ML20052G638

Evaluation of Midland Nuclear Power Plant Borated Water Storage Tank for Nonuniform Support Loading Resulting from Ring Wall Settlement.

ML20052G638
Person / Time
Site:	Midland
Issue date:	03/31/1982
From:	Banon H, Campbell R, Hardy G STRUCTURAL MECHANICS ASSOCIATES
To:
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ML20052G625	List: ML20052G624 ML20052G638
References
NUDOCS 8205180530
Download: ML20052G638 (93)

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! SPA 13704.01-R001.1 EVALUATION OF MIDLAND NUCLEAR POWER PLANT BORATED WATER STORAGE TANK FOR NON-UNIFORM SUPPORT LOADING RESULTING FROM RING WALL SETTLEMENT l prepared for CONSUMERS POWER COMPANY Jackson, Michigan March, 1982 gf STRUCTURAL mECHRnKS RSSOCIATES W- a c i., c or e 5160 Bech Street, Newport Beach, Cahf. 92660 (714) 833 7552 8205180 630

!2 SMA 13704.01-R001.1 EVALUATION OF MIDLAND NUCLEAR F0WER PLANT BORATED WATER STORAGE TANKS FOR NON-UNIFORM SUPPORT LOADING RESULTING FROM RING WALL SETTLEMENT prepared by R. D. Campbell G. S. Hardy H. Banon Approved by:

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R. P. Kennedy / T. R. Kipp President Manager of Quality Assurance prepared for CONSUMERS POWER COMPANY Jackson, Michigan March, 1982 g g STRUCTURAL mECHAnlCS

-T ASSOCIATES A Coht Coro 516O BachStreet,NewportBeach, Calif.92660 (714)833 7552

DOCUMENT REVISION RECORD l

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Revision Approval Approval Date Description January,1982 y g%

SMA Report No. /~ M' 13704.01-R001 Initial Printing fg/fz, March, 1982 a) Addition - Pg. 1-1 M l./

b) Correction - Pg. 2-3 ~

> 3 04 -RO '.01 c) Correction - Pg. 2-6 d) Correction - Pg. 3-1 ff/g g , z.( -f "L--

e) Addition - Pg. Al-1 to AR-1 May, 1982 SMA Report No.

a) Correction - pages A2-1, A2-2, A2-3. A2-4, AS-3, M [h 13704.01-R001.02 AS-4 and AR-1 to reflect final SME loading on ggfg E~ fy n anchor bolts s

Note: New features, as well as changes, deletions, and addition to infonnation within this report, are indicated by bars in the margins or by a bar under the page number if the entire page is affected.

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TABLE OF CONTENTS Section Title Paj2t

r LIST OF TABLES .................. iii

" iv LI ST OF FI GU RES . . . . . . . . . . . . . . . . . .

EVALUATION OF BWST FOR PRE OPERATIONAL SETTLEMENT 1 INTF0 DUCTION . . . . . . . . . . . . . . . . . . . . 1-1 j-

~' 1-1 l .1 Statement of Problem .............

1.2 Description of BWSTs and Ring Walls . . . . . . 1-1 1.3 Purpose of Study . . ............. 1-4 ll 1.4 Scope of Work . . . . . . . . . . . . . . . . . 1-4 1.5 General Approach . . ............. 1-4 ll

2

SUMMARY

AND CONCLUSIONS , ............. 2-1

i; 2.1 Summary ................... 2-1 3 ACCEPTANCE CRITERIA . . . ............. 3-1

3.1 Governing Codes and Standards . . . . . . . . . 3-1 3.2 Stress Criteria for Settlement Loading . . . . 3-2 4 ANALYTICAL MODELS AND ANALYSIS METHODS . . . . .. . 4-1 4.1 Fini te Element F' del . . . . . . . . . . . . . . 4-1 4.2 Finite Element Model Loading . . . . . . . . . 4-3 if 4.3 Bolt Chair Model . . . . . . . . . . . . . . . 4-6 5 AN ALYTICAL RESULTS . . . . . . . . . . . . . . . . . 5-1 5.1 Results from Finite Element Model . . . . . . . 5-1 5.2 Bolt Chair Top Plate . . . . . . . . . . . . . 5-2

', 5.3 Tank Wall at Bolt Chair Location . . . . . . . 5-3 REFERENCES i

e-- , ~ . , _ _ , , . ,

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. . 1 TABLE OF CONTENTS (Continued)

Section Title Pajgt ADDENDA

' EVALUATION OF BWST FOR FUTURE SETTLEMENT AND SEISMIC MARGIN EARTHQUAKE Al INTRODUCTION . . . . . . . . . . . . . . . . . Al-1 A1.1 Statement of Problem . . . . . . . . . . Al-1 A1.2 Pu rpo s e o f S t udy . . . . . . . . . . . . Al-1 A1.3 Scope of Work ............. Al-1 A1.4 General Approach . . . . . . . . . . . . Al-2 A2

SUMMARY

AND CONCLUSIONS ........... A2-1 A3 ACCEPTANCE CRITERIA ............. A3-1 A3.1 Criteria For Bolt Chair And Vessel Wall In Tension . . . . . . . . . A3-1 A3.2 Criteria For lessel Wall In Com-pression ............... A3-2 i

A3.3 Criteria For Anchor Bolts ....... A3-3 A4 ANALYIICAL MODELS AND ANALYSIS METHODS . . . . A4-1 A4.1 Finite Element Model . . . . . . . . . . A4 A4.2 Finite Element Model Loading . . . . . . A4-2 A4.3 Bol t Chai r Model . . . . . . . . . . . . A4-4 A4.4 Seismic Loading Conditions . . . . . . . A4-4 A5 ANALYTICAL RESULTS A5.1 Results From The Future Settlen.ent Analysis . . . . . . . . . . . . . . . . A5-1 AS.2 Results From The Seismic Margin Study Earthquake Analysis .......... A5-2 A5.3 Total Response Calculations ...... A5-4 ADDENDA REFERENCES APPENDIX A ii 1

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l LIST OF TABLES

. Table Title Page EVALUATION OF BWST FOR PRE OPERATIONAL SETTLEMENT f

I-1 Pre-erection Elevation Data for the Borated Water Storage Tanks (1T60 & 2T60) Foundations . . . . . . . . 1-6 1-2 Ring Wall Elevations Taken on June 15, 1981 . . . . . . 1-7 1-3 Measured Loads in Bolts Anchoring Tank 1T-60 ..... 1-8

, 1-4 Measured Loads in Bolts Anchoring Tank 2T-60 ..... 1-9 5-1 Stresses at Center of Elements, Bottom Three Rows - PSI ...................... 5-6 5-2 Stresses at Center of Lower Edge of Element, Bottom Row - PSI ................... 5-9 5-3 Vertical Reactions into Shell - lbs. ......... 5-10 ADDENDA EVALUATION OF BWST FOR FUTURE SETTLEMENT AND SEISMIC MARGIN EARTHQUAKE Al-1 Predicted End-of-Life Settlement for BWST # 1T-60 . . ................... Al-3 A2-1 Tank Model With Celotex Stress Combinations - SME + DW + Settlement . . . . .. A2-3 A2-2 Tank Model Without Celotex Stress Combinations - SME + DW + Settlement . . . . . . A2-4 A5-1 Bolt Force and Gap Resultants for BWST Configuration Without Celotex Layer . . . . . . . . . . A5-5 A5-2 Bolt Force and Gap Resultants for BWST Configuration with Celotex Layer ........... AS-6 iii

e,r LIST OF FIGilRFS Figure Title Page EVALUATION OF BWST FOR PRE OPERATIONAL SETTLEMENT l-1 Plan View of Tank 1T-60 Identifying Bolt Numbers and Location Angle e . . . . . . . . . . . . . . . 1-10 1-2 Plan View of Tank 2T-60 Identifyf ag Bolt Numbers and Location Angl e e . . . . . . . . . . . . . . . 1-11 1-3 Comparison of Tank 1T-60 Ring Wall Relative Elevations before and after the Ground Settlement . . . . . . . . . . . . . . . . . . . . 1-12 2-1 Plan View of Tank Model ............. 2-8 2-2 Elevation View of Tank Model . . . . . . . . . . . 2-9 2-3 Displacement of Tank Bottom Relative to Ring Wall and Celotex Contours after Ground Settlement . . . 2-10 2-4 Compressive Loads at Tank Bottom . . . . . . . . . 2-11 4-1 Beam-on-Elastic-Foundation Model . . . . . . . . . 4-8 4-2 Linearization of Boundary Springs ........ 4-9 4-3 Effective Water Annulus . . . . . . . . . . . . . . 4-10 4-4 Effective Water Force / Unit Width of Circunfer-ence Vs Gap ................... 4-11 4-5 Force Vs Deflection at Boundary Elements . . . . . 4-12 4-6 Beam Model for Bolt Chair Design . . . . . . . . . 4-13 4-7 Yield Line Model for Bolt Chair .........

4-14 4-8 Analysis Model for Local Membrane Stresses in Shell Due to Anchor Bolt Loading . . . . . . . . . 4-15 ADDENDA EVALUATION OF BWST FOR FUTURE SETTLEMENT AND SEISMIC MARGIN EARTHQUAKE Al-1 Ringwall Settlement Curve for Tank 1T-60 End of Life Condition ... ........... Ai-4

,' Al-2 Elevation View of Tank Model (End of Life l Condition) ................... Al-5 i A5-1 Displacement of Tank Bottom Relative to Ringwall l and Celotex Contours After Ground Settlement . . . AS-7 i

t iv l

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c . . .... . . . . . . - -

1. INTRODUCTION f

1.1 STATEENT OF PROBLEM Soil settlement at the site of the Midlands Nuclear Power Plant I has resulted in deformation of the ring walls that serve as a supporting base for the Units 1 and 2 Borated Water Storage Tanks (BWSTs).

' Survey measurements of the ring walls indicate that the too surf aces have distorted from their original position. Visual examination has also indicated that some cracking of the ring walls has occurred. The ring i

wall deformation has resulted in a non-uniform support condition for the BWSTs. Examination of the tanks indicated that some of the anchor bolts

,l connecting the BWSTs to the ring walls were unloaded and some were loaded. During the initial installation, all bolts were lightly loaded.

Concerns have been raised that uneven support of the BWSTs may have resulted in yielding of the tank walls or that increased anchor bolt loading could have yielded the bolt chairs. Bechtel Corp. has designed a retrofit for the ring walls which would stiffen the walls and prevent further distortion. Shims would then be installed between the ring walls and tank bottoms to provide uniform support for the tanks. Prior to initiating the ring walls retrofits, the condition of the BWSTs must be assessed. Evaluation of the BWSTs in their current condition is the subject of the main body of this report.

Subsequent to the release of the main body of this report, an additional BWST study was undertaken. This additional study consisted of an evaluation of the tanks when subjected to projected end-of-life soil settlement conditions combined with Seismic Margin Earthquake loads. The description and results of this additional study have been integrated into this report as an Addenda.

1.2 DESCRIPTION

OF BWSTs AND RING WALLS There are two BWSTs in the Midland Nuclear Power Plant complex, one each for Units 1 and 2. The two tanks are indentical and are l l-1 l

v -

l cylindrical flat bottom storage tanks with umbrella-shaped roofs. They are 52 feet in diameter and 32 feet in height. The roofs are welded to ring girders at the top of the tanks. The tank walls are 0.375 inches thick for a height of 8 feet from the flat bottom and are 0.25 inches thick for the remaining 24 feet. The bottom is 0.25 inches thick. All materials are type 304 L stainless steel.

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The two tanks are located outdoors in the tank farm area, north of the Auxiliary Building. The Unit 1 tank, IT-60, is located on the west side of the tank farm and the Unit 2 tank, 2T-60, is located on the east side of the tank f arm. Tank details are shown on Graver drawings NL12046, Rev. 3, NL-12047, Rev. 2, and NL-12051, Rev. 2.

i Ring walls for the two tanks are identical except in the valve pit area. Unit I has a larger valve pit than Ur.it 2. The ring walls are detailed on Bechtel Drawings C-127 (Q), Rev. 6, and C-128 (Q), Rev. 7.

Figures 1-1 and 1-2 show a plan view of the two ring walls with anchor bolt locations identified. Prior to erection of the tanks, elevations at the top of each ring wall were verified by Graver, the tank fabricator. Table 1-1, taken from Reference 1, tabulates the measured elevations relative to a bench mark. Design elevations according to the Bechtel construction drawings were 635.04 feet for IT-60 and 635.12 feet for 2T-60. Af ter erection of the tanks and filling with water, soil settlement has occurred resulting in distortion of the top surf aces of the ring walls. Table 1-2, taken from Reference 2, depicts elevations measured at the top of each ring wall on 16 June 1981. Note that Table 1-2 depicts actual elevations wherein Table 1-1 provides elevations relative to a bench mark. From Tables 1-1 and 1-2 it is evident that the deviation from a plane surf ace is greatest on IT-60. Maximum deviation from a plane surf ace is about 1.2 inches for IT-60 compared to about 0.36 ,

l inches for 2T-60. Figure 1-3 shows relative caflections of the ring wall top surface from the initial top surf ace contour for IT-60. Note that the deflections are relative since the ring walls have settled to a lower elevation since initial contruction i.e., the surveyed contours have been l l

l 1-2 i

g . ._

adjusted in elevation so that at a conrion point, the elevations are i dentical . Relative deflections for 2T-60 are considerable less and are not the governing case.

During initial erection of the tanks, the anchor bolts were tightened snugly. Torque was not specified. Settlement of the ring walls has resulted in distortion of the ring wall top surfaces which support the BWSTs. As a result, several of the anchor bolts were observed to have gaps between the bolt chairs and nuts while several appeared to be loaded. In order to determine actual bolt loads, strain gages were applied by another contractor to the loaded anchor bolts and the nuts l were backed off to zero load. Tables 1-3 and 1-4 from Reference 3 indicate acutal bolt loads prior to backing off the nuts. As would be deduced from the measured e16vations of the ring walls supporting both tanks, the bolt loads are much higher in Tank 1T-60. Three of the bolt loads in Tank IT-60 exceed the f aulted condition design load of 20.43 kips from Reference 7. All other loads are within the original design load.

The tank walls are very stiff eith respect to remaining in a horizontal plane along the bottom surf ace and at points where the ring walls have settled out of the horizontal plane, the anchor bolts are loaded, trying to close up the gap between the tank walls and ring walls. At points where the tanks are resting on the ring walls, bolt loads are zero.

1 There is a 1/2 inch thick asphalt impregnated fiberboard

- (Celotex) between the tanks bottoms and the ring walls. The material is compressible and tends to distribute the tank wall loading to the ring wall in a more unifonn manner than if there were no compressible material at the interf ace.

The tanks are currently full of water and are resting on the ring walls with the anchor bolts all unloaded. The current unloaded anchor bolt condition is much less critical to the tanks than the prior condition with some of the bolts loaded, trying to force the tanks walls to the contours of the ring wall surfaces.

1-3

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Visual observations of the BWSTs with the bolt loads still applied did not reveal any obvious damage. Distortion from welding was apparent in the bolt chairs and there was no apparent difference between chairs that were loaded vs. chairs that were not.

1.3 PURPOSE OF STUDY The purpose of this study is to evaluate the present or worst condition for each of the tanks in order to determine if any yielding,

[ buckling or permanent damage has occurred due to soil settlement and ring wall distortion. The objective is to verify the BWSTs structual

, integrity prior to retrofiting the ring walls.

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' l 1.4 SCOPE OF WORK l

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The scope of work consists of performing a finite element analysis of the worst tank condition to determine stress conditions in

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the tank walls and bolt chairs caused by soil settlement and consequent ring wall distortion. Calculated stresses are to be compared to code based acceptance criteria or material yield strength to determine if pemanent defomation has occurred and to cylinder buckling criteria to determine if elastic or plastic buckling has occurred. In the event that yielding or buckling is predicted to occur, an assessment is to be made as to the possible detrimental effects that may result due to the anticipated level of inelastic strain.

1.5 GENERAL APPROACH A three-dimensional finite element model was constructed to represeni. the BWSTs cylindrical walls. The model was constructed of flat plate elements posessing both bending and membrane stiffness. Vertical boundary elements are utilized at the bottom surf ace to represent the

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nonlinear behavior of the asphalt impregnated fibe board between the tank and ring wall. The boundary elements have gap capability so that non-uniform support conditions, including gaps, can be properly represented.

Actual load-deflection relationships were determined by conducting material compression tests on samples of asphalt impregnated fiberboard used in the tank installation. Additional horizontal boundary elements 1-4

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k at the tank wall lower edge represent the stiffness and restraint offered by the flat tank bottom. The upper edge of the cylindrical tank is stiffened by a ring girder with properties chosen to represent the stiff ness of the tank roof and the restraint that it offers to the tank at the roof / cylinder intersection.

Loading conditions on the tank cylindrical shell, which are reacted by the ring wall, include dead weight of the tank roof, weight of the cylindrical shell and weight of an effective annulus of water plus measured anchor bolt loads. Other loading imposed on the model is the hydrostatic radial pressure acting on the tank wall.

Chapter 2 sunmarizes the overall results and conclusions of this study. Chapter 3 presents acceptance criteria used for evaluating settlement-induced tank stresses. Chapter 4 describes the tank model, boundary conditions, loading conditions and method of solution in detail. Detailed results are presented in Chaoter 5. Only the worst case condition representing IT-60 was analyzed since the worst case clearly did not result in any predicted excessive loading or permanent deformations in the tank walls, bolts or bolt chairs.

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TABLE 1-1 PRE-ERECTION ELEVATION DATA FOR THE BORATED WATER STORAGE TANKS (lT60 & 2T60) FOUNDATIONS

~-

ANGLE O ELEVATION (FEET) ELEVATION (FEET)

BOLT NUMBER (CLOCKWISE FROM NORTH) TANK 1T-60 TANK 2T-60 1 355.5 4.48 5.09 2 4.5 4.46 5.07 3 13.5* 4.a7 5.09 4 22.5' 4.47 5.09 5 31.5' 4.48 5.08 6 40.5' 4.48 5.07 7 49.5* 4.48 5.10 8 58.5 4.47 5.09

9 67.5* 4.48 5.07 10 76.5 4.47 5.07 11 85.5* 4.48 5.09 12 94.5 4.47 5.08 13 103.5* 4.48 5.08 14 112.5 4.48 5.07 15 121.5 4.48 5.10 16 130.5 4.48 5.08 17 139.5 4.48 5.09

., 18 148.5* 4.48 5.08

! 19 157.5 4.47 5.09 20 166.5" 4.48 5.08 21 175.5* 4.48 5.07 22 184.5 4.48 5.08 23 193.5" 4.48 5.09

, 24 202.5' 4.47 5.08 25 211.5* 4.48 5.08 26 220.5* 4.47 5.09 27 229.5 4.49 5.09 l 28 238.5* 4.48 5.08

- 29 247.5 4.47 5.07 30 256.5 4.47 5.08 31 265.5' 4.49 5.09

! 32 274.5' 4.47 5.08 33 283.5 4.47 5.08 34 292.5' 4.46 5.07 35 301.5' 4.48 5.10 36 310.5* 4.47 5.08 37 319.5* 4.48 5.09 38 328.5' 4.48 5.08 39 337.5* 4.47 5.09 40 346.5' 4.47 5.08 1-6

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TABLE 1-2 RING WALL ELEVATIONS TAKEN ON JUNE 15, 1981 ANGLE O ELEVATION (FEET) ELEVATION (FEET)

(CLOCKWISE FROM DUE NORTH) TANK iT-60 TANK 2T-60 0 634.86 634.94 30 634.87 634.93 60* 634.86 634.93 90* 634.86 634.93 120* 634.87 634.94 150* 634.85 634.96 180* 634.83 634.96 210' 634.78 634.95 240* 634.77 634.96 270* 634.79 634.96 300* 634.82 634.96 330* 634.86 634.95 C

i 1-7

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TABLE 1-3 MEASURED LOADS IN BOLTS ANCHORING TANK 1T-60 BOLT NUMBER LOAD (KIPS) BOLT NUMBER LOAD (KIPS) 1 0.0 21 0.0 2 0.0 22 0.0 I 3 0.0 23 0.0 4 0.0 24 0.0 i

5 0.0 25 0.0 6 0.0 26 16.44 7 17.83 27 31.31

, 8 14.02 20 16.10 9 21.32 29 10.13 10 22.51 30 0.02 11 16.46 31 2.32 12 0.0 32 0.0 13 0.0 33 -

0.0 14 0.0 34 0.0 -

15 0.0 35 0.0 16 0.0 36 0.0 17 0.0 37 0.0 18 0.0 38 0.0 19 0.0 39 0.0 20 0.0 40 0.0 I

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• BOLT LOCATIONS CORRESPONDING WITH THESE BOLT NUMBERS ARE LISTED IN TABLE -

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TABLE 1-4

!. MEASURED LOADS IN BOLTS ANCHORING TANK 2T-60 r

BOLT NUMBER LOAD (KIPS) BOLT NUMBER LOAD (KIPS)

-l 1 0.0 21 0.0

'l 2 0.0 22 0.0 3 0.0 23 0.0

.! 4 0.0 24 0.0 5 0.0 25 0.0 6 1.19 26 0.0 7 2.82 27 0.0 8 0.24 28 0.0 9 1.15 29 0.0 10 0.0 30 0.0 i 11 0.0 31 0.0 12 0.16 32 0.0 i

13 0.0 33 0.0 14 0.0 34 0.0 4

15 0.0 35 0.0 16 0.0 36 0.0 17 0.0 37 0.0 18 0.0 38 0.0 19 0.0 39 0.0 I 20 0.0 40 0.0

• BOLT LOCATIONS CORRESPONDING WITH THESE BOLT NUMBERS ARE LISTED IN TABLE 1-1.

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FIGURE 1-1: FLAN VIEW OF TANK 1T-60 IDENTIFYING BOLT NUMBER 3 AND LOC.ATION-AN.GLE'e:. , ,

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+0 f' TANK 2T-60 34 9 ae = 9' TYPICAL p,

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t FIGURE 1-2: PLAN VIEW OF TANK 2T-60 IDENTIFYING BOLT NUMBERS .

AND LOCATION ANGLE e 1-11

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• ANGLES ARE MEASURED CLOCKWISE FROM DUE NORTH PRE-ERECTION RELATIVE ELEVATION DATA FOR THE TANr' IT-60 RING WALL

- - - RING WALL ELEVATION DATA TAKEN JUNE 15, 1981 FIGURE 1-3. COMPARISON OF TANK IT-60 RING WALL REl.ATIVE ELEVATIONS BEFORE AND AFTER THE GROUND SETTLEMENT

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2. SUMt%RY AND CONCLUSIONS 2.1 SIM ARY The most critical case was that of tank 1T-60 where the ring wall top surf ace supporting the tank had the greatest deviation from a planar surf ace, see Figure 1-3. Since results of the evaluation of BWST IT-60 were positive, only the one tank was analyzed. Figure 2-1 is a I

plan view of the BWST finite element model and Figure 2-2 is an elevation q- view of the model. The BWST was modeled on the ANSYS computer program using 40 flat shell elements around the circumference and 8 elements f

along the vertical, resulting in elements that are approximately four feet scuare. The becm type elements shown in Figures 2-1 and 2-2 are boundary elements to represent the restraint of the cylindrical shell j afforded by the tank roof and the tank bottom. The gap-type elements shown in Figure 2-2 represent the nonlinear compressibility afforded by the asphalt impregnated fiberboard at the interface between the tank bottom and ring wall. ,

Figure 2-3 shows the resultant displa:ement of the tank bottom relative to the uncompressed fiberboard positions, ie, the 1/2 inch thick fiberboard resting on top of the distorted ring wall surface. The displacement plot incorporates the nonlinear deflection of the fiberboard and the linear deflection of the tank wall. Maximum compression in the fiberboard is 0.19 inches.

Figure 2-4 plots the compressive loads at the node points along the tank wall lower end. Zero load indicates a gap between the ring wall and the tank bottom plus fiberboard thickness. Note that a gap exists from about 31 degrees to 103 degrees and from about 193 degrees to 283 degrees measured clockwise from north. The tank is being supported'over about 198 degrees and being forced downward by anchor bolt loads in the regions of the gaps.

2-1 l

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.t i Maximum primary membrane stress intensity occurs in element number 392 which is in the first row of elements above the shell thickness change from 0.375 inches to 0.25 inches. Maximum stress intensity is calculated to be 12495 psi. Stress intensity is defined as two times the maximum shear stress in accordance with the applicable ASME code, Reference 5. Stress intensity is the appropriate value to compare to code allowables or yield strength. The tanks are constructed of SA 240. Type 304 L stainless steel and the allowable stress intensity for design and normal operating conditions is 15,700 psi. The minimum specified yield strength is 25,000 psi; thus the maximisn stress intensity is only 0.5 of yield. The components of stress that make up the maximum stress intensity are 10,571 psi hoop stress from hydrostatic pressure.

-1,923 psi axial stress from deadweight and support reactions and -85 psi shear stress due to the nonlinearity in support reactions. The most significant contributor to stress intensity is the hoop component which i results from hydrostatic pressure. Stresses are sununarized in Tables 5-1 and 5-2.

Now that the anchor bolts have been unloaded, the compressive stress component is reduced and the stress intensity is reduced. .

Checks were made for buckling using NAM developed buckling formulas, Reference 6 for axially and moment-loaded thin cylinders. The NASA formulas predict lower bound values of buckling stress considering experimental data as well as theory. Calculated critical buckling stresses for cylinders stressed uniformly along the axis and around the circumference are 5050 psi axial compression in the 0.375 inch thick wall and 2690 psi axial compression in the 0.25 inch thick wall. In the situation at hand, axial stresses are not unifonn around the circumference nor along the length. Each of these factors would increase the critical buckling stress. From Reference 6, critical axial stresses for buckling i of a cylinder in bending would be 7950 and 4750 psi for the 0.375 and 0.25 inch thick walls, respectively. These are considered to be more representative values for the BWST condition where axial stress varies from positive to negative around the circumference and is not uniform ,

along the length. l 2-2

=. _ _ _ . . _ _ _ . _ _ _ _ . _ _ _ .. . . ._ ___ . . _ . _ . ._

41 Maximum compressive stress occurs in element 392, a 0.25 inch j thick element, and is -1930 at the bottom of the element. At this point.

l the thickness changes to 0.375 inches for elements below. This is less than the calculated buckling stress for uniform axial r,ompression and I much less than the predicted buckling stress level for bending. The I buckling f actor of safety, based on elastic buc> ling for a shell in

? bending, is 2.46. In the current condition, the bolt loads have been relieved and compressive stresses have been reduced such that there is no l imediate danger of buckling in the event that further settlement and j ring wall distortion occur prior to retrofit of the ring walls. l Maximum anchor bolt loading measured in the field via strain f gaging was 31,313 pounds, Tables 1-3 and 1-4. Review of the tank design y report, Reference 7 reveals that maximum service loading is predicted for the safe shutdown earthquake event and is 20,427 pounds. Three anchor p bolt loads (Table 1-3) exceeded this value. The bolts are fabricated from ASTM A-36 steel with a minimum specified yield strength of 36 ksi.

• The anchor bolts are l' 1/2 inches nominal diameter. Maximurn anchor bolt stress is 22.32 ksi in the threaded area. The factor of safety on bolt i t yield is 1.61. It is therefore concluded, that no permanent defomation or damage has occurred in the anchor bolts. Anchor bolt pullout was I

checked and the capacity was calculted to be greater than 136 kips. The l

f actor of safety on pullout is 4.31. Note that the bolts are currently l unloaded. When the ring walls are retrofit, the ta:ks are shimed and bolts are retightened, the loading conditions cited above will not exist.

Original bolt chair design analysis was conducted by methods j contained in Reference 8. These methods am very conservative design methods and use beam theory as opposed to plate theory. Maximum design

i. stress occurs in the top plate of the bolt chair. Scaling the calcula-tions in the design mport for the maximum measured bolt load of 31,313 pounds and the actual thickness of the top plate, the elastically calcu-lated stress is 47.76 ksi and exceeds the minimum specified yield stress 1

of 25 ksi. This is a bending stress and a plastic hinge will form when yield is exceeded by 50%. Actually for the case of a strain hardening U

2-3

6L '

material-like stainless steel, the effective plastic hinge will not occur I until the elastically calculated load controlled stress is greater than 150% of yield. In addition, stainless steel yield strength is typically much greater than minimm specified yield.

A yield line (limit) analysis was conducted to determine the limit load for a fully plastic condition to occur. The calculated yield

line analysis limit load based on minimisn yield properties is 39,950

'j pounds. None of the measured bolt loads exceed this value. The minimtsn f actor of safety for the maximum measured bolt load is 1.28. The

calculated anchor bolt load at first yield of the top plate, using the design fomula of Reference 8, is 16.38 kips. There are four bolts that

.. exceeded this value by more than a few percent and three bolts that am

[ essentially at this value. In Chapter 3, an acceptance criteria is

} developed that would allow settlement induced stresses to be comparable to stresses normally encountered in hydrostatic tests of pressure vessels without further consideration of inspection. For the case at hand, the

$ stress in the top plate of the bolt chair is classified as primary bending and in accordance with the acceptance criteria of Chapter 3, the allowable stress could approach 1.35 Sy when calculated on an elastic i basis. The allowable bolt load based on this criterion would be 1.35 f (16.38) = 22.11 kips. Only bolts 10 and 27 exceed this value, bolt

. number 10 exceeding the value by only 1.8%. The only concern should then I be addressed to bolt location 27 where the elastic analysis acceptance criterion is exceeded by 41.6%. In this case, limit analysis acceptance

{ criteria are applied.

The yield line (lower bound limit analysis capacity) was deter-mined to be 39.95 kips. As discussed in Chapter 3, the limit anlaysis acceptance criterion was derived from ASME code philosophy ar.J is 0.8 of

the lower bound limit analysis capacity. The maximum measured bolt load of 31,313 pounds is 0.78 of the calculated limit load capacity and the bolt chair is considered acceptable as is.

2-4

s .r ..  ;

t i

l' i

Since scrne yielding could have occured in the holt chair at bolt location 27, it is suggested that a dye penetrant examination be condu,,;ed of welds that attach the top plate to the gussets and tank wall c to assure that plastic straining has not opened up any surface cracks.

If cracks are not found and/or repaired, the bolt chairs should be considered acceptable.

i Local stresses in the tank wall due to bolt chair reactions were

.g evaluated and found to be acceptable. The original design analysis, Reference 7. utilized a method contained in Reference 8 to compute local j membrane plus bending stresses in the tank wall resulting from anchor bolt loading on the bolt chairs. Using the design formula in Reference 7, the combined membrans plus bending stress in the tank wall was I computed to be 1.93 times the minimm specified yield strength for the maximum measured bolt loading. The governing ASME code design criteria,

Reference 5, does not limit secondary stresses. Since local shell bending stresses are considered secondary by the ASME Code, only local membrane stresses are of consequence for evaluating whether any gross yielding has occurred. Reference 8 does not provide a method for calcu-lating local membrane stresses only; thus, the methods of Reference 9 were utilized. The top plate of the bolt chair was conservatively assmed to transmit an outward radial load from the shell equal to the applied bolt chair moment divided by the height of the bolt chair (Figure 4-3). This is a conservative assumption since contribution from the bolt chair gussets are ignored in the analysis. The resulting local membrane stresses are 13.2 ksi in the hoop direction and 12.16 ksi in the axial direction. Both stresses are tensile. Other stresses resulting from hydrostatic pressure and downward bolt chair loading are 6384 psi hoop tension,1614 psi axial tension and 159 psi shear stress. Combining all the stress components, the maximum stress intensity is 19.59 ksi. This is below the 25 ksi minimum specified yield strength of the material and significantly below the 33.75 ksi acc.eptance criterion of Chapter 3 for primary local membrane stress intensity. 3 1

2-5

h=, ,

I l

t

.I Secondary stresses are not limited by the governing code, Reference 5. Primary plus secondary stresses should be limited to 2 S y

for cyclic load applications to demonstrate shake down. In the case at hand, the loading is a single applied load and primary plus secondary 1 stresses calculated by the design fomula of Reference 8, when combined with stresses induced by other loading are 1.93 yS ; thus, shell stresses are within acceptable limits.

l The following summarizes the most severe stress condition in tank 1T-60 prior to relieving the anchor bolt loads.

a. Maximum primary membrane stress intensity occurs in the 0.25 inch thick shell about 10 feet from the base and is 12.5 ksi compared to a yield strength of 25 ksi and'an ASME code primary membrane stress allowable of 15.7 ksi for design conditions. The factor of safety on yield is 2.0.

b. Maximum membrane compressive stress occurs 8 feet above the base in the 0.25 inch thick shell and is 1930 psi compared

, to a lower bound elastic buckling stress of approximately 4750 psi. The resulting factor of safety on elastic buckling is 2.46. l

c. One bolt chair at bolt location 27 may have yielded to a small degree. The f actor of safety on collapse, based on a lower bound limit analysis, is 1.28. This meets the accept-tance criterion derived in Chapter 3. For this one bolt location, it is recomended that a dye penetrant inspection be conducted for welds attaching the bolt chair top plate to the gussets and tank wall to ensure that localized inelastic strains have not resulted in cracking at the weld joints.

l 2-6

.r -

i

/

i

d. Local membrane stress intensity in the tank wall at the maximum loaded bolt chair is 19.E9 ksi. The factor of safety on ASME code minimum specified yield is 1.28.

Primary plus secondary stress intensity is 1.93 Sy which i

is less than the shakedown acceptance criterion of two times yield.

It should be concluded from this study that tank 1T-60 is ti acceptable as is subject to dye penetrant examination of bolt chair top plate welds at bolt location 27. Ring wall deformation and resulting anchor bolt loading in tank 2T-60 is significantly less than IT-60 and tank 2T-60 is considered acceptable by comparison.

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• DISPLACEMENT RELATIVE TO THE AS BUILT TOP OF THE CELOTEX IN THE UNLOADED CONDITION RINGWALL CONTOUR AFTER GROUND SETTLEMENT

• * * *** TOP OF CELOTEX CONT 0UR IN THE UNLOADED CONDITION AFTER GROUND SETTLEMENT (UNLOADED CONDITI0fl =

1/2 INCH ABOVE RIllG .

---

TANK BOTTOM CONT 0UR IN LOADED CONDITION AFTER GROUND SETTLEMENT WALL CONTOUR)

FIGURE 2-3. DISPLACEMENT OF TANK BOTTOM RELATIVE TO RINGWALL AND CELOTEX CONT 0URS AFTER GROUND SETTLEMENT

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FIGURE 2-4. COMPRESSIVE LOADS AT TANK BOTTOM (SEE TABLE 5-3)

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3. ACCEPTANCE CRITERIA r

!k 1

I 3.1 GOVERNING CODES AND STANDARDS Governing codes and standards are delineated in the design

,l specification, Reference 4. The BWSTs are designed and code stamped to the ASME code,Section III, Nuclear Power Plant Components, Subsection NC, Class 2 Components, Paragraph NC3300, Design of Vessels. The 1974 l{

code with no addenda are applicable and Code Case 1607-1 is applicable fo: upset, emergency and f aulted condition stress allowables. The API 650 code is also specified for design, Reference 10. In cases of

., conflict, the ASE Code governs.

)

The basic design is conducted using API-650 criteria since

-4 NC3300 of the A':E code does not specifically address flat bottom storage tank designs. NC3800 does provide criteria for flat bottom storage tanks

'l and is essentially identical to API-650. The ASME code stress acceptance criteria from code case 1607-1 is used for evaluation of the OBE and SSE events.

, Under the governing criteria the following stress intensities t are allowed.

} Loading Primary Local Membrane plus Condition Primary Membrane, Pm Primary Bending, PL+Pb Design and Nomal S 1.5S i Upset 1.15 1.65S Emergency 1.55 1.8S Faulted 2.05  : 2.4S

{;

Testing 1.255* 1. 87S**

Not to exceed 0.9 Sj

• • Not to exceed 1.35 Sy 0

3-1

i No. .

e f

S is the allowable stress intensity of 15.7 ksi for 304 L l stainless steel. Secondary stresses do not require evaluation for Class

! 2 components designed by rule (NC3300 criteria). Minimum specified yield

, strength is 25 ksi. It can be seen from the allowable stress criteria

that any stress in excess of 1.59 S exceeds the minimisn specified yield strength and that primary local membrane and pricary bending stress allow-t ables all exceed the yield strength for all service conditions except design and normal.

f -

3.2 STRESS CRITERIA FOR SETTLEMENT LOADING Loading due to soil settlement is considered a one time application event that will be relieved when the ring wall foundations

. are retrofit and the tanks are shintned to provide uniform support. There

} is no clear analogy to operating service conditions, however, the testing event would be the most applicable as acceptance criteria if we were to choose a similar stress category for which no further examination would be mquired. If the higher emergency condition allowables were used, some j additional inspection may be warranted to assess deformation.

Flat bottom storage tanks are not hydrotested, but the basic ASME code philosophy regarding stress and permanent strain limits during i pressure vessel hydrotesting is considered a rational philosophy for I setting a conservative stress acceptance criteria for comparison to elastic stress analysis results. If elastically calculated stresses due to soil settlement do not exceed code acceptance criteria for testing conditions, then the component should be considered acceptable with no

[ additional examination mquired.

i

[

ASME code philosophy for hydrostatic testing is to test the  ;

vessel at 1.25 times design pressure (fomerly the hydrostatic test pressure was specified to be 1.5 times design pressure). Primary

, membrane stress intensity for design pressure is limited to S for Class 2 vessels and the primary local membrane plus primary bending stress intensity is limited to 1.55. At a hydrostatic test pressure of 1.25 times the design pressure, the implied limit is 1.255 for primary l

l 3-2 l

I

~. .

9 membrane and 1.875 for primary local membrane plus primary bending. A

- limit is placed on yield for those cases where the allowable stress intensity. S, can be as high as 0.9 times yield. The limit is set at 0.9

., times yield for primary membrane stress intensity and 1.35 times yield

i for primary local membrane plus primary bending stress intensity for cases where the primary membrane stress intensity is less than 0.67 times yield.

l

~j The ASME code philosophy in placing stress limits on hydrostatic testing induced stresses is to guard against gross deformation. The same

( philosophy is consi^ red applicable to the BWST settlement problem. If I stress intensities from settlement combined with hydrostatic pressure, weight and anchor bolt loading do not exceed hydrostatic testing

f allowables, then no detrimental effects am considered to have occurred.

l' Secondary stresses are not considered to be detrimental due to the single loading applications and are not considered to govern for

acceptance criteria. This is consistant with ASME code design pholosophy I

for class 2 and 3 components and even for Class I components for infmquent (emergency) events. Note also, that if secondary stresses ,

I were limited to the shakedown mgime (2 yS ) that during hydrostatic testing at 1.25 times the design pressure, primary plus secondary stress

I intensity could reach 2.5 S y. Thus, code philosophy would allow primary plus secondary stress intensity to exceed the shakedown limit for single i or limited numbers of events.

t l Amas where the above stress acceptance criteria may be exceeded should be treated on a case-by-case basis by applying ASME code limit

!; analysis acceptance criteria, going to emergency condition allowables with same additional inspection requirements or limiting inelastic strains, i

l 3-3

s, ,

l 1

i Them is only one case where calculated stress intensities

, exceed the recommended allowable stress limits. This is in the bolt chair top plate for the most severely loaded bolt (bolt location 27). In

, this case, limit analysis concepts are employed. The ASME code for Class ,

i 1 components and component supports, References 12 and 13, allow limit analyses to be used in lieu of meeting elastic stress criteria for l primary local membrane and primary bending stress intensities (Paragraphs NB 3228.2 and NF 3224 (a)). The bolt chair top plates are considemd to be plate and shell-type component supports and limit analysis concepts lf from the component support code, (NF 3224 a) am considered to be appli-j cable.

s

, Under Level C Service (Emergency Conditions), Reference 13 allows 1 0.8 of the lower bound collapse load. The same allowable is also specified for Class I components (Reference 12). Design philosophy for j( Level C Service Condition allowables is that some pemanent defomation may be experienced but the component is still serviceable. Deformation i limits may be specified if defomation is a controlling f actor for function. In the case of BWST bolt chairs, deformation is not a limiting I f actor and a limit analysis acceptance criteria analogous to that for

-'>l Level C Service for Class 1 plate and shell-type component supports is considered a valid concept for evaluation of settlement induced loading i on bolt chairs. As long as the bolt loads do not exceed 0.8 times the lower bound collapse load of the bolt chairs, they are considered to be f acceptable without further analysis or mtrofit being required. It is suggested, however, that for the one chair where the maximum bolt load occurs and is close to the recomended acceptance criteria, that a dye penetrant examination be conducted of fillet welds that attach the bolt

( chair top plates to the tank wall and gussets. This will ensure that any plastic defomation that may have occurred did not initiate any cracking.

Because of uneven support conditions, axial compressive stresses exist in the tank wall. For the large diameter thin wall storage tank, buckling will occur in the elastic range. The ASME code buckling criterion for axially loaded cylinders nominally contains a safety factor 3-4

. . 1 I

or 3 f or sustained design loads. Since tne condition under consioeration is local and is more of a strain controlled condition than load f, controlled, a more liberal buckling criteria is reconinended.

Reference 6 provides thin shell buckling formulae modified to reflect extensive test data. Fomulae are provided for the case of l

uniformly axially loaded cylinders and cylinders subjected to bending moment wherein the axial compressive stress peaks at one location. The formulae contain correction factors based upon experimental data and are

'i considered to be lower bound fomulae. Since the axial compressive stresses are local and vary from compression to tension around the circum-f ference, the bending fomula for buckling is considered to be more appro-pri ate. For the geometry under consideration, the applicable formula is:

0.6 y Et cr R where 1 Y = correction f actor for cylinders in bending E = Young's modulus

} t = shell thickness

• R =

mean radius of shell 1

For the bending case and the geometry under consideration:

i

! i =

0.35 for the 0.25 inch thick shell y =

0.39 for the 0.375 inch thick shell i

f Critical buckling stresses are: _

4750 psi for the 0.25 inch thick shell 7950 psi for the 0.375 inch thick shell 3-5

1 4 , ,

For strain controlled conditions the allowable axial compressive stress should be limited to acr/1.67 resulting in allowable axial I, cunpressive stresses of:

f 2845 psi for 0.25 inch thick wall 4760 psi for 0.375 inch thick wall.

In the event that the above allowables are exceeded, a geometry l check should be made to determine if elastic buckles have actually I

occurred. It should be noted that if elastic buckles do occur in a strain controlled condition, they will spring back upon removal of the applied loading. After the anchor bolt nuts have been loosened, part of

, the loading condition that could potentially have caused buckling was

} removed. Currently, the anchor bolts are unloaded. During a site visit

! in September, while the anchor bolts were still loaded, there was ne visual evidence of any buckling in the most critically loaded tank (IT-60). Subsequent analysis also demonstrates that buckling would not i occur under the prior conditions.

I

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1 3-6

.l

4. ANALYTICAL MODELS AND ANALYSIS METHODS i A finite element model was contructed to represent the cylindrical shell portion of the BWSTs. Boundary conditions were applied

.f at the top and bottom of the model to represent the restraint offered by the unbrella roof, the tank bottom and the asphalt impregnated fiberboard i

between the ring wall and tank bottom. Analysis of bolt chairs and the tank wall adjacent to the bolt chairs was conducted by hand methods using empirical design analysis approaches and classical stress analysis techniques.

j 4.1 FINITE ELEMENT MODEL The tank wall finite element model was constructed for the ANSYS

, computer program (Reference 11). Plan and elevation views of the model are shown in Figures 2-1 and 2-2. ANSYS element 63 was used to represent

-i the tank wall. Element 63 is a quadrilateral flat thin shell element with both membrane and bending stiffnesses. There are 40 elements around the circumference and 8 elements along the vertical axis. The elements

,I are almost square being 49 inches X 48 inches. Each tank has 40 anchor bolts and, thus, the choice of 40 elements around the circumference.

-I Anchor bolt locations correspond to nodal points along the bottom of the model .

i it The first two rows of elements from the bottom are 0.375 inches thick and the remaining six rows are 0.25 inches thick to represent the nominal tank wall thickness. At the top of the tank, beam elements are

placed around the circumference of the tank to keep the tank round at the i

roof / tank wall junction. The unbrella roof is welded to a ring girder at the tank top and would keep the tank wall from ovaling under asynenetric loading.

At the tank bottom two types of elements are used to represent boundary conditions. Beam elements (element 4) are placed radially outward from the tank to represent the radial restraint of the tank wall 4-1

bs .

afforded by the 0.25 inch thick tank bottom, the rotational stiffness of the tank bottom and to restrain the model from rotation about the vertical centerline of the tank. Vertical gap elements (element 10) are placed along the tank bottom to mpresent the stiffness of the asphalt impregnated fiberboard and any gaps that may exist between the tank l bottom and ring wall. The gap elements have linear force displacement f

characteristics in the compression direction and zero load capacity in

}.

tension.

The actual asphalt impregnated fiberboard (Celotex) is

,g nonlinear. Force-displacement properties were determined by laboratory b

testing of material taken from the site. Two tests were conducted to

{. detemine force-displacement relationships. One specimen was 3" X 3" and j one was 3" X 6". The laboratory test report is included as Appendix A.

( Results from both tests were similar and 3" X 3" specimen results were l used in developing an appropriate model for the boundary elements.

e In order to model the force-displacement properties of compressible asphalt impregnated fiberboard, a beam on an elastic founda-tion model was used to find an appropriate equivalent linear stiffness

{ for the boundary elements. The beam on an elastic foundation model is shown in Figure 4-1. Figure 4-2 compares a plot of the beam on an elastic

foundation model results versus the equivalent linear stiffness used in the computer model for boundary elements. In developing the besn on an elastic foundation model, the resulting stiffness was biased on the conservative (stiff) side. Stiffer boundary elements result in the 3

reaction loads being concentrated in more localized regions. Asstanptions used in the beam on elastic foundation models are described as follows.

Referring to Figure 4-1, a rigid boundary is asstsned on the tank bottom at the junction of the Celotex and the oil impregnated sand. This tends to be conservative in that the oil impregnated sand is assumc.1 to be rigid, thus, forcing all defomation in the tank bottom to occur in a relative short span of 6.81 inches, the distance from the tank inside wall to the inner edge of the Celotex.

4-2

Several beam on elastic foundation solutions were carried out in order to include the nonlinear force-deflection characteristics of the Celotex. Equivalent linear springs were asstsned for the Celotex material for discrete displacement ranges. For larger displacements, the equivalent spring rates were increased to approximately follow the force-deflection relationship determined in laboratory tests (Appendix A). For each displacement increment, the equivalent linear spring stiffness was selected to be the best fit of the nonlinear force-displacement relation-

' ship of the Celotex.

i

Referring to Figure 4-1, for a given force P, water pressure, w, i

shell rotational stiffness, k e, and Celotex effective stiffness, k, a dis-placement directly under the load P was computed. Figure 4-2 plots four p points determined by this method. Maximum displacement was determined to

{ be less than 0.26 inches; therefore, a best fit linear stiffness for the

! three beam on an elastic founation solutions for 0.26 inches and less dis-placement was found. This equivalent linear stiffness of 3800 lbs/in/in of circumference was used to develop appropriate stiffnesses for the vertical boundary elements that represent the Celotex force-displacement characteristics. ,

4.2 FINITE ELEMENT MODEL LOADING Loading conditict present in the BWSTs are:

Dead weight of the tank roof Dead weight of the tank wall Anchor bolt loading Hydrostatic pressure acting radially on the tank Hydrostatic pressure acting downward on an annulus of the bottom plate adjacent to the tank wall.

Loads were applied to the model at nodal points that would best represent the physical forces acting on the tanks.

4-3

hs Tank roof weight was applied at the top of the model and was uniform around the circumference, i.e., equal loads at each of 40 nodal points. Dead weight of the tank was applied by specifying a lg downward gravitational load. The program computes the proper nodal forces from the specified geometry, material density and g loading. Anchor bolt

, loading, tabulated in Table 1-3 was applied at nodal points along the bottom of the tank wall finite element model. P.sdial hydrostatic

, pressure was specified to be acting on the face of each shell element.

Pressure must be specified as constant on element f aces and the average pressure was used for each row of elements. In order to eliminate any fictitious bending stresses that could be computed due to use of flat plate elements to mpresent a curved shell, an option in the ANSYS code

., was used to transfer pressure loads on the element f ace to nodal forces.

c A pressure analysis check case was conducted to verify the model geometry I

and assure that hoop stresses and displacements were properly computed.

Most of the tank bottom rests on soil. There is an effective

, annulus of water adjacent to the tank wall that is not supported by soil.

The resultant shear force on the tank bottom that is transferred to the tank wall and ultimately to the ring wall was derived and applied as nodal forces along the tank bottom.

5

l. In areas where gaps exist between the tank bottom and ring wall, I the weight of an effective annulus of water that extends radially inward beyond the inside ring wall boundary is carried partially by the tank wall. Figurk 4-3 shows the effective radial length, or, of this annulus of water. The effective annulus radial length, ar, and the resulting vertical reaction load, P, at the tank bottom /shell junction are a function of the gap,6, between the tank bottom and the free surface of the Celotex. Figure 4-4 plots the reacton, P, as a function of 6. The P versus 6 plot was made using classical equations developed for seismic design of flat bottom storage tanks. Loading due to effe :tive water weight is, thus, nonlinear, depending upon the gap between the tank bottom and its support. For the range of gaps calculated in the final 4-4

.e- ...

solution, a weighted average of water weight was less than 110 lbs/in/ inch of circumference. The final computer run conservatively used a water weight equal to 120 lbs/in/ inch of circumference.

In regions where the tank bottom is supported, the effective annulus of water carried by the tank support is less than at gap locations . In order to avoid several iterative solutions that would mquim changing the effective water weight at points where the tank bottom was stoported, the vertical spring boundary effective stiffness included the effect of the 120 lb/in/ inch of circumference water weight.

[ Consequently, a conservative constant value of water weight could be used without requiring iteration. Figure 4-5 portrays this pictorially. At

. points where the tank is in contact with the Celotex, the effective water

( weight of 120 lbs/in/in of circumference corresponds to a compression in the Celotex of 0.010 inches. Using a linear boundary element stiffness of 3800 lbs/in/in of circumference, the 2ero load point corresponds to a downward displacement of -0.021 inches. In other words, the starting

] point of compression in the Celotex is shifted 0.021 inches and all gaps are shifted 0.021 inches. The computer program boundary element mactions carried into the shell are correct and resulting shell stresses -

are correct. However, computed displacements m1ative to the starting point are 0.021 inches greater than actual and the sisn of the vertical i

reactions is greater than actual reactions due to fictitious water weight being applied at points where the tank is supported, i.e., the excess water weight and excess mactions at the points of tank support cancel each other out and am not reacted by the shell.

Actual effective water weight is computed after completion of the finite element analysis and is equal to the unsupported circumfer-ential distance times the water weight of 120 lbs/in plus a fraction of this weight at c. few points where the displacement is so small that the full water weight cannot develop. This occurred at four boundary elements. At points where the tank is supported, the water weight is reacted directly into the soil and ring wall and them is no effective loading on the tank wall.

l l 4-5

9-I s

Weights of the tank, roof and internal bracing were taken from r tank drawings. Tank shell weight applied to the model was checked against weights on the tank drawings and a total force balance was conducted to assure correct input for loadings. The relative e

contributions of downward loading from dead weight, anchor bolt loads and

, vertical water pressure loading macted by the shell are:

l.

Tank Roof 36,000 lbs Tank Shell and Hardware 67,090 lbs j Effective Water Weight Reacted by Shell 115,045 lbs

' Total of Bolt Loads 168,470 lbs

,_ 386,605 T5s y '

, Froh the vertical force sunnary, it can be seen that the anchor

, ( bolt loads are 43.5% of the total vertical load. Now that the anchor bolt loads have been released, the total downward forces are considerably lower than the case evaluated.

i 4.3 BOLT CHAIR MODEL Bolt chairs were evaluated by hand calculations. Initially, the l method of Reference 8 used in design of the bolt chairs, was used to i

determine stress response in the top plate for a maximurn bolt load of 31,310 pounds. The method assumes that a beam of width equal to the edge

! distance from the hole to the plate outside edge carries 1/3 of the total bolt load. Figure 4-6 shows the analytical model.

L Since this model resulted in yielding of the top plate for some I

of the measured bolt loads in Tank 1T-60, a yield line analysis (limit analysis) was conducted to develop a limit load capacity. Figure 4-7 is the yield line model. Minimum limit load capacity was calculated to be l

39.95 kips which is above the maximum measured bolt loading of 31.3 kips by a f actor of 1.28. The acceptance criterion developed in Chapter 3 requires a safety f actor of 1.25 on the lower bound collapse load, thus,

, the criterion is satisfied.

l 4-6

-y 1500 -

k = variable m Beam on elastic f

$ foundation respons

[C 1000 - E = 3800 f/in l 5 '

/

b ' k = 1000 #/in/in E E = 800 f/in/in l 500 - E = 800 f/in/in 2

l /

i

. . i 0.1 0.2 0,3 DEFLECTION-INCHES I

E= Average spring rate at asphalt impregnated fiberboard over finite deflection range, lbs/in/in/ unit width E= Average spring rate of tank support, lbs/in/ unit width FIGURE 4-2. LINEARIZATION OF BOUNDARY SPRINGS 4-9

W.

e 3

N-

= or  ;

i .

P C Tank Wall Tank Bottom

, w II  %

r-

'l if y ip 1,

, p

'f if p p Asphalt Impregnated Fiberboard 011. . Impregnated Sand Ring Wall

--e I

w = water force, lbs/in/ unit width P = tank and roof weight, lbs/ unit width 6 = gap between tank bottom and asphalt impregnated fiberboard, in.

o = effective width of water innulus, in.

r FIGURE 4-3. EFFECTIVE WATER ANNULUS 4-10 i

Q

, g ..

s ., ~

4

., s , c v. .

s .  : '

i -

1,' -

n I -

,ss -

s

? ,

s, 4 ,

150- - ,

s' l s, <

, .s b, 120 . _ __ ._ ,

F

, e-.  :

z 100 .

4 D

N v,

m ai -

, i ww Ck:

4 o

' 50 -

a:

W *

>= '

3 -

w  :

M

~* '

D w

I I _ s e I ' '

g 0.2 0.4 ~0.6 0.8 1:0 -l w .

~

- Gap - Inches s l

> , s i

~'

l l FIGURE 4-4. EFFECTIVE W/JER FORCE / UNIT WIDTH OF CIRCUMFERENCE VS GAP 1

s s s k s

(

• i, 4 g l

l 4-11 l

. t.,

l l

l 800 - -

E!

a

• l .

$ 600 -

3E Y = 3800 #/in t;

E

? D U E 400 -

i e 5

p U

a:

200 -

120 . __

l -

/  :

e a a e a a

-0.021 0.01 0.05 0.1 0,15 0.2 0.25 Compression of Boundary Elements, Inches FIGURE 4-5, FORCE VS DEFLECTION AT BOUNDARY ELEMENTS

't

. . l

\ TANK WALL l I

'M I

\

,  % }

n f

4 6 d+

1/3 ANCHOR BOLT LOAD j

U g u i.- vv v

\

c 1

\

h Ak AL I l PARTIALLY FIXED I

I ls, W z 9  :

FIGURE 4-6: BEAM MODEL FOR BOLT CHAIR DESIGN 4-13 '

, . ~ . , ,

( ( ( '

b__ , -. b - ,

1 3 3

. i, ,

+ $

p

/\ s ij

/

s'  %,

\

l a  :.

l nc l 3

?

,= =

=

)

\

/

1. eff g

/

/

s s /

% /

k l

P P l 2 i =-r,ff T

-6 V "

Gusset l y e 3 r 6 Load I h ] r l Plate I i T N L ~u#

t ~. l ~ -a FIGURE 4-7. YIELD LINE MODEL FOR BOLT CHAIR 4-14

P 1 e =

W' U 1 I Le -

h A

', 4, t

._ 0.625" h

1. + 0.375" 312"R h

P = ANCHOR BOLT LOAD b ' s If h -

l FIGURE 4-8. ANALYSIS MODEL FOR LOCAL MEMBRANE STRESSES IN SHELL DUE TO ANCHOR BOLT LOADING 4-15

I

5. ANALYTICAL RESULTS 2

Results from the finite element model analysis and hand calculatior.s are sumarized in this chapter. Computer output is too voluminous to include, consequently, only important highlights of the output are sumarized.

5.1 RESULTS FROM FINITE ELEMENT MODEL Important stress results from the finite element model are from the bottom 3 rows of elements representing the lower 12 feet of the tank. The highest loading occurs in the bottom row of elements where the i shell thickness is 0.375 inches. Since the bottom elements are restrained frm radial movement by the tank bottom, hoop stresses in the second row of elements are greater than for the bottom row. At the junction of the second to third row of elements, the tank wall thickness changes from 0.375 inches to 0.25 inches. Maximtrn stress intensity occurs in the third row of elements at element 392 where the hoop stress is tensile and negative axial compressive stress exists. Above row 3, element stresses diminish with decreasing hydrostatic pressure and a more even distribution of the asymetric vertical reaction loads.

Because of the wave front solution techniques used in ANSYS, i

elements in a column are in numerical suequence while elements in a row (around the circumference) are numbered in steps of 22 (Figure 2-2).

Table 5-1 tabulates the three stress components and resulting stress intensity for the first three rows of elements beginning at row 3 and working down, row 1 being the bottom row.

Stress output is given at the center of each element and represents the average stress in the element. In addition to output at the center of each element, stress output at the bottom edge of each element was requested. Table 5-2 sumarized stress components for each 5-1 l

i

element along the tank sall bottom. Along the bottom, tank hoop stresses are low due to radial restraint offered by the tank bottom. Axial stresses are more concentrated along the tank bottom resulting from concentrated anchor bolt loading and uneven vertical mactions from the distorted ring wall.

i Vertical reactions carried into the shell along the bottom row of nodes (nodes 321 through 360) are tablulated in Table 5-3 and were visually displayed in Figure 2-4. Displacements of the tank bottom and surf ace of the asphalt impregnated fiberboard were previously plotted in Figure 2-3.

5.2 BOLT CHAIR TOP PLATE Figure 4-7 shows the yield line model used to evaluate limit load capacity of the top plate of the bolt chairs. In this model,

{ plastic hinges were assmed to form in the 0.5 inch thick gusset plates,

[ the 0.375 inch thick tank wall and the 0.625 inch thick top plate. For a 25,000 psi minimum yield material, and assning elastic-perfectly-plastic material behavior, the plastic hinge moment capacities are computed to be:

1 2440 in-lbs/in. in the .625 inch thick top plate 1560 in-lbs/in. in the 0.50 inch thick gussets 879 in-lbs/in. in the tank wall.

Work-energy relationships were used to solve for the limit load capacity. The angle a in Figure 4-2 was varied to determine a minimum value of capacity. Resulting values at minimum capacity are:

.l, a = 56.80 P

limit = 39.95 kips

. The limit load is a factor of 1.28 greater than the maximum measured bolt load of 31.3 kips and meets the acceptance criteria for limit analysis developed in Chapter 3. -

5-2

l l

5.3 TANK WALL AT BOLT CHAIR LOCATION '

Design calculations contained in Refemnce 7 indicate tank wall stresses in excess of yield for the f aulted condition loading (SSE plus normal operating loading). Calculations are based on a formula given in Reference 8. The formula in Reference 8 is empirical and combines membrane and bending stresses. For local loadings on shells, the ASE j, code defines the resulting membrane stress as a primary local membrane

] stress and the bending stress as secondary. The code stress acceptance criteria, Reference 5, places a limit on primary local membrane stress 3

) intensity but ignores secondary stress intensity. This is justifiable if

]. the components do not undergo a large neber of cycles of loading. For

} the case under consideration, the loading is a one-time-only event and

! secondary stresses can justifiably be ignored.

1 3 In order to compute the membrane components of stress in the 1 tank wall due to bolt chair loading, the methods of Reference 9 were

[ .

applied in a conservative manner. The top plate was assumed to be loaded

] radially outward, resisting the applied moment from anchor bolt loading.

Figure 4-8 shows the analytical nodel. This model is very conservative j, since no credit is taken for load distribution into the shell from the

! gusset plates.

i

, In applying Reference 9 to the problem, non-dimensional membrane

] forces Rm Nt/P and mx R N /P am given for R/t ratios up to 300. The R/t

[ ratio of the tank wall is 832. Consequently, non-dimensional membrane F

-l forces were extrapolated via log-log plots of RgN/P vs y. The value of 8.

.j which is a function of the ratio of the lug dimensions to the shell radius, is small and was conservatively taken to be near vero for the extrapolation process.

d

. For a maximm experimentally detennined anchor bolt load of l 31,310 pounds, the local primary membrane stress components were detennined to be:

c e

= 13.20 ksi e

x

= 12.16 ksi 5-3 i

g .- -- . - . .

st These stress components were combined with stress components from the finite element analysis. The finite element analysis resultant stresses account for all loading conditions except for the bolt chair moment. Maximum bolt chair loading occurs between elements 130 and 152.

i Membrane stresses were averaged between these elements and added to local membrane stresses computed for the bolt chair moment.

. The finite elemelt model stresses are:

Element c e x *0x i- 130 6385 1640 331 152 6383 1589 -13

Avg. 6384 1514 159 I

i i Resulting local membrane stress components are:

= 9,584 psi e

o, = 13,744 psi T = 159 psi ex Primary local membrane stress intensity derived from the stress components is:

5 = 19,588 psi The derived stress intensity is conservative for cwo reasons:

1. Finite element stresses are taken at the center of the element. Hoop stress diminishes toward the bottom of the element due to the radial restraint offered by the tank bottom. The point of application of the radial bolt chair reaction load is about midway between the tank bottom anr' f the mid point of the element; thus, actual hydrostatic pressure induced hoop stresses would be lower than assinied e above. ,

5-4

) . .. . .

2. Computed local membrane stresses are conservative since the total moment reaction is assuned to be taken out by a concentrated radial load on the tank wall at the bolt chair top plate. Load distribution into the tank wall from the _

gusset plates is ignored.

s Using the design stress formula of Reference 8, the maximum primary local membrane plus secondary stress components in the axial tank wall direction is computed to be 41.94 ksi. Since a major portion of the i stress is bending the resultant stress is assuned for design purposes to be negative and when canbined with the positive hoop stress of 6.38 ksi

from hydrostatic pressure loading, the maximum primary plus secondary stress intensity is

S= 48.32 ksi This stress combination is not restricted by the governing i

design criteria, Reference 5. ASME code philosophy regarding shakedown to elastic action suggests that primary plus secondary stress intensities be limited to 2 Sy (2x25,000 psi) for cyclic loading. The above stress intensity meets the 2 S yshakedown philosophy and is of no concern.

4 i

5-5 i

- - -a ---

TABLE 5-1 STRESSES AT CENTER OF ELEMENTS BOTTOM,THREE R0WS - PSI

{ R0W 3 ELEMENT NO. T 3 O x x0

! 7 10581 -1429 544 12060 18 10600 -1106 761 11805 40 10619 -406 918 11176 l 62 10633 127 858 10703 i

84 10642 199 799 10703 106 10661 768 721 10713 l 128 10685 1708 460 10708 1 150 10689 1743 38 10689 172 10669 1056 -208 10674 194 10650 523 -404 10666 216 10646 361 -495 10670 238 10648 448 -599 10683

. 260 10634 193 -694 10680 l 282 10621 -536 -757 11258

' 304 10615 -509 -532 11174 326 10614 -457 -649 11146 l 348 10608 -894 -434 11536 1 370 10587 -1223 -314 11826 392 10571 -1923 -85 12495 414 10582 -1778 301 12374 I 436 10604 -575 543 11232 -

T 0 x xe S ELEMENT N0.

29 10573 -1885 332 12475

.l 51 10516 -1807 84 12385 73 10590 -1189 281 11792 95 10596 -1093 459 11726 l

117 10593 -1262 -638 11924 139 10606 -893 -851 11624 161 10624 -116 -912 10894 l 183 10641 188 -814 10704 l 205 10657 650 -790 10719 227 10676 1431 -526 10706

, 249 10684 1594 -197 10689 l 271 10679 1337 107 10681 293 10670 1199 337 10682 315 10659 824 619 10698 337 10645 311 702 10692 359 10632 -19 786 10766 381 10614 -432 743 11146 403 10607 -974 733 11674

425 10613 -605 425 11250 5-6 f

e t

t TABLE 5-1 (Continued)

R0W 2 ELEMENT NO. T O "x x6 S 8 10866 -1081 349 11966 19 10872 -895 502 11810

( 41 10830 -309 627 11210 63 10801 129 564 10831 85 10805 153 529 10831 107 10781 598 484 10804

! '129 10731 1376 318 10741 1 51 10737 1375 7 10737 173 10767 830 -126 10768 195 10783 434 -278 10790 217 10796 288 -325 10806 239 10796 347 -404 10812 261 10792 201 -448 10811 283 10848 -459 -544 11360 305 10831 -378 -299 11225 327 10827 -324 -484 11192 1

349 10870 -752 -258 11632 371 10854 -908 -215 11770

, 393 10897 -1503 -67 12400 i 415 10908 -1450 196 12364

~437 10882 -381 404 11232 g

R0W 2 ELEMENT N0. T 3 O x xe I 30 10900 -1488 248 12396 52 10899 -1433 -80 12332 74 10862 -907 -177 11774 j 96 10864 -857 -310 11738

. 118 10869 -985 -417 11882 140 10858 -714 -573 11628

, 162 10808 -52 -618 10930 1

184 10807 146 -520 10832 206 10787 503 -547 10817

-1 228 10742 1156 -344 10755 1 250 10741 1263 -133 10743 l 272 10757 1043 79 10758 294 10753 968 207 10758 316 10774 659 434 10793 338 10800 246 451 10819 360 10816 -19 539 10888 382 10819 -290 469 11148 404 10867 -799 539 11716 426 10845 -479 220 11332 5-7

(s

. o TABLE 5-1 (Continued)

R0W 1 ELEMENT ND. 0 T b 0 #x xe 9 5891 -1173 332 7096 20 5925 -1092 498 7086 42 6042 -341 642 6510 64 6135 190 554 6186 86 6164 159 526 6210 108 6252 664 485 6294 130 6385 1640 331 6408 l 152 6383 1588 -13 6383 174 6297 944 -109 6300 196 6230 522 -290 6245 218 6188 329 -317 6205 240 6176 378 -412 6205 262 6155 319 -425 6186 t

284 6029 -615 -594 6750 306 6039 -405 -231 6461 328 6033 -312 -546 6438 350 5937 -964 -218 6914 372 5922 -954 -227 6892 394 5824 -1733 - 71 7558 416 5822 -1784 184 7614 438 6008 -307 456 6380 R0W 1 -

ELEMENT f0. 0 T b 0 x xe 31 5812 -1739 280 7572 53 5818 -1683 -109 7504 75 5905 -997 -162 6910 97 5918 -995 -317 6942 119 5920 -1130 -408 7097 141 5972 -863 -578 6932 163 6010 6 -630 6165 185 6150 149 -493 6190 207 6229 553 -571 6286 229 6350 1381 -333 6372 251 6372 3466 -137 6376 l 273 6338 1175 91 6339 295 6322 1151 1 84 6329 317 6255 765 459 6294 l 339 6173 281 428 6204 1 361 6109 -43 561 6254 383 6065 -251 433 6375 405 5961 - 995 601 7058 g 427 5994 -568 145 6568 i

5-8 i

i

s

, e TABLE 5-2 STRESSES AT CENTER OF LOWER EDGE OF ELEMENT, BUTIOM ROW - PSI R0W 1 ELEMENT f0. a g o, t xe 9 2987 -1180 332 20 3044 -1098 498 42 3233 -348 641 64 3384 184 554 86 3433 151 526 108 3576 657 484 130 3788 1633 331 152 3785 1582 -13 174 3650 937 -109 196 3541 515 -290 218 3472 322 -316 240 3452 371 -412 262 3414 312 -425 284 3216 -622 -594 306 3228 -412 -231 328 3217 -318 -546 350 3064 -971 -21 8 372 3038 -961 -227 394 2881 -1740 -71 416 2879 -1790 184 438 3174 -314 456 R0W 1 ELEMENT NO. T O x xe 31 2861 -1746 280 53 2871 -1690 -109 75 3010 -1004 -162 97 3032 -1001 -317 119 3036 -1136 -408 141 3120 -869 -577 163 3327 0 -630 185 3410 143 -493 207 3539 546 -571 229 3732 1374 -332 251 3769 1459 -137 273 3715 1168 90 295 3687 1144 184 317 3580 758 459 l 339 3447 274 428

! 361 3343 -50 561 383 3269 -258 432 405 3104 -1001 601 427 3154 -574 145 5-9 l

WLE 5-3 VERTICAL REACTIONS INTO SHELL- LBS.

'i. ^

N0DE REACTION l 321 0

~!

322 -29163 323 -4671 324 -7183 325 0 326 0 327 0 328 0 329 0 330 0 331 0

332 0 333 0 f' 334 -8701 335 -22443 336 -19812 337 -19971

'! 338 -22206 339 -40898 340 -24302 341 -23464

342 -16538 343 0

, 344 0

345 0

. 346 0 347 0

'i 348 0 349 0 350 0 351 0 352 0 353 0 354 -22081 355 0' 1

356 -21393 357 -14061 358 -26501 359 -37680 360 -25537 Sum of vertical reactions into shell = 386605 lbs 5-10 F

REFERENCES

1. Graver memo to L.H. Curtis, Bechtel, Ann Arbor from G.M. Ault, 1

Graver - East Chicago, " Foundation Elevations for T-60s, Midland Project, Graver Lead Order 61590",13 April,1981

'I I 2. Bechtal Corp letter, L. H. Curtis, Bechtel, Ann Arbor to G. M. Ault, Graver Energy Systems, " Midland Plant Units 1 and 2, Constners Power

. Company, Bechtel Job 7220, Borated Water Storage Tank, (IT-60 and l 2T-60)" File C-18,15 July,1981.

. 3. Stress Relief Testing of Strain Gaged Anchor Studs for Bechtel Fower Corporation, Bechtel Job 7220-WJE81681 Q, Wiss-Janney-Elstner &

q Associates, November 20, 1981.

4

4. Bechtel Specification 7220-C18 (Q) Rev.14. " Technical Specification for Subcontract for Field Erected Storage Tanks f or the Consumers

,- Power Company, Midland Plant Units 1 and 2, Midland, Michigan."

5. ASME Boiler and Pressure Vessel Code, Sections III, " Nuclear Power Plant Components", Subsection NC,1974, with no addenda, Paragraph NC3300, and Code Case 1607-1.

6. NASA SP-8007, " Buckling of Thin-Walled Circular Cylinders", National

Aeronautics and Space Adrdnistration, Sept.1965.

1

7. Graver Design Calculation, Bechtel/ Consumers Power, Midland, Michigan, Borated Water Tanks - 1T-60 and 2T-60, Rev. 4 dated 11 Feb, 1980.

B. AISI Steel Plate Engineering Data - Vol. 2 "Useful Information on the Design of Plate Structures", Part VII, Anchor Bolt Chairs, Feb.

1979, American Iron and Steel Institute, Wash. D.C. -

l

)

9. Welding Research Council Bulletin 107, " Local Stresses in Spherical and Cylindrical Shells due to External Loadings", Welding Research Council, New York.

10. API 650, " Welded Steel Tanks for Oil Storage", Fifth edition and Supplement 1, Oct.1973. American Petroleum Institute.

1

11. ANSYS, Engineering Analysis System User Manual, Rev. 3 Swanson Analysis Systems Inc., Houston, PA.

R-1

~. - - -

y.,; , , , ,

REFERENCES (Continued) l 12. ASME Boiler and Pmssure Vessel Code,Section III, " Nuclear Power Plant Components", Su'section a NB Class 1 Components,1980.

l 13. ASME Boiler and Pmssure Vessel Code,Section III, " Nuclear Power Plant Components" Subsection NF, Component Supports,1980.

f

)i 4

l R-2

w .

i 4

i i

ADDENDA

{

BORATED WATER STORAGE TANK ANALYSIS f'

FOR END-OF-LIFE S0IL SETTLEMENT

, AND SEISMIC MARGIN EARTHQUAKE

! LOADING CONDITIONS i

l f

e t

I h

- w -

.9 . . . .

A1. INTRODUCTION

{

, A1.1 STATEENT OF PROBLEM Ring wall's supporting the Borated Water Storage Tanks (BWSTs) i are being reinforced by Bechtel Power Corporation and the tanks will be releveled to a condition of uniform support. Subsequent to the prooosed

-( ring wall retrofits and tank releveling, further soil settlement is

'I projected to occur. This addenda evaluates the effects of the projected future soil settlement on the BWSTs. Bechtel Power Corporation has i

undertaken a study to quantify the ring wall distortion resulting from soil settlement over the 40-year life of the plant. Tabulated distortion

't.'

data were provided to SMA by Bechtel personnel (Reference A1) and are presented within Table Al-1. Figure Al-1 from Reference A2 is a graphical presentation of predicted future settlement and distortion.

A1.2 PURPOSE OF STUDY The purpose of this study is to evaluate the effects of future

.; so'il settlement loading conditions combined with the seismic margin earthquake and to demonstrate compliance with applicable design "ades for the combined loading. <

A1.3 SCOPE OF WORK The scope of work consists of performing a finite element analysis of the BWSTs to determine the stress conditions in the tank t wall, anchor bolts and bolt chairs caused by soil settlement and consequent ring wall distortion. Tank models, with and without a inch layer of asphalt impregnated fiberboard (Celotex) between the tank bottom and ring wall, are analyzed in order to assess support stiffness effects on stress levels in the tanks. Calculated stress from the static load cases are combined with seismically induced stress to produce a combined response. The seismic loads are taken from the Midland Seismic Margin Study as described in Section A4.4. The combined stresses due to the l seismic and settlement response are compared to 1974 ASME Code accentance criteria as described in Section A3.

Al-1

e-

. A1.4 GENERAL APPROACH The general approach to analyzing the effects of the end-of-life settlenent condition on the BWST is very similar to the analysis performed on the BWSTs for current settlement condition analysis, as described in Section 1.5. The finite element model constructed to analyze the end-of-life soil settlement condition is shown in i Figure Al-2. This model is nearly identical to the model which was

'I constructed for the analysis of the current soil settlement condition (Figure 2-2), with the exception of the addition of 40 three-dimensional i

spar elements to simulate anchor bolts. Spar elements within the ANSYS program are defined as uniaxial tension-compression elements with three t

, displacement degrees of freedom at each node. These spar elements have been placed in parallel to the gap elanents. Each set of nodes which formerly defined the endpoints for the gap elements, now define the endpoints for a parallel set of a gap element and a spar element. The I

gap element and spar element combinations model the anchor bolt, Celotex and ring wall interf ace, as described in Section A4. '

The loading conditions on the tank cylindrical shell, which are

, reacted by the ring wall, include the weight of the tank roof, the weight j of the cylindrical shell, the weight of an effective annulus of water, and the bolt loads which are caused by the ring wall deflection. Another l'

loading imposed on the model is the hydrostatic radial pressure acting on the tank wall, f

Section A2 summarizes the overall results and conclusions of l this study. Section A3 presents acceptance criteria used for evaluating t

seismic and settlement-induced tank stresses. Secticn A4 describes the tank model, boundary conditions, loading conditions and methods of

, solution in detail. Detailed results are presented in Section AS. Only one tank was analyzed since future settlement and seismic loading are applicable to both BWSTs.

i Al-2

TABLE A 1-1 PREDICTED END-OF-LIFE SETTLEMENT DATA FOR BWST #1T-60

!f Angle e Settlement Angle e Settlement from North (ft.) from North (ft.)

~'

}

O.22* 0.0745 180.22* 0.0622 7.72' O.0754 187.72* 0.0648 lf 15.22' O.0764 195.22' O.0676 22.72* 0.0773 202.72' O.0706 30.22' O.0782 210.22' O.0732

.f 37.72* 0.0789 117.72* 0.0752

-} 45.22' O.0794 225.22' O.0765 52.72* 0.0796 232.72* 0.0776

, 60.22' O.0794 240.22' O.0780 67.72* 0.0787 247.72* 0.0781 75.22" 0.0774 255.22' O.0779 82.72' O.0756 262.72* 0.0774 90.22' O.0729 270.22' O.0767 97.72' O.0699 277.72' O.0759 105.22' O.0669 285.22* 0.0751 i 112.72* 0.0640 292.72* 0.0743 120.22' O.0616 300.22' O.0736 127.72* 0.0597 307.72' O.0731 135.22' O.0583 315.22' O.0727 142.72' O.0575 322.72* 0.0725 150.22' O.0573 330.22' O.0725 157.72* 0.0576 337.7 2' O.0727 165.22' O.0586 345.22' O.0732 172.72* 0.0601 352.72* 0.0738 i Al-3

'N' ~ % ~ ~~

~ ~

~~ ~ ~-

~'~-~~.-- _ ._ . _ _ _ _ _ _ . . - - -

p C +0 .

^

F ANGLE 9 FROM DUE NORTH

" ~ ~ w ' .

- E E0 Eo Eo EO 8O 8 O E

F

.06 -3L

.10 -

l

.12

> LEGEND:

1. - DEVELOPED VIEW OF RING FOUNDATION PROFILE CONSUMERS POWER COMPANY MIDLAND PLANT UNITS 1 & 2 ,

== = DEVELOPED VIEW OF VALVE PfT PROFILE t

i BORATED WATER STORAGE TANK FOUNDATION FOUNDATION SETTLEMENTFROM FINITE ELEMENT ANALYSIS FIGURE BWST-12 FIGURE A l-1: RINGWALL SETTLEMENT CURVE FOR TANK IT-60 END OF LIFE CONDITION 1

kE

^8z 5 x @

u W>

E 85 E zg a

! M8 E w

s=-

s

- 2 o

N e

8 8 e

2 SS h b o 8 7  ? 7T A E 4 E

_M i

- A, e -

G 5

~

G

~

E

~

E m ,s g 8 '

3 0 *

0

0 m = 5 0 s E

$ 8 8 E 8 8 8 8 g I l

v E z z = E z z =

8

u

, o

o

. S

• E8t ~ f 5

. S

• 26E ~

d o

. S

• l0E - " E

.S*0lt -

E! - m

, z .9* 6 tE - " o 2 @ @ . m 2 .s 8zE- - e s:

e D .S* LEE -

" {

z s

C a @ @ t 8

- m y "' ~ @ @ @ @ @ @ @ @ @ pp g w  ;

s .s ssE- - t -

.s , -

e e e e e s e e e #9. ie g

- 's t t -

c E @ @ @

o 6 E .9*23 E 1 .s lE -

/ 6

.S

• 0# " j p ,'

9* 6V ~ z g , nj =

.S

• 89 - s / at

/ r

-=

'd -  ;

,q

1. M $me - 1 5 w gW s '

5 @ g w$ .--5^

• l d w .5 :  !

y d

w

/* 5J,da, 5 .

i

• d&tn l a. d *

• I SOfi?O Al-5

A2.

SUMMARY

AND CONCLUSIONS l A2.1

SUMMARY

OF RESULTS The five critical areas of the Borated ',later Storage Tank for combined seismic and settlement type loading coreditions have been determined to be:

1) Compression in 3/8 inch tank wall

2) Compression in 1/4 inch tank wall

3) Local membrane stress in tank wall at the bolt chair

4) Bolt chair top plate [

5) Anchor bolts Table A2-1 contains a stamary of the stress resultants for the tank model configuration with a inch layer of Celotex, along with the appropriate allowable stress limits for each of these five critical areas. As shown in Table A2-1, each of the total response stresses falls within the

~

allowable stress limit. The factor of safety relative to the allowable response for ~each critical area is given below:

. 1). Factor of Safety on Allowable Buckling Stress in the '

. 3/8 inch shell = 2.22.

2) Factor of Safety on Allowable Buckling Stress in the 1/4 inch shell = 1.53.

3) Factor of Safety on Allowable Local Membrane Stress in the Shell at the Bolt Chair =! 1.71

4) Factor of Safety on Allowable Bending Stress in the Bolt Chair Top Plate =i 1.08 -

5) Factor of Safety on Allowable Anchor Bolt Loading =l 2.65 Note that there is an additional factor of safety included in the design code allowable stress limits.

A2-1 ,

Table A2-2 contains a Stress Resultant Sumary for the tank model configuration without the Celotex layer along with the appropriate allow-able stress limits for the five critical areas of the tank. Table A2-2 shows that the maximum compressive stress in the 1/4 inch shell is 3,657 psi compared to an allowable stress of 2,834 psi and the bolt chair top plate bending stress is 41,386 psi compared to an allowable of 37,680 psi.

Thus, the use of Celotex is necessary in order for combined settlement and seismic margin loading to remain within design code allowables. The stress resultants at the three remaining critical areas on the BWST are below the allowable stresses and would be acceptable for the case of no Celotex.

Section A5 of this addenda contains a more detailed descriptionfof the results presented within Table A2-1 and Table A2-2.

A

2.2 CONCLUSION

S The following conclusions are made frons the study presented within this addendu a) The BWST configuration without the Celotex fiberboard layer results in the allowable buckling stress and bolt chair top plate stress being exceeded when subjected to the combined Seismic Margin Study Earthquake and future settlement

, loadings.

b) The EWST configurat4on with a 1/2 inch Celotex layer between the tank bottom and ring wall is acceptable, and all design code streus allowables are met for combined Seismic Margin Earthquake and settlement loading..

A2-2 l

l

Table A2-1: TANK MODEL WITH CELOTEX STRESS COMBINATIONS - SME + DW + SETTLEMENT Allowable Seismic Response from SME DW + Settlement Total Response Parameter (Fpu)tedCondition) Response Response (SeeSectionA3) 853 psi 1.066 psi 1,919 pst 4,252 psi p Compression in 3/8" Shell Compression in 1/4" Shell 684 psi 1.164 psi 1,848 psi 2.834 psi Local Manbrane in Shell ..

at Bolt Chair 8,332 psi 13,720 psi 22,052 psi 37,680 pst Bolt Chair Top Plate ..

Bending 25,120 psi 9,637 psi 34,757 psi 37,680 psi

~

Anchor Bolt Load 15.7 kips 6 kips 21.7 kips 57.4 kips SME = Seismic Margin Earthquake DW = Deadweight + Hydrostatic Pressure Loads

Table A2-2 TANK MODEL WITHOUT CEL0 TEX STRESS COMBINATIONS - SME + 0W + SETTLEMENT

~

! Allowable DW + Settlement Total Response Seismic Response from SME Response Response (See Section A3)

Parameter (Faulted Condition)

'l l l e

853 psi 2,661 psi 3,514 psi 4,252 psi Compression in 3/8" Shell

, 3, 684 psi 2,973 psi 3,657 psi 2,834 pst

[ Comhression in 1/4" Shell Local Membrane in Shell 8,332 psi 15,469 psi 23,801 psi 37,680 psi 4

at Bolt Chair Bolt Chair Top Plate .

25.120 psi 16,266 psi 41,386 psi 37,680 psi Sending 15.7 kips 10 kips 25.7 kips 57.4 kips Anchor Bolt Load

)

SME = Seismic Margin Earthquake

.)

DW = Deadweight + Hydrostatic Pressure Loads j .

4 I

)

i 1

1

- j A3. ACCEPTANCE CRITERIA The governing codes and standards for design of the Midland BWSTs are delineated within Section 3.2 of this report. Section 3.2 contains the ASE code stress acceptance criteria. The ASE code does not specify loading combinations. In order to assure satisf actory perfomance of the BWSTs in a seismic event, response to a load i combination of dead weight plus settlement plus the seismic margin earthquake was compared to f aulted condition allowable stress limits.

The critical areas of the tank for stress analysis purposes are:

1) Local Membrane Stress in Vessel Wall

2) Vessel Wall in Compression (Buckling)

3) Bolt Chair Top Plate

4) Anchor Bolts i

The specific acceptance criteria for each of these areas is addressed below, a A3.1 CRITERIA FOR BOLT CHAIR AND VESSEL WALL IN TENSION The basic design criteria for the bolt chair and the vessel wall in tension is the 1974 ASE Code,Section III (Nuclear Power Plant

) Components), Subsection NC (Class 2 Components). In addition, ASE Code Case 1607-1 is applicable for upset, emergency and f aulted condition stress allowables. Under these governing criteria the following principal stresses are allowed.

Loading Primary Local Membrane plus Condition Primary Membrane, o, Primary Bending, L* b Design and Normal 1.05 1.5S Upset 1.1S 1.65S Emergency 1.5S 1.8S Faulted 2.05 2.45 A3-1

sp . . - .

S is an allowable stress and has a value of 15.7 ksi for 304L stainless steel. The Seismic Margin Study earthquate loads represent safe shutdown (SSE) loads, and, thus, must be compared to the faulted allowable stress limits. Secondary stresses do not require evaluation 7

for Class 2 components designed by rule (NC 3300 criteria).

A3.2 CRITERIA FOR VESSEL WALL IN COMPRESSION-For the large diameter thin wall storage tank, buckling will occur in the elastic range. The ASE Code buckling criterion for axially t

loaded cylinde s nominally contains a safety f actor of 3 for sustained design loads. The ASE Code specifies in Article NC-3000 that the maximum allowable compressive stress to be used in the design of cylindrical shells shall be the lesser of:

a) The allowable S value given in Tables I-7.0.

i b) The value of B determined from the applicable chart in Appendix VII.

For the case under consideration, the latter criterion governs.

This B value for elastic buckling in Appendix VII can be obtained with a higher degree of accuracy by using the design formula shown below, taken from the 1977 ASE Code.

B = 0.0625 E T R

where E = modulus of elasticity - ,

T = shell thickness R = inside radius of shell A3-2

e a. . -

This f ormula applies to the linear portion of the buckling curves in Appendix VII and is applicable for the BWST analysis. The buckling i

allowables for the design and normal loading conditions are therefore:

o cr = 1417 psi (for 1/4" shell) ,

er = 2126 psi (for 3/8" shell)

Based on Code Case 1607-1, the Faulted Condition allowables can be increased by a factor of 2.0 for primary membrane stresses. Thus, the i buckling allowables for faulted conditions are:

r o cr = 2834 psi (1/4" shell) o cr = 4252 psi (3/8" shell) i i

~

A3.3 CRITERIA FOR ANCHOR BOLTS The allowable load criteria for anchor bolt pullout'of the i concrete is based on the provisions of ACI 349-80 (Reference A7). An allowable bolt load of 136 kips has been conservatively calculated for concrete pullout. The allowable load for the failure of the bolt itself 3 is based on the AISC criteria (Refercqce A6). Section 1.5.2.1 of the b AISC Code states that for tension or the nominal bolt area, the allowable i

. stress is 1/3 of the ultimate tent:1e stress. Further, Part 2 of the ,

l AISC Code states that an addition f actor of 1.7 may be applied for the safe shutdown earthquake. Thus, for 1 inch diameter A36 bolts with an ultimate tensile strength of 58 ksi, the allowable load is 57.4 kips.

This bolt failure allowable of 57.4 kips will govern for loading of the anchor bolts.

A3-3

o .

A4. ANALYTICAL MODELS AND ANALYSIS METHODS A finite element model was constructed to represent the cylindrical shell portion of the BWSTs. Boundary conditions were applied at the top and bottom of the model to represent the restraint offered by the umbrella roof, the tank bottom and the asphalt impregnated fiberboard between the ring wall and tank bottom. Analysis of bolt chairs and the tank wall adjacent to the bolt chairs was conducted by hand methods using empirical design analysis approaches and classical stress analysis techniques.

l A4.1 FINITE ELEENT MODEL The basic tank model has already been described in Section 4.1.

The finite element model generated for the end-of-life settlement loading

( differs from this original model only in the vertical boundary elements l

p1 aced along the bottom of the tank. As Figure Al-2 illustrates, the interface between the tank bottom, the asphalt impregnated fiberbcard (Celotex) and the ringwall have been modeled as a gap element and a spar element acting vertically at each of the 40 nodes located at the bottom of the tank. The spar elements model the 40 anchor bolts as linear springs. The gap elements mpresent the rigid ringwall for the case where the Celotex layer is not present, and they model the Celotex layer f, for the case where it is present.

The equivalent linear spring stiffness for the forty anchor bolts was calculated by adding the flexibilities of the bolt, the bolt chair top plate and the tank wall out-of-plane rotation due to the moment created by bolt chair loading. These three items act as springs in series, and their individual stiffnesses were calculated by classic strength of material equations and by the methods given within Reference A8. An equivalent spring stiffness of 1.1 x 105 lb/in was calculated and used for each of the 40 spar elements in the model. As a A4-1

check on the derived bolt tiffective stiff Mss, the stiftnesses of each of the loaded bolts within the viginal BWST analysis were evaluated by dividing the measured bolt lead by the calculated bolt stretching caused by the ring wall deflection relative to the tank bottom. All but one of l these derived bolt stiffnesses were lower in magnitude than the l 1.1 x 105 lb/in value used in the current model. This higher stiffness value is conservative since it results in higher bolt stresses and tank wall stresses.

The gap element stiffnesses, which represent the behavior of the inch thick asphalt impregnated fiberboard, were determined using a beam on an elastic foundation analysis as reported in Section 4.1 of this mport. For the model configuration without a Celotex layer, the gap elements were modeled as rigid once the gap has closed and the tank bottom meets the ring wall.

1 A4.2 FINITE ELEENT MODEL LOADING Loading conditions present in the BWSTs are:

1) Dead weight of the tank roof

2) Dead weight of the tank wall

3) Anchor bolt loading

4) Hydrostatic pressure acting radially on the tank

5) Hydrostatic pressure acting downward on an annulus of the botts plate adjacent to the tank wall.

I Loading conditions 1, 2 and 4 above are identical to the loads applied to the original tank settlement analysis, and their descriptions can be found in Section 4.2 of this mport. The anchor bolt loadings had to be applied differently for this end-of-life settlement analysis since the actual bolt loads had to be computed rather than measured by use of strain gages as they were for the original tank settlement analysis. The A4-2 L

1 ,

tank is initially modelled with a planar bottom, and 40 preloaded springs attached at each of the anchor bolt locations. The preload value in each of these springs was determined by multiplying the bolt stiffness by the distance from the tank bottom to the deflected ring wall position beneath it. After these preloads, along with the deadweight loads and hydrostatic pressure loads, have been applied to the tank, the computer will iterate the model into an equilibrium position. Actual bolt loads can be obtained from this computer solution by subtracting the amount of relieving force occurring due to tank deflection from the original spring preload.

The deadload attributed to the hydrostatic pressure acting downward on an annulus of the bottom plate was determined from Figure 4-4, as in the previous analysis. For the model configuration without the Celotex layer, a conservative value of 100 lbs/in/in. of circumference was used for the water annulus weight. For the model configuration with the Celotex layer, the resultant gaps are smaller than for the case without Celotex. Thus, a conservative water annulus load of 90 lb/in/in.of circumference was used. This deadweight load was placed on those nodes where a gap'still exists after equilibrium has been satisfied. The water load will be carried directly by the ring wall and soil beneath the tank bottom for those nodes which have contacted the Celotex layer.

The relative contributions of downward leading from deadweight, anchor bolt loads and vertical water pressure loading reacted by the shell are:

i With Celotex Layer Tank Roof 36,000 lbs Tank Shell and Hardware 67,090 lbs Effective Water Weight Reacted by Shell 70,560 lbs Total of Bolt Loads 60,340 lbs TOTAL 233,990 lbs A4-3

a- .

s

• Without Celotex Layer Tank Roof 36,000 lbs Tank Shell and Hardware 67,090 lbs Effective Water Weight Reacted by Shell 132,300 lbs Total of Bolt Loads 149,760 lbs TOTAL 385,150 lbs g

From the vertical force summary, it can be seen that the anchor bolt loads are 26% and 39% of the total vertical load for the "with Celotex" and the "without Celotex" cases, respectively. Thus, a considerable amount of the load for each of these configurations could be avoided by periodically backing off the load developed on these bolts.

This is not necessary, however, as the analysis results indicate that end of settlement loading combined with the seismic margin earthquake result

,, in responses within the design code allowable limits.

A4.3 BOLT CHAIR MODEL The bolt chair model is identical to that presented within i

Section 4.3 of this report.

t A4.4 SEISMIC LOADING CONDITIONS A seismic analysis of the Borated Water Storage Tank was conducted in addition to the soil settlement analysis. The earthquake excitation used for this seismic analysis is the ground response spectra developed for the Midland Seismic Margin Earthquake Structural Evaluation Program. The Seismic Margin Study response spectra consist of an envelope of the site specific spectra develootd by Weston Geophysical Corporation (Reference A3) for structures founded at the top of fill, and Housner response spectra (Reference A4) which are anchored to a 0.12 peak ground acceleration. The seismic analysis for this loading condition is contained within Reference A5. A Seismic Margin Earthquake (SME) overturning moment of 8154 ft-kios and an SME base shear of 539 kios were taken from Reference A5 to develop the seismic stresses.

A4-4 0

l i A5. ANALYTICAL RESULTS l

Results from the ring wall future settlement analysis and the Seismic Margins Study Earthquake analysis are sunmarized in this chapter. Computer output is too voluminous to include, consequently, only important highlights of the output are sunmarized.

I A5.1 RESULTS FROM THE FlITlRE SETTLEENT ANALYSIS Because of the wave front solution techniques used in ANSYS, tank model elements in a column are in numerical sequence, while elements in a row (around the circumference) am numbered in steps of 24 (see Figure Al-2). The bottom 3 rows of elements representing the lower

l

- 12 feet of the tank contain the important stress results from the fin te i element model. The highest loading occurs in the bottom row of elements where the shell thickness is 0.375 inches. Since the bottom elements are restrained from radial movement by the tank bottom, hoon stresses in the second row of elements are greater than for the bottom row. At the junction of the second to third row of elements, the tank wall thickness l! changes from 0.375 inches to 0.25 inches. Above row 3, element stresses diminish with decreasing hydrostatic pressure and a more even distribution of the asymnetric vertical reaction loads.

The maximum compressive stresses in the tank wall can be taken directly from the computer printout. These maximum compressive stresses in both the 1/4 inch wall and the 3/8 inch wall are listed below for each of the tank model configurations.

Compressive Stress, psi Shell Thickness No Celotex With Celotex

(

1/4" -2973(Element 55) -1164 (Element 55) i 3/8" -

-2661 (Element 57) -1066(Element 57)

A5-1

Element 55 is located on the bottom row of the 1/4 inch elements, and l

element 57 is located on the bottom row of the 3/8 inch thick elements.

Both of these elements are directly above the highest point on the settled ring wall contour (Node f338). .

I The bolt force and gap results from the computer analysis are tabulated in Table A5-1 and Table A5-2. Table AS-1 contains data for the tank configuration without Celotex and Table A5-2 contains data for the tank configuration with Celotex. The maximtsn bolt loads for the "with" i

and "without" Celotex configurations are 6 kips and 10 kips, respectively, as shown in these two tables. Figure AS-1 shows the resultant displacement of the tank bottom relative to the uncompressed fiberboard position for the tank model with the Celotex layer. The displacement plot incorporates the deflections of both the fiberboard and the tank wall. Maximum compression in the fiberboard is 0.09 inches.

The local membrane stress in the shell at the bolt chair location is the combined hoop stress due to hydrostatic pressure and anchor bolt loads. The hydrostatic pressure hoop stress is calculated at a depth of 31 ft to be 11,177 psi.

The localized hoop stress induced by anchor bolt loading was

, derived by linear scaling from previous analysis, i.e., future settlement stress is equal to the bolt load for future settlement divided by the bolt load for current settlement times the computed stress for current settlement from Section 5.0. The local membrane hoop stress due to deadweight plus settlement was then obtained by absolute summation.

The bending stress located in the top plate of the bolt chair is a linear function of the bolt load. Thus, the bending stresses in Table A2-1 and Table A2-2 were scaled from bolt load ratios times the stresses derived in Section 5.0.

A5.2 RESULTS FROM THE SEISMIC mRGIN STUDY EARTHOUAKE ANALYSIS

)

The complete results of the BWST seismic analysis using Seismic -

Margin Study loading conditions is documented in Refemnce AS. A brief A5-2 l

su-mry of the results from Reference A5 which were utilized in this addenda will be included in this section.

Seismic Margin Earthquake (SME) overturning moments from Reference A5 were utilized to calculate compressive stresses in the tank shell. The maximtsn overturning moment is reported to be 8154 f t-kips in the 3/8 inch shell, and 4358 f t-kips in the 1/4 inch shell. Using standard bending stress equations from beam theory, stresses of 853 psi and 684 psi are calculated for the 3/8 inch shell and the .

1/4 inch shell, respectively. >

The reported local membrane stresses in the tank wall at the "

bolt ch' air are hoop stresses due to 2 different sources. The first source of hoop stress is the seismic induced bolt loads acting on the '

bolt chair which tend to want to stretch the shell in hoop membrane.

This stress is scaled from Section 5.0 for the Seismic Margin bolt load The second computed in Reference A5 and is computed to be 6629 psi.

source of hoop stress is made up of three pressure induced components.

These three components are:

the hydrostatic pressure due to the 0.lg vertical earthquake which produce a hoop stress of 1118 psi. This 1) stress was calculated by taking one tenth of the 11,177 psi

- hydrostatic pressure induced hoop stress calculated for a water coltsnn 31 fqet ahove the bolt chair top plate.

2) sloshing induced pr' essure of 12.9 lb/ft2 The (Reference sloshingA5)

~

which produces a 74.5 osi hoop stress.

induced pressure is caused by that. portion of the tank water mass which moves in a sloshing motion.

3) convective induced pressure of 222 lbs/ft2 (Reference A5) which produce a 1283 psi hoop stress. The convective induced pressure is caused by that portion of the tank water mass which moves as a rigid body.

These three pressure induced hoop stresses were combined by SRSS since they will generally be out of phase with each other and the resulting stress level is 1703 psi. The total hoop stress at the bolt chair location was conservatively calculated to be 8332 osi by absolute sumation of the 6629 psi bolt load induced stress and the 1703 psi pressure induced stress.

AS-3 l

l l

The seismic induced bending stress of 25.12 ksi in the bolt chair top plate was calculated by scaling results in Section 5.0 for predicted future settlement bolt loading.

A maximisn anchor bolt load of 15.7 kips for the Seismic Margin Earthquake was taken directly from the calculations in Reference AS.

AS.3 TOTAL RESPONSE CALCllLATIONS The total stress response for the f aulted condition was conservatively obtained by absolute stenation of the SE induced stresses and the ring wall settlement (including deadweight and hydrostatic pressure) induced stresses. .

E e

an W

e S

AS-4

Table A5-1 BOLT FORCE AND GAP RESULTANTS FOR BWST CONFIGURATION WITHOUT CELOTEX LAYER de Number Gap (Inches) j Bolt Force (lbs) 321 0 0 322 0.0037 407 323 0.0169 1,863 324 0.0312 3,432 325 0.0462 5,100 326 0.0619 6,81 0 I 327 0.0756 8.322 328 0.0863 . 9,499 329 0.0926* 10,190*

i 330 0.0906 9,974 3 31 0.0798 8,785 332 0.0597 4,566 333 0.0373 4,103 334 0.0134 1,473 335 0 0 336 0 0 i

337 0 0 338 0 0

, 339 0 0 l 340 D 0 i

341 0.0002 24 342 0.0174 1 ,916 l.

343 0.0396 4,354 344 0.0622 6,850 345 0.0789 8,690 346 0.0872 9,600 l 347 0.0875 9,636 348 0.0827 9,096 349 0.0730 8.033 350 0.0600 6,606 3 51 0.0449 4,937 352 0.0297 3.269

, 353 0.0164 1,799

,l 354 0.0039 426 1 355 0 0

, 356 0 0

~'

357 0 0 358 0 0 359 0 0 360 0 0

• Maximum Value Sum of the Bolt Force = 149,760 lbs.

A5-5 i

7 _

I Table AS-2 i'

i BOLT FORCE AND GAP RESULTANTS FOR BWST CONFIGURATION WITH CELOTEX LAYER I

Node

• Gap (Inches) BoltForce(1bs) 4 0 0 l 321 0 0 322 0 0 323 0 324 0 0.0094 975

! 325 0.0252 2818-326 0.0399 - 4383 327 0.0500 5501 328 l

0.0549* 6037*

329 0.0506 5567 330 0.0368 4028

? 331 0.0122 1393

! 332 0 0 333 0 0 334 0 0

! 335

' 0 0 336 0 0 337 0 0 338 0 339 0 g 0 340 0 0 0 341 0 0 l l 342 0 0 1 343 0.0158 1771 344 0.0366 4032 l 345 0.0478 5292 346 0.0507 5577 347 0.0470 5194 348 0.0381 4185 349 0.0246 2700 350 890 351 0.0080 0 0 l ll l 352 0 0 353 0 354 0 0 0 355 0 0 356 0 357 0 0 0 358 0 359 0 0 0 360

• Maximum Valve Sum of the Bolt Forces = 60,343 lbs.

A5-6

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REFERENCES _

A1.

Tabulated Data presented to R. O. Campbell during January 13, 1982 meeting between Consuners Power Company, Bechtel Power Corportion, Nuclear Regulatory Comission and Structural Mechan.ics Associates held at NRC Headquarters, .

Phillips Building in Bethesda, Maryland.

A2. " Testimony of Alan J. Boos and Dr. Robert D. Hanson on behalf of l the Applicant Regarding Remedial Measures for the Midland Plant f Borated Water Storage Tank", Atomic Safety and Licensing Board i

Docket Numbers 50-329 OM, 50-330 OM, 50-329 OL, and 50-330 OL, February 16, 1982.

t A3. " Site Specific Response Spectra Midland Plant - Units 1 and 2, '

Part II - Response Spectra Allicable for the Top of Fill Material at the Plant Site", Weston Geophysical Corporation, Westboro, Massachusetts, May 1,1981.

A4. " Nuclear Reactors and Earthquakes", TID-7024, prepared by Lockheed Aircraft Corporation and Holmes & Narver, Inc., for the Division of Reactor Developnent, U.S. Atomic Energy Comission, Washington, D. C., August,1963.

AS.

"Miul'and Safety Margin Earthquake Structural Evaluation of the Borated Water Storage Tank", SMA Report #13701.01-R002, prepared for Consuners Power Company,l April,1982.

A6. Manual of Steel Construction (Seventh Edition), American Institute of Steel Construction, Inc.,1973.

A7. ACI 349-80. Code Requirements for Nuclear Safety Related Concrete Structures, American Concrete Institute, April,1981.

A8. Bijlaard, P. P., Stresses From Radial Loads and External Moments in Cylindrical Pressure Vessels, Welding Journal Research Supplement, December,1955.

AR-1 i

t .- _. -

bd e

+ .. ...... ,,

9 t

APPENDIX A 6

TEST LABORATORY REStlLTS FROM COMPRESSION TESTING OF ASPHALT IMPREGNATED FIBERBOARD (CELOTEX) e i

%A

, Ostee eoseo e.sais . o.** *a e-e t 6 SMITH-ENIERY CONIPANY CHEn :STS

• TESTeNG * $N5PEc,Teoad . ENGaNEERS 731 EAST W A S ** t N G T O N ROutEVAAn .

L nr. A N G E L E S. C A LIF OR NI A 90021 . (213) 749 3411 3540- 0 LA PALMA AVFNUE .

A N A 64 F I M . C AllF ORNI A 92806 . (714) 630 esto

.....-...................................... .............,...,..............................c..........

.......... . ......... ...... ....... .. ... . ....... . .... ... ..... ..... ...............~...e..............

fre6s No.: 8246 0.vs October 16, 1981 s.. .. . ,n. . N o. : 81-809 Structural Mechanics Associatos 5160 Birch Street RECEIVED OCT 2 019b Newport Beach, California 92660

} y q,d C gO}

Attention: Creg S. Hardy t

n SUBJECT : COMPRESSION TESTS ON 1/2" THICK CEIOTEX

~F 1 l FIBERBOARD MATERIAL

!kN SOURCE: Submitted to our Laboratory (

W; REPORT OF TESTS l

In compliance with your request, we have conducted compression tests on 3-inch by 3-inch, and 3-inch by 6-inch specimens of Celotex fiberboard cut from a 6-inch by ill-inch by 1/2-inch thick sample submitted to our laboratory.

The tests were conducted in a universal testing machine with the specimen laid flat. Rigid steel plates slightly larger than each specimen were ~

placed on the top and bottom of the specimen to assure full and uniform loading over the surface of the specimen. Dial indicators having 0.001 inch accuracy were placed on each side of the specimen, evenly spaced from the center of the specimen to measure the deformation of the specimen.

A load was applied to the specimen and the corresponding average deformation --

reading recorded. The load was then released and the deformation again recorded. The load was increased and the deformations again recorded, followed by unloading. This sequence was continued with increasing loads until a reasonable curve could be established and the deformation exceeded 50% of the original thickness.

Load deflection readings are shown on the attached Plates A and B with accompanying load-defomation curves.

l Page 1 of 2 B-1

g . . - . _ . .

, a- ..

, SMITil-EMERY COMPANY r

File No. 8246 Laboratory No.81-809 I.

Structural Mechanics Associates CEIDTEX FIBERBOARD COMPRESSION TESTS October 16, 1981 l

l We are also enclosing a copy of the most recent calibration of our i

Tinius-Olsen universal testing machine used for these tests, f

I i

Respectfully submitted, r

SMITH-EMERY COMPANY

.f W ...-L  ;'

ByY.1 * - ( x r. 5 #. -1 m _

PAUL LINSTROM Civil Engineer 2-Addressee Attachments PL:ks l

Page 2 of 2 B-2 L

.g.

SMITH-EMERY COMi'ANY File No. 824G Laboratory No.81-809 STRUCTURAL MECHANICS ASSOCI ATES October 16, 1981 i

i CEIOTEX COMPRESSION TEST Sample Size: 3" x 3" x 1/2" Loa d , Lbs . Deficction, In. Load , Ibs. Deflection, In.

0 .000 1000 .151 240 .050 0 .083 0 .017 1500 .184 300 .066 0 .108 j O .026 2000 -

.210 350 .068 0 .129 0 .029 2500 .232

- 400 .075 0 .151 O .034 3000 .251 450 .083 0 .173 0 .036 3500 .264 l' 500 .090 0 .190 l 0 .039 4000 .274 550 .097 0 .200 t 0 .046 4500 .295 y 600 .104 0 .235 0 .048 , 5000' .307

, 650 .110 0 .243

'l 0 .055 .r 5500 .312 700 .120 0 .249 0 *

.057 6000 .316 750 .124 0 .251 0 .063 7000 .323 800 .129 0 .267 0 .070 8000 .334

! 850 .134 0 .277 O .070 9000 .342 900 .144 0 .291 l 0 .078 Set Af ter 5-Minutes:

L. 950 .147 0 .274 0 .081 i

PLATE "A" B-3

r- SMITH-EMERY COMPANY e File No. 8246 Laboratory No.81-809

,j-STRUCTURAL MECHANICS ASSOCIATES October 16, 1981 CEIDTEX COMPRESSION TEST r Sample Size: 3" x 6" x 1/2" Load, Lbs. Deflection, In. Load, Lbs. Deflection, In.

- 0 .000 5,500 .'234 0 .000 0 .172 500 .041 6,000 .241 O .017 0 .179 1000 .084 6,500 .250 0 .045 0 .189 i 1500 .120 7,000 .255

} O .060 0 .198 1750 .130 7,500 .260

O .072 0 .203

'. .' 2000 .143 8,000 .267 O .083 0 .212 2250 .158 8,500 .271 0 .095 0 .217

, - 2500 .166 9,000 .277 j 0 .102 0 .223 I', 2750 .173 10,000 .283 O .111 0 .231 3000 .184 11,000 .290 0 .118 0 .237 3250 .186 12,000 .295 0 .126 0 .245 3500 .196 14,000 .303 0 .131 0 .253 le 3750 .201 16,000 .311 0 .139 0 .262 l '4000 .208 18,000 .319

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0 .148 O_ .278 4500 .218 20,000 .325 g 0 .153 0 .284 l 5000 .225 set A f ter 5-Minutes:

O .162 0 .270  ;

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. UNITED CAllBRATION CORPe 12761 MONARCH STREET

• GARDEN GROVE, CALIFORNIA 92641 PHONE (714) 893-1821

• 638 2322 PAGE 1 OF 2  ;

CERTIFICATE OF CALIBR ATION FOR: SMITH EMERY TYPE OF MACHINE OR APPARATUS DATE February 26, 1981 CAP: 1b,00 S 8 700-4 I' ' "O' 68'F P 0. 40 AMBIENT TEMPER ATURE o, O O O V V --

_ i .,, -is 300 600 1,200 1,800 2,400 3,000 Lbs.

0-12,

. 04 DIAL : M W. ,,g

, ,3 g O Lbs.

O A O -

--- O v

o, V V V o

.,3 - i ~.

1,200 2,400 4,800 7,200 9,600 12,000 Lbs.

• " ' * ** + 8 **

eig0-30,000 Lbs.

4 A V

A V

A V h

_ig -i s 3,000 6,000 12,000 18,000 24,000 30,000 Lbs.

CAllBRATION APPARATUS USED Method Used O ASTM E4-79 Machine Moels M E 4-79 N.B.S. TRACEABLE MOREHOUSE PROVING RINGS calibratedO X3 I 15 by National Standards Test Lab: .

N

~

100,000 Lbs., Serial 41431, 3/29/79, 1/20th of it, NBS #SJT.01/101468. &

20,000 Lbs., Serial 92820, 3/21/79, 1/20th of it, NBS GSJT.01/101469 g*$emge Egneer 2,000 Lbs. Serial 93010, 3/22/79, 1/20th of 1%, NBS GSJT.01/101469.

b *- .

n Veynnical Duetter Jitfe H. Watson

_ _ _ _ _ - - _ _ _ _ _ _ _ _ _ _ _ _ - _ .