ML20040C573
| ML20040C573 | |
| Person / Time | |
|---|---|
| Site: | Midland |
| Issue date: | 01/31/1982 |
| From: | Banon H, Campbell R, Hardy G STRUCTURAL MECHANICS ASSOCIATES |
| To: | |
| Shared Package | |
| ML20040B738 | List: |
| References | |
| SMA-13704.01, NUDOCS 8201290033 | |
| Download: ML20040C573 (58) | |
Text
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SMA 13704.01-sfoo/
EVALUATION OF MIDLAND NUCLEAR POWER PLANT B0 RATED WATER STORAGE TANKS FOR NON-UNIFORM SUPPORT LOADING RESULTING FROM RING WAl.L SETTLEMENT prepared by R. D. Campbell G. S. Hardy H. Banon Approved by:
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R. 'P.
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T. R. Kipp President Manager of Quality Assurance prepared for CONSUMERS POWER COMPANY Jackson, Michigan January,1982 g g STRUCTURAL mECHRnKS
- RSSOCIATES Y
A Caht Coes 516o Birch Street, Newport Beach, Cahf. 92660 (714) 833 7552 8201290033 820118 i
PDR ADOCK 05000329 l
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TABLE OF CONTENTS Section Title Page LIST OF TABLES.................
ii LIST OF FIGURES iii 1
INTRODUCTION..................
1-1 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 S u mma ry..................
2-1 3
ACCEPTANCE CRITERIA 3-1 3.1 Governing Codes and Standards.......
3-1 3.2 Stress Criteria for 5ettlement 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 Locatior 5-3 REFERENCES APF6NDIX A i
i
6 LIST OF TABLES Table Ti tle Page 1-1 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 l
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o LIST OF FIGURES Figure Title Page 1-1 Plan View of Tank 1T-60 Identifying Bolt Numbers and Location Angle e..............
1-10 1-2 Plan View of Tank 2T-60 Identifying Bolt Numbers 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 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-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 Menbrane Stresses in Shell Due to Anchor Bolt Loading 4-15 iii
e 1.
INTRODUCTION, t
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 supporting 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 indicated.that some cracking of the ring walls has occurred. The ring wall deformation has resulted in a non-uniform support condition for the BWSTs. Examination of the tanks indicated that some of the anchor bolts 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 this report.
1
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 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.
1-1 l
e The two tanks are located outdoors in the tank farm area, north of the Auxiliary Building. The Unit 1 tank, 1T-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 I 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, 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 surfaces of the ring walls. Table 1-2, taken from Reference 2, depicts elevations measured at tne 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 surface 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 surface 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 common point, the elevations are identical. 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 1-2
e 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 cont actor to the loaded anchor bolts and the nuts 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 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 plane along the bottom surface 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.
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 rino 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 ring walls with the anchor bolts all unloaded. The current unloaded anchor bolt condition is much less critical to the tanks than the prior conditica with some of the bolts loaded, trying to force the tanks walls to the contours of the ring wall surfaces.
l Visual observations of the BWSTs with the bolt loads still l
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 l
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. Tha 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 permanent deformation 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 cccur, 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 represent the BWSTs cylindrical walls. The model was constructed of flat plate elements possessing both bending and membrane stiffness. Vertical boundary elements are utilized at the bottom surf ace to represent the nonlinear behavior of the asphalt impregnated fiberboard 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 lcad-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 properties chosen to represent the stiffness of the tank roof and the restraint that it offers to the tank at the roof / cylinder intersection.
i 1-4
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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 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.
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1-5
TABLE 1-1 PRE-ERECTION ELEVATION DATA FOR THE BORATED WATER STORAGE TANKS (1T60 & 2T60) FOUNDATIONS ANGLE e ELEVATION (FEET)
ELEVATION (FEET)
BOLT NUMBER (CLOCKWISE FROM NORTH)
355.5 4.48 5.09 2
4.5*
4.46 5.07 3
13.5*
4.47 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 1
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
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27 229.5*
4.19 5.09 28 238.5*
4..! S 5.08 29 247.5 4.47 5.07 30 256.5*
4.47 5.08 l
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
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36 310.5 4.47 5.08 37 319.5*
4.48 5.09 38 328.5 4.48 5.08 39 737.5*
4.47 5.09 40 346.5*
4.47 5.08 i
1-6
4 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 18'0*
634.83 634.96 210*
634.78 634.95 240*
634.77 634.95 270*
634.79 634.96 300 634.82 634.96 330' 634.86 634.95 1-7
4 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 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 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
- BOLT LOCATIONS CORRESPONDING W:TH THESE BOLT NUMBERS ARE LISTED IN TABLE 1-1 l
1-8
TABLE 1-4 MEASURED LOADS IN BOLTS ANCHORING TANK 2T-60 BOLT NUM3ER LOAD (KIPS)
BOLT NUMBER LOAD (KIPS) 1 0.0 21 0.0 2
0.0 22 0.0 3
0.0 23 0.0 i
4 0.0 24 0.0 5
0.0 25 0.0 6
1.19 26 0.0 7
2.82 27 0.0
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8 0.24 28 O.0 9
1.15 29 0.0 10 0.0 30 0.0 11 0.0 31 0.0 12 0.16 32 O.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
- BOLT LOCATIONS CORRESPONDING WITH THESE BOLT NUMBERS ARE LISTED IN TABLE 1-1.
1-9
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NORTH l l 40 1
2 3
39 q
4 38 g
6 37 7
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36 as 8
w 35 a0 = 9* TYPICAL TANK iT-60 10 33
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PLAN VIEW OF TANK IT-60 IDENTIFYING BOLT NUff8ERS AND LOCATION ANGLE e 1-10
NORTH 0
40 3
2 3
39 4
38 37 7
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34 a0 = 9 TYPICAL 33 10 32,
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PLAN VIEW OF TANK 2T-60 IDENTIFYING BOLT NUMBERS AND LOCATION ANGLE e 1-11
0.10 ANGLE O b
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- ANGLES ARE HEASURED CLCCKWISE FROM DUE NORTH PRE-ERECTIO.*4 RELATIVE ELEVATION DATA FOR THE TANK IT-60 RING WALL
- - - RING WALL ELEVATION DATA TAKEN JUNE 15, 1981 i
FIGURE l-3.
COMPARISON OF TANK IT-60 RING WALL RELATIVE ELEVATIONS BEFORE AND AFTER THE GROUND SETTLEMENT 4
2.
SUMMARY
AND CONCLUSIONS 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 along the vertical, resulting in elements that are approximately four 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 roof and the tank bottom. The gap-type elements shown in Figure 2-2 represent the nonlinear compressibility aff orded by the asphalt impregnated fiberboard at the interface between the tank bottom and ring wall.
Figure 2-3 shows the resultant displacement 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 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 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 gaos.
2-1
Maximum pri. nary membrana stress intensity occurs in element number 392 which is in the first row of elements above the shell thickness change from 0.375 incnes 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 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 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 results from hydrostatic pressure. Stresses are summarized ir. 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 NASA developed buckling fomulas, Reference 6, for axially and noment-loaded thin cylinders. The NASA formulas predict lower bound values of buckling stress considering experimental data as well as theory. Calculated critical buckling i
stresses for cylinders stressed uniformly along the axis and around the circumference are 5050 psi axial compression in the 0.375 inch thick wall
'nd 2690 psi axial compression in the 0.25 inch thick wall.
In the situation at hand, axial stresses are not uniform around the circumference nor along the length. Each of these factors would increase the critical l
buckling stress. From Reference 6, critical axial stresses for buckling l
of a cylinder in bending would be 7950 and 4750 psi for the 0.375 and l
0.25 inch thick walls, respectively. These are considered to be more representative values for the 3WST condition where axial stress varies from positive to negative around the circumference and is not uniform along the length.
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l 2-2 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 much less than the predicted buckling stress level for bending. The buckling f actor of safety, based on elastic buckling 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 immediate danger of buckling in the event that further settlement and ring wall distortion occur prior to retrofit of the ring walls.
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 report, Reference 7 reveals that maximum service loading is predicted for the safe shutdown earthquake 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 deformation or damage has occurred in the anchor bolts. Anchor bolt pullout was 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 shimmed and 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-1 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 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
material-like stainless steel, the effective plastic hinge will not occur until the elastically calculated load controlled stress is greater than 150% of yield.
In addition, stainless steel yield strength is typically much greater than minimum 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 minimum 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 tco 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 are essentially at this value.
In Chapter 3, an acceptance criteria..
developed that would allow settlement induced stresses to be comparable to stresses nomally 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 (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 criterion is exceeded by 41.6%.
In this case, limit analysis acceptance criteria are applied.
l The yield line (lower bound limit analysis capacity) was deter-mined to be 39.95 kips. As discussed in Chapter 3, the limit anlaysis I
acceptance criterion was derived from ASME code philosophy and is 0.8 of i
the lower bound limit analysis capacity. The maximum measured bolt load i
of 31,313 pounds is 0.78 of the calculated limit load capacity and the bolt chair it cansidered acceptable as is.
l r
2-4
Since some 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 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.
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 minimuin 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 assumed 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 acceptance criterion of Chapter 3 for primary local membrane stress intensity.
2-5
Secondary stresses are not limited by the governing code, Reference 5.
Primary plus secondary stresses shculd 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 secondary stresses calculated by the design fonnula of Reference 8, when combined with stresses induced by other loading are 1.93 S ; thus, shell y
stresses are within acceptable limits.
The following sumarizes the most severe stress condition in tank 1T-60 prior to relieving the anchor bolt loads, a.
Maximun 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 f actor of safety on elastic buckling is 246.
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 recommended 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.
2-6 I
d.
Local membrane stress intensity in the tank wall at the maximum 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 S which y
is less than the shakedoko 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.
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NOTE:
INNER CIRCLE OF NUMBERS REPRESENT N0 DES ON THE TOP 0F THE TANK.
OUTER CIRCLE OF NUMBERS REPRESENT N0 DES AT THE OUTER END OF THE B0UNDARY ELEMENTS AT THE TANK BOTT0fi.
FIGURE 2-1.
PLAN VIEW 0F TANK MODEL 2-8 m.
i B0 RATED WATER STORAGE TANKS AT MIDLANDS 99 9 9 k
k k
k k
k k9 9 99 E@ @
E U
C
~ " '
CLOCKWISE ANGLE FROM HORTH I 1 I
I I
I i
l~
~1 I
i e
i i
) i BEAM ELEMENTS N x N
63)
(fp)
N0 DES 1-40 NODES 41-80 N0 DES81-120
_.p
/
N0 DES 121-160 SHELL ELEMENTS
+-
\\
N0 DES 161-200 7
o N0 DES 201-240 N0 DES 241-280 N0 DES 281-320 ELEVENTS
@ @ b b 9
NODES 321-360 (TANK N0 DES )
N0 DES 401-440 ( BOUNDARY NODES)
GAP ELEMENTS *h h
+
tiODES 361-400 (BOUNDARY N0 DES) in,, i,,,,,,
m z7 zr
- r.,
rr zr er nr mi a r a rs,.,
@ - Represent Element Numbers FIGURE 2-2.
ELEVATION VIEW Of TANK MODEL
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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 faulted 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 NC3300 of the ASME code does not specifically address flat bottoa 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 folicwing stress intensities are allowed.
Loading Primary Local Membrane plus Condition Primary Membrane, P Primary Bending, PL+Pb m
Design and Normal S
1.5S Upset 1.1S 1.65S l
Emergency 1.55 2.05 l
/
Faulted 1.8S 2.4S l
Testing 1.255*
1.875**
Not to exceed 0.9 Sy
- Not to exceed 1.35 S l
y 3-1
i 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). Mininen 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 minimum specified yield strength and that primary local membrane and primary bending stress allow-ables al*, 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 shinned to provid 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 required.
If the higher emergency condition allowables were used, some additional inspection :nay be warranted to assess deformation.
Flat bottom storage tanks are not hydrotested, but the basic ASME code philosophy regarding stress and pemanent strain limits during pressure vessel hydrotesting is considered a rational philosophy fur setting a conservative stress acceptance criteria for comparison to elastic stress analysis results.
If elastically calculated stresses due to soil settlement do not exceed codc acceptance criteria for testing conditions, then the component shtuld 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 (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 3-2
membrane and 1.87S for primary local meinbrane 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 primary local membrana 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 deformation. 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 stress?s 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 infnaquent (emergency) events. Note also, that if secondary stresses were limited to the shakedown regime (2 S ) that during hydrostatic y
testing at 1.25 times the design pressure, primary plus secondary stress intensity could reach 2.5 S. Thus, code philosophy would allow primary y
plus secondary stress intensity to exceed the shakedown limit for single or limited 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 some additional inspection requirements or limiting inelastic strains.
3-3 l
There is only one case wheie calculated stress intensities exceed the recommended allowable stress limits. This is in the bolt chcir too plate for the most severely loaded bolt (bolt location 27).
In this case, limit analysis concepts are employed. The ASME code for 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 0.8 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 pemanent def amation may be experienced but the component is still serviceable. Def emation 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 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 cn 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 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 comprec:ive stresses exist in the tank wall. For the large diameter thin wall storage tank, beckling will occur in the elastic range. The ASME code Duckling criterion fcr axially loaded cylinders nominally contains a safety f actor l
3-4 l
of 3 for sustained design loads. Since the condition under consideration is local and is more of a strain controlled condition than load controlled, a more libe.'al buckling criteria is recommended.
Reference 6 provides thin shell buckling formulae modified to reflect extensive test data. Formulae are provided for the case of uniformly axially loaded cylinders and cylinders subjected to bending moment wherein the axial canpressive stress peaks at one location. The formulae contain correction factors based upon experimental data and are considered to be lower bound fonnulae. 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-pri ate. For the geometry under consideration, the applicable formula is:
0.6 y Et where o
=
cr R
correction f a,ctor for cylinders in bending Y
=
E Young's modulus
=
shell thickness t
=
R mean radius of shell
=
For the bending case an'd the gesnetry under consideration:
i 0.35 for the 0.25 inch thick shell
=
0.39 for the 0.375 inch thick shell y
=
Critical bu:kling stresses are:
4750 psi for the 0.25 inch thick shell 79E0 psi for the 0.375 inch thick shell 3-5
For strain controlled conditior', the allowable axial compressive stress should be limited to acr/' 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 deterndne 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.
3-6
4.
ANALYTICAL MODELS AND ANALYSIS METHODS A finite element model was contructed to represent the cylindrical shell portion of the BWSTs.
Boundary conditions were applied at the top and botton of the model to represent the restraint offered by the unbrella 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.
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 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 t wnents along the vertical axis. The elements are almost square being 49 inches X 48 inches. Each tank has 40 anchor bolts and, thus, the choice of 40 elenents around the circumference.
Anchor bolt locations corre.: pond to nodal points along the bottom of the model.
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 thic'< ness. 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 umbrella roof is welded to a ring girder at the tank top and would keep the tank wall from ovaling under asymmetric 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 i
4-1
afforded by the 0.25 i1ch 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 elenents (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 canpression direction and zero load capacity in tension.
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-displacenent relationships. One specimen was 3" X 3" and 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 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 elenents.
In developing the bean on an elastic foundation model, the resulting stiffness was biased on the conservative (stiff) side. Stiffer boundary elements result in the reaction loads being concentrated in more localized regions. Assumptions used in the beam on elastic foundation models are described as follows.
Referring to Figure 4-1, a rigid boundary is assumed on the tank i
bottan at the junction of the Celotex and the oil impregnated sand. This l
tends to be conservative in that the oil impregnated sand is assumed to 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 tr,nk inside wall
}
to the inner edge of the Celotex.
4-2 l
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 assumed for the Celotex material for discrete disp'acement ranges. For larger displacements, the equivalent spring t ates 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 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 conditions present in the BWSTs are:
Dead weight of the tank roof Dead weight of the tank wall Anchor bolt loading l
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 thit would best represent the physical forces acting on the tanks.
1 i
4-3 l
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. f +3d weight of the tank was applied by specifying a 1g 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 moael. 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 represent 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 modal geometry 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.
In areas where gaps exist between the tank bottom and ring wall, the weight of an effective annulus of water that extends radially inward i
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 j
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 l
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 l
weight is, thus, nonlinear, depending upon the gap between the tank bottom and its support. For the range of gaps calculated in the final i
4-4
l t
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 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 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 shifted 0.021 inches. The computer program boundary element reactions carried into the shell are correct and resulting shell stresses are correct. However, computed displacements mlative to the starting point are 0.021 inches greater than actual and the stan of the vertical 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.
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 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 there is no effective loading on the tank wall.
4-5
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 c9aducted to assure correct input for loadings. The relative contributions of downward loading from dead weight, anchor bolt loads and vertical water pressure loading niacted by the shell are:
Tank Roof 36,000 lbs Tank Shell and Hardware 67,090 lbs Effective Water Weight Reacted by Shell 115,045 lbs Total of Bolt Loads 168,470 lbs 386,605 T5s From the vertical force sumnary, 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.
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 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 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.
4-6
Stresses in the tank wall due to bolt chair loading were calculated usirg the methods of Reference 9.
The original design analysis used methods in Reference 8 to compute stresses in the shell f rom bolt cha',r loading. Reference 8 does not, however, separate primary 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 attachment 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 assuued 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.
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.
4-7
P r~
Tank Wall ke 7 [
Rigid Boundary Tank Bottom w
I Y
i V
U Jr y
y P
1r t
U U
'r 1
L 2 f.
1 5YT55555555 N
Ring Wall Asphalt Impregnated Fiberboard Oil Impregnated Sand E
average spring rate of asphalt impregnated fiberbeird over
=
finite deflection range, lbs/in/ unit width water force, lbs/in/ unit width w
=
Load into tank / unit width P
=
Rotational stiffness of tank wall ke
=
FIGURE 4-1.
BEAM-ON-ELASTIC-FOUNDATION MODEL 4-8
1500 k = variable Beam on elastic j
m foundation respons m
j$ 1000 K = 3800 #/in e
E k = 1000 #/in/in t-g E = 800 #/in/in
_k = 800 #/in/in 500 E
/
/
e e
1 0.1 0.2 0.3 DEFLECTION-INCHES E = Average spring rate at asphalt impregnated fiberboard over finite deflection range, lbs/in/in/ unit width K = Average spring rate of tank support, lbs/in/ unit width FIGURE 4-2.
LINEARIZATION OF BOUNDARY SPRINGS 4-9
_a
~
l t
=
tr
=
P C Tank Wall Tank Bottom W
U y
y 9
- N_
q NQt
(
6 7
V u
v ]-
k j
Asphalt Impregnated Fiberboard 011.. Impregnated Sand Ring Wall I
water force, lbs/in/ unit width w
=
tank and roof weight, lbs/ unit width P
=
gap between tank bottom and asphalt impregnated fiberboard, in.
a
=
effective width of water annulus, in.
a
=
r l
FIGURE 4-3.
EFFECTIVE WATER ANNULUS l
l 4-10 l
I 150 g
y 120 sy 100 S
J M2 50 5
ti 2
sp i
i i
S 0.2 0.4 0.6 0.8 1.0 g
Gap - Inches FIGURE 4-4.
EFFECTIVE WATER FORCE / UNIT WIOTH OF CIRCUMFERENCE VS GAP 4-11
a 800 E
600 3
K = 3800 #/in
~
b E
P D
U 400 m
5p 4u 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 BOUNDARY ELEMENTS
l
\\
TANK WALL I
I i
\\NJ A
f 7
+d+
1/3 ANCHOR BOLT LOAD
\\)l y
(
sr u - rv v c
1 h
h h
\\
i PARTIALLY FIXED I
I lyI c
9 FIGURE 4-6: BEAM MODEL FOR BOLT CHAIR DESIGN 4-13
= YIELD LIllES
[
( (
/
(
g J U;_ _ J
)
3 3
+
/
I l
3) l=
= -l c 's l
l 6
- r
,1 r,ff s
\\
p
% /
P PT 2
-=- r, f f
-6 U
Gusset 1 f 6
I h
] f l
s' late Load i
i x
i l~~. l T
w FIGURE 4-7.
YIELD LItiE MODEL FOR BOLT CHAIR 4-14
P e,,
9 l
Pe m
h A I 4 s_ 0.625" 0.375"
+
312"R h
P = ANCHOR BOLT LOAD
.P e, i f h
w
' FIGURE 4-8.
ANALYSIS MODEL FOR LOCAL MEMBRANE STRESSES IN SHELL DUE TO ANCHOR BOLT LOADING 4-15
5.
ANALYTICAL RESULTS Results from the finite element model analysis and hand calculations are summarized in this chapter. Computer output is too voluminous to include, consequently, only important highlights of the output are sumnarized.
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 t ank. 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, 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 asymmetric vertical reaction loads.
l Because of the wave front solution techniques used in ANSYS, elements in a column are in numerical suequence while elements in a row (around the circumference) are riumoered 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 elenents 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 raquestad. Table 5-2 summarized stress components for each 5-1
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 stresses are more concentrated along the tank bottom resulting from concentrated anchor bolt loading and uneven vertical reactions from the 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 nodel used to evaluate limit load capacity of the top plate of the bolt chairs.
In this model, plastic hinges were assumed 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 assuming elastic-perfectly-plastic material behavior, the plastic hinge moment capacities are computed to be:
t 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
' 4 P
= 39.95 kips s
limit
~ '
The limit 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 l
5-2 N
5.3 TANX WALL AT BOLT CHAIR LO, CATION Design calculations contained in Reference 7 indicate tank wall stresses in excess of yield for the f aulted condition loacing (SSE plus normal operating loading). Calculations are based on a formula given in Reference 8.
The fomula 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 second'sry stress intensity. This is justifiable if the components do not undergo a large number of cycles of loading. For the case under consideratf on, the loading is a one-time-only event and secondary stresses on 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 assumed 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 fx load distribution into the shell from the gusset plates.
In applying Reference 9 to the problem, non-dimensional membrane forces R Nt/P and R N /P are given for R/t ratios up to 300. The R/t m
mx 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 detemined anchor bolt load of 31,310 pounds, the local primary membrane stress components were determined to be:
13.20 ksi c
=
e 12.16 ksi
=
x 5-3
\\
These stress components were combined with stress components from the finite element analysis. The finite element analysis resultant stresses account for all loading conditions excep. for the bolt chair momen t.
Maximum bolt chair loading occurs between elements 130 and 152.
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:
Element c e x
'ex 130 6385 1640 331 152 6383 1589
-13 Avg.
6384
'1614 159 Resulting local membrane ' stress components are:
19,584 psi c
=
e o,
13,744 psi
=
159 psi
=
r i
ox 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 reaction load is about midway between the tank bottom and the mid point of the element; thus, actual hydrostatic pressure induced hoop stresses would be lower than asstrned above.
5-4
2.
Computed local membrane stresses are conservative since the total moment reaction is assumed 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.
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 assuned 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.
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 shakedown philosophy and is of no concern.
y l
5-5
TABLE 5-1 STRESSES AT CENTER OF ELEMENTS BOTTOM THREE R0WS - PSI R0W 3 ELEMENT NO.
o o
T 3
g x
xe 7
10581
-1429 544 12060 18 10600
-1106 761 11805 40 10619
-406 918 11176 62 10633 127 858 10703 84 10642 199 799 10703 106 10661 768 721 10713 128 10685 1708 460 10708 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 282 10621
-536
-757 11258 304 10615
-509
-532 11174 326 10614
-457
-649 11146 348 10608
-894
-434 11536 370 10587
-1223
-314 11826 392 10571
-1923
-85 12495 414 10582
-1778 301 12374 436 10604
-575 543 11232 R0W 3 e
"x Txe-S ELEMENT NO,.
29 10573
-1885 332 12475 l
51 10516
-1807 84 12385 73 10590
-1189 281 11792 95 10596
-1093 459 11726 117 10593
-1262
-638 11924 139 10606
-893
-851 11624 161 10624
-116
-912 10894 183 10641 188
-814 10704 205 10657 650
-790 10719 227 10676 1431
-526 10706 249 10684-1594
-197 10689 271 10679 1337 107 10681 293 10670 1199 337 10682 l
315 10659 824 619 10698 l
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
TABLE 5-1 (Continued)
R0W 2 ELEMENT NO.
o 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 349 10870
-752
-258 11632 371 1085A
-908
-215 11770 393 10897
-1503
-67 12400 415 10908
-1450 196 12364 437 10882
-381 404 11232 R0W 2 ELEMENT N0.
c T
3 e
x xe 30 10900
-1488 248 12396 52 10899
-1433
-80 12332 74 10862
-907
-177 11774 96 10864
-857
-310 11738 118 10869
-985
-417 11882 140 10858
-714
-573 11628 162 10808
-52
-618 10930 184 10807 146
-520 10832 206 10787 503
-547 10817 228 10742 1156
-344 10755 250 10741 1263
-133 10743 272 10757 1043 79 10758 294 10753 968 207 10758 316 10774 559 434 10793 338 10800 246 451 10819 360 10816
-19 539 10888 382 10819
-290 469 11148 404 10867
-799 539 11715 426 10845
-479 220 11332 5-7
TABLE 5-1 (Cont'nued)
R0W 1 EL EME NT NO.
o
- x T
3 0
x9 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 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 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 PO.
c "x
- x0 e
31 5812
-1739 280 7572 53 5818
-1683
-109 7504 75 5905
-997
-162 6910 97 5913
-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 1466
-137 6376 273 6338 1175 91 6339 295 6322 1151 184 6329 317 6255 765 459 6294 339 6173 281 428 6204 361 6109
-43 561 6254 383 6065
-251 433 6375 7
405 5961
- 99 5 601 7058 427 5994
-568 145 6568 5-8
TABLE 5-2 STRESSES AT CENTER OF LOWER EDGE OF ELEMENT, BOTTOM ROW - PSI R0W l ELEMENT NO.
o I
O X
X0 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
-971 218 372 3038
-961 227 394 2881 1740
-71 416 2879 1790 184 438 3174
-314 456 i
R0W 1 ELEMENT NO.
o s "2
The 31 2861 1746 280 53 2871 1690 109 75 3010 1004 162 97 3032 1001 317 119 3036 1136 408 l
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 j
317 3580 758 459 339 3447 274 428 l
361 3343
-50 561 383 3269
-258 432 405 3104 1001 601 427 3154
-574 145 5-9
.w..-..
-,c
_..__.-._m y _, --, - _ ___
TA8LE 5-3 VERTICAL REACTIONS INTO SHELL-LBS.
NODE REACTION 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
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
348 0
349 0
350 0
351 0
352 0
353 0
354
-22081 355 0
356
-21393 357
-14061 358
-26501 359
-37680 36 0
-25537 Sum of vertical reactions into shell =
33%C5 lbs 5-10
REFERENCES 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 2.
Bechtal Corp letter, L. H. Curtis, Bechtel, Ann Arbor to G. M. Ault, Graver Energy Systems, ' Midland Plant Units 1 and 2, Consumers Power 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 for Subcontract for Field Erected Storage Tanks for the Consumers Power Company, Midland Plant Units 1 and 2, Midland, Michigan."
S.
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 Administration, Sept.1965.
7.
Graver Design Calculation, Bechtel/ Consumers Power, Midland, Michigan, Borated Water Tanks - 1T-60 and 2T-60, Rev. 4 dated 11 Feb, 1980.
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. 0.C.
9.
Welding Research Council Bulletin 107, " Local Stresses in Spherical and Cylindrical Shells due to External Loadings", Welding Research Council, New York.
- 10. API 550, " Welded Steel Tanks for 011 Storage", Fif th edition and i
Sepplement 1, Oct.1973, American Petroleum Institute.
- 11. ANSYS, Engineering Analysis System User Manual, Rev. 3 Swanson Analysis Systems Inc., Houston, PA.
R-1 I
~. -
gs w
2 O
e I
t APPENDIX A 1
TEST LABORATORY RESULTS FROM COMPRESSION TESTING OF ASPHALT IMPREGNATED FIBERBCARD (CELOTEX)
?
i i
1 1
- w,
,4-.w.
- ~.- -,.,
,,e...
.------,y....,,.
_,. - -. - - - - _.. _. - -,. - - _. _. ~ _.. -.
.r,_
__,m...__-.-,-,.--,
Ma u aun e.use SMITH-EM ERY COMPANY CM gadlSTS
- TESTING
- tNss.ECTION. ENGINEERS 741 EAST W A S w s N r. T o N R o u t. E v a n n t F.S ANGELES C A LIFORNI A 90021. f 213) 749 3411 A N A w F t 8.9 Ca t IF ORNI A 92806 3:43 0
- t. A P A L eA A A v F NisE 7:41 s304sto
_.s.
_. _ _ _.. _ = _ _
........................-.e..........-...
=n
.....u-..
...a.--
....-...-...e=*cu...eae.....-
.-.-....~..~a...
.......-.........-c..........u.ue............
. mi.e No.:
8246 0.,r O?tober 16, 1981
- i...... v N o. :
81-809 Structural Mechanics Associate 5160 Birch Street RECENED OCT 2 01981
!!cwport Beach, California 9660 g y g g gO}
A ttention : Greg S. fla rdy
SUBJECT:
COMPRESSION TESTS Oti 1/2" TifICK CEIOTEX FIBERBOARD MATERIAL
!kN kN SOURCE:
Submitted to our Laboratot*/
REPGRT OF TESTS In compliance with your requent, we have conducted compression tests on 3-inch by 3-inch, and 3-inch by 6-inch specimens of Celotex fiberboard cut frcm a 6-inch by Ill-inch by 1/2-inch thick sampic 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 unifom 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 defomation
~
reading recorded. The load was then released and the defomation again recorded. The load was increased and the deformations again recorded, t
followed by unloading. This acquence was continued with increasing loads until a reasonable curve could be established and the defomation exceeded 50% of the original thickness.
Load deflection rcadings are shuwn on the attached Plates A and B with accompanying load-dofomation curves.
l l
Page 1 of 2 i
A-1
t S M I T 'l-EME RY COMPANY Filo No.
8246 Laboratory No.81-809 Structural Mechanics Associates CEIDTEX FIBERBOARD COMPRESSION TESTS October 16, 1981-We are also enclosing a copy of the most recent calibration of our -
Tinius-01sen universal testing machine used for these tests.
Respectfully submitted, SMi*H-EMERY COMPANY m
/
j.e ByI/'I '
M II.
PAUL LINSTROM Civil Engineer i-2-Addressee Attachments PL:ks Page 2 of 2 A-2
o S M ITII-E -iER Y CO M PA!!Y File No.
824G T.doentory No.81-809 STRUCTURAL MLCilANICS ASSuCI ATES October 16,'1981 CEIOTEX COMPRESSION TEST Sampic Size 3" x 3" x 1/2" Load, Lbs.
De ficc t ion, In.
Load, I.b s.
Deflection, In.
0
.000 1000
.151 240
.050 0
.083 0
.017-1500
.184 300
.066
, 0
.103 0
.026 2000-
.210 350
.068' O
.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 l
0
.039 4000-
.274 550
.097 0
.200 0
.046 4500
.295 600
.104 0
.235 O
.048 5000
.307 650
.110 0
.243 0
.055 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 0
.070 9000
.342 900
.144 0
.291 0
.078 Set Af ter 5-Minutos 950
.147 0
.274 0
.081 PLATE "A"
A-3
D S M I TII-E M E R Y COMPANY File No.
8246 Laboratory No.81-809 STRUCTURAL MECf!ANICS ASSOCIATI2s Octobcr 16, 1981 CEIDTEX COMPRESSION TEST 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 0
.017 0
.179 1000
.084 6,500
.250 0
.045 0
.189 1500
.120 7,000
.255 0
.0GO O
.198 1750
.130 7,500
.260 0
.072 0
.203 2000
.143 8,000
.267 0
.083 0
.212 2250
.158 8,500
.271 0
.095 0
.217 2500
.166 9,000
.277 0
.102 0
.223 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 3750
.201 16,000
.311 0
.139 0
.262 4000
.208 18,000
.319 O
.148 0
.278 4500
.218 20,000
.325 0
.153 0
.284 5000
.225 set Ar_ter 5-Minutes:
0
.162 0
.270 PIATE "B"
A-4
Fila Ho.
8246 Lab Ho.81-809 SMITH-EMERY COMPANY.
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U NITED 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: SMITil EMERY TYPE OF MACHINE OR APPARATUS DATE February 26, 1981 Anaheira, Ca.
T.O. SUPER "L" U.T.M.
CAP: 120,000 Lbs. S/N 80700-4 68'F P.O.NO.
AMBIENT TEMPER ATURE I
N I s.
D-3,0 s.
,i o,
o
-is
-is 300 600 1,200 1,800 2,400 3,000 Lbs.
2 olv 0" ol^' = 1 LbS-0-12,gtbs.
p
.is
.,s Q
g n
V V
o o,
- i a.
.i.
1,200 2,400 4,000 7,200 9,600 12,000 Lbs.
i DIV. ON DI AL = 25 Lbs.
, iso-30,000 Lbs, m
h O
A A
V V
V
-is
_is 3,000 6,000 12,000 18,000 24,000 30,000 Lbs.
i Call 8 RATION APPARATUS U5ED Method Used O ASTM E4-79 Machine Meets M
E4-79 1 It N.B.S. TRACEABLE HOREi!OUSE PROVING RINGS calibratedO XD by National Standards Test Lab:
g 100,000 Lbs., Serial #1431, 3/29/79,1/20th of 14, NBS #S.IT.01/101468.
'N g/ g) 3 20,000 Lbs., Serial #2820, 3/21/79, 1/20th of it, NBS #SJT.01/101469 g grvige E g g 2,000 Lbs. Serial #3010, 3/22/79, 1/20th of 14, NBS #SJT.01/101469.
A
}efhrucal Desector
.14rn H. Watann i
UNITED CALIBR ATION CORP.
lfo G) i h
12761 MONARCH STREET
- GARDEN GPOVE. CALIFORNIA 92641 PHONE (714) 8931821
- 638-2322 CERTIFICATE OF CALIBR ATION FOR:- SMITH EMERY TYPE OF MACHitBE OR APPARATUS IATE February 26, 1981 T.O. SUPER L'* U.T.M.
f.naheim, Ca.
CAP:
120,000 Lbs. S/N 80700-4 5.0. NO.
AMBIENT TEMPER ATURE 68*F 1 DIV. ON DI AL : 100 Lbs.
,0-120,000 Lbs.
N A
O 4
Q 6
-t1
~l%
12,000 24,000 48,000 72,000 96,000 120,000 Lbs.
1 DIV. ON DIAL D
co O
U
_g g
- t ".
1 DIV. ON DIAL :
,,g o
0 1%
~8 %
CAtl8RATIOle APPARATUS USED Method Used 9 ASTM E4-75 Machine Meets S
E4 -73 N.B.S. TRACEABLE MORE!!OUSE PROVING RINGS calibratedO 911*
by National Standards Test Laba h"-
100,000 Lbs., Serial #1431, 3/29/79, 1/20th of it, NBS #SJT.01/101468.
400,000 Lbs., Serial #3448, 12/26/79, 1/20th of 14, NBS #SJT.01/101637.
Field Service Engineer Walter I. Brown
- UL, N
Ili'. h*Nson
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