ML20040F458
| ML20040F458 | |
| Person / Time | |
|---|---|
| Site: | Midland |
| Issue date: | 01/31/1982 |
| From: | Banon H, Campbell R, Hardy G, Kennedy R, Kipp T STRUCTURAL MECHANICS ASSOCIATES |
| To: | |
| Shared Package | |
| ML20040F449 | List: |
| References | |
| SMA-13704.01, SMA-13704.01-R, SMA-13704.01-R0, SMA-13704.01-R001, SMA-13704.01-R1, SMA-13704.01-RA, NUDOCS 8202090220 | |
| Download: ML20040F458 (69) | |
Text
{{#Wiki_filter:b [ SMA 13704.01 -Ecol I EVALUATION OF MIDLAND NUCLEAR POWER PLANT B0 RATED WATER STORAGE TANKS FOR NON-UNIFORM SUPPORT LOADING RESULTING FROM RING WALL SETTLEMENT l l . I prepared by R. D. Campbell G. S. Hardy H. Banon 1 Approved by:
- e. a ft is ,; $- as Y /b
[ E R. P. Kennedy President
/ T. R. Kipp Manager of Quality Assurance I prepared for CONSUMERS POWER COMPANY I Jackson, Michigan i January, 1982 8202090220 B20204 PDR ADOCK 05000329 T PDR g
g g g STRUCTURRL mECHRnlCS [ -T RSSOCIRTES A C a lif Coro 5160 Bech Street, Newport Beach. Calif. 92660 (714) 833 7552
I TABLE OF CONTENTS Section Title 'Page I LIST OF TABLES . . . . . . . . . . . . . . . . . LIST OF FIGURES ................ ii iii 1-1 I 1 INTRODUCTION . . . . . . . . . . . . . . . . . . 1.1 Statement of Problem ........... 1-1 1.2 Description of BWSTs and Ring Walls . . . . 1-1 1.3 Purpose of Study . . . . . . . . . . . . . . 1-4 1.4 Scope of-Work . . . . . . . . . . . . . . . 1-4 1.5 General Approach ............. 1-4 2
SUMMARY
AND CONCLUSIONS ............ 2-1 2-1 I 3 2.1 Summary . . . . . . . . . . . . . . . . . . ACCEPTANCE CRITEP.IA .............. 3-1 3.1 Governing Codes and Standards . . . . . . . 3-1
- 5. 3.2 Stress Criteria for Settlement Loading .. 3-2 4 ANALYTICAL MODELS AND ANALYSIS METHODS . . . . . 4-1 4.1 Finite Element Model ........... 4-1 4.2 Finite Element Model Loading ....... 4-3 4.3 Bolt Chair Model ............. 4-6 5 ANALYTICAL 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 APPENDIX A I
I I l
1 I l E \ LIST OF TABLES l i Table Title Page I 1-1 1-2 Pre-erection Elevation Data for the Borated Water Storage Tanks (IT60 & 2T60) Foundations . . . . . . Ring Wall Elevations Taken on June 15, 1981 . . . . 1-6 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 I 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 i s I 'I I I I I I . I ii
E I LIST OF FIGURES I Eigure Title Page Plan view of Tank IT-60 Identifying Bolt Numbers I 1-1 1-2 and Location Angl e e . . . . . . . . . . . . . . Plan View of Tank 2T-60 Identifying Bolt Numbers 1-10 and Location Angle 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 Circumfer-4-11 1 ence Vs Gap . . . . . . . . . . . . . . . . . . . 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 I I ' a I I I I
I 1. INTRODUCTION l l I l 1.1 STATEMENT OF PROBLEM Soil settlement at the site of the Midlands Nuclear Power Plant has resulted in deformation of the ring walls that serve as a sopporting l base for the Units 1 and 2 Borated Water Storage Tanks (BWSTs). ! Survey measurements of the ring walls indicate that the top surfaces have distorted from their original position. Visual examination has also j indicated that some cracking of the ring walls has occurred. The ring wall deformation has resulted in a non-uniform support condition for the g BWSTs. Examination of the tanks indicated that some of the anchor bolts 'E 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 ould 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 I initiating the ring walls retrofits, the condition of the BWSTs must be 1- assessed. Evaluation of the BWSTs in their current condition is the i subject of this report.
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 cylindrical flat bottom stcrage 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 ) I thick for the remaining 24 feet. The bottom is 0.25 inches thick. All materials are type 304 L stainless steel. ) 1 1-1
1 E The two tanks are located outdoors in the tank fann area, north of the Auxiliary Building. The Unit I 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 farm. Tank details are shown on Graver drawings NL12046, Rev. 3, NL-12047, Rev. 2, and NL-12051, Rev. 2. Ring walls for the two tanks are identical except in the. valve pit area. Unit 1 has a larger valve pit than Unit 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, elev'ations at the top of each ring wall were verified by Graver, the tank I 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 I deviation from a plane surface 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 inches for 2T-60. Figure 1-3 shows relative deflections of the ring wall top surf ace 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 adjusted in elevation so that at a comon point, the elevations are identi cal . 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 I 1-2
I I 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 g applied by another contractor to the loaded anchor bolts and the. nuts 5 were backed off to zero load. T' ables 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 elevations 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 faulted 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 with respect to remaining in a horizontal I 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 I where the tanks are resting on the ring walls, bolt loads are zero. I l There is a 1/2 inch thic'k 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 l I wall in a more uniform manner than if there were no compressible material at the interface. The tanks are currently full of water and are resting on the l ring walls with the anchor bolts all unloaded. The current unloaded anchor bolt condition is much less critical to the tanks than the prior l condition with some of the bolts loaded, trying to force the tanks walls to the contours of the ring wall surfaces. Visual observations of the BWSTs with the bolt loads still applied did not reveal any obvious damage. Distortion from welding was l apparent in the bolt chairs and there was no apparent difference between chairs that were loaded vs. chairs that were not. L u F 1-3 H
I I t 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, I buckling or permanent damage has occurred due to soil settlement.hnd ring wall distortion. The objective is to verify the BWSTs structual integrity prior to retrofiting the ring walls. 1.4 SCOPE OF WORK The scope of work consists of performing a finite element analysis of the worst tank condition to determine stress conditions in 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 I permanent defonnation 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 fintte element model was constructed to I represent the BWSTs cylindrical walls. The model was constructed of flat plate elements possessing both bending and membrane stiffress. Vertical boundary elements are utilized at the bottom surface to represent the I nonlinear behavior of the asphalt impregnated fiberboard between the tank and ring wall. The boundary elements have gap caoability 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 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 pr perties chosen to represent the
'I a stiffness of the tank roof and the restraint that it offers to the tank at the roof / cylinder intersection. ,
1-4
I I 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 (s the hydrostatic radial pressure acting on the tank wall. Chapter 2 summarizes 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 Chapter 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 defonnations in the tank walls, bolts or bolt chairs. I .I I I I
- I lI I
I I 1-5
I I TABLE l-1 PRE-ERECTION ELEVATION DATA FOR THE BORATED WATER STORAGE TANKS (lT60 & 2T60) FOUNDATIONS I ANGLE e BOLT NUMBER (CLOCKWISE FROM NORTH) ELEVATION (FEET) TANK 1T-60 ELEVATION (FEET) TANK 2T-60 I 1 2 3 355.5 4.5 13.5 4.48 4.46 4.47 5.09 5.07 5.09 I 4 5 6 22.5 31.5 40.5 4.47 4.48 4.48 5.09 5.08 5.07 7 49.5 4.48 5.10 8 58.5 4.47 5.09 I 9 67.5" 4.48 5.07 10 76.5 4.47 5.07 I 11 12 13 85.5 94.5 103.5 4.48 4.47 4.48 5.09 5.08 5.08 4.48 5.07 I 14 15 16 112.5 121.5 130.5 4.48 4.48 5.10 5.08 17 139.5 4.48 5.09 148.5 4.48 5.08 1 18 19 157.5 4.47 5.09 20 166.5 4.48 5.08 I 21 22 23 175.5 184.5 193.5 4.48 4.48 4.48 5.07 5.08 5.09 4.47 5.08 I 24 25 26 202.5 211.5 220.5 4.48 4.47 5.08 5.09 5.09 27 229.5 4.49 l 28 29 30 238.5 247.5 256.5 , 4.48 4.47 4.47 5.08 5.07 5.08 I 31 32 33 265.5 274.5 283.5 4.49 4.47 4.47 5.09 5.08 5.08 292.5 4.46 5.07 I 34 35 301 .5 4.48 5.10 36 310.5 4.47 5.08 37 319.5 4.48 5.09 I 38 39 40 328.5 337.5 346.5 4.48 4.47 4.47 5.08 5.09 5.08 I 1-6
I I TABLE 1-2 RING WALL ELEVATIONS TAKEN ON JUNE 15, 1981 g I ANGLE O ELEVATION (FEET) I ELEVATION (FEET) (CLOCKWISE FROM DUE NORTH) TANK 1T-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 5 150 634.85 634.96 634.83 634.96 I 180 210' 634.78 634.95 240 634.77 634.96 270 634.79 634.96 300 634.82 634.96 330 634.86 634.95 I I I I I I 1-7
l I TABLE 1-3 l 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 3 0.0 23 0.0 4 0.0 24 0.0 5 0.0 25 0.0 6 0.0 26 16.44 7 17.83 27 31.31 8 14.02 28 16.10 10.13 I 9 10 21.32 22.51 29 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 0.0 0.0 I 18 19 0.0 38 39 0.0 20 0.0 40 0.0 I OBOLT LOCATIONS CORRESPONDING WITH THESE BOLT NUMBERS ARE LISTED IN TABLE 1-1 I I i-e I I
I TABLE l-4 MEASURED LOADS IN BOLTS ANCHORING TANK 2T-60 BOLT NUMBER LOAD (KIPS) BOLT NUMBER LOAD (KIPS) 1 0.0 21 0.0 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 0.24 28 0.0 I 8 9 1.15 29 0.0 s 10 0.0 30 0.0 11 0.0 31 0.0 12 0.16 32 0.0 l 13 14 0.0 0.0 33 34 0.0 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 CBOLT LOCATIONS CORRESPONDING WITH THESE BOLT NUMBERS ARE LISTED IN TABLE 1-1.
I 1-9
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, ANGLES ARE MEASURED CLOCKWISE FROM DUE NORTH 4 PRE-ERECTION RELATIVE ELEVATION DATA FOR THE TANK IT-60 RING WALL
- - - RING WALL ELEVATION DATA TAKEN JUNE 15, 1981 i
l I i FIGURE l-3. COMPARISON OF TANK IT-60 RING WALL RELATIVE ELEVATIONS BEFORE AND AFTER THE GROUND SETTLEMENT 1
I I 2. SUMMRY AND CONCLUSIONS I 2.1
SUMMARY
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 plan view of the BWST finite element model and Figure 2-2 is an elevation view of the model. The BWST was modeled on the ANSYS computer program using 40 flat shell elements around the circumference and 8 elements i al ng the vertical, resulting in elements that are approximately four E W feet square. The beam type elements shown in Figures 2-1 and 2-2 are boundary elements to represent the restraint of the cylindrical shell afforded by the tank reaf 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 interf ace between the tank bottom and ring wall. I Figure 2-3 shows the resultant displacement of the tank bottom
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relative to the uncompressed fiberboard positions, ie, the 1/2 inch thick I fiberboard resting on top of the distorted ring wall surf ace. 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 I 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. I 2-1
1 I I Maximum primary membrane stress intensity occurs in element number 392 which is in the first row of elements above the shell l I. 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 strsss 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 nomal operating conditions is 15,700 psi. The minimum specified yield strength is 25,000 psi; thus the maximum 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 ax1al 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 results from hydrostatic pressure. Stresses are summarized in Tables 5-1 and 5-2.
I 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 NASA developed buckling I fomulas, 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 uniform around the circumference I nor along the length. Each of these f actors would increase the critical buckling stress. From R/ference 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 .alues for the BWST condition where axial stress varies from positive to negative around the circumference and is not uniform along the length, 2-2
l l I I Maximum compressive stress occurs in element 392, a 0.25 inch thick element, and is -1930 at the bottom of the element. At this point, the thickness changes to 0.375 inches for elements below. This is less than the calculated buckling stress for uniform axial compression *and I much less than the predicted buckling stress level for bending. The buckling factor of safety, based on elastic buckling for a shell in f bending, is 2.46. In the current condition, the bolt loads have been l relieved and compressive stresses have been reduced such that there is no imediate danger of buckling in the event that further settlement and ring wall distortion occur prior to retrofit of the ring walls. 5 Maximum anchor bolt loading measured in the field via strain gaging was 31,313 pounds, Tables 1-3 and 1-4. Review of the tank design I report, Reference 7 reveals that maximum service loading is predicted for the safe shutdown carthquake event and is 20,427 pounds. Three anchor 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 11/2 inches nominal diameter. Maximum anchor bolt stress is 22.32 ksi in the threaded area. The factor of safety on bolt yield is 1.61. It is therefore concluded, that no permanent defonnation or damage has occurred in the anchor bolts. Anchor bolt pullout was I checked and the capacity was calculted to be greater than 90 kips. The f actor of safety on pullout is 2.87. Note that the bolts are currently unloaded. When the ring walls are retrofit, the tanks are shimed and I bolts are retightened, the loading conditions cited above will not exist. Original bolt chair design analysis was conducted by methods contained in Reference 8. These methods are very conservative design methods and use beam theory as opposed to plate theory. Maximum design stress occurs in the top plate of the bolt chair. Scaling the calcula-tions in the design report for the maximum measured bolt load of 31,313 pounds and the actual thickness of the top plate, the elastically calcu-lated stress is 47.78 ksi and exceeds the minimum specified yield stress 'E .m 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 2-3
I I I material-like stainless steel, the effective plastic hinge will not occur ; until the elastically calculated load controlled stress is greater than l 150% of yield. In addition, stainless steel yield strength is typically 1 much greater than minimum specified yield. ? A yield line (limit) analysis was conducted to determine the I limit load for a fully plastic condition to occur. The calculated yield line analysis limit load based on minimtsn yield properties is 39,950 pounds. None of the measured bolt loads exceed this value. The minimum 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 fonnula 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 are essentially at this value. In Chapter 3, an acceptance criteria is developed that would allow settlement induced stresses to be comparable I to stressec nonnally 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 basis. The allowable bolt load based on this criterion would be 1.35 t (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 be addressed to bolt location 27 where the elastic analysis acceptance I 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 aerived from ASME code philosophy and 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. I 2-4 l
I I Since same yielding could have occured in the bolt chair at bolt location 27, it is suggested that a dye penetrant examination be conducted of welds that attach the top plate to the gussets and tank wall I to assure that plastic straining has not opened up any surf ace 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 evaluated and found to be acceptable. The original design analysis, Reference 7, utilized a method contained in Reference 8 to compute local 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 membrane plus bending stress in the tank wall was computed to be 1.93 times the minimum specified yield strength for the maximum measured bolt loading. The governing ASME code design criteria, I 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 pl' ate of the bolt chair was conservatively assumed to trarismit 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 1 ;nificantly below the 33.75 ksi acceptance criterion of Chapter 3 for primary local membrane stress intensity. I 2-5
I I Secondary stresses are not limited by the governing code, Reference 5. Primary plus secondary stresses should be limited to 2 Sy for cyclic load applications to demonstrate shake down. In the case at hand, the loading is a single applied load and primary plus secortdary I stresses calculated by the design fomula of Reference 8, when combined with stresses induced by other loading are 1.93 S y ; thus, shell I stresses are within acceptable limits. The following summarizes the most severe stress condition in tank 1T-60 prior to relieving the anchor bolt loads. I a. Maximum primary membrane stress intensity occurs in the 0.25 inch thick shell about 10 feet from the base and is I 12.5 ksi compared to a yield strength of 25 ksi and an ASME code primary membrane stress allowable of 15.7 ksi ) or 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 f actor of safety on elastic buckling is 246.
- c. One bolt chair at bolt location 27 may have yielded to a I small degree. The f actor of safety on collapse, based on a lower bound limit analysis, is 1.28. This meets the accept-
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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. lI lI 2-6
I I d. Local membrane stress intensity in the tank wall at the maximtsn loaded bolt chair is 19.59 ksi. The factor of safety on ASME code minimum specified yield is 1.28. Primary plus secondary stress intensity is 1.93 Sy .*hich I is less than the shakedown acceptance criterion of two times yield. It should be concluded from this study that tank 1T-60 is 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. I I 'I I I I I I I I I 2-7
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a y e A g 16 4
$, Pg Ps IB 19 e
3 22 21
+, 1 g S#
b?p @ f2p 22P
\
ot LEV I NOTE: INNER CIRCLE OF NUMBERS REPRESENT NODES ON THE TOP OF THE TANK. l OUTER CIRCLE OF NUMBERS REPRESENT N0 DES AT THE OUTER END OF THE BOUNDARY ELEMErlTS AT THE TANK BOTT0:1. I PLAN O FIGURE 2-1. I 2-8
, g g g g g e e m W M 'M M .M M E B0 RATED WATER STORAGE TANKS AT MIDLANDS hh h h k k k k k k k k9 9 99 OO E $ E 0 ENE CLOCKWISE ANGLE FROM NORTH I ) i I I I i i 1 I I i I i 1i BEAM ELEMENTS y s, s, (3 @ (fy)
N0 DES 1-40 h hb h b b N0 DES 41-80 b b b N0 DES 81-120 y b b b N0 DES 121-160 SHELL ELEMENTSi-~ A N0DESi61-200 { b b b N0 DES 201-240 b N0 DES 241-280 b @ b N0 DES 281-320 ELEMENTS h hb b NODES 321-360 (TANK N0 DES ) d N0 DES 401-440 ( B0UNDARY NODES) GAP ELEMENTS ' *:( h in,,, ,, n, ,i , w r. , rr r. . 7r n ra m u n are, N0 DES 3'61-400 (BOUNDARY NOD
@ - Represent Element Numbers FIGURE 2-2. ELEVATION VIEW 0F TANK MODEL
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_ . _ _ _ _ _ _ _ _ _ , . _ - - - - - - - - - - . - - - - , - .--- . - -- h-- - -- - -----------h- ------s-4 " - + ' - - *----.w+-- -_u - mm .a. goe 20 ' - 355.5" 31.5* 67.5 121.5* 166.5* 211.5* 256.5* 301.5* 396.5* ANGLES 321 3,25 330 335 340 345 350 35) 360 N0DE NOS. 0 . , 4 M O l " s w
-20
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-60 _
FIGURE 2-4 COMPRESSIVE LOADS AT TANK BOTTOM (SEE TABLE 5-3)
E
- 3. ACCEPTANCE CRITERIA 3.1 GOVERNING CODES AND STANDARDS Governing codes and standards are delineated in the design 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 code with no addenda are applicable and Code Case 1607-1 is applicable for upset, emergency and f aulted condition stress allowables. The API 650 code is also specified for design, Reference 10. In cases of conflict, the ASME Code governs.
The basic design is conducted using API-650 criteria since NC3300 of the ASME code does not specifically address flat bottom storage tank designs. NC3800 does provide criteria for flat bottom storage tanks 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 are allowed. Loading' Primary Local Membrane plus Condition Primary Membrane, P m Primary Bending, Pt+Pb Design and Normal S 1.5S Up' set 1.15 1.65S Emergency 1.55 2.05 Faulted 1.85 2.45 Testing 1.25S* 1. 87S**
- Not to exceed 0.9 Sy
** Not to exceed 1.35 Sy 3-1
S is the allowable stress intensity of 15.7 ksi for 304 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 crjteria that any stress in excess of 1.59 S exceeds the minimum specified yield strength and that primary local membrane and primary bending stress allow-ables all exceed the yield strength for all service conditions except design and normal. 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 shimed to provide unifonn support. There I 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 required. If the higher emergency condition allowables were used, some 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 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 required. ASME code philosophy for hydrostatic testing is to test the vessel at 1.25 times design pressure (fonnerly the hydrostatic test pressure was specified to be 1.5 times design pressure). Primary I, membrane stress intensity for desiga 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 1.25S for primary 3-2
1 I 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 for pr' mary local membrane plus primary bending stress intensity for cases where the primary membrane stress intensity is less than 0.67 times yield. The ASME code philosophy in placing stress limits on hydrostatic testing induced stresses is to guard against ' gross defomation. The same philosophy is considered applicable to the BWST settlement problem. If stress intensities from settlement combined with hydrostatic pressure, weight and anchor bolt loading do not exceed hydrostatic testing allowables, then no detrimental effects are considered to have occurred. 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 for class 2 and 3 components and even for Class 1 components for infmquent (emergency) events. Note also, that if secondary stresses vere limited to the shakedown mgime (2 yS ) that during hydrostatic testing at 1.25 times the design pressure, primary plus secondary stress incensity could reach 2.5 Sy. Thus, code philosophy would allow primary l plu.; secondary stress intensity to exceed the shakedown limit for single or lia ted numbers of events. Areas 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 sme additional inspection requirements or limiting inelastic strains. I 3-3
I There 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 fer Class 1 components and component supports, References 12 and 13, allow limit analyses to be used in lieu of meeting elastic stress criteria for primary local membrane and primary bending stress intensities (Paragraphs NB 3228.2 and NF 3224 (a)). The bolt chair top plates are considered to be plate and shell-type component supports and limit analysis concepts from the component support code, (NF 3224 a) are considered to be appli-cable. Under Level C Service (Emergency Conditions), Reference 13 allows O.8 of the lower bound collapse load. The same allowable is also specified for Class 1 components (Reference 12). Design philosophy for Level C Service Condition allowables is that some permanent defomation may be experienced but the component is still serviceable. Deformation limits may be specified if deformation is a controlling f actor for function. In the case of BWST bolt chairs, deformation is not a limiting f actor and a limit analysis acceptance criteria analogous to that for Level C Service for Class 1 plate and shell-type component supports is considered a valid concept for evaluation of settlement induced loading 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 acceptable without further analysis or retrofit being required. It is suggested, however, that for the one chair where the maximum bolt load occurs and is close to the recommended 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 defonnation that may have occurred did not initiate any cracking. I 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 f actor 3-4
I
'I of 3 for sustained design loads. Since the condition under consideration is local and is more of a strain controlled condition than load i -
controlled, a more liberal buckling criteria is recomended. I Reference 6 provides thin shell buckling formulae modified to reflect extensive test data. Formulae 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 ] considered to be lower bound formulae. Since the axial compressive stresses are local and vary from compression to tension around the circum-ference, the bending formula for buckling is considered to be more appro-q pri ate. For the geometry under consideration, the applicable formula is:
- where cr R y = correction f actor for cylinders in bending E = Young's modulus t = shell thickness R = mean radius of shell For the bending case and the geometry under consideration:
i = 0.35 for the 0.25 inch thick shell l y = 0.39 for the 0.375 inch thick shell I Critical buckling stresses are: 4750 psi for the 0.25 inch thick shell 7950 psi for the 0.375 inch thick shell 3-5
E For strain controlled conditions the allowable axial compressive stress should be limited to cr/1.67 resulting in allowable axial compressive stresses of: 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 check should be made to determine if elastic buckles have actually 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 no visual evidence of any buckling in the most critically loaded tank (IT-60). Subsequent analysis also demonstrates that buckling would not occur under the prior conditions. I ig 1I l l i 3-6
I 4. ANALYTICAL MODELS AND ANALYSIS METHODS i
. i A finite element model was contructed to represent the l l
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 I techniques. 4.1 FINITE ELEMENT MODEL The tank wall finite element moAl 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 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. Anchor bolt locations crrrespond to nodal points along the bottom of the model . .I 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 roof / tank wall junction. The unbrella roof is welde6 to a ring girder at tP tank top and would keep the tank wall from ovaling under asyninetric I 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
I 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 represent the stiffness of the asphalt .- impregnated fiberboard and any gaps that may exist between the tank bottom and ring wall. The gap elements have linear force displacement characteristics in the compression direction and zero load capacity in tension. I The actual asphalt impregnated fiberboard (Celotex) is nonlinear. Force-displacement properties were determined by laboratory testing of material taken from the site. Two tests were conducted to determine force-displacement relationships. One specimen was 3" X 3" and I 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 used in developing an appropriate model for the boundary elements. 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 stiff" ness 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 I foundation model results versus the equivalent linear stiffness used in , the computer model for boundary elements. In developing the beam on an elastic foundation model, the resulting stiffness was biased on the conservative (stiff) side. Stiffer boundary elements result in the l reaction loads being concentrated in more localized regions. Asstnptions used in the beam on elastic foundation models are described as follows, l Referring to Figure 4-1, a rigid boundary is assumed on the tank bottom at the junction of the Celotex and the oil impregnated sand. This l l tends to be conservative in that the oil impregnated sand is assumed to i be rigid, thus, forcing all deformation 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 1
I I 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 asslined for the Celotex material for discrete displacement ranges. For larger displacements, the ,- I 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. Referring to Figure 4-1, for a given force P, water pressure, w, shell rotational stiffness, ke, and Celotex effective stiffness, k, a dis-placement directly under the load P was computed. Figure 4-2 plots four I 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 3300 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 I Loading conditions present in the BWSTs are: Dead weight of the tank roof Dead weight of the tank wall Anchor holt loading Hydrostatic pressure acting radially on the tank Hydrostatic pressure acting downward on an annulus of the bottom plate adjacent to the tank wall. I Loads were applied to the model at nodal points that would best represent the physical feces acting on the tanks. 4-3
I I Tank roof weignt 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 I. bottom of the tank wall finite element model. Radial 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. A pressure analysis check case was conducted to verify the model geometry and assure that hoop stresses and displacements were properly computed. Most of the tank bottom msts 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. I In areas where gaps exist between the tank bottom and ring wall, 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. Figure 4-3 shows the effective radial length, ar, 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 surf ace 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 effective water ! weight is, thus, nonlinear, depending upon the gap between the tank I bottom and its support. For the range of gaps calculated in the final 4-4
.E i i
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. l In regions where the tank bottom is supported, the effective annulus of water carried by the tank support is less than at gap locati ons . In order to avoid several iterative solutions that would require changing the effective water weight at points where the tank bottom was supported, 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 I 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 zero load point corresponds to a downward displacement of -0.021 inches. In other words, the starting point of compression in the Celotex is shif ted 0.021 inches and all gaps are shif ted 0.021 inches. The computer program boundary element reactions carried into the shell are correct and resulting shell stresses are correct. However, computed displacements relative to the starting point are 0.021 inches greater than actual and the sisn of the vertical l 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 reactions at the points of tank support cancel each other out and are not reacted by the shell. l 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 a few points where the cisplacement is so small that the I full water weight cannot develop. This occurred at four boundary elements. At points where the tank is supported, the water weight is l reacted directly into the soil and ring wall and there is no effective loading on the tak wall. 4-5
I Weights of the tank, roof and internal bracing were taken from 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 .. I contributions of downward loading from dead weight, anchor bolt loads and vertical water pressure loading reacted by the shell are: Tank Roof 36,000 lbs Tank Shell and Hardware 67,0 % lbs I Effective Water Weight Reacted by Shell Total of Bolt Loads 115,045 168,470 tts lbs 386,605 lbs From the vertical force sumary, it can be seen that the anchor I 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. 4.3 BOLT CHAIR MODEL Bolt chairs were evaluated by hand calculations. Initially, the method of Reference 8 used in design of the bolt chairs, was used to I determine stress response in the top plate for a maximum 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. Since this model resulted in yielding of'the top plate for some 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 I 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 I requires a safety f actor of 1.25 on the lower bound collapse load, thus, the criterion is satisfied. I 4-6
l I I Stresses in the tank wall due to bolt chair loading were calculated using the methods of Reference 9. The original design
. analysis used methods in Reference 8 to compute stresses in the shell from bolt chair loading. Reference 8 does not, however, separatg. primary I local membrane stresses from secondary bending stresses. Since we are only concerned with primary stresses, Reference 9 was used to calculate primary local membrane stresses in the hoop and axial direction.
Reference 9 has cases for a rectangular lug attcchment on a cylindrical shell subjected to radial and bending loading. Since the bolt chairs are at the bottom of the tank, the bending case in Reference 9 is not applicable. A conservative model was asstned wherein the applied moment on the bolt chair is reacted by radial forces at the bottom of the tank and at the top of the bolt chair. Figure 4-8 shows the simple model. I Radial load R is applied to a rectangular lug whose geometry is that of the bolt chair top plate. Local membrane stresses in the shell are computed using the empirical methods of Reference 9. It should be noted that this analysis is conservative since no credit is taken for the gusset plates distributing load into the shell. The conservative model resulted in shell stresses, including pressure loading, below the material yield, thus no further refinement is warranted. 'I I I I I I 4-7
I ! I ~ I P r~ I ke If [ Tank Wall Rigid Boundary Tank Bottom W
~
y y y y p If y v 1r IM y 1r p YYTYYYYYYYY "9
- Asphalt Impregnated Fiberboard Oil Impregnated Sand I
E = average spring rate of asphalt impregnated fiberboard over finite deflection range, lbs/in/ unit width w = water force, lbs/in/ unit width l P = Load into tank / unit width ke = Rotational stiffness of tank wall I FIGURE 4-1. BEAM-0N-ELASTIC-FOUNDATION MODEL I 4-8
il '
!I d
~
!i i 1500 - k = variable i Beam on elastic / y foundationresponsh
/
$ 1000 - = 3800 #/in o
/
$ /
' r E = 1000 #/in/in j E = 800 #/in/in - D E = 800 #/in/in ! g 500 - l 2 / I /
, , i l O.' 1 0.2 DEFLECTION-INCHES 0.3
!I I E = Average spring rate at asphalt i.mpregnated 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 g . d 1 ll 4-9
I I l l l l ~- e or w i P C Tank Wall 1 f Tank Bottom ! W N 'I U y N-
, p
.a 6 5 EQ - 1 y p p (q - 3 + , Asphalt Impregnated Fiberboard l 011. Impregnated Sand .I 1 w = water force, lbs/in/ unit width l P = tank and roof weight, lbs/ unit width 6 = gap between tank bottom and asphalt impregnated fiberboard, in. op = effective width of water annulus, in. I FIGURE 4-3. EFFECTIVE WATER ANNULUS
- I -
iI 4-10 I
lI l !I 4I
- I j 150 .
, E y 120 .._ _ __ l ! r
- y 100 -
- M l uI
, M i 2 50 5 ll l Y ! C I u I I I I I
$ 0.2 0.4 0.6 0.8 1.0 w
Gap - Inches I I J t FIGURE 4-4. EFFECTIVE WATER FORCE / UNIT WIDTH OF CIRCUMFERENCE VS GAP B I 4-11 I
M M 4 W M M M M M M M M M M M M M M
\
800 - E a 600 - 52 Y = 3800 #/in M E P D U g 400 - u. 5 C N u 200 - 120 . _
-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 B0UNDARY ELEMENTS
I . I !q l I
\ TANK WALL I I 1
\
j .
' +
! x ) NJ l n ,i l f I,I 1/3 ANCHOR BOLT LOAD
- I
/
if g 't o < rv v \ c ; j j h i Ak Ak
\
I I PARTIALLY g FIXED 1 l 4 W ll , vi .
= e =
ia FIGURE 4-6: BEAM MODEL FOR BOLT CHAIR DESIGN B 4-13
l Tank Wall ( ( ( t__ g , 2 c' _J t
- s 3 - 3
. .t
- l
- s.
~
/
'l s' s I I O O]
, 6 = =
i ,
,1 ll r,ff s
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x /
% /
I O I P P 2 -=- r gff , T
-6 Y
" Gusset Plate i
Dj, 8 lf x 6 Load I
~-
h 1r l t ~) i~~-a FIGURE 4-7. YIELD LINE MODEL FOR BOLT CHAIR 4-14
P
- e --
W E U l I
- h 2
~
h A I h __ 0.625" 1 I
#. + 0.375" I e 312"R I
h P = ANCHOR BOLT LOAD I \ I I a h v
" l l
1 I FIGURE 4-8. ANALYSIS MODEL FOR LOCAL MEMBRANE STRESSES l l IN SHELL DUE TO ANCHOR BOLT LOADING 4-15 I
E i f l
- 5. ANALYTICAL RESULTS 1
I l Results from the finite element model analysis and hand ! calculations are sumarized in this chapter. Computer output is too l 1 voluminous to include, consequently, only important highlights of the 1 output are sumarized. I 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 from 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. Maximum 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, i
i i Because of the wave front solution techniques used in ANSYS, elements in a column are in numerical suequence while elements in a row l (around the circumference) are numbered in steps of 22 (Figure 2-2). Table 5-1 tabulates the three stress components and resulting stress l intensity for the first three rows of elements beginning at row 3 and working down, row 1 being the bottom row. g Stress output is given at the center of each element and g 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 I element was mquested. Table 5-2 stmnarized stress components for each 5-1
I I element along the tank wall bottom. Along the bottom, tank hoop stresses are low due to radial restraint offered by the tank bottom. Axi al l 7 stresses are more concentrated along the tank bottom resulting from concentrated anchor bolt loading and uneven vertical reactions from the I distorted ring wall. 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. Far a
".000 psi minimum yield material, and assuning elastic-perfectly-plastic material behavior, the plastic hinge moment capacities are computed to be:
I 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. l Work-energy relationships were used to solve for the limit load I capacity. The angle a in Figure 4-2 was varied to determine a minimum value of capacity. Resulting values at minimum capacity are: 'W a P
= 56. 80
= 39.95 kips limit The limi.t load is a f actor 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.
l 5-2 l
I
" 5.3 TANK WALL AT BOLT CHAIR LOCATION Design calculations contained in Reference 7 indicate tank wall I stresses in excess of yield for the f aulted condition loading (SSE plus normal operating loading). Calculations are based on a formula given in I Reference 8. The fonnula in Reference 8 is empirical and combines membrane and bending stresses. For local loadings on shells, the ASME 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 intensity but ignores secondary stress intensity. This is justifiable if the components do not undergo a large number of cycles of loading. For the case under consideration, the loading is a one-time-only event and secondary stresses can justifiably be ignored.
In order to compute the membrane components of stress in the tank wall due to bolt chair loading, the methods of Reference 9 were applied in a conservative manner. The top plate was rssumed to be loaded radially outward, resisting the applied moment from anchor bolt loading. Figure 4-8 shows the analytical model. This model is very conservative 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 N?/P andmx R N /P am given for R/t ratios up to 300. The R/t I ratio of the tank wall is 832. Consequently, non-dimensional membrane forces were extrapolated via log-log plots of RmN /P vs y. The value of 8, which is a function of the ratio of the lug dimensions to the shell radius, is small and was conservatively taken to be near zero for the extrapolation process. For a maximtsn experimentally determined anchor bolt load of 31,310 pounds, the local primary membrane stress components were determined to be: o = 13.20 ksi e x
= 12.16 ksi 5-3
I I sq These stress components were combined with stress components from the finite element analysis. The finite element annlysis msultant 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 element model stresses are: , I Element c e x 'ex I 130 152 6385 6383 1640 1589 331
-13 Avg. 6384 1614 159 Resulting local membrane stress components are:
c = 19,584 psi e o = 13,744 psi T = 159 psi ex Primary local membrane stress intensity derived from the stress components is: S = 19,588 psi The derived stress intensity is conservative for two 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 a
reaction load is about midway between the tank bottom and the mid point of the element; thus, actual hydrostatic h pressure induced hoop stresses would be lower than assumed above. 5-4
- 2. Computed local membrane stresses are conservative since the total moment reaction is assumed to be taken out by a I 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.
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 stress is bending the resultant stress is asstaned for design purposes to be negative and when combined 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 design criteria, Reference 5. ASE code philosophy regarding shakedown to elastic action suggests tnat 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. l i I i \ I l 5-5 i _ - - _ . . . .
i I TABLE 5-1 l STRESSES AT CENTER OF ELEMENTS 7 B0TTOM THREE R0WS - PSI R0W 3 ELEMENT N0. T 3 O x xC 7 10581 -1429 544 12060 18 10600 -1106 761 11805 40 10619 -406 918 11176 I 62 84 106 10633 10642 10661 127 199 768 858 799 721 10703 10703 10713 I 128 150 172 10685 10689 10669 1708 1743 1056 460
-208 38 10708 10689 10674 194 10650 523 -404 10666 I 216 238 10646 10648 10634 361 448 193
-495
-599
-694 10670 10683 10680 260 I 282 304 326 10621 10615 10614
-536
-509
-457
-757
-532
-649 11258 11174 11146
-434 I 348 370 392 10608 10587 10571
-894
-1223
-1923
-314
-85 11536 11826 12495 414 10582 -1778 301 12374 436 10604 -575 543 11232 R0W 3
# T
'I ELEMENT NO. C 10573 x
-1885 xe 332 s
12475 29 I 51 73 95 10516 10590 10596
-1807
-1189
-1093 84 281 459 12385 11792 11726
-638
'I 117 139 161 10593 10606 10624
-1262
-893
-116
-851
-912 11924 11624 10894 183 10641 188 -814 10704 205- 10657 650 -790 10719 227 10676 1431 -526 10706 249 10684 1594 -197 10689 l 271 10679 1337 107 10681 l
293 10670 1199 337 10582 l 31 5 10659 824 61 9 10698 337 10645 311 702 10692 l 359 10632 -19 786 10766 ! 381 10614 -432 743 11146 403 10607 -974 733 11674 425 10613 -605 425 11250 5-6
E TABLE 5-1 (Continued) R0W 2 ELEMENT N0. T xC *S I 0 , x 8 10866 -1081 349 11966 19 10872 -895 502 11810 I 41 63 85 10830 10801 10805
-309 129 153 627 564 529 11210 10831 10831 598 484 10804 I 107 129 1 51 10781 10731 10737 1376 1375 318 7
10741 10737 173 10767 830 -126 10768 I 195 217 239 10783 10796 10796 434 288 347
-278
-325
-404 10790 10806 10812 I 261 283 305 10792 10848 10831 201
-459
-378
-448
-544
-299 10811 11360 11225 327 10827 -324 -484 11192 I 349 371 10870 10854
-752
-908
-258
-215 11632 11770 393 10897 -1503 -67 12400 I 415 437 10908 10882
-1450
-381 196 404 12364 11232 R0W 2 ELEMENT N0. T b 6 x x0 30 10900 -1488 248 12396
. 52 10899 -1433 -80 12332 74 10862 -907 -177 11774 I 96 118 140 10864 10869 10858
-857
-985
-714
-310
-417
-573 11738 11882 11628 162 10808 -52 -618 10930
'I 184 206 10807 10787 146 503
-520
-547 10832 10817 228 10742 1156 -344 10755 250 10741 1263 -133 10743 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
TABLE 5-1 (Continued) R0W 1 ELEMENT NO. T b O x x0 , f, 9 5891 -1173 332 7096 a 20 5925 -1092 498 7086 4? 6042 -341 642 6510 I 64 86 108 6135 6164 6252 190 159 664 554 526 485 6186 6210 6294 I 130 152 174 6385 6383 6297 1640 1588 944 331
-13
-109 6408 6383 6300 196 6230 522 -290 I 218 240 262 6188 6176 329 378
-317
-412 6245 6205 6205 6155 319 -425 6186 E 284 6029 -615 -594 6750 E 306 6039 -405 -231 6461 328 6033 -312 -546 6438 I 350 372 394 5937 5922 5824
-964
-954
-1733
-218
-227
- 71 6914 6892 7558 416 5822 -1784 184 7614 l 438 6008 -307 456 6380 R0W 1 ELEMENT f0. 0 T 0 #x x0 31 5812 -1739 280 7572 53 5818 -1683 -109 7504 75 5905 -997 -162 6910 I
4 97 5918 -995 -317 6942 119 5920 -1130 -408 7097 141 5972 -863 -578 6932 3 163 6010 6 -630 6165 E 185 6150 149 -493 6190 207 6229 553 -571 6286 229 6350 1381 -333 6372 I 251 273 295 6372 6338 6322 1466 1175 1151
-137 91 184 6376 6339 6329 I
317 339 361 6255 6173 6109 765 281
-43 459 428 561 6294 6204 6254 ;- E 383 6065 -251 433 6375 g 405 5961 -995 601 7058 427 5994 -568 145 6568 I
5-8
I TABLE 5-2 3 STRESSES AT CENTER OF LOWER EDGE OF ELEMENT, q R0W 1 BOTTOM ROW - PSI ELEMENT NO. T 0 #x x0 9 2987 -1180 332 I 20 42 64 3044 3233 3384
-1098
-348 184 498 641 554 I 86 108 130 3433 3576 3788 151 657 1633 526 484 331 152 3785 1582 -13 I 174 196 218 3650 3541 3472 937 515 322
-109
-290
-316 I 240 262 284 3452 3414 3216 371 112
-622
-412
-425
-594 306 3228 -412 -231 I 328 350 3217 3064
-318
-971
-546
-218 372 3038 -961 -227
'I 394 416 438 2881 2879 3174
-1740
-1790
-314 184 456 71 R0W l ELEMENT NO. o s % *x0 31 2861 -1746 280 lE i
lB 53 75 2871 3010
-1690
-1004
-109
-162 97 3032 -1001 -317 lg 119 3036 -1136 -408 g 141 3120 -869 -577 163 3327 0 -630 185 3410 143 -493 I 207 229 251 3539 3732 3769 546 1374 1459
-571
-332
-137 273 3715 1168 90 1 295 317 3687 3580 1144 758 184 459 339 3447 274 428 361 3343 -50 561 383 3269 -258 432 405 3104 -1001 601 427 3154 -574 145 5-9
TABLE 5-3 VERTICAL REACTIONS INTO SHELL- LBS. N0DE REACTION I 321 0
-29163 322 323 -4671 I 324 325 326 -
-7183 0
0 327 0 1 328 0 329 0 330 0 331 0
- I -
332 0 l 333 0 I 334 335 336
-8701
-22443
-19812 1
1 337 -19971 338 -22206 339 -40898 340 -24302 l 341 -23464 , u 342 -16538 l i 343 0 l I 344 345 346 0 0 0 347 0 I 348 349 350 0 0 0
;E 351 0 E
352 0 i 353 0 ) a 354 -22081 l g 355 0 ; 356 -21393 357 - -14061 i 358 -26501 1 359 -37680 360 -25537 Sum of vertical reactions into shell = 386605 lbs B I 5-10
REFERENCES I 1. Graver memo to L.H. Curtis, Bechtel, Ann Arbor from G.M. Ault, Graver - East Chicago, " Foundation Elevations for T-60s, Midland Project, Graver Lead Order 61590", 13 April, 1981 I 2. Bechtal Corp letter, L. H. Curtis, Bechtel, Ann Arbor to G. M. Ault, Graver Energy Systems, " Midland Plant Units 1 and 2, Constrners Power l Company, Bechtel Job 7220, Borated Water Storage Tank, (IT-60 and 2T-60)" File C-18,15 July,1981.
- 3. Stress Relief Testing of Strain Gaged Anchor Studs for Bechtel Power Corporation, Bechtel Job 7220-WJE81681 Q, Wiss-Janney-Elstner &
Associates, November 20, 1981.
- 4. Bechtel Specification 7220-C18 (Q) Rev.14, " Technical Specification I f or Subcontract f or Field Erected Storage Tanks for 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 Czse 1607-1.
- 6. NASA SP-8007, " Buckling of Thin-Walled Circular Cylinders", National Aeronautics and Space Administration, Sept.1965. .
Graver Design Calculation, Bechtel/ Consumers Power, Midland, I 7. Michigan, Borated Water Tanks - 1T-60 and 2T-60, Rev. 4 dated 11 Feb, 1980. I 8. 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.
- 9. Welding Research Council Bulletin 107, " Local Stresses in Spherical I and Cylindrical Shells due to External Loadings", Welding Research Couhcil, New York.
I 10. API 650, " Welded Steel Tanks for Oil Storage", Fifth edition and Supplement 1, Oct.1973, American Petroleum Institute. l
- 11. ANSYS, Engineering Analysis System User Manual, Rev. 3 Swanson 1 Analysis Systems Inc., Houston, PA.
R-1
REFERENCES (Continued) I
'g 12. ASME Boiler and Pressure Vessel Code, Section III, " Nuclear Power iE Plant Components", Subsection NB Class 1 Components,1980.
, 13. ASME Boiler and Pressure Vessel Code, Section III, " Nuclear. Power 4 Plant Components", Subsection NF, Component Supports,1980.
- I
.i lI 1 I I I 'I l R-2 j
- I
' APPENDIX A 1 1
'l
'I TEST LABORATORY RESULTS FROM j l l COMPRESSION TESTING OF ASPHALT l IMPREGNATED FIBERBOARD (CELOTEX) I !I
- I
'I 4 i d 4 I i I
pecuuswa e.sses . o.**
- a t .e s 6 SNnTn-ENIERY CONIIhNY CH E astSTS
- TE STING
- INSPECTION
- ENGINEE8tS 7 fi t EAST W A S ** 1 N .'", 7 0 N R O t; t. ( V a n n . g n". A *4 G E L F S C ALIF ORNI A 90021 . 1213) 749-3411 314 3 - Q LA P ALM A A V F NtJE . A N A 6. F I S.9 Cat:5 0RNI A 92806 . a714 saa.4s t o
....................................-....-.. .....---.....,..............................<..cs...........
....................................-..............e.............
r.La No.: 8246 D.,r Oytober 16, 19g1 t . . . . ,,, . . u n . : 81-809 Structural Mechanics Associ.ites 5160 Birch Street RECENED OCT 2 01 Newport Beach, California 'R660
]y gg g gO]
A ttention : Greg S. lia rd y I S UBJECT : CO!!PRESSION TESTS ON 1/2" TilICK CEIDTEX FIBERBOARD MATERIAL 7 Qh kid l l SOURCE: Submitted to our Laboratory k , i hc41 REPORT OF TESTS In compliance with your requent, we have conducted ccepression tests on 3-inch by 3-inch, and 3-inch by 6-inch specimens of Celotex fiberboard cut from a 6-inch by lil-inch by 1/2-inch thick sample submitted to our I laboratory. The tests were conducted in a universal testing machine with the specimen I 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 I inch accuracy were placed on each side of the specimen; evenly spaced frcm the centor of the specinen to measure the deformation of the specimen. A load was applied to the specimen and thc corresponding average defor a tion -- reading recorded. The lcad was then released and the defomation again recorded. The load was increased and the defomations again recorded, I followed by unloading. This acquence was continued with increasing loads until a reasonable curve could be established and the defomation exceeded 50% of the origin.tl thickness. Load deflection readings are shown on the attached Plates A and B with acconpanying load-deforination curves. I Page 1 of 2 A-1
I SMITII-EMERY COMPANY File No. 8246 I.aboratory No. 81-809 I
~
Structural Mechanics Associat.cs CEIOTEX FIBERBOARD COMPRESSION TESTS October 16, 1981 I We are also encloaing a copy of the most recent calibration of our Tinius-01sen universal testing machine used for these tests. I Respectfully submitted, SMITH-EMERY COMPANY m ^
/
/
Byli" ' M I. . _ PAUL LINSTROM Civil Engineer 2-Addressee
- Attachments PL
- ks i
i I Page 2 of 2 A-2
I SMITil-EMERY COMPANY File No. 8246 f.iboratory No. 81-809 STRUCTURAL MDCHANICS AUSvCI ATES October 16, 1981
~
1 CEIOTEX COMPRESSIOff TCST Sample Size: 3" x 3" x 1/2" I Lo1d , Lbs. Deflection,, In. Load, Ibs. De flec tion , In. 0 .000 1000 .151 240 .050 0 .083 0 .017 1500 .184 300 .066 0 .108 0 .026 2000 .210 l W 350 .068 0 .129 0 .029 2500 .232 400 .075 0 .151 0 .034 3000 .251 450 .083 0 .173 0 .036 3500 .264 500 .090 0 .190 O .039 4000 .274 550 .097 0 .200 0 .046 4500 .295 600 .104 0 .235 0 .048 5000 .307 650 .110 0 .243 I 700 0 .055
.120 5500 0
.312
.249 O .057 6000 .316 I 750, 800 0
.124
.063
.129 0
7000 0
.251
.323
.267 i
I 850 O 0
.070
.134
.070 8000 0
9000
.334
.277
.342 900 .144 '
O .291 I 950 0 0
.078
.147
.081 Set Af ter 5-Minutes:
0 .274 1 PLATE "A" A-3 I
I i I. - i S M I Til- E f1E R Y COMPANY File No. 8246 Laboratory No. 81-809 STRUCTURAL MECitANICS ASSOCI ATI'i l October 16, 1981 CELOTEX COMPRESSION TEST I 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
.241 I .041 500 6,000 0 .017 0 .179 1000 .084 6,500 .250 0 .045 0 .189 1500 .120 7,000 .255 0 .060 0 .198 1750 .130 7,500 .260 0 .072 0 .203 2000 .143 H,000 .267 0 .083 0 .212
.158 8,500 .271 I 2250 2500 0 .095
.166 9,000 0 .217
.277 0 .102 0 .223 2750 .173 10,000 .283 0 .111 0 .231 3000 .184 11,000 .290 O .118 0 .237
$I 3250 0
.186
.126 12,000 0
.295
.245 3500 .196 14,000 .303 I 3750 0 .131
.201 16,000 0
0
.253
.311
.262 0 .139 4000 .208 18,000 .319 0 .148 0 .278 4500 .218 20,000 .325 0 .153 0 .284 5000 .225 .: set A f tpr 5-Minuter -
0 *
.162 0 .270 B
PIATE "B" A-4
File No. 8246 i I Lab No.
-;:g-81-809 S M IT H.> E M E RY COMPANY
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g _ UNITED CALIBR ATION CORP. 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 Anaheim, Ca. T.O. SUPER "L" U.T.M.
CAP: 120,000 Lbs. S/N 80700-4 6W U AMBIENT TEMPER ATUR E 1 DIV. ON DI AL = 2.5 Lbs.
,,D-3,0 s. en o, O___ O o a v l
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> 1 DIV. ON DI AL = 10 W s. ,n
, , ,,0 - 1 2 Lbs.
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,is,0-30,000 Lbs.
O V A V O V h
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.is
- 3,000 6,000 12,000 18,000 24,000 30,000 Lbs.
CAllSRATION APPARATUS USED Method Used O ASTM E4-79 Machine Meets M E4-79 N.D.S. TRACEABLE MOREllOUSE PROVING RINGS calibratedO )QD 1 l' by National Standards Test Lab: - , 100,000 Lbs., Serial #1431, 3/29/79, 1/20th of it, NBS NSJT.01/101468. ~N_ 20,000 Lbs., Serial #2820, 3/21/79, 1/20th of 14, NBS #SJT.01/101469 g g mge E g g 2,000 Lbs. Serial #3010, 3/22/79, 1/20th of 14, NBS #SJT.01/101469.
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- U NITED CALIBR ATION CORP. .
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@ PHONE (714) 893-1821
- 638 2322 GE 2 & 2 :
12761 MONARCH STREET
- GARDEN GRO /E, CALIFORNIA 92641 CERTIFICATE OF CALIBR ATION DATE February 26, 1981 TYPE OF MACHINE OR APPARATUS FOR:-SMITH EMERY T.O. SUPER "L" U.T.M.
Anaheim, Ca. 120,000 Lbs. S/N 80700-4 5.0. No. CAP: AMBIENT TEMPER ATURE 68"P 1 Div. ON otAL = 100 Lbs. ,,,, 0-120,000 Lbs.
+s.
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Mothed Used S ASTM E4-79 Machlae Meets S E 4 -79 CAllBRATION APPARATUS USED N.D.S. TRACEABLE MOREHOUSE PROVING RINGS calibratedO E 1 '* by National Standards Test Lab: -- 100,000 Lbs., Serial #1431, 3/29/79, 1/20th of 1%, NBS #SJT.01/101468. F k8 5"vde Enguwe 400,000 Lbs., Serial #3448, 12/26/79, 1/20th of 1%, NBS #SJT.01/101637. Wa'*lter I. Brown h _ m.oa.. ,}}