7200 5J 885 2165 8 a 6 a a 197-205 418.426 436-444 454-482 472-480 508-516 526-534 553-S6 1074 13775 98 2454 900 72 1l1 271 537 6888 49 1227 450 36 55 135 21650 1000 15000 16000 1600 20700 20100 a 571-571 589-197 200 100 ____1 0 388200

6. 0......

... 350 21 U 150 38820a 862000 34U43350 73U800 __3.... 3 U4343C 73880 3___..._ 3J434S 923S 4618 LOAD TO BE APtIED IN VERTICAL DIRECTION ONLY. MISCELLANEOUS LOADS INCLUOD 616000 LO WATER WE16HT AT 94.75 FT. ELEVATION 100000 LB EQUIPMENT DOO VfEIHT AT 30.25 FT. ELEVATION AID WELD PAD LIVt LOADS OF 24000. ZOOOD AKD 20000 AT 58, 60 AND 65 FT. ELEVATIONS REF6T.WKI 3-15

QRF# 00664 INDEX 9-4, REV. 0 Table 3-7 Oyster Creek Drywell Additional Weiqhts - Post-Accident Condition ELEVATION (feet) I5.55 16 20

• ' 15-20 226
• 3 21-25 26 30 30.25 et 26-30 31 32 33 34 35

" 31-35 36 40

• - 36-40 50#
• 4S-50S 54
• t 51-55 SI-S55 s0
• - 55-60 65 to 81-65

~'66-70 73

• ' 71-7S 82.17

'7 5590 8*6-90 93.75 95.15

• 91-SO TOTALS:

DEAD WE 16Ht1 l0bf) 50000 PENETR. WEIGHT Ilbf) mISC. LOADS flbf) 168100 11200 156000 64100 105000 41000 1102000 56400 95200 52000 11100 51500 16500 750 15450 26050 1500 1550 43350 7s50 700 TOTAL LOAD (lbf) 50000 168100 11200 556000 11100 115600 105l00 16S00 750 15460 28O50 1500 1550 84350 1102000 7850 5400 95500 52000 S750 8850 2150 1000 15000 20700 20100 5 FOOT NAKG( LOAO 229300 556000 LOAD PER 36 DEG. (Tbf) 22930 55600 I OF NODES OF ELEMENTS APPLICATION 6 116-1i9 6 161-169 8 179-187 18-196 LOAD PER FULL MODE (lbf) LOAD PER HALF NODE (ib!) (w____ 3UZ2 1911 6950 3475 Z31700 23170 8Z250 8225 2896 144 776 363 85500 1102000 7850 152o30 52000 5750 "s50 21650 8590 110200 785 a a 5750 UlSO 15230 S200 57$ 685 2165 8 a a 197-205 418-426 436-444 454-462 472-480 508-Sl1 526-534 553-561 1904 650 72 111 271 1074 13775 98 531 6888 45 952 325 36 135 21650 100D 1500 1600 1600 20700 20100 J 571-579 I 589-597 290 lOO SID 2Ss .40800 2164150 388200 0 2572350 2572350 I - LOA TO BE APPLIED IN VERTICAL ODRECTION ONLY. NO MISCELLANEMW LOADS FOR THIS COSDITION. 4080 257235 FLOMT.bAI 3-16

DRFf 00664 INDEX 9-4. REV. C0 Table 3-8 Hydrostatic Pressures for Post-Accident, Flooded Condition WATER DENSITY: FLOODED ELEV: 62.32 lb/ft3 0.03605 lb/in3 74.5 ft 894 inches ELEMENTS ABOVE NODES 2;' 40 5'1 6(i 79 9;! 10;! 1012 111; 121) 1241 131) 131 1413 161 171) 173' 18;3 197 40'1 409 41B 427 436 445' 454 463 472 481 490 499 SC8 517 526 ANGLE ABOVE EQUATOR (degrees) -53.32 -51.97 -50.62 -49.27 -47.5SO -46.20 -44.35 -41.89 -39.43 -36.93 -34.40 -31.87 -29.33 -26.80 -24.27 -20.13 -14.38 -8.63 -2.88 2.88 8.63 14.38 20.13 25.50 30.50 35.50 40.50 45.50 50.50 54.86 ELEVATION (inch) 110.2 116.2 122.4 128.8 137.3 143.9 153.4 166.6 180.2 194.6 209.7 225.2 241.3 257.6 274.4 302.5 342.7 384:0 425.9 468.1 510.0 551.3 591.5 627.8 660.2 690.9 719.8 746.6 771.1 790.5 805.6 820.7 835.7 850.8 885.3 DEPTH (inch) 783.8 777.8 771.6 765.2 756.7 750.1 740.6 727.4 713.8 699.4 684.3 668.8 652.7 636.4 619.6 591.5 551.3 510.0 468.1 425.9 384.0 342.7 302.5 266.2 233.8 203.1 174.2 147.4 122.9 103.5 88.4 73.3 5B.3 43.2 8.7 PRESSURE (psi) 28.3 28.1 27.8 27.6 27.3 27.1 26.7. 26.2 25.7 25.2 24.7 24.1 23.5 23.0 2.2.3 21.3 19.9 18.4 16.9 15.4 13.8 12.4 10.9 9.6 8.4 7.3 6.3 5.3 4.4 3.7 3.2 2.6 2.1 1.6 0.3 ELEMENTS 1-12 13-24 25-36 37-48 49-51, 61-66,55-57 52-54, 138-141,58-60 142-147, 240-242, 257-259 148-151, 243, 256 152-155, 244, 255 156-159, 245, 254 160-165, 246, 2S3 166-173, 247, 252 174-183, 248-251 184-195 196-207 208-215 216-223 224-231 232-239 430-437 438-445 446-453 454-461 462-469 470-477 478-485 486-493 494-501 502-509 510-517 518-525 526-533 534-541 542-549 550-557 187.3 706.7 25.5 340-399 (Ventlinia) FLOOOP.WlI 3-17

DRFF 00664 INDEX 9-4, REV. 0 Secti c Table 3-9 Meridional Seismic Stresses 2-0 Shell Elevation Model an (inches) Node Sandbed 119 32 low Equator 323 302 te Equator 489 461

kle 1037 1037 at Four Sections Meridional Stresses Refuel ing Post-Accident (psi )

(psi) 1258 1288 295 585 214 616 216 808 A) Middle of B) 17.25' Bel C) 5,75' AboN D) Above Knuw 3-18

?NDEX 9.4, REV. 0) Table 3-10 Application of Loads :o Match Seismic Stresses - Refueling Case COMPAESSIVE STRESSES

t~i~ 2-D ANALYSIS O.058' SEISMIC DEFLECT:I' HORIZ. PLUS VERTICAL SE:SMIC INERTIA:

TOTAL SEISMIC COMPRESSIVE STRESSES: SECTION: 2-0 NOOE CLEV: SECTION: 3-0 MODES: ELMY: 3-0 INPUT LOAD SECTION A C 0 2-0 SEISMIC STRESSES AT SECTION il;si) 1 2 3 4 32 302 461 1037 119.3" 322.5" 469.1" 912.3' 788.67 155.54 103.46 85.31 468.55 139.44 110.13 130.21 1258.22 294.98 213.59 215.52 3-0 STRESSES AT SECTION (pst) 1 2 3 4 53-65 170-176 400-408 S2-134 119.3* 322.S' 469.1" 912.3* 85.43 37.94 34.94 55.23 89.88 39.9t 35.17 O.CO 97.64 43.37 0.00 0.Co 88.85 0.00 0.00 M.Do 12St.22 294.98 213.59 21S.!,2 RESULTINB STRESSES AT SECTION filsi) INPUT 3-0 UNIT LOA DESCRIPTION 1000 lbs at PodXS 563 tnrough 569 500 lbs at 4271435. 1WCO lbs at 428-434 500 lbs at 197&205. IO lb. at 198-204 500 lbs at 161MMS. 10C: lbs at 162-1U DESIRED OMPRESSIVE STRESSES (psi): 3-0 INPUT LO10 SECTION A C 0 LOAD TO BE APPLIED TO PATCH 2-0 STRESSES 3902.2 2101.4 1453.1 5611.6 333.37 I&4.8 141.93 594.05 SU14: 1256.22 143.05 63.89 63.04

0. 0 294.93 138.34 77.ZS 0.00 0.00

_..3.__ 213.59 215..52

0. 0 0.00 IS.52 SEISUKPL.WK1 3 19

DMRF#X00664 INDEX 9-4, REV. 0 Table 3-11 Application of Loads to Match Seismic Stresses - Post-Accident Case SECTION: 2-D NOCE: ELEY; COMPRESSIVE STRESSES FROM 2-0 ANALYSIS _^_..........___............. 0,058" SEISMIC DEFLECTION: HMRIZ. PLUS VERTICAL SEISMIC INERTIA: TOTAL SEISMIC COMPRESSIVE STRESSES: 3-0 INPUT LOAD SECTION A B C 0 SECTION: 3-0 NODES: ELEV: INPUT 3-0 UNIT LOAD oESCRIPTION 2-0 SEISMIC STRESSES AT SECTION (psi) 1 2 3 4 32 302 461 1037 119.3" 322.5 489.1" 912.3" 788.67 1SS.S4 103.46 IS.31 499.79 429.31 512.76 723.14 1Z88.46 584.93 616.22 808.45 3-0 STRESSES AT SECTION (psi) 1 2 3 4 53-65 170-178 400-401 525-534 119.3" 322.5-44.1" 912.3" 85.43 37.54 34.54 55.23 89.38 39.9Z 36.76 0.00 97.64 43.37 0.00 0.00 3.85. 0.00 0.00 0.00 1288.46 584.93 616.22 805.45 RESULTING STRESSES AT SECTION (psi) 1250.51 555.36 511.4S 803.,4S Z5S.17 113.75 104.71 0.00 -169.54 -54.21 0.00 0.00 -28.64 0.00 0.00 0.00 1288.46 584.93 616.22 808.45 3-0 INPUT LOAD SECTION A C 0 1000 lbs at nodes 563 through 569 500 lTb at 4276435, 1000 lbs at 42U-434 500 lbs at 1971205. 1000 lbs at 191-204 500 lbs at 161t169. 1Q00 lbs at 162-168 DESIRED COMPRESSIYE STRESSES (psi): LOAD TO BE APPLIED T7 MATCH 2-0 STRESSES 14837.9 2850.2 -1941.7 -318.8 SUN: SEISFL.WK1 I -in

RF# 006A4 INDEX 9-:. REV. 0 DRYWELL ELEV.51'0* THK.67r Oyster Creek Drywell Geoietry Figure 3-1. 3-21

I ot CW LLIS-I OWStZX CatX DIMfIC ANALYSIS - OWC-RLO (NO SAND, POST-ACC. hNSVS 4.4 DEC 4 1990 PkOT NQ.~ M pKP KL MENTS UXAL MUM UV =1 M374=27AM-79.I Xr =13.O31 ZF =629.498 ar4c2-BSo C&zT9ID HIDDEN A' p. Ficurm 3-2. n.vctr rrm.I noveali ifn rtnitn riamnmt uA.,

I A"Wat. DEC 4 1990 AEAL "UKN WU =1 1S1-2G376 xr =42a.h5a 2r =216.528 Cx=lTOID HIDDZI U& 14 W I I I OYSTER CR DRYWELL AINLYSIS OYCJI+/-O CHO SAND, POST-ACC4 Figure 3-3. Closeup of Lower Orywell Section of FEM (outside vfawi

1 ANSVS 4.4 DEC 4 l990 in:S?:6 LOT MO 3 UKlP7 U+/-qEKTrS REAL MA Xv =-1 Xr =429.43a 2r =21X.52B ANG2z96 CIMRO1D H"tDPU 4*b

• tk I

I oySTER CMM D RYELL AMLYSIS OYCR10 CHO SAMD, POST-ACC. CC-..-- I A rl-r^ Arc a ---- n-^ bn ^- CCU f rrXUt-..l

1 AI SS

4. 4 DXC 4 1996 13:10:37 Q;7 IMNTS ITYPE tIUr DC SYNBOLS kVU

=1 YU =-a. a DKST=718.786 XF =3

• .1 Zr

=6319.498 I "tO1 oIDDay up O cS w DRashLL ANALYSIS -CRuo NO SA ND. POST-AC 4 l FiallrA t 4, nAnIdafw ri^nditinn-nf Finitp Fluenfit IoduI

w A. ct I I OYSTER CEIC DRYWELL ANALYSIS OYCIAd (SAHD UIT LOAD CA ONSYS 4.4 OCT L5 199. 09:32:26 TYPE "UK EC SYMBoLs RVi =I YU =o@. 8 DrSi=718.786 Xr =3Q3.031 Zr =639.498 Ad+:=-0o CENROID HIDDEN 9E) ~~. A - _ _ -l I l__JR X-C. v_em.: -_l_

L tOU L6990 STXP=1 SY C*UGC1 uL5S D"X -. 221779 SF =-BI?4 SPIX =655.047 XU =1 YU =-D... DI ST=710. 706 XJr =383.931 CVIROID HIXDE -7198 -3247 -127' 4 _a & 5 69S7 s--...-." g.s.a's. s it

• f.. v*tv.

.. a mred atlf Y V WE C I*ruin Si I IESlet} OYSTSX CMX DARY1l:LL ANALYSIS - OYCRIS CHO 5$AD. RE"EL!1G Flaure 3-7. F 3eridional Stresses - Refuelinq Case

.-NSS 4.4 mo4J 16 1~9S 8G:57:a9 ETZR=l r __-SY (AU,) M I F I I -s UI I1 00 0DM( =8. 139473 MaSMt =-0174 SMC =69S.947 XU -1 YU =-ai. a DT ST=280 376 XIr =429. 452 CEMTROID HIDDIN -8174 Mt. -7198 -3247 -1276 @WSTZ3M CXEEX DRYIIELL ANALYSIS -OYCH+/-S C?@ SAND, RKFUCLINCO I.

Amsys 4.4 a"O S 6 1990 S9:5:39 STEP--I SX (AUC) DX =9. 121779 oM =- 3 47 SNX =6?S4 .XU =1 DIST=710.766 XI =33~.31. E e tjg. 497@C CJINTOID HIDDZI -3547 - 246 4465 '754 tALi C maD OrYSTE CRY1J DPEUrL AItLYSIS QYCRLS CNOt SAND. RLFUEI 4 l C.r---g -___a 4-r.ls r._ r_.__ ft a

A-SVS 4.4 1 OU 1X6 190 KTEP=1 Sy < AUG) DX =89.139473 SMH =-3547 SPIX 6754 MU =1 Y =-U.S DIST=28 .376 Xr =429.452 _~RA =_2J8.328 C_1TROID HIDADZ E -3547 -240 2 L7 6 4465 6754 OYSTER CZERX DRYIPELL ANALYSIS -OVCHIS C(tO SAKD. REFUELING Figure 3-10. Lower Orywell Circumferential Stresses - Refueling Case

1 rm~ rmw nrm H-v _ 1.1 11 lI Vvy I U UIf I so r +rr Ii ,.9"~ s9ib1.

• MSYS 4.4 NOW 19 1990 IA :3U: 8 STEP--I I TE]tR1 ST C AUC)

UEDLES BDX =9.479734 Sm =-13LZ5 SMIX =3894 XU =1 YU =-B.8 DI S=710.176 MY =3 M3. 83. &ZE-. 498 CZNTROID HIXDDE 1315S E -11260 -3683 519. 136 _ 294 L OYSTER CRS DRYE AtML9YSIS SAND. POST-ACC.? Fiaure 3-11. eridional Stresses - Post-Accident Case

hNSYS 4.4 NOU L9 199G 16:32:46 r--r--r-r .SY (AbUG) I S I S a i 4 L 3 I S It I It It ~ 2 .4 a_ _ s 6_ 2 _ _; w_ mu¢ =216 *:731 a a a S 3 U" Z _ a a t t r 5 xu = 1RODHt I S I

• S

-311 FLO 24f M2 a a&aSYI ( C A.453 a a.. - -0 a -J.5. -11 .5. .- A..?D X =0 9 3 _ X d -3683 _iS J OYS C 4 DRYL ALYSIS - OYO CO SAND, nST-gCC. Figure 3-12. Lower Orywell Heridlonal Stresses - Post-Accident Case

I qSV 4.4 16:39:42 ALSO In& RE=1 $TEP=i T1ST]=1 SX ( AUC) DC( =6.479734 Sl9 =-5295 SMX =27791 XU =1 YU =-G.. DIST=71.786 XH =3in.931 CEITROID NIDIDK -52 -1539 13126 20459 27791 OVSTER cREM3 DRYWIELL ANALYSIS SANDS POST-ACC.~ Figure 3-13. Circumferential Stresses - Post-Accident Case

-.aSVS 4.4 1 0U 19 1993 srii'=i sTXr=l I TXFRJ ELE CS _WX =6.479734 SoM =-325 SMX =27791 )(U =1 YV =-G. 0 DI S=286.376 _Xy =426.452 CEMTROKP HIDDZN 153 13126 -0459 27791 OYSTER CRtEZE DRYI4ELL ANALYSIS OYlCH1O (NO $AND. Posr-Acc. Figure 3-14. Lower Drywell Circumferential Stresses - Post-Accident Case

NUEX0T4 REV. 0 Center of-_* Drywel ,SN Sphere P! Si A. -36o---\\ / anes of vmmetry Unbudded Shape Bucdded Shape Vent Radial Displacerment No Rotation ) Symmetric Buckleng of Drywegl - Ubucdded Shape Budded Shape ( Rotation i kNo Radai Dip. ) Vent Asymmetric Budding of Drywel SYM.DRW Figure 3-15. Symnetric and Asymetric Buckling Modes 3-35

• HSVS 4.4 1~N 1

l 990 13:23:92 >°T~tsi^ STEIP-- t TER-1 __C7=7.665 _U aoce81S Snx =9. m14 X -P YU =-0. a DIST196.059 = =32a7. 322 -S. 963147 -SA WQ tab OYSTER CMEW DRYN-LI. ANALYSIS @C!LLT CNO SArND ItETUECLING Fiure 3-16. Syffetric Buckling Mode Shape - Refueling Case

I Eddie~p~ a-4 DEC 11 1996 P8:21:S9 POST Ii5m STEP~l ITEA=1 FACT=1U.134 uxD NO4 L DSX =9 1B744 SHU =-g60152 SVU =0.S744 xV =1 IDISI=186. 6S9 XT =327.422 RT 277128 -0._. 01a2 E -_.35OR-03 09.0574 U. 66744 Ca Z.1 OYSTER CREEX DRYWIELL ANALYSIS - OYCRICC NO, SMN.

REJUELU, ASyH.

t I ^.

• _ £.-

_4_ D....C.1*.. B _U. Chi DA..1 4...

ANSYS 4.4 NOU 29 1G9* 01l:27 :14 STI=l iFO=5. *1L DMC =. 9U68 51..m =-a. "1173 SNX =U.Au67 XU =1 YU =-e.s DISI=196.959 Xr =327.422 -. 691.73 B41 724K-9a E 275K-Ba a a.01969

9. 9MS67 w

Finitra I_11q lReil Irine, Mn,4 Thi2ng. P>nt Arritint r3n

ORF= 00664 INDEX 9-4, REV. 0

4. ALLOWABLE BUCKLING STRESS EVALUATION Applying the methodology described in Section 2 for the modification of the theoretical elastic buckling stress, the allowable compressive stresses are now calculated.

Tables 4-1 and 4-2 summarize the calculation of the allowable buckling stresses for the Refueling and Post-Accident conditions, respectively. The modified capacity reduction factors are first calculated as described in sections 2.2 and 2.3. After reducing the theoretical instability stress by this reduction factor, the plasticity reduction factor is calculated and applied. The resulting inelastic buckling stresses are then divided by the factor of safety of 2.0 for the Refueling case and 1.67 for the Post-Accident case to obtain the final allowable compressive stresses. The allowable compressive stress for the Refueling case is 10.65 ksi. Since the applied compressive stress is 7.58 ksi, there is a 41% margin. The allowable compressive stress for the Post-Accident, flooded case is 13.77 ksi. This results in a margin of 15% for the applied compressive stress of 11.96 ksi. 4-1

DRF:- 00664 INDEX 9-4, REV. 0 Table 4-1 Calculation of Allowable Buckling Stresses - Refueling Case Parameter Value Theoretical Elastic Instability Stress, Oie (ksi) 58.10 Capacity Reduction Factor, a; 0.207 Circumferential Stress, ac (ksi) 4.49 Equivalent Pressure, p (psi) 15.74 "X" Parameter 0.173 AC 0.118 Modified Capacity Reduction Factor, aimod 0.402 Elastic Buckling Stress, ae = limod aie (ksi) 23.34 Proportional Limit Ratio, A - aelay 0.614 Plasticity Reduction Factor, Ai 0.913 Inelastic Buckling Stress, a; = T7i~e (ksi) 21.30 Factor of Safety, FS 2.0 Allowable Compressive Stress, aall c aj/FS (ksi) 10.65 Applied Compressive Meridional Stress, am (ksi) 7.58 Margin - [((i/am) - 1] x 1007 41% 4-2

NDEX 9-4, REV. 0 Table 4-2 Calculation of Allowable Buckling Stresses - Post-Accident Case Parameter Value Theoretical Elastic Instability Stress, aie (ksi) 61.95 Capacity Reduction Factor, a; 0.207 Circumferential Stress, a. (ksi) 20.08 Equivalent Pressure, p (psi) 70.38 "'X" Parameter 0.774 AC 0.195 Modified Capacity Reduction Factor, almod 0.529 Elastic Buckling Stress, ae = *l,mod aie (ksi) 32.74 Proportional Limit Ratio, A a ae/oy 0.862 Plasticity Reduction Factor, ni 0.702 Inelastic Buckling Stress, a c nriae (ksi) 22.99 Factor of Safety, FS 1.67 Allowable Compressive Stress, call i ac/FS (ksi) 13.77 Applied Compressive Meridional Stress, am (ksi) 11.96 Margin a [(ai/qm) 1] x 100% 15% 4-3

ORF.4 0O5&4 INDEX 9-4, REV. 0 S.

SUMMARY

AND CONCLUSIONS The results of this buckling analysis for the refueling and post-accident load combinations are summarized in Table 5-1. The applied and allowable compressive meridional stresses shown in Table ;-1 are for the sandbed region which is the most limiting region in terms of buckling. This analysis demonstrates that the Oyster Creek drywell has adequate margin against buckling with no sand support for an assumed sandbed shell thickness of 0.736 inch. This thickness is the 95% confidence projected thickness for the 14R outage. 5-1

QRF# 00664 INOEX 9-4, RCV, a Table 5-1 Buckling Analysis Summary -Load Combination Refueling Post-AcCident Service Condition Design Level C Factor of Safety Applied 2.00 1.67 Applied Compressive Meridional Stress (ksi) 7.58 11.96 Allowable Compressive Meridional Stress (ksi) 10.65 13.77 Buckling Margin 41% 15% 5-2}}