ML20053A441
| ML20053A441 | |
| Person / Time | |
|---|---|
| Site: | Oyster Creek |
| Issue date: | 05/20/1982 |
| From: | Crutchfield D Office of Nuclear Reactor Regulation |
| To: | Fiedler P JERSEY CENTRAL POWER & LIGHT CO. |
| References | |
| TASK-03-07.B, TASK-3-7.B, TASK-RR LSO5-82-05-044, LSO5-82-5-44, NUDOCS 8205260101 | |
| Download: ML20053A441 (39) | |
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May 20,1982 i
Docket No. 50-219 i
LS05 05-044 Mr. P. B. Fiedler Vice President and Director - Oyster Creek j
Oyster Creek Nuclear Generating Station Post Office Box 388 Forked River. New Jersey 08731
Dear Mr. Fiedler:
SUBJECT:
SEP TOPIC III-7.B. DESIGN CODES. DESIGN CRITERIA.
AND LOADING C0fBINAT10NS - OYSTER CREEK l
Enclosed is a revised analysis of the drywell subjected to combined l
seismic and LOCA loads. This analysis supersedes the one sent to you in the April 30, 1982 SER as enclosure 2.
The revised analysis reflects 4 change to the structural model to more accurately reflect the as-built drywell and it incorporates a higher metal temperature due to a main steam line break.
It is concluded that the drywell will adequately resist the combined seismic LOCA loads.
This evaluation will be a basic input to the integrated safety assess-ment of your facility unless you identify changes needed to accurately reflect the as-b6112 conditions at your facility or if NRC criteria are modified before the integrated assessment is completed. We encour-age you to supply any other material that might affect the staff's evaluation or be significant to the integrated assessitent of your t
facility.
Sincerely.
Dennis M. Crutchfield Chief
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Operating Reactors Branch No. 5 O
Division of Licensing to C3
Enclosure:
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Docket flo. 50-219 LS0582 Mr. P. B. Fiedler Vice President and Director - Oys+ar Creek Oyster Creek Nuclear Generating atation Post Office Box 388 Forkdd River, New Jersey 08731
Dear Mr. Fiedler:
SUBJECT:
SEP TOPIC III-7.B. DESIGN CODES, DESIGN CRITERIA, AND LOADING COMBINATIONS - OYSTER CREEK Enclosed is a revised analysis of the drywell subjected to combined seismic and LOCA loads. This analysis supersedes the one sent to you on April 30, 1982.
The revised analysis reflects a change to the structural model to more accurately reflect the as-built drywell and it incorporates a higher metal temperature due to a main steam line break.
It is concluded that the drywell will adequately resist the combined seismic LOCA loads.
This evaluation will be a basic input to the integrated safety assess-ment of your facility unless you identify changes needed to accurately reflect the as-built conditions at your facility or if NRC criteria are modified before the integrated assessment is completed. We encourage to supply any other material that might affect the staff's evaluation or be significant in the integrated assessment of your facility.
Sincerely, Dennis M. Crutchfield, Chief Operating Reactors Branch No. 5 Division of Licensing
Enclosure:
As stated SEPB:Dt W SEPB:DL cc w/ enclosure:
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Mr. P. B. Fiedler cc G. F. Trowbridge, Esquire Resident Inspector Shaw, Pittman, Potts and Trowbridge c/o U. S. NRC 1800 M Street', N. W.
Post Office Box 445 Washington, D. C.
20036 Forked River, New Jersey 08731 J. B. Lieberman, Esquire Commissioner Berlack, Israels & Lieberman New Jersey Department of~ Energy 26 Broadway 101 Commerce Street New York, New York 10004 Newark, New Jersey 07102-
, i aid C. Haynes, Regiooal Administrator Nuescar Regulatory Commission, Region I o
631 Park Avenue King of Prussia, Pennsylvania 19406 J.. Knubel BWR Licensing Manager GPU Nuclear 100 Interplace Parkway Parsippany, New Jersey 07054 Deputy Attorney General
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State of New Jersey Department of Law and Public Safety 36 West State Street - CN 112 Trenton, New Jersey 08625
' Mayor Lacey Township 818 Lacey Road Fo.rked 3iver, New Jersey 08731 U. S. Environmental Protection Agency Region II Office ATTN:
Regional Radiation Representative 26 Federal Plaza
. New York, New York 10007 Licensing Supervisor Oyster Creek Nuclear Generating Station Post Office Box 383 Forked River, New Jersey 08731 I
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Systematic Evaluation Program STRUCTURAL REVIEW 0F THE
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OYSTER CREEK NUCLEAR POWER PLANT DRYWELL. CONTAINMENT STRUCTURE UNDER COMBINED LOADS i
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March 1982 A. G. Debeling C. Y. Liaw EG&G/ SAN RAMON OPERATIONS
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.nW DR&n~1 FOREWORD The U.S.
Nuclear Re>Julatory Conmission (NRC) is conducting a Systematic Evaluation Program (SEP), which is a' plant-by-plant reassessment of the safety of 11 operating nuclear reactors that received construction permits between 1956 and 1967.
Because many safety criteria have changed since these plants were licensed, the purpose of the SEP is to develop a current, documented basis for the safety of these older facilities.
Seismic analyses for a Safe Shutdown Earthquake (SSE) for the
~
Oyster Creek Nuclear Power Plant had been performed in a previous study of selected plant structures, systems, and components from generic groups of equi pment.
The results of these analyses were reported in an earlier SEP report, NUREG/CR-1981.
In the study reported on here, the containment structure was selected for further evaluation of the combined effect of the SSE and the design. basis accident (DBA), including evaluation of both the lnts of coolant accident (LOCA) and the main steam line break (MSLB).
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ACKNOWLEDGEMENTS The authors wisn to thank P.
Y.
Chen and S.
Brown, technical -
monitors of this work at the Office of Nuclear Reactor Regulation (NRR),
for their continuing support, and T. A. Nelson and T. Y. -Lo of Lawrence Livermore National Laboratory (LLNL), who provided project management support and reviewed the report.
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ABSTRACT A reassessment of the drywell containment structure of the Oyster Creek Nuclear Power Plant was performed for the Nuclear Regulatory Comis-sion as part of the Systematic Evaluation Program.
Cortclusions about the ability of the drywell containment structure to withstand the combined effect of a Safe Shutdown Earthquake (SSE) and a Design Basis Accident (DBA) which include either a double ended recirculation line break (LOCA) or main steam line break (MSLB), are presented.
Both the SSE and DBA loads employed in this reassessment came from previous studies conducted by the U.S. Nuclear Regulatory Commission (NRC).
This reassessment focused mainly on the overall structural in-tegrity of the drywell containment to withstand an SSE plus DBA.
The stress intensity limits of Subsection NE of the 1977 ASME Boiler and Pres-sure Vessel Code,Section III, Division 1 for the Design and Service Con-ditions were used as the compliance criteria for structural reassessment.
Design Co.nditi on is described by the combination deadweight, SSE, and pressure loads from a DBA.
Service Condition is described by the Design Condition combination plus the thermal loads from the DBA.
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a CONTENTS 3
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i Page j
i CHAPTER 1:
INTRODUCTION.
1-1 l
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1-1 1.1 Scope of Work.
1.2-Structure Description.
1-2 1.3 Loads and Load Combinations.
1-4 1.3.1 Deadweight Loads.
1-4 1.3.2 Seismic Loads.
1-4 1.3.3 Design Basis Accident (DBA) Loads 1-7 1.4 Material Properties.
1-12 Previous Analyses of Drywell Structure 1-13 1.5 9
CHAPTER 2: ANAlkSIS OF DRYWELL.
2-1 2.1 Assumptions 2-1 2.2 Mathematical Model.
2-1 2.3 Method of Analysis.
2-3 2.4 Results of Analysis.
2-4 CHAPTER 3:
SUMMARY
AND CONCLUSION.
3-1 REFERENCES 4
4 6
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LIST OF ILLUSTRATIONS dV on s2 Figure Page 1
Drywell structure 1-3 2
Drywell bending moment due to SSE.
1-5,
3 Drywell shear load due to SSE 1-6 4
Drywell pressure response to a double-ended recirculation line break 1-8 2
5 Drywell pressure response to 0.75 ft MSLB 1-3 2
6 Drywell atmosphere temperature response to a 0.75 ft MSLB.
1-10 7
.Drywell atmosphere temperature response to a double-ended recirculation line break 1-11 8
Mathematical model of the drywell structure.
2-2 9
Drywell gross structural discontinuity and local primary membrane regions 2-5 10 Dead load results (psi).
2-6 11 Pressure results (psi)..
2-7 12 Thermal results (psi) 2-8 13 Seismic results (psi) 2-9 LIST OF TABLES Table Page I
1 Drywell metal temperatures...
1-12 2
P values summary, Design Condition.
2-12 m
3 P values at Region A, Design Condition 2-12 t
4 P values at Region A, Design Condition 2-13 b
5 Total stress at base location due to MSLB, Service Condition 2-13 6
Total stress at base location due to double ended recirculation line break, Service Condition.
2-14
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a CHAPTER 1 ~
INTRODUCTION 1.1 SCOPE OF WORK Structural reassessment of nuclear power plants -is one facet of the Systematic Evaluation Program (SEP) being conducted by the Nuclear Regulatory Comission (NRC).
This report is a structural review of the Oyster Creek drywell containment structure.
The overall structural in-tegrity of the drywel-1 was evaluated based on the stress intensity limits
~ 'for the design and service conditions defined in the ASME Boiler and Pres-sure Vessel' ~ Code,Section III, Subsection NE,1977.
In this reassessment, the design conditions included a Safe Shutdown Eart.hquake (SSE), as repre-sented by the Reg. Guide 1.60 response spectra, which is scaled to 0.179 peak ground acceleration, and the pressure due to a postulated Design Basis Accident (DBA).
The DBA consists of either a loss of coolant a.ccident (LOCA) or a main steam line break (MSLB).
Service conditions were defined as the design condition loads plus thermal loads caused by the DBA.
In this analysis, the seismic event and the worst of the main steam line breaks (MSLB) and double ended recirculation line breaks were applied as if they occurred simultaneously.
The thermal and pressure loading conditions due to the DBA are based on the analysis discussed in Reference 1.
The seismic stress evaluation is an extension of the analysis presented in " Seismic Review of the Oyster Creek Nuclear Power Plant as part of the SEP Program" (Reference 2).
i l
Because the primary objective of this analysis is to determine the drywell's ability to perform containment functions when subjected to the combinad action of the DBA, deadweigh', and seismic loads, no local t
loads or flooding conditions were considered.
Reinforcement around pene-i trations in the drywell were assumed to be as strong as the drywell shell; 1-1
therefore, the local discontinuity effect of the penetrations on the over-all shell stresses was neglected.
Hence, an axisymmetric shell model was used in this analysis.
1.2 STRUCTURE DESCRIPTION The drywell containment structure of the Oyster Creek plant houses the reactor vessel, reactor coolant recirculating loops, and other The structure is a combina-components associated with the reactor system.
tion spherical, cylindrical, and 2:1 ellipsoidal dome that resembles an inverted light bulb. The spherical section has an inside diameter of 70 ft which. reduces down to 33 ft at the cylindrical portion connecting the sphere to the dome.
The structure is approximately 99 ft high.
The plate thickness varies from a maximum of 2.56 in. at the transition of the sphere and cylindrical section down to a minimum of 0.640 in. at the cylindrical section.
The dome wall thickness is 1.18 in.
Figure 1 illustrates the drywell structure along with pertinent dimensions.
The top closure, which in diameter, is made with a double tongue and groove seal which is 33 ft permits periodic checks for tightness.
Ten vent pipes, 6 ft in diameter, are equally spaced around the circumference to connect the drywell and vent headers in the pressure absorption chamber (Torus).
A 3-inch gap between the drywell and concrete biological shield is filled with foam material that provides insulation but no significant structural support.
An upper lateral seismic restraint attached to the cylindrical portion of the drywell at an elevation of 82.17 ft, allows for thermal, deadweight, and pressure deflection, but not for lateral movement due to seismic excitation.
All penetration for piping, instrumentation lines, vent ducts, electrical lines, equipment accesses, and personnel A sand entrance have expansion joints and double seals where applicable.
entrenchment surrounds the drywell from an elevation of 8.93 ft (foundation-steel interface) to 12.25 ft.
The entrenchment acts as a transition buffer for loads generated at the shell-foundation region.
1-2
[_2:1 ELLIPTICAL 00ME 1.188 in.
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El. 94'-9
.640 in.
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.722 in.
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.770 in.
gr El. 37'-3 1.154 in.
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Drywell structure.
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The peak drywell stresses due to these loads were combined algebraically, but with the worst combination of signs possible (which is a conservative approach).
1.3.1 Deadweight Loads Deadweight loads were generated by multiplying the steel weight 3
density (0.238 lbw/in ) by the drywell vol ume.
The dimensions used to determine the drywell volume were based on the structural drawings con-tained in the Oyster Creek FSAR (Reference 6).
1.3.2 Seismic Loads Figures 2 and 3 illustrate the recommended moment and shear loads that came from the SSE analysis of the - drywell, which is contained in Reference 2.
A Reg. Guide 1.60 spectra with peak ground acceleration of 0.22 g's was used to generate these values.
In the time span between this analysis and the SSE analysis presented in Reference 2, a site-specific spectra with a peak ground acceleration of 0.17 g's was approved by the
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NRC.
The information from the NRC contained in Reference 8 allows two options if reanalysis is needed: (1) to scale down the previous analysis (which 4 sed a Reg. Guide 1.60 spe'ctra) by the ratio of the new site-l specific peak ground acceleration to that of the peak ground acceleration l
used in the original analysis, and (2) to use the new site-specific spectra l
for reanalysis.
For this analysis the scaling method using the ratio 0.17/0.22 was used to determine the.eismic loads that were applied to the d rywell.
It should be noted that a Reg. Guide 1.60 spectra with peak ground acceleration of 0.17 g's envelops the new Oyster Creek site-specific spectra; therefore the seismic loads used in this analysis are conserva-tive.
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@ La a Ft r.m scir 1.3.3 Design Basis Accident (DBA) Loads A spectra of possible recirculation and main steam line breaks were analyzed to determine the peak post-accident pressure and temperature response of the drywell.
The maximum calculated peak pressure response was due to a double ended recirculation line break and the temperature response was caused by a 0.75 ft main steam line break.
Reference 1 contains detailed information on the mass energy release analysis for these post ulated breaks and formed the basis for determining the temperature and pressure loads used in this analysis.
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Pressure Lads Figures 4 and 5 illustrate the drywell pressure history response 2
predicted for a double ended recirculation line break and a 0.75 ft main steam line break, respectively.
A maximum pressure of 37 psig at 4.5
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seconds, caused by a double ended recirculation line break, was applied to the drywell to determine structural adequacy.
Thermal Loads Figures 6 and 7 illustrate the drywell atmospheric temperature response generated by the 0.75 ft main steam line break and the double ended recirculation 1ine break, respectively.
The metal temperatures will actually be less and will lag in time behind the atmospheric temperatures.
The peak metal temperature caused by the 0.75 ft main steam line break was actually applied to the finite element model to calculate drywell thermal stresses.
Table 1 shows the meridional and radial metal temperature 2
values caused by the 0.75 ft main steam line break.
The values shown in the table were extracted from results generated by the CONTEMPT-LT/028 computer code used in the post-accident mass and energy release analysis (Reference 1).
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Table 1.
Drywell metal temperatures.
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INNER SURFACE OUTER SURFACE
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ELEVATION PLATE THICKNESS TEMPERATURE TEMPERATURE (ft)
(in.)
(*F)
( F) 21.35 1.154 271 270 49.15 0.770 275 275 64.97 0.772 275 275 69.25 2.563 265 265 97.31 0.640 276 276 105.59 1.188 280 280 1.4 MATERIAL PROPERTIES Reference 4 identifies the A 212 GR "B" plate and its equivalent ASME/ ASTM designation as the material used jn fabrication of the drywells.
Young's Fbdulus. and the coefficient of thermal expansion were extracted from Reference 3, Tables I-6.0 and I-5.0.
Material property values used in this analysis are as follows:
MATERIAL PROPERTY VALUE A 212 GR "B" is equivalent to SA 515 GR "65" Plate 6
i (Young's Modulus) 30 x 10 psi (Poisson ratio) 0.3 Minimum Yield Stress 35,000 psi
-6 Coefficient thermal expansion 6.07 x 10 in/in/*F
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1-12
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1.5 PREVIOUS ANALYSES OF DRYWELL STRUCTURE
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Many analyses have been performed for various load conditions.
It is not our purpose to review all. earlier work; therefore, only those analyses which deal with load combinations similar to those considered in this report will be reviewed.
Preliminary Description and Safety Analysis Report, Amendment 1, discusses the design-accident condition including deadweight, thermal, internal pressure, and earthquake loading.
The drywell was designed for two loading scenarios: (1) internal pressure of 62 psig and 175*F metal
_ temperature, and (2) S5 psig at metal temperature of 281 F.
The maximum design pressure and temperature exceed the Design / Service conditions used in this report. A Housner spectrum with 0.229 peak. ground acceleration was used i.n the analysis.
It is not clear from the report how these loads were combined, but it is -suspected that the peak stresses due to these loadings were not combined algebraically.
It does not appear that a computer analysis was performed for any phase of the drywell stress analysis.
9 a
e i
1-13
CHAPTER 2
~
ANALYSIS OF DRYWELL 2.1 ASSUMPTIONS The drywell was analyzed using a finite element model for thermal, pressure, and deadweight loading conditions.
The following as-sumptions were made in constructing the finite element model.
~
1.
The structure and the loading were assumed to be axi-symmetric; the penetrations and their effects were not consid'ered.
2.~
The drywell was modeled assuming that the base of the shell was compl etely fixed to the foundation for translational and rotational movement.
3.
The sand entrenchment was. assumed to be linear elastic.
4.
Thermal gradients were assumed to be linear.
5.
The foam material between the drywell and the building was assumed to have no structural stiffness and was not included in the model.
6.
A stress free temperature of 70*F was assumed.
7.
The tongue and groove seal section at the top closure is considered a integral part of the structure; there-fore, it was not modeled.
- 8. ' Pressure loading was calculated to be static, i.e.,
no dynamic effects of sudden pressurization were deemed necessary.
2.2 MATHEMATICAL MODEL The finite element mathematical model of the drywell illustrated in Figure 8 was constructed based on the assumptions described in Section 2.1.
A general purpose structural analysis program (ANSYS) which is out-lined in Reference 5, was used to perform the stress computations.
Axi-symmetrical conical shell elements were used to model the structure.
Sand entrenchment at the base of the drywell was modeled using spring elements.
2-1
I i. yyg
~
1266
~
118 '.
1136 -
'g 1005 -
1 t
875 o 98 57 744 614
~.
T 483
.! = N0DE number 1
8 "33 m...;
N 353 f
sy.
n c
N u
223 1
374226 2
378849 s
K 3
353904 9
/-
4 358185 5
332029
-92 K
6 274706 1
7 923052 8-663317 9
297923
-38
-42 88 219 349 480 Figure 8. itathematical model of drywell structure.
i 2-2
.n
The spring constraints that were used were based on data extract-ed from Reference 7, which used similar drywell geometry in all respects, except for a larger sphere radius.
Constraints on the model included fixing the shell base at the 8.93 ft elevation for translation and rotational movement and fixing the top of the dome at the center line in
~
the horizontal direction for symmetry.
Verification of the model was performed by comparing deflection and stress results predicted by the finite element model against hand calculated values.
In all cases the results compared were acceptable.
2.3 METHOD OF ANALYSIS
~
Th'e a'bove described finite element analysis presents directly meridional and circu[nferential membranc arid bending stresses for the dead-weight, pressure, and thermal loads.
The seismic loads were extracted from the results of the stick model analysis presented in Reference 2.
The moment and shear due to SSE are depicted in Figures 2 and 3, respectively.
These data were used to calculate the meridional membrane and shear stresses in this analysis.
It was necessary to determine local bending stresses at the base of the structure due to seismic loading.
This was accomplished by applying a pseudo-seismic horizontal. load to the finite element model to produce the same fixed base global bending moment as that of the stick model.
l 1
Because a new site-specific spectrum with a peak ground accelera-tion of 0.17 's was approved, all seismic icadings were decreased by the 9
ratio of 0.17/0.22 (Reference 8).
Meridional membrane stresses were cal-culated by taking the _ bending moment values from' Figure 2 and applying them to the appropriate drywell cross-section using the beam theory equation (M/nR t).
No hoop stresses were calculated for seismic loading.
The shear stress was assumed to have a distribution of sin 0 along the circumferential direction, where 0 is measured from the positive direction
+-
2-3 i'
-s--
-,-+-J
,w,
$f) i.
of seismic motion.
Therefore, the maximum shear stresses in the "drywell were calculated by taking the total shear loads presented in Figure 3 and dividing by nRt.
All membrane and bending stresses were combined by absolute sum for the appropriate code compliance criteria.
The acceptance criterion used to establish structural adequacy for the drywell wasSection III, of the ASME Boiler and Pressure Vessel Code - Subsectiorr NE, July, 1977.
Figure 9 illustrates locations on the drywell where the structure was considered to be in compliance with ASME definitions of Gross Structural Discontinuity and Local Primary Membrane stress regions.
The remaining areas fall under General Membrane stress sections.
2.4 RESULTS OF ' ANALYSIS Figures 10 through 12 illustrate the membrane and bending stresses generated by the finite element analysis for deadweight, maximum pressure, and maximum thermal loading, respectively.
Figure 13 shows the stress results from the seismic analysis.
Verification of the finite element model was performed by com-paring membrane stresses at the mid-section of the sphere and at the top of the dome against the closed form solution of Pr/2t for pressure loading.
Radial deflection 'due to thermal loading at the mid-section of the. sphere was compared t'o the closed form solution of Raat.
In both cases the results compared were acceptable.
A study was performed to determine the effects of varying the spring stiffness used to model the sand entrenchment around the dryw i' the base location.
The thermal loading condition produced the hight u stresses in this region; therefore, it was chosen as the loading condition to be studied.
Three cases were analyzed using spring stiffnesses class-ified as strong, medium, and zero.
The meridional bending stress was most 2-4
U i
h h'/fL x R'N v5 do$.
t
~
t f.
a-o i
/
i A
i
{
Gross structural f
discontinuity
' Regions of local primary L.
j membrane stresses s
B.
j Figure 9. Drywell gross structural discontinuity and local primary membrane regions.
P
=
.~
2-5 e
.__________-_R___________..
. J*
Ds d$3l(Ni m f 9 ea,'
~
l 5
(
\\,
N
-322
\\
-355 N
~
,hI0
,,.- 53 422 Meridian Membrane Stress Hoop Membrane Stress a
f D
D e
t 354
'1I6 e
4
/
907
-60 272
-228'
?
Meridian Bending Stress Hoop Bending Stress Figure 10. Dead load results (psi).
2-6 i.
. 6171 6171 E
-3122 I
5671.
1M41 d
\\
13219 N-g.
~
.i0014 10028
~~~-
~
~.
7152 2146 Meridio~nal Membrane Stress Hoop Membrane Stress a
f 4
6 8 2042
-522d 156 8 il 4
13250 3975 sg h_7[61
-22.63 4
/
i
-2163 4427 1328 Meridional Bending Stress Hoop Bending Stress Figure 11.
Pressure results (psi).
2-7 l-g
E r :,
t
-1667
\\
928
\\
.-819 g
1
-35366
-857 3816 Mer'idional llembrane Stress Hoop Membrane Stress o
f f
I 4
E,2269
~.
~
?
)
I
/
-11417 _
36570
-2991 11062 Meridional Bending Stress Hoop Bending Stress s'
' Figure 12. Thermal results (psi).
2-8
4 s
446 16 19 O
1172 1106 605 l
~
~
l l
\\
1
)
791 5
665 l
'? -
/
e 2450 1275 757 Mdridi6n.al Membrane Stress Local Bending Stress Maximum Shear Stress Fig 0rer13. Seismic.r,esults (psi).
g prh e..
A e
i
I, -
affected by the ' stiffness changes; the other membrane and bending stress results were relatively insensitive to stiffness change.
The meridional bending stresses at the shell base were 62 ksi for zero stiffness, 36 ksi for medium stiffness, and 6 ksi for a strong spring stiffness.
~
It should be noted that even though the zero stiffness case would' produce combined stress intensities that exceed the Service Condition requirements, there is relief in the Code under sub-paragraph NE 3228.3 that might allow the drywdll design to be judged acceptable.
In addition, the zero stiffness is actually a rather extreme condition.
Any small stiffness in the sand will improve the bending stress a great deal.
The stress results used to determine ASME code compliance are all based on a medium spring stiffness being used in the finite element model; this is virtually the same stiffness value as that presented in Reference 7.
Because the pressure loading rises saarply during a section of the transient (approximately 4.5 seconds), it was deemed prudent to verify that dynamic effects of this sudden pressurization could be neglected.
A frequency analysis for Fourier Harmonic number = 0 was performed using the finite element model described in Section 2.1.
The first mode frequency was determined to be approximately 22 Hz, which corresponds to a period of about 0.05 seconds.
As. a general rule, if the rise time of the loading (tr) divided by the natural period cf the structure (T) is greater than 4, the loading can be considered to be essentially static (Reference 9).
In this case, the ratio t /T is approximately 90, assuming t = 4.5 sec. This r
r implies that the pressure loading can be considered as static.
Meridional buckling compressive loads produced in the drywell at the lower portion of the sphere were caused by deadweight and seismic loading. These compressive loads are not axisymmetric because seismic loads produce compression on one side and. tension on the opposite side.
It is difficult, therefore, to determine critical stresses because most publish-ed buckling analyses pertain to axisymmetric loading on simple geometric structures.
2-10
The method used to determine the critical load for meridional buckling of the drywell consisted of employing two different axisymmetric buckling analyses: externally pressuri zed sphere analysis and axially loaded truncated shell analysis.
Both of these analyses were similar in geometry to that of the drywell.
In both cases, the critical meridian buckling stresses exceeded the drywell stress by a factor of five.
If-thermal or pressure loads, were superimposed on these compressive loads they would put the shell in tension and therefore eliminate the need to check for meridional buckling loads.
This is not the case for the pre-dicted circumferencial critical buckling stress at the base, which is stresses.
The l
estimated at -15.00 ksi versus - 35.4 ksi for the thermal I
l
. critical buckling stress was calculated assuming that the sphere had a unifonn external pressure applied to its ' outer surface.
This loading condition would produce a uniform membrane stress in the sphere, which is not what would take place in the drywell sphere because the compressive circumferential stress is confined to a narrow ring at the base location.
For this reason the calculated critical circumferential buckling stress is judged to be extremely conservative; however, for lack of a better solution it was used for comparison.
It is possible that some amount of circum-ferencial buckling could occur at the base location due to thermal loading.
It should be noted that no credit was taken for the structural effect of the concrete that extends upward from the foundation around the inside of the drywell.
This concrete will surely help structural stability in this lower region, and greatly reduce the chance of circumferential buckling.
After combining maximum stresses prod'uced at various locations of the drywell, it was determined that the critical location for Design Con-l inwards to the cylindrical section ditions was where the sphere curves (Region A on Figure 9).
The critical location for Service Conditions is at the base of the drywell.
Table 2 includes maximum general membrane stresses; Tables 3 and 4 present the stress values used for ASME Code Tables 5 Design condition compliance for local primary membrane regions.
and 6 are for Service Level compliance for the MSLB and double-ended re-circulation line break respectively.
2-11
)
a eMk Table 2.
P Values Sumary, Design Condition R D
Loading Element c4 08 T
Condition Number (psi)
(psi)
(psi)
Deadweight 10'1
-151 0
0_
Pressure 101 5671 11341 0
Seismic 101 1037 0"
1488 Sum I 6557 11341 1488
.}
+T
= 11766 psi s S = 17,800 psi P
m 2
m Table 3.
P Values, Design Condition L
Loading Element of as T
Condition Number.
(psi)
(psi)
(psi)
~
Deadweight 71
-73
-344 0
Pressure 71
' 2126 13219 0
Seismic 71 327 0
343 Sum E 2380 12875 343 2
c4 os 6fa
+T2 = 12886 psi s 1.5 S = 26,700 psi P
=
t m
4 w
2-12
or P ) + Pb Values,
,P y Myp Table 4.
(P L
m i
b Design Condition s
~
Membrane Bending Stress Stress Loading Element Condition Number c4 as of as T
Deadweight 96
-176
-30
-151
-45 0
Pressure 96 5604 5572 13250 3975 0
Seismic 96
+296 0
6 0
372 Sum I 5724 5542 13105 3930 372
~
(o4) Total = 5724 + 13105 = 18829 (ce) Total = 5542 + 3930 = 9472
~
T' = 372 psi r
(
) Total + ( e) Total *
( *) Total + ( e) Total
+T2 S.I. =
}
= 18844 psi s 1.5 S, = 26.,700 psi Table 5.
Total stress at base location due to MSLB,
' Service Condition Membrane Bending Loading Stress Stress Description c4 as of os T
Deadweight
-51 0
-153 907 272 0
Pressure 3908 1173 2419 726 0
Seismic 2450 0
1275 0
757 Thermal 3816
-35366 36570 11062 0
Sum I 9664
-34346 41171 12060 757 (oc) Total = 9664 + 41171 = 50835 psi (ae) Total =l-343461 + 12060 = 46406 psi T = 757 psi I
) Total + ( e) Total
( *) Total - ( e) Total
+ 72
~~
S.I* =
2 2
j
=.50961 psi 35m = 53400 psi
-- 13
_g
1 l
Table 6.
Totalstressatbaseiocationduetodoubleended recirculation line break; Service Conditions.
Membrane Bending Stress Stress Loading Description o4 as of os T
s Deadweight
-51 0
-153 907 272 0
Pressure 7152 2146 4427 1328 0
Seismic 2450 0
1275 0
757 Thennal 3112
-28840 29845 9045 0
Sum I 12204
-26847 36454 10645 757 (ot) Total = 12204 + 36454 = 48658 psi (ae) Total =l-26847l+10645.=37492 psi T = 757 psi
('
32
- Total +(
) Total -[(
) Total - ( e) Total
+ h2 S.I. -
2
)
= 48709 psi s 35,= 53400 psi
,ar 2-14
_a
~
Deadweight, pressure, and seismic loading apply for " Design Con-ditions Only"; thermal loading is not considered.
The following conditions l
must be satisfied for code compliance:
PmISm P f 1.5 S, t
(P[ or P ) + Pb i 1.5 S,
m Deadweight, pressure, seismic, and thermal loading apply for
_ Service Conaitions.
The following conditions must be satisfied for code compl.iance:
(P, or P ) + Pb+Qf3S, L
P is the general primary membrane stress.
P is the local m
L primary membrane stress.
P is the bending stress.
Q is the secondary b
stress.
The definitions for P, P, P and Q are described in the ASME m
L b
code.
In this study, Q is the bending stress at gross structural dis-continuities.
The values of S can be found in Table I-10.0, Appendix A of m
the ASME code.
The sum of general membrane stresses (P,) listed. in Table 2 meet the general primary membrane stress intensity requirements of Pm i 1.0 Sm.
The local membrane stress intensity, primary membrane stress intensity, and primary plus secondary stress intensity shown in Tables 3, 4, 5, and 6 respectively all meet the stress intensity limits specified by the ASME Code.
o 2-15
r
"%#'"'4be"$D%
fi A
CHAPTER 3 ~
SUMMARY
AND CONCLUSION i
~
The drywell containment structure was analyzed for four combined-load conditions:
deadweight, accident pressure (P
= 37 psig), thermal max loads (Tmax = 280*F), and seismic loads from a 0.179 peak ground accelera-tion, Reg. Guide 1.60 spectra. The maximum stress intensity produced in the drywell was 50.9 ksi at the base location.
All stresses in the drywell were within the stress intensity limits of the Design and Service Condi-tions given by the ASME-Boiler and Pressure Vessel Code,Section III,1977,
' subsection NE.
The meridional critical buckling stress is estimated to be five times ' the meridional compressive stress calculated for deadweight and seismic loading.
The calculated circunferencial critical buckling stress is less than the 35.4 ksi 'circumferencial compressive stress produced at the base due to thermal loading.
There could be some localized circum-ferencial buckling at the base of the structure, but the self-limiting nature of thermal stresses suggest that no adverse structural degradation will appear.
It should be noted that no credit was taken for the structural effect of the concrete.that extends upward from the drywell base.
This would certainly lower any chance that buckling would occur.
e e
~
3-1
REFERENCES-1.
D.
G.
Vreeland, " Mass and Energy Release for Possible Pipe Break Inside Containment, Containment Pressure and Heat Removal Capability for Oyster Creek Nuclear Power Plant," Lawrence Livermore National Laboratory, Livermore, California, Letter report addressed to I. R.
Finfrock, Jr.
2.
R.
C.
Murray, T.
A.
Nel son, S.
M. Ma, J.
D.
Stevenson, " Seismic Review of the Oyster Creek Nuclear Power P1 ant,as Part of the Systematic Evaluation Program," Nureg/CR-1981.
3.
ASME Boiler and Pressure Vessel Code,Section III, 1977 Edition, Appendices.
4.
S. W. Chous, Burns and Roe Inc., letter addressed to Y. Nogai, Oyster Creek Nuclear Station SPE Seismic Evaluation Program.
5.
- ANSYS Engineering Analysis System User's Manual, Swanson Analysis System Inc., Houston, Pennsylvania.
6.
Jersey Central Power and Light Company, Facility Description and Safety Analysis Report, Volume 1, NRC Docket Item 50219-1.
7.
" Seismic Safety Margin Evaluation Reactor Building Primary Coolant Loop, Big Rock Point Nuclear Power Plant," D' Appolonia for Consumer Power.
8.
" Site Specific Ground Response Spectra for SEP Plants Located in the Ea stern United States," Letter to all SEP Owners from Dennis Crutchfied, NRC.
- 9.,
Biggs, J.
M., Introduction to Structural 9ynamics, McGraw-Hill Book Company, New York, New York, 1964.
e w -