ML20238F396

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Evaluation of Seismic Capacity of Condensate Storage Tanks (CST) at CCNPP Using Local Nonlinear Finite Element Model for Bolt-Chair-Tank Shell Interaction
ML20238F396
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Site: Calvert Cliffs  Constellation icon.png
Issue date: 02/08/1996
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STEVENSON & ASSOCIATES
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NUDOCS 9809030330
Download: ML20238F396 (98)


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ENCLOSURE (C)

" Evaluation of Seismic Capacity of Condensate Storage Tank (CST) at the Calvert Cliffs NPP Using a Local Nonlinear Finite Element Model for Bolt-Chair-Tank ShellInteraction," prepared by Stevenson & Associates l

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Baltimore Gas and Electric Company Calvert Cliffs Nuclear Power Plant

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l EVALUATION OF SEISMIC CAPACITY OF CONDENSATE STORAGE TANKS (CST) AT j

THE CALVERT CLIFFS NPP USING A LOCAL NONLINEAR FINITE EL5 MENT MODEL FOR BOLT-CHAIR - TANK SHELL INTERACTION

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Prepared by l

STEVENSON & ASSOCIATES 9217 Midwest Ave., Cleveland, OH Phone:(216)587-3805 Fax:(216)S87-2205 l

Prepared for Attn: Mr. Mark Wright Baltimore Gas & Electric Company Calvert Cliffs NPP 1650 Calvert Cliffs Parkway i

Lusby, MD 20657 l

February 8,1996 l

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l TABLE OF CONTENT l

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1. 0 I N TRO D U CTI O N......................................................'..........

1 1.1 S cope an d Obje ctive................................................................... 1 l

1. 2 B a ck g ro u n d.................................................................................... 1 2.0 NONLINEAR FINITE ELEMENT MODEL FOR BOLT CHAIR EVALUATION 2 1

1 2.1 Description of the ANSYS Computational Model............................ 2 2.2 Nonlinear Finite Element Analysis Results................................... 2 3.0 TAN K S EISMIC CAPACITI ES............................................................

3 3.1 Tenk Seismic Response Using SUPER SASSI/PC........................ 3 3.1.1 Design Ground Spectrum for 0.15g (for USI-A46)......................... 4 3.1.2 Uniform Hazard Spectrum for 0.40g (for IPEEE)........................... 4 l

3.2 Safety Factors and Tank Seismic Capacities..................................... 4 3.2.1 Deterministic Capacity with Respect to Design Basis (USI-A46)..... 5 3.2.2 Median and HCLPF Capacities Using Fragility Analysis (IPEEE)... 5 4.0 CON C LU DIN G R EMARKS.$.................................................................... 6

5. 0 R E F E R E N C E S............................................................................................. 7 APPENDIX A: Description of the Local Tank Shell-Bolt Chair ANSYS Model APPENDIX B: Plastic Strains and Displacements in the Tank Shell and the Bolt Chair Based On Nonlinear Finite Element Analysis (ANSYS Files) l B.1 Tank Response for A 42.5 ksi Tensile Stress in the Anchor Bolt B.2 Tank Response for A 56.0 ksi Tensile Stress in the Anchor Bolt l

l l

APPENDIX C: Tank Seismic Response Using SUPER SASSI code C.1 Design Ground Spectrum anchored to 0.15g ZPGA (USl A-46)

C.2 Uniform Hazard Spectrum anchored to 0.40g ZPGA (IPEEE)

APPENDIX D; Evaluation of Tank Seismic Capacities Using TANKV code D1 Deterministic Capacity (USI A-46)

D2 Median Capacity for 44 ksi tensile stress in bolt - yielding (IPEEE)

Median Capacity for 56 ksi tensile stress in bolt - ultimate ii

1.0 Introduction 1.1 Scope The scope of this evaluation is to perform a more detailed tank capacity analysis of.th.e Condensate Storage Tanks (CST # 11) using an advanced nonlinear finite element model for the tank shell - bolt chair interaction. Previous analyses indicated that the tank capacity was greatly reduced by the bolt chair capacity. The bolt chair and tank wall supporting the chair were unable to develop the full strength of the anchorage for the tank. This evaluation was requested by BG&E for further resolution of the outlier status of the CST and other flat bottom tanks (by qualitative comparisons) at the Calvert Cliffs NPP. The evaluation was also used to modify the fragility of the CST and other flat bottom tanks (by qualitative comparison) at Calvert Cliffs by obtaining a more accurate Median and High Confidence Low Probability of Failure (HCLPF) capacity for the tanks.

The results of the capacity evaluation were combined with the results of a rigorous seismic response evaluation of the tank including soil-structure effects. Two parallel analyses were performed to establish the safety margins and seismic capacities of the CST in accordance to (i) the USI-A46 requirements and (ii) IPEEE (fragility) requirements. For USI-A46 a Deterministic tank seismic capacity was determined using the Design Basis conditions and the results of the refined structural analyses. For

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IPEEE the HCLPF and Median tank capacities were computed using fragility analysis based on current PRA methodolgy (Ref.1) and the results of the refined structural analyses. The SUPER SASSIA3p code was used for seismic response evaluation and the ANSYS/PC code was used for capacity evaluation.

1.2 Background

Previous investigations performed by Gilbert & Associates (Ref. 2) and EQE (Ref. 3 and 4) indicated that the CSTs were outliers for both the USI-A46 and IPEEE programs. The critical failure mode of the CSTs was indentified to be tank overtuming.

The CST failure is caused by the buckling of the tank shell on the compression side combined with the yielding in the tank shell at the bolt chair on the tension side. Based l

on the previous calculations using GIP metiiodolgy for A-46 evaluation (Ref. 3) the reduction factor of the anchor bolt capacity due to the stress in the tank shell at the bolt l

chair was 0.17 (effective capacity of 17,279 K versus nominal capacity of 53.46 K for a minimum yielding stress of 36 ksi). For IPEEE evaluation (Ref. 4) the median reduction factor of anchor bolt capacity due to the stress levelin the tank shell was 0.36 (effective median capacity of 23.7 K versus median nominal capacity of 65.34 K for a median l

yielding stress of 44 ksi).

The resulting safety factors in terms of overtuming moments were computed in Ref. 3 and 4 to be 0.77 (determinsitic safety factor) with respect to the Design Basis Eathquake with 0.15g ZPGA and 1.21 (median safety factor) with respect to the Reference Median Uniform Hazard Spectrum (UHS) Earthquake with 0.40g ZPGA.

Using these safety factors the Deterministic seismic capacity of the CST was 0.129, which is below the Design Basis condition of 0.15g, and the HCLPF seismic capacity 1

IJ was 0.18g, which below the screening limit for IPEEE of 0.30g. The Median seismic capacity of the CST was 0.48g. Based on these results the CSTs were considered as outliers for both the USI-A46 and IPEEE programs.

The anchor bolt failure due to the pullout of J-bolts from the concrete (30 in embedmont) which was a concern in Ref. 3 when using the GIP methodolgy it is not a realistic issue. The pullout capacity of the anchor bolts as shown in Ref. 4 and Ref. 5 is considerably higher than the anchor bolt tensile capacity. Therefore, in the nonlinear finite element analysis performed herein to investigate in detail the tank shell-chair-bolt

- interaction, the bolt pullout failure was not considered.

L The main scope of the the nonlinear analysis was to investigate the potential reduction l

due to failure of the bolt chair and the bolt chair to tank interface.

L

- 2.0 Nonlinear Finite Element Model for Bolt Chair Evaluation L

L 2.1 Description of ANSYS Computational Model i

The nonlinear finite element ANSYS model of the CST bolt chair is shown in Figures 2.1 through 2.4. The tank wall and bolt chair were modeled using the plastic shell element SHELL43. The anchor bolt was modeled using plastic truss element, LINK 8.

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The contact with the concrets foundation was idealized by contact elements which l

permit the tank bottom edge uplift, CONTACT 52. The continuity of the tank in the hoop l

direction was simulated using axial spring elements, COMBIN14. The boundary L

conditions were selectsd to reflpct the symmetry conditions of the model with respect l

with a vertical plane containing th'e anchor bolt.

9 The plastic behavior of the tank shell, bolt chair and anchor bolt were introduced by a nonlinear stress-strain relationship for each material. The multilineal stress-strain curves assumed for the tank wall and the anchor bolt are shown in Figures 2.5 and 2.6.

l They correspond to the median estimations. Thus, for the tank shell material, SA-240 l

SS (304L) the median yield stress and ultimate stress were considered to be 37 kai and 84 ksi, respectively. For anchor bolt material, A-36, the median yield stress and ultimate stress were considered to be 44 ksi and 64 ksi. It should be noted that these values are equal to those used in Ref. 4 for tank fragility evaluation. The median ultimate stresses l

were defined for an axial strain of 10%. The nonlinear finite element model is described l.

In more detailin Appendix A.

2.2 Nonlinear Finite Element Analysis Results l

The nonlinear analysis of the bolt chair was performed in two steps: (1) first the seismic water pressure was applied and then (2) an uniformly distributed tension at the top edge of the tank shell model was gradually increased until failure. Material failure criterium was initially defined when a 10% peak axial strain or the material ultimate tensile stress is reached.-

Figures 2.7 through 2.16 show the bolt chair behavior for two levels of tank wall tension due to global overtuming moment. Figures 2.7 through 2.11 illustrate the bolt chair i

2

l behavior for a tensile force which produces a 42.5 ksi tensile stress in the anchor bolt (close to the median yielding stress). For comparison, Figures 2.12 through 2.16 illustrate the bolt chair behavior for a tensile force which produces a 56.02 ksi tensile stress in the anchor bolt (close to 0.90 times ultimate tensile stress).

For the first tension level in the tank wall, i.e. 42.5 ksi in the bolt, the tank uplift is 0.20 in and the maximum peak axial strain is 1.5% in the top plate. At weld locations the peak axial strains are less than 1%. For the second tension level, i.e. 56.02 ksi, the maximum peak axial strain is 10% in the chair top plate. At weld locations the maximum peak axial strains are less than 4% and in the tank shell near chair top plate less than 1%. The corresponding stress distributions to the two tension levels,42.5 ksi and 56.02 ksi, in the tank shell-bolt chair model are given in Figures 2.11 and 2.16. It should be noted from these figures that the yielded zone in the tank shell at above the bolt chair largely increases from one tension level to the other one, maybe twice, showing that a significant redistribution in the tank shell stresses occur. The increase in the tank shell stresses is less significant around 10%.

Based on the nonlinear analysis results it has been decided that 10% axial strain criterium accepted initially to estimate the median capacity of the tank shell-bolt chair model maybe unconservative due to the high peak strains close to 4% in the weld (largest strain is in the bottom of tank wall weld). Therefore, a 5% axial strain criterium (2% strain in the weld) was accepted to reflect the median capacity of the tank shell-bolt chair model. The resulting tensile stress in the anchor bolt was 51.04 ksi, i.e.

around 0.80 times the ultimate tensile stress.

Nonlinear analysis results have.ghown that the failure of the tank on the tension side is govemed by the failure of the* bolt chair top plate due to large local bending strains near the bolt studs. At the same time nonlinear analysis has shown that the tank shell stresses at bolt chair are not a real concem if plastic redistribution of stresses is considered. Based on the nonlinear analysis results the the tank overtuming moment capacity has been recomputed without introducing any reduction factor to the anchor bolt nominal capacity due to the tank shell stress as shown in Section 3.2. The results of the nonlinear finite element analyses are given in Appendix B.

3.0 Tank Seismic Capacities i

3.1 Tank Seismic Response Using SUPER SASSIIPC The seismic analysis of the CST was performed using 3D foundation models to quantify the for soil-structure interaction (SSI) effects. The SUPER SASSl/PC code (Ref. 6) was used for this purpose. The structural model of the tank was the same stick model which has been used by EQE and S&A in some previous investigations (Ref. 7 l

and 8). The SSI tank model is shown in Figures 3.1 and 3.2. The SUPER SASSI/PC results are given in Appendix C.

3.1.1 Design Ground Acceleration for 0.15g ZPGA (for USI-A46)

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The earthquake input was an accelerogram compatible with the Design ground spectrum which is a Housner spectrum anchored to a 0.15g ZPGA. The accelemgram I

was defined at the ground surface. The earthquake input was the same with that used I

/ by S&A in a previous investigation (Ref. 8). Figure 3.3 shov'vs the computed versus design ground spectra. The soil deposit was considered as a visco-elastic halfspace having a shear wave velocity of 1072 fps (4164 ksf shea? modulus) which is 2 times i

larger the reduced equivalent shear wave velocity used in the original soil-structure interaction analyses (2057 ksf shear modulus and 40,000 psi Young's modulus). This value appears to be an appropriate estimate for an upper bound of soil shear wave velocity on the site based on the Design Basis condition. The structural damping ratio of the tank (impulsive mode) was considered to be 3%, which is more conservative than the 4% damping used by S&A in Ref. 8.

The computed overtuming moment at the CST #11 base including SSI effects was 6,276 k-ft, which sligthly larger than the results given in Ref. 8, but below the tank overturning capacity of 6,292 k-ft evaluated extremely conservative by EQE in Ref.7. As shown in Section 3.2, based on refined nonlinear finite element analyses, a higher tank overturning moment capacity was computed, and the CST has a significantly larger safety margin than was previously estimated using the GlP methodolgy (Ref. 7).

3.1.2 Uniform Hazard Spectrum for 0.40g ZPGA (for IPEEE)

The earthquake input was an accelerogram compatible with the Median UHS ground spectrum anchored to a reference level of 0.40g ZPGA. The accelerogram was defined at the ground surface. The earthquake input was the same with that used by S&A in a previous investigation (Ref. 9).hgure 3.4 shows the UHS ground spectra anchored to 0.10g ZPGA. The soll deposit was considered as a visco-elastic halfspace having.a shear wave velocity of 1480 fps. This value was based on a SHAKE 91 site response analysis performed using a Median UHS reference spectrum (anchored to 0.40g ZPGA) compatible accelerogram. This value of soil shear wave velocity is an appropriate best-estimate for the equivalent soll shear wave velocity under the UHS reference l

earthquake. The median structural damping ratio (impulsive mode) was considered to be 5%

The computed median overtuming moment at the CST #11 base including SSI effects was 12,870 k-ft. It should be noted that the 12,870 k-ft overtuming moment computed by S&A using SUPER SASSI/PC code compares favorably with the 12,152 k-ft overturning moment computed by EQE in Ref. 8 (page 9/25), using CLASSI code a more refined CST structural model for the same UHS reference earthquake with 0.40g ZPGA (3,038 x 4.0 = 12,152 k-ft).

3.2 Safety Factors and Tank Seismic Capacities Using TANKV code developed by S&A (Ref.10) the CST # 11 capacity were computed. The TANKV code peforms seismic evaluation for flat bottom tanks using Kennedy's approach (Ref.10, see also Appendix D). Based on the nonlinear analysis results, no reduction factor was applied to the nominal tensile bolt capacity due to the stresses in the tank shell at the bolt chair. No reduction factor for pullout capacity was introduced. It has been shown in Ref. 4 and 5 that for a 13/8" J-bolt with a 30 in 4

L embedment in concrete the pullout capacity for is considerably larger than the nominal bolt capacity.

3.2.1 Deterministic Capacity with Respect to Design Basis (A46) l The tank overtuming moment capacity computed using TANKV code and assuming a full capacity of the anchor bolts (36 A-36 bolts around circumference with 13/8" diameter) was 27,780 k-ft (Appendix D1). The computed overtuming moment for the 0.15g Housner spectrum input using SSI analysis was 6,276 k-ft (see Section 3.1.1).

Thus, it follows that the safety factor against failure due to tank overtuming is 4.43.The tank seismic capacity with respect with a Housner spectrum type earthquake follows to be 0.66g. This indicates a very high seismic margin of the CST, more than six times larger than that computed using GIP methodology (Ref. 7) which exaggerates the shell stress reduction factor for bolt capacity evaluation. When interpreting this result it should be also noted that the reference seismic input was defined by the Housner spectrum which has a lower spectral amplification than Newmark or RG 1.60 spectra.

Roughly speaking, this margin may drop to 0.459 if the RG 1.60 spectrum earthquake j

is considered as the reference earthquake input.

3.2.2 Median and HCLPF Capacities Using Fragility Analysis The fragility analysis was performed following the guidelines given by Reed and Kennedy (Ref.1). The tank failure criteria were assumed based on the strain levels in the tank shell and bolt chair. Using the TANKV code the median overtuming capacity of the CST was determined assur'r)(ng three strain levels which were associated to three i

tensile stress levels in the anchor bolt:

(1) MODERATE STRAIN LEVEL = BOLT YlELDING: 44 ksi tension in the bolt, which corresponds to a median yielding stress level (A-36 steet), and peak axial strains smaller than 1.5% in the tank shell, bolt chair plates and 1% in the welds and (2) HIGH STRAIN LEVEL = BOLT CHAIR " CONSERVATIVE" FAILURE: 51 ksi tension in the bolt, which corresponds to a 0.80 times the median ultimate stress level

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(64 ksi for A-36) and 5.0% peak axial strains in the tank shell, bolt chair plates and 2%

in the welds.

t (3) ULTIMATE STRAIN LEVEL = BOLT CHAIR FAILURE: 56 ksi in the bolt, which corresponds to a 0.90 times the median ultimate stress level for A-36 and 8.0 %

l peak axial strains in the tank shell, bolt chair plates and 4.0% in the welds.

The computed median overtuming capacities for the three strian levels were 30,068 k-

)

ft, 34,500 k-ft and 39,603k-ft, respectively. The computed overtuming moment using SUPER SASSI/PC code for the 0.40g Median UHS earthquake input (see Section l

3.1.2) was 12,870 k-ft. For evaluating the median capacity the following median safety factors and random variabilities were considered as shown in Table 3.1. These values lead to the following safety factors:

i 5

- _ _-____ _ L

(1) MODERATE STRAIN LEVEL:

Median Safety Factor = 2.10 (2) HIGH STRAIN LEVEL:

Median Safety Factor = 2.41 (3) ULTIMATE STRAIN LEVEL:

Median Safety Factor = 2.78 I

J Table 3.1 Median Safety Factors and Logarithmic Standard Deviations (Ref.13)

Parameter Safety Factor p,

pu Spectral Shape 1.00 0.15 0.20 Peak Acceleration 0.90 0.10 0.00 Frequency 1.00 0.00 0.05 Damping 1.00 0.00 0.15 Modeling 1.00 0.00 0.10 Fluid Pressure 1.00 0.00 0.02 C, Variation 1.00 0.00 0.03 T Variation 1.00 0.00 0.10 nc Inelastic Factor 1.00 0.00 0.05 The total logarithmic standard deviations are px = 0.18 and pu=0.30. These values lead to the following final tank seismic capacities:

(1) MODERATE STRAIN LEVEL:

Median Capacity = 0.84 g HCPLF Capacity = 0.38 g (2) HIGH STRAIN LEVEL:

Median Capacity = 0.96 g HCLPF Capacity = 0.43 g (3) ULTIMATE STRAIN LEVEL:

Median Capacity = 1.11 g HCLPF Capacity = 0.50 g The HIGH STRAIN LEVEL was used to determine the safety margin of the CST.

4.0 Concluding Remarks Based on a rigorous evaluation of bolt chair behavior and soil-structure interaction effects the seismic capacity of the CST # 11 was reevaluated. A refined nonlinear finite element ANSYS model was used to quantify the tank shell-bolt chair interaction. The tank seismic response including soil-structure interaction effects was determined using a three-dimesional model of the tank foundation via SUPER SASSI/PC code. The tank l

capacity analyses have shown that the critical failure mode of the CST is the shell buckling at the compresion side due the global overtuming. The nonlinear analysis has also shown that the stress in the tank shell at the bolt chair at the tension side is not a critical parameter for tank failure, due to the significant plastic stress redistribution (under limited strain). This result confirms that the current practice based on the elastic j

stress evaluation in the tank shel! at the bolt chair for determining the reduction in the anchor bolt capacity is overly conservative.

6

7 l

For the USl A-46 program the deterministic seismic capacity of the CST was estimated to be 0.66, which 4.4 times the SSE ZPGA of 0.i5 9

9 (for Housner spectrum input).

However, such a high seismic capacity may be only a theoretical margin if the tank

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failure is induced at a ZPGA lower than 0.66g by soil liquefaction (apparent there is no study to quatify the liquefaction for Housner spectrum input anchored to ZPGA higher than 0.15g).

For the IPEEE program the Median and HCLPF capacities were estimated to be 0.96g and 0.43g, respectively (for Median UHS input). The HCLPF capacity is 2.87 times the SSE ZPGA of 0.15g. At the HLCPF capacity level of 0.43g the soil liquefaction is un3kely as shown in a previous investigation of S&A (Ref. 9)

Based on the results presented herein, the outlier concems identified for both the USl A-46 and IPEEE programs for the CST and the other flat bottom tanks (by qualitative comparison) may be considered resolved.

5.0 References 1.

Reed, J.W., Kennedy, R.P. (1994). Methodology for Developing Seismic Fragilities, TR-103959, RP, EPRI 2.

Gilbert & Associates (1994,1995). Seismic Verification of Vertical Tanks at the Calvert Cliffs NPP, including Technical Reports and Memos between Mr. Steve Cowne and P. Baughm n.

3.

EQE Engineering (1995). Calvert Cliffs A-46 Tank and Hx Outlier Resolution, CST 11 and 21, Job 42111, Calc. No. C-013, May 4.

EQE Engineering (1994). Calvert Cliffs PRA for IPEEE, Fragility Derivation for Condensa;e Water Storage Tanks, Job 42111.19, Calc. No. 0-010, December.

5.

S&A (1995). HCLPF calculations for the Turkey Point RWCST and CST.

6.

S&A (1995). SUPER SASSI/PC - User, Example and Theoretical Manuals.

7.

EQE Engineering (1995). Calvert Cliffs Tank Analysis, SSI Analysis of CST, Job 42115.01, April.

8.

S&A (1995). Seismic SSI Evaluation for CST and RWST at the Calvert Cliffs NPP, Job 94C11842b.

9.

S&A(1995). IPEEE SPRA Study for Calvert Cliffs NPP, Seismic Probabilistic Structural Response and Soil-Structure Interaction Analyses, Task ll, Job 94C1842.

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10.

Methodology for Assessment of NPP Seismic Margin, Rev.1, Appendix 8 (1991), EPRI, Report NP-6041-SL 7

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APPENDIX D Evaluation of Tank Seismic Capacities Using TANKV codef l ' 'c

TANKV Code 9, Description l Stevenson & Associates, Inc. 9217 Midwest Ave. Cleveland, OH 44125 Ph: (216)587-3805 Fax- (216)587-2205 i-

{ ) TABLE OF CONTENTS i 1.0 I NTRO D U CTI ON..............................,.................. A-1 ) 2.0 OPERATION AND CAPABILITIES _OF THE PROGRAM................. A-1 3.0 TECHNICAL ASSUMPTIONS LIMITATIONS OF THE PROGRAM......... A-1 4.0 THEORETICAL BACKGROUND OF THE PROGRAM................... A-3 5.0 SEISMIC CAPACITY OF METAL FLAT-BCTTOM VERTICAL CYLINDRICAL LIQUID STORAGE TANKS....................................... A-4 5.1 Ge ne ral Rem a rks................................................. A-4 _ 5.2 Compressive Buckline Capacity of the Tank Shell.......................... A 5 5.3 Bolt Hold-Down Capacity........................................... A-6 5.4 Liquid Hold-Down Forces........................................... A-6 5.5 Base Overturning Moment Tank Capacity............................... A-7 6.0 DATABASES USED IN THE PROGRAM.............................. A-8

7.0 REFERENCES

......... Tr........................................ A-10 l 4 i 1 !w _ - - - _____________ ___ ________________ - __ _ ____ _

1.0 ' INTRODUCTION The purpose of this program, TANKV, is to calculate the seismic response of vertical cylindrical liquid storage tanks, and to estimate their seismic capacity in terms of seismic margins.-These calculations are intended to be conservative, and they give a High-Confidence-Iow-Probability of-Failure Seismic Margin Earthquake (HCLPF SME) capacity of the tank following the general Conservative Deterministic Failure Margin (CDFM) approach as required in the Seismic Margin - - Assessment (SMA) methodology /1/. It is also discussed in Section 9 how to use this program for other purposes outside of SMA 2.0 OPERATION AND CAPABILITIES OF THE PROGRAM The program is designed to perform aninteractive analysis of vertical cylindrical liquid storage tanks subjected to base seismic excitations. The user may: input the data interactively on the screen using program options and a simple help system, execute the analysis, check results on the screen, repeat the current tank seismic 9nalysis for interactively changed input data, print input and output data which are stored permanently and during the run on a disk for documentation or further analysis. 3.0 TECHNICAL ASSUMPTIONS LIMITATIONS OF THE PROGRAM 'Ihe main technical assumptions of the program are: The tank is assumed to be a vertical cylindrical cantilever with an unconstrained top, a circular cross section, with or without a roof, and also with or without an internal floating roof. The tank may be anchored to the foundation, or may be unanchored. Effects of the tank bottom uplift and liquid hold-down forces are considered. Interaction between the tank and its flexible foundation may be considned approximately.. Various tank materials (e.g. carbon steels, stainless steels, aluminum) may be used from the program database which is consistent with the ASME BPVC material specification, or material properties may be specified by the user. A-1

_ _ = _ _ _ _ . Also various anchorage materials and anchorage details mcy be considered using either the program database or user specified data. Various seismic response spectra (RG 1.60, Newmark, Housn tr or user defined) may be applied to the tank base. The tank may be filled by any liquid from the program database or as specified by the user. There are, however, severalimportant limitations when using this program which may be expressed as follows: 0.4 < h, / r 14.6 1 < h / h, s 1.25

01. h, / h 10.4 0.0005 5 t. / r 5 0.004-where: h, = Height of, Liquid in the Tank r = Radius of the Tank Shell h = Height of the Tank Shell h, = Height of the Tapl; Roof
t. = Average thickness of the Tank Shell resulting mostly from the applied theoretical solutions, and t, > 0, -

t, > 0 - 1 1 t, / t. $ 4 0 $ h, / h 5 0.15 I 0 < h, s 30 nu h-h,5 f $ h - h, + 0.4 h, where: t, = Thickness of the Tank Roof t = L Thickness of the Tank Bottom t, = Thickness of the Tank Shell near the Bottom

t. = Average 'Ihickness of the Tank Shell h, = Height of the Anchor Bolt Chairs i

A-2 l l-

h = Height of the Tank Shell h, = Effective Embedment Depth of the Anchor Bolts nm = Nominal Diameter of Anchor Bolts 11, = Height of the 1.iquid in the Tank f., = Freeboard Height h, = Height of the Tank Roof the character of which is rather constructional. Allitems used in (1) and (2) are shown in Figure 1. I l 4.0 THEORETICAL BACKGROUND OF THE PROGRAM A seismic analysis of vertical cylindrical liquid storage tanks usually consists of two phases: a seismic response analysis and a seismic capacity assessment. The program performs the seismic response analysis (first part) primarily based on works of Veletsos /20,21,22,23/ and Haroun and Housner /6,7/. The capacity assessment (second part) is based on the papers of Kennedy /11,12/. The following response phenomena are estimated in the first part of the program: the fundamental horizontal, vertical, and sloshing natural frequencies of the tank liquid system without and with (if requested) consideration of foundation-tank interaction, ' the response spectral accelerations for these frequencies, . the impulsive mode base shear, moment, and hydrodynamic pressure, which represent the effect of the part of the liquid that maybe considered to move in synchronism with the tank wall as a rigidly attached mass, the convective (sloshing) mode base shear, moment, hydrodynamic pressure, and theoretical sloshing height, which represent the action of the part of the liquid near the free surface that experiences sloshing or rocking motions, the hydrodynamic pressure by the vertical component of seismic motion, the combined seismic base shear, overturning moment, static liquid pressure and total liquid pressure both near the tank bottom, seismic liquid pressure at the tank bottom, and pressure at the tank bottem which loads the tank foundation. The topic of response evaluation of liquid storage tanks has been extensively described in the L literature during the past few years /4,6,20,21,22,23/. These responses are then used in the second part for evaluation of tank seismic capacities: The combined overturning base moment determined in the tank shellimmediately above the tank bottom is compared to the nominal overturning tank moment capacity which is governed by a combination of shell buckling, liquid hold-down capacity, and anchor yielding or failure, A-3 -I {

and generally governa the seismic capacity of the tank. The combined base shear is compared to the nominal sliding s, hear tank capacity, which only l / seldom controls the seismic capacity of the tank. The totalliquid pressure near the tank bottom is compared to the hoop membrane capacity l of the tank shell, which essentially never governs the seismic capacity of properly designed tanks. L I The fluid sloshing height is compared to the freeboard height above the top of the liquid to ( estimate whether roof damage is likely. Because capacities to withstand horizontal responses are slightly influenced by vertical responses, a small amount of non-linearity will develop when computing the Seismic Margin Earthquake (SME) capacity of the tank. Therefore,it is preferable first to estimate an SME capacity of the tank, SME,, I and to compute the seismic response, SEISMIC, for this SME, Then the actual SME capacity can be estimated from SME= CAPACITY-STATIC (SME,) (3)

SEISMIC, where CAPACITY is the HCLPF capacity, and STATIC is the portion of this capacity used up by

~ static loads. If the resultant SME differs substantially from SME,, then iteration of the procedure is necessary because of slight non-linearities. When SME and SME, are close (within an acceptable defined tolerance), no iteration is necessary. 5. SEISMIC CAPACITY OF METAL FIAT BOTTOM VERTICAL CYLINDRICAL LIQUID STORAGE TANKS 5.1 General Remarks The seismic capacity approach for such anchored or unanchored tanks is mostly based on works of Kennedy /11,12/. Flat bottom vertical cylindricalliquid storage tanks have sometimes failed with loss of contents during strong earthquakes. For tanks with radius-to-wall thickness ratios greater than about 600, or tanks with minimal or no anchorage, failures have often been associated with rupture of the tank wall near its connection to the bottom, due to excessive tank wall buckling or bolt stretching, and excessive bottom uplift. Both these failure modes are primarily due to dynamic overturning moment at the tank base due to impulsive mode liquid pressure on the tank wall. 1 Another common failure mode has been breakage of piping connected to a tank. Breakage of a pipe between the tank wall and shutoff valve is one of de most prevalent causes of loss of liquid from l a storage tank. ] ) Other failure modes which are of much lesser importance are: distortion of the tank bottom near i A-4

the tank side wall due to a soil failure under the tank bottom, tank sliding, excessive hoop tensile stresses due to hydrodynamic pressures on the tank wall, and damage of the roof and internal attachments due to liquid sloshing and insufficient freeboard height., Die effective yield stress of the tank materialis defined as follows: a, = F/P (for carbon steel, e.g. SA-36, and stainless steels) a, = 2.4 S. (for materials without a well defined yield point, e.g. aluminum) o, = greater from Fjmc and 2.4 S. (when user specified) where: Fjec and S, are the specified minimum yield point and basic allowable stress respectively according to ASME BPVC /3/. 5.2 Compressive Buckling Capacity of the Tank Shell The most likely way for tank shells to buckle is in " elephant foot" buckling near the base. The onset of such " elephant foot" buckles can be estimated using elastic-plastic collapse theory /16,18/. To estimate the HCIEF SME capacity of tanks, the onset of " elephant foot" buckling is judged to represent the limit to the compressive bu:kling capacity of the tank shell, C. However, because such buckling does not define actual tank failure, no significant conservatism needs to be introduced when estimating C.,. The " elephant-foot" buckling axial stress capacity, a, of the tank shell can be p acurrately estimated by (Ref.16,18): ?$ (T / t) [1-[ pr ]2 (t_ M4 5) [ d(a, / 36ksi)] 0.6E. 1 S (4) a = P a t, Sp1 1 p where: S, = (r/t,/400) t, = sidewall thickness near the shell base p = tank internal pressure near the base a, = effective yield stress of the tank shell material E, = Modulus of Elasticity of the tank shell material and C, = 0.9 a (5) p Although not expected to govern for overall seismic capacity, the program checks for classical clasic " diamond" buckling of the tank shell under combined axial bending and internal pressure using: (Ref. 25): A-5 t l l L_-____--_-_____________-____________________

E' cs = (0.6y + Ay) /t,- (6) o r / where: y = 1 - 0.73(1 - e+) 1 _ r 4 _ 16$t, If ace < C, then C, is set equal to nes. 5.3 Bolt Hold-Down Capacity The bolt hold-down capacity, TBC, is governed by the weakest of the following: Bolt tensile capacity Anchorage of the bolt into the concrete foundation Capacity of the top plate of the bolt chairs to transfer bolt loads to the vertical chair ~ gussets Attachment of the toltrplate and vertical chair gussets to the tank shell Capability of the tank shell to withstand ' concentrated loads imposed on it by bolt chairs. At this time, the user must check these items outside the program. For situations _where the bolt tensile capacity does not govern, the program allows a reduction factor to be applied to the calculated bolt tensile capacity. 5.4' Uquid Hold-Down Forces For tanks with minimum anchorage and for unanchored tanks, hold-down forces resulting from fluid pressure acting on the tank bottom will contribute significantly to the overturning moment capacity of the tank. De situation in the region of axial tension in the tank shell is illustrated in Figure 18 for small uplift displacements. l For anchored tanks, it is recommended that the small displacement theory be used to compute the liquid hold-down forces. He corresponding relationships used in the program have been obtained by Kennedy /11,12/. The small displacement theory is strictly applicable under the following conditions: A-6 $L_-________-_-__________________________

(L / r ) s 0.15 ' (6, / t,) s 0.6 (M / M ) s 0.9, (M, / M,) s 0.9, and (M.,/ M ) s 0.9 p g l where Mpb and MP. are the plastic moment capacity of the base plate and tank shell respectively. 'Ihe small displacement theory is excessively conservative for unanchored tanks. In accordance with Kennedy /11,12/, a more appropriate upper bound theory is used for such tanks. This upper bound theory has been implemented in the program. 5.5 Base Overturning Moment Tank Capacity These limit stresses are used in the seismic capacity assessment: effective yield stress of the tank material a,, = k F,, where F,, = the minimum yield stress specified in ASME BPVC, The overturning moment capacity is,then determined as(Ref.1, App. H): j 1 l 1 M, = C,'C r2 + f(T ;rcosa) + T r (2siq$) + AT,C r 2 2 (7) 2 3 i o 4 i=1 where: C.' = maximum shell compression D W.rs + { T8i (8) C,'=( + T,f)C + AT,C 3 3 2r COSa - COS$) s T (9) g T;=T3p + K ( sc 3 3 1 - cosS i A-7 L_____.____.__m___

.i 1+cosS C' = sin # + (w-g)cosS ,. Sin #CosS + w-S 1+cosS sin # - Scos# 1+cosS sins + (w-S)cosS 1 - cosS) C, = # - sWco$ 1-cos# The program iterates on S until convertcace of C, and C ' is achieved. Included in this iteration process is the adjustment of the fluid hold-down force in accordance with convergence to the neutral axis. 6.0 DATABASES USED IN THE PROGRAM V. 'Ihe following databases are used in the program: a) TankMaterials (based on the ASME BPVC specification): carbon steel: stainless steel: aluminum: i SA-36 SA-240(304L) SB-209 5052/5652-0 SA-285(A) SA-240(304LN) SB-209 5052/5652-H32 SA-285(B) SA-240(304H) SB-209 5052/5652-H34 SA-285(C) - SA-240(304N) SB-209 5083-0 SA-414(C) SA-240(316L) SB-209 5086-0 SA-442(55) SA-240(316H)~ SB-209 5086-H32 SA-515(55) SA-240(316LN) SB-209 5086-H34 SA-516(55) _ SA-240(316N) SB-209 5154/5254-0 SA-442(60). SB-209 5154/5254-H32 SA-515(60) SB-209 5154/5254-H34 SA-283(A) SB-209 5454-0 SA-283(B) SB-209 5454-H32 SA-283(C) SB-209 5454-H34. user specified : A-8 _ - _ _ _ _ - _ _ = _ _ _ - _ - _ _ _ _ _ - _ _ _ _ _

b) Tank Liquids: water, demiwater seawater dieselfuel gasoline - kerosine mineral oil vegetable oil alcohol user specified c) Anchor Bolt Steels (based on the A3ME BPVC specification): SA-36 SA-307 SA-325 l user specified d) Response Spectra RG 1.60 Tr, Housner Newmark - median, competent soil Newmark - median, rock Newmark - median + 1 sigma, competent soil Newmark - median + 1 sigma, rock l l-user specified i 1 l l i A-9 l f 1 L_---__-

~

7.0 REFERENCES

/1/ A_ Methodology for Assessment of Nuclear Power Plant Seisn$ic Margin. Report NP-6041, Rev.1. EPRI, Palo Alto,1991. /2/ ASCE Standard 4-86 and Commentasy, " Seismic Analysis of Safety Related Nuclear Stnictures," American Society of Civil Engineers, New York, September 1986. /3/ ASME Boiler and Pressure Vessel Code, Section 111, " Nuclear Power Plant Components," Division 1 and Appendices. American Society of Mechanical Engineers, New York,1989. /4/ Fischer, F.D., Rammerstorfer, F.G, Scharf, K, " Earthquake Resistant Design of Anchored and Unanchored Tanks Under Three-Dimensional Earthquake Excitation," Structural Dynamics. Recent Advances. Edited by G.I. Schueller, Springer-Verlag, Berlin,1991. /5/ " Generic Implementation Procedure (GIP) for Seismic Verification of Nuclear Power Plant Equipment," Section 7, Tanks and Heat Exchangers Review, SQUG, Rev. 2,1991. /6/ Haroun, M.A-and Housner, G.W., " Seismic Design of Liquid Storage Tanks," Journal of the Technical Councils of ASCE. Vol.107, No. TC1,1981, pp.191-207. f1/ Haroun, M.A-and Housner, G.W., " Complications in Free Vibration Analysis of Tanks," ~ Journal of the Ennineerinn-Mechanics Division. Vol.108, No. EMS. ASCE,1982, pp. 801-818. ?L /8/ Haroun, M A. and Badawi, H.S., " Nonlinear Axisymmetric Uplift of Circular Plates," Dynamic g.[ Structures, ASCE, August 1987, pp. 77-89. ] /9/ Hashimoto, P.S. and Tiong, LW., " Earthquake Experience Data on Anchored, Ground Mounted Vertical Storage Tanks," Report NP-6276, EPRI, Palo Alto,1989. /10/. Kausel, E. et al., " Dynamic Stiffness of Circular Foundations," Journal of the Engineering Mechanics Divis. ion. ASCE, Vol.101, No. EM6, Dec.1975, pp. 771-785. /11/ Kennedy, R.P. et al., " Assessment of Seismic Margin Calculation Methods," Report NUREG/CR-5270, UCID-21572, Lawrence Livermore National Laboratory,1988. /12/ Kennedy, R.P. and Kassawara, R.P., " Seismic Evaluation of Large Flat-Bottomed Tanks," Second Symoosium on Current Issues Related to Nuclear Power Plant Structures. Eauioment and Pininn With Emohacia on Resolution of Seismic Issues in Iow-Seismicity R i j Report NP-6437D. EPRI, Palo Alto,1989. i /13/. Luco, J.E. and Westman, R., " Dynamic Response of Circular Footings," Journal of the l. Engineering Mechanics Division, ASCE, Vol. 97, No. EMS, Oct.1971, pp.1381-1395. /14/ Manos, G.C., " Earthquake Tank-Wall Stability of Unanchored Tanks," Journal of Structural A-10 l L-

Ennineerine. ASCE, Vol.112, No. 8, Aug.10986, pp.1863-1880. NovaN, M. and L El Hifnawy, "Effect of Foundation Flexib'ility on Dynamic Behavior of /15/ Buildings," Proc. 8th WCEE. July 21-28,1984, San Francisco. Publ. by Prentice-Hall, New Jersey,1984. /16/- Priestley, M.J.N. et al., " Seismic Design of Storage Tanks," Recommendations of a Stu@d _Qr2UD of the New Zealand National Society for Earthauake Engineering, December,1986. /17/. Richard, F.E. et al., " Vibrations of Soil and Foundations," Prentice-Hall, New Jersey,1970. /18/ Rotter, J.M., " local Inelastic Collapse of Pressurized Thin Cylindrical Steel Shells Under Axial Compression," Research Report No. R502, School of Civil and Mining Engineering, University of Sydney,1985. /19/ " Seismic Design and Evaluation Guidelines for the Department of Energy High Level Waste Storage Tanks," [ Draft) Prepared by Tanks Seismic Experts Panel, January 1992. 1 /20/ Veletsos, A-S. and Tang, Yu., " Dynamics of Vertically Excited Liquid Storage Tanks," Journal of Structural Ennineerine. Vol.112, No. 6, ASCE, June 1976, pp.1288-1246. /21/ Veletsos, A-S. and Yang, J.Y.," Dynamics of Fixed Base Liquid Storage Tanks," Proceedings ~ of the U.S. - Japan Seminar for Earthauake Engineering Research with Emohasis on Ufeline Systems. Tokyo, November,1976. 9 /22/ Veletsos, A S., " Seismic Response and Design of Uquid Storage Tanks," Guidelines for the Seismic Design of Oil and Gas Pioeline Systems. ASCE, New York,1984. /23/ Veletsos, A-S., Tang, Yu., and Tang, H.T., " Dynamic Response of Flexibly Supported Uquid Storage Tanks," Journal of Structural Engineering. Vol.118, No.1, January 199'2, pp. 264-l 283. l /24/ ' Wickman, ICR. et al., " local Stresses in Spherical and Cylindrical Shells due to External l: Imadings," Welding Research Bulletin 107, August,1965. /25/ Buckline of Thin-Walled Circular Cylinders. NASA SP-8007, National Aeronautics and Space Administration, August 1968. A-11 l L--- - - - - - - - - - - - - ~

) APPENDlX D1 Deterministic Capacity (USl A-46) l l "?r i f l l

i p /' s l APPENDIX D2 Median Capacity for 44 ksi tensile stress in bolt - yielding Median Capacity for 56 ksi tensile stress in bolt - ultimate i ' f, i' l-i l

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