ML13333B109
| ML13333B109 | |
| Person / Time | |
|---|---|
| Site: | San Onofre  |
| Issue date: | 05/31/1982 |
| From: | Lo T LAWRENCE LIVERMORE NATIONAL LABORATORY |
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| CON-FIN-A-0436, CON-FIN-A-436 UCID-19141-DRFT, NUDOCS 8207300222 | |
| Download: ML13333B109 (45) | |
Text
J FR CUMMET UCID 19141 Systematic Evaluation Program Structural Review of the San Onofre Nuclear Generation Station Unit 1 Containment Structure Under Combined Loads Ting-Yu Lo May 1982 8207300222 820723 ADOCK 05000206
UCID 19141 Distribution Category:
Systematic Evaluation Program Structural Review of the San Onofre Nuclear Generation Station Unit 1 Containment Structure Under Combined Loads Ting-Yu Lo Manuscript date: May 1982
ABSTRACT This report is a structural assessment of the containment structure of the San Onofre Nuclear Generation Station Unit 1, performed for the Nuclear Regulatory Commission as part of the Systematic Evaluation Program (SEP).
The SEP is a plant-by-plant reassessment of the safety of 11 operating nuclear reactors that received construction permits between 1956 and 1967.
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 older facilities.
The San Onofre assessment focused on the overall structural integrity of the containment structure under a safe shutdown earthquake and a postulated design basis accident. The safe shutdown earthquake was represented by the Housner Spectra, scaled to 0.67 g peak ground acceleration. The postulated design basis accident was either a loss of coolant accident or a main steam line break. No structural details, such as the equipment hatch and personnel air lock, or other penetrations, were evaluated. In accordance with the
. intent of the SEP, current licensing criteria were used as scales to assess the safety margins of the containment structure, but the containment structure was not evaluated to demonstrate compliance with these criteria.
Several combined stresses were evaluated for their adherence to the 1980 edition of the American Society of Mechanical Engineers Boiler and Pressure Vessel Code allowables. All the calculated stress intensities were found to be acceptable according to this code except the general primary membrane stress due to combined dead and pressure loads under level A service limits.
Because the containment structure was previously tested under combined dead and pressure loads for a higher peak pressure than the one used here, this study concluded that it was acceptable.
The compressive hoop stress of the containment structure in the sand filled transition zone was found to be rather high, due to the thermal load induced by a postulated main steam line break.
Further detailed study is needed to reach a final conclusion about possible shell buckling in the circunferential direction.
111i
CONTENTS Abstract.
List of Figures.
List of Tables Preface
- 1. Introduction
- 2. Description of the Containment Structure.
- 3. Loads and Load Combinations 3.1.
Dead Load.
3.2. Seismic Loads....................................
3.3. Design Basis Accident Loads 3.4. Load Combinations
- 4. Analysis of the Containment Structure.
4.1.
Material Properties..............
4.2.
Modeling 4.3. Results of the Analysis
- 5. Summary and Conclusions.
References V
~
LIST OF FIGURES
- 1. Horizontal safe shutdown earthquake.
- 2.
Schematic drawing of the San Onofre Nuclear Generation Station Unit 1 containment structure.
- 3. Containment pressure during a loss of coolant accident due to a double-ended pump suction break
- 4. Containment temperature during a loss of coolant accident due to a double-ended pump suction break
- 5. Containment pressure during a main steam line break.
- 6. Containment temperature during a main steam line break.
- 7. Finite-element model of the containment structure showing (a) the overall containment structure and (b) detail of the sand filled transition zone.
- 8.
Stresses or force/moment resultants of a shell
) 9. Results of the analyses using dead load and the higher estimate of the static sand stiffness.
- 10.
Results of the analyses using a peak metal temperature of 268 oF and the higher estimate of the static sand stiffness.
- 11.
Results of the analyses using an internal pressure of 1.0 psi and the higher estimate of the static sand stiffness
- 12.
Results of the analyses using dead load plus peak pressure load and the higher estimate of the static sand-stiffness
- 13.
Results of the analyses using dead load plus thermal load plus pressure load and the higher estimate of the
- static sand stiffness.
- 14.
Results of the analyses using dead load and the lower estimate of the dynamic sand stiffness.
vi
- 15.
Results of the analyses using a peak metal temperature of 268 OF and the lower estimate of the dynamic sand stiffness.
- 16.
Results of the analyses using an internal pressure of 1.0 psi and the lower estimate of the dynamic sand stiffness.
- 17.
Results of the analyses using dead load plus peak pressure load and the lower estimate of the dynamic sand stiffness.
vii
LIST OF TABLES
- 1. Membrane and bending stresses caused by the safe shutdown earthquake.
- 2.
Peak containment pressure and temperature caused by design basis accidents 3a. Maximum general primary membrane stress (PM) for the level A service limits 3b.
Assessment based on the stress intensity of Table 3a 4a. Maximum local membrane stress (PL for the level A service limits 4b. Assessment based on the stress intensity of Table 4a Sa. Maximum primary general or local membrane stress plus primary bending stress (PL
+b for the level A service limits 5 Sb. Assessment based on the stress intensity of Table 5a 6a. Maximum primary general or local membrane stress plus secondary stress intensity (PL
+ b +
for the level A service limits 6b. Assessment based on the stress intensity of Table 6a 7a. Maximum general primary membrane stress (PM for the level C service limits.
7b. Assessment based on the stress intensity of Table 7a 8a. Maximum local membrane stress (PL) for the level C service limits.
8b. Assessment based on the stress intensity of Table 8a 9a. M.ximum primary general or local membrane stress plus primary bending stress (PL +
b) for the level C service limits.
viii
9b. Assessment based on the stress intensity of Table 9a ix
RA FT F
~fEN PREFACE This report is part of the Systematic Evaluation Program (SEP) being conducted by the Nuclear Regulatory Commission (NRC).
The SEP is a plant-by-plant reassessment of the safety of 11 operating nuclear reactors that received construction permits between 1956 and 1967. 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 older facilities.
The author wishes to thank T. A. Nelson for reviewing this report and P. Y. Chen and S. Brown, technical monitors of this work at the Office of Nuclear Reactor Regulation, for their continuing support.
We also wish to thank C. D. Judd of LLNL for publication support.
xi
- 1.
INTRODUCTION
)
As part of the Systematic Evaluation Program (SEP), Lawrence Livermore National Laboratory (LLNL) is conducting a safety evaluation of some older operating nuclear power plants for the Nuclear Regulatory Commission (NRC).
This report presents the LLNL assessment of the safety margins of the San Onofre Unit 1 containment structure to withstand the combined effects of a safe shutdown earthquake (SSE) and a design basis accident (DBA).
The effect on containment due to a DBA alone was also evaluated. The DBA was either a loss of coolant accident (LOCA) or a main steam line break (MSLB).
The SSE was represented by the Housner spectra, scaled to 0.67 g peak ground acceleration in the horizontal direction (Fig. 1).
The vertical spectra were equal to two-thirds of the horizontal spectra.
The seismic analysis of the containment structure was previously performed by the Bechtel Power Corporation.1 The stresses due to the SSE at two representative locations in the containment structure are available in Ref. 2.
The thermal and pressure loads due to a DBA were based on an NRC review of transmittals by the licensee. 3 Since we focused on the overall structural integrity, no details, such as the equipment hatch and personnel air lock or other penetrations, were evaluated. We assumed that the thickening around the penetration is strong enough to negate the discontinuity effect of the penetration. Hence, we developed an axisymmetric shell-of-revolution finite-element model to represent the containment structure.
In accordance with the intent of the SEP, the structural review was not conducted to demonstrate compliance with the current licensing criteria, such as the Standard Review Plan and the Regulator Guide. They were used only as scales to assess the safety margin of the containment structure.
10 6:
Damping %
L.
~0.6j 0.14__
02 00 0.00 25 O
- 0.11.21 Period (sec)
Fig. 1. Horizontal safe shutdown earthquake.
- 2.
DESCRIPTION OF THE CONTAINMENT STRUCTURE The San Onofre Unit 1 containment structure (Fig. 2) is a spherical shell with an inside diameter of 140 ft and a thickness of slightly greater than 1 in. The shell extends 40 ft below grade and is supported by a concrete cradle, which is a spherical segment with its rim extending only a few inches above the ground surface. A sand filled transition zone around the sphere is provided between the concrete cradle and the free-standing steel shell.
This sand filled cavity is approximately 6 ft deep and an average of 2 ft across from top to bottom. It provides a stiffness transition to reduce stress concentration in the steel shell.
Inside the containment sphere is the concrete reactor building, which provides support and shielding for a three-loop PWR nuclear steam supply system (NSSS). The main function of the containment sphere is to contain any possible release of radioactive material from the NSSS. Therefore, it is an airtight steel vessel. The equipment hatch and personnel air lock are the major containment penetrations. Other penetrations are much smaller.
The containment sphere was originally designed for a peak internal pressure of 46.4 psig and a peak temperature of 271.2 OF. 4 Assuming the stress free temperature of the sphere is 71.2 OF, the maximum temperature rise is, therefore, 200 0F. The containment sphere was also designed to withstand an internal vaccuum of 2.0 psig to cover possible extreme ambient weather conditions.
It was designed to the requirements of 1963 edition of the American Society of Mechanical Engineers (ASME) Boiler and Pressure Vessel (B&PV) Code.
3
-el. 120 ft,
.f 1 in. thick steel spheritcal shell Equator-el. 50 ft Personnel air lock-el. 47 ft 6 in.'
Top of interior concrete
- e.
143 ft Door,-el.47 ftel14f 15 ft diameter oGrade-el. 20 ft Sand filled transition zone Fig. 2. Schematic drawing of the San Onofre Nuclear Generation Station Unit 1 containment structure.
4
- 3. LOADS AND LOAD COMBINATIONS Only loads affecting overall integrity of the structure were condsidered in this analysis. These were the dead load, the seismic loads, and the pressure and thermal loads due to a postulated DBA.
3.1 DEAD LOAD The dead load was generated by multiplying the steel weight density (0.284 lb/in 3) by the containment steel volume.
3.2 SEISMIC LOADS Bechtel Power Corporation performed the seismic analysis of the containment structure. The horizontal SSE seismic input was the Housner response spectra normalized to 0.67 g peak ground acceleration. The vertical response spectra were two-thirds of the horizontal spectra. The stresses due O
to the SSE at two locations in the containment structure were taken from Ref. 2 and are shown in Table 1.
Aside from the locations close to the penetrations, these two locations, at the northern or southern most points of the sphere along the horizontal loop direction, were considered to be typical. They were also considered to be the most critically stressed areas of the sphere. The location representative of the continuous shell is at a 31-ft elevation, and the location representing the gross structural discontinuity at grade is at a 20-ft elevation. Since the shell was considered to be axisymmetric in this study, these stress results were used to represent any location in the horizontal hoop direction at the same elevation.
3.3 DESIGN BASIS ACCIDENT LOADS The licensee has performed the mass and energy analyses of the containment pressure and temperature responses due to both a LOCA and an MSLB. References 5-9 document the results of these analyses.
LLNL reviewed 3
these analyses for the NRC as part of the SEP.
The design basis LOCA was a double-ended cold leg (pump suction) break at 102% of full power. The 5
resulting post-LOCA pressure and containment atmosphere temperature profile
- are shown in Figs. 3 and 4.
The licensee presented four cases of MSLB. 7 Table 2 summarizes the peak values of containment pressure, containment atmosphere temperature, and containment metal temperature due to LOCA and MSLB cases.
The four MSLB cases showed almost the same peak containment atmosphere temperature-s, which all exceeded the peak LOCA temperature by a substantial margin. Case 3 had a higher peak metal temperature than case 1. Cases 3 and 4 had higher pressure responses than the LOCA and cases 1 and 2. The peak pressure of 52.2 psig for case 3 was only slightly less than 53.0 psig for case 4.
Case 3 pressure and temperature transients were selected for our assessment purposes because of their high peak values and the closeness of peak pressure and peak temperature on the time scale. In addition, this case was based on nominal and better estimated blowdown data as stated in Ref. 7.
The pressure and temperature profiles of case 3 are shown in Figs. 5 and 6.
The peak metal temperature was used, and the peak pressure and peak metal O
temperature were assumed to occur at the same time.
3.4 LOAD COMBINATIONS We used the load combinations specified under level A and level C service limits of the Standard Review Plan, Section 3.8.2.10 The stress intensity limits specified in this Plan are consistent with those of ASME Code,Section III, Division 1, Subsection NE, 1980 edition. The peak stresses due to dead, seismic, and DBA loads were combined algebraically with the worst combination of signs possible. Our load combinations were:
Level A service limits = 0 + T + P,
a a
Level C service limits = D + T + P + E',
a a
where D
Dead load, P = Pressure load generated by the postulated pipe break accident, Ta = Thermal load under thermal conditions generated by the postulated pipe break accident, El =Load generated by the SSE.
6
y FOR COV The live load was neglected because it was considered insignificant; the pipe reactions were not considered, since they are localized loads.
- =7
Table 1. Membrane and bending stresses caused by a safe shutdown earthquake.
Stresses are measured in psi.
Membrane stress resultants Bending stress resultants Hoop or Hoop or Location longitudinal Meridional Shear longitudinal Meridional Shear Continuous shell 1172.0 1483.0 1517.1 115.1 33.5 4.1 Gross structural discontinuity at grade 2034.4 6685.2 1849.3 7921.7 2404.3 25.8 8
n n
Table 2. Peak containment pressure and temperature caused by design basis
. accidents. The mass/energy release assumptions for MSLB cases 1 and 3 are based on better estimates. Those for cases 2 and 4 are based on standard calculation procedures.
Main steam line break accident Mass/energy release assumptions LOCA 1
2 3
4 Reactor power (%)
102 100 103 0
0 Peak pressure (psig) 49.4 48.4 50 52.2 53.0 Peak atmosphere temperature (OF) 291 405 406 403 404 Peak metal temperature (OF)
Not 239 Not 268 Not available available available 9
60 a
j I
II I
I I
i max - 49.4 psic at 62 sec!
50 40' C.
830i vij 20 10 102 103 1041 1.o5 10 l.
Time (sec)i Fig. 3. Containment pressure during a loss of coolant accident due to a double-ended pump suction break.
T (vapor)= 291. F at 47 secl ma Containment vaporl 250 200 C.
E 50 I
10l
- 0 I1OI 10J 105l 1061 lo0l Time (sec)
Fig. 4. Containment temperature during a loss of coolant accident due to a double-ended pump suction break.
10
60 -
P 52,2 psig at 355 sec' 50 40 30 20 10 1
- 10.
100.
- 1000, 10000 Fig. 5. Containment pressure during a main steam line break.
(vapor) 403 0F at 31 see 400 350 E-0 200
%In'side surface containment sphere 150 -
/e 1
10 100 1000 10000 Time (seconds)
Fig. 6. Containment temperature during a main steam line break.
7,- -1
- 4.
ANALYSIS OF THE CONTAINMENT STRUCTURE
. 4.1 MATERIAL PROPERTIES The containment shell material was supplied to the ASME specification SA-212, Grade B Fire Box Quality produced to SA-300 specification. This corresponds to current ASME Specification SA-516, Grade 70 Plate Material.
The 1980 edition of the ASME code appendices specify the following material properties corresponding to the peak sphere metal temperature of 268 0F:
o Minimum yield strength, S = 34,000 psi, y
o Minimum ultimate tensile strength, S = 70,000 psi, 0
Allowable stress intensity, Smc = 19,250 psi, 0
Design stress intensity, Sm1 = 22,700 psi, o
Young's modulus, E = 29.2 x 10
- psi, o
Coefficient of thermal expansion, a = 6.14 x 10-6 in/in/ oF.
Bechtel Power Corporation studied documented test data consisting of 12 tensile strength tests that included welds as well as the material itself.
These test data indicate that the Su of the sphere material exceeds the code specification considerably. A statistical approach was used to establish the S u. Its value was selected as 78,572 psi to ensure that 95% of the statistical strength values would exceed the S with a 90% confidence level.
This value is about 1,800 psi less than the minimum of the 12 test values, but it exceeds the ASME code specified value of 70,000 psi by 12%.
The calculated S mc is 21,600 psi.
The S used in Bechtel's analysis of the containment sphere was 38,000 psi.
y The Poisson's ratio and density for steel were assumed to be 0.3 and 3
0.284 lb/in, respectively. The sand in the transition zone was assumed to have the following best-estimate material properties 11 o
Young's modulus, E Static = 48,600 psi (or 7,000 ksf),
Dynamic = 20,800 psi (or 3,000 ksf),
o Poisson's ratio 0.35.
Both steel and sand were assumed homogeneous and isotropic.
- s 12
4.2 MODELING An axisymmetric finite-element model was constructed to analyze the containment structure using the MODSAP computer program. The steel sphere was modeled by 83 shell-of-revolution finite elements. Twenty-four solid-of-revolution finite elements were used to simulate the sand filled transition zone. In the meridional direction the length of the elements in the sand filled transition zone and the region right above grade were made relatively small compared to the rest of the shell in order to capture the bending behavior of the shell in this region. This decision was made as the result of a sensitivity study of finite-element mesh sizes in this region.
Figure 7 shows the finite-element model, with the detail of the sand filled transition zone. We believed that the sand close to the surface was not capable of developing the full stiffness.
Hence, we arbitrarily reduced the Young's moduli for the top three layers of finite elements to 1/6E, 3/6E and 5/6E, respectively.
4.3 RESULTS OF ANALYSIS To cover possible uncertainty involved in modeling the sand, including the reduction of stiffness for the top three sand layers, the Young's moduli were increased by 100% and reduced by 50% to get higher and lower estimate values.
The analyses were done using these estimated values together with the best-estimate sand stiffness.
The most critical stress values from these three sand stiffnesses were then used in the assessment of the containment structure.
Two sets of finite-element analyses were performed: one for each of the service levels considered. For level C service limits, the Young's modulus corresponding to the dynamic condition for the sand fill was used since the load combination included the dynamic load of the SSE. The static Young's modulus was used for level A service limits.
Each set of finite-element analyses included three computer runs using the best estimate, higher estimate, and lower estimate of the sand stiffness.
The two extreme cases are the lower estimate case of dynamic sand stiffness, which has the lowest sand Young's modulus, and the upper estimate case of 13
AFT FO COMMENT static sand stiffness, which has the highest sand Young's modulus.
The
. notations and sign conventions for the stresses or force/moment resultants of a shell are shown in Fig. 8. The results for the analyses are presented in Figs. 9-17.
The abcissa indicates the location on the spherical shell along its meridional direction. This is represented by the angle, measuring from the vertical axis as shown in Fig. 7a.
The ordinate represents the response quantities, such as the force/moment resultants or the normal stresses.
Figures 9-11 are the force/moment resultants due to the dead load, the peak thermal load, and the pressure load of 1.0 psi using the higher estimated sand stiffness. Figure 12 shows the normal stresses at the mid-surface and the two extreme fibres of the steel shell for combined dead and peak pressure loads. Figure 13 shows the normal fibre stresses for combined dead load, peak thermal load, and peak pressure load.
Corresponding results (except for those of Fig. 13) using the lower estimated sand stiffness are shown in Figs. 14-17.
The results presented in these two sets of figures allow one to get a measure of the sensitivity of
. shell stresses to the possible variations in sand stiffness because the results of other cases are bounded by these two extreme cases.
In general, the variation of sand stiffness affects the shell stress distribution in the sand filled zone and the area immediately above it. As one would expect, it has little effect on the stresses far away from the sand filled transition zone.
Neither does it affect the membrane stresses much.
However, it does have a significant effect on the bending resultants (Ms' M ), especially on the bending resultants close to the gross structural discontinuity at grade.
The capability of the containment sphere to withstand the postulated DBA and the combined effect of a DBA and SSE was assessed based on the criteria specified by the ASME code and the Standard Review Plan as stated in Section 3.4. Tables 3-6 present the assessment for level A service limits; Tables 7-9 show the assessment for level C service limits.
To measure the structural integrity of the containment structure under variou: loads or load combinations, the term Safety Margin (SM) was adopted from Ref. 1. The SM is defined as the ratio of the allowable stress intensity O
limit specified by the code to the calculated stress intensity. The larger 14
L AFT FOR CO PEI the SM, the better the structure is able to withstand the specified loads.
The "a" part of Tables 3-9 show the stresses due to various loads and load combinations, along with the maximum stress intensity for the specific load combinations. The "b" part of each table presents the SM based on the metal strengths established by Bechtel or specified by the ASME code.
In these tables all SMs are greater than 1.0 except for the case of maximum general primary membrane stress (P m) under level A service limits for a dead load plus pressure load combination. In this case the SM values are equal to 0.94 and 0.84 for metal strength established by Bechtel and specified by the code, respectively. The structure is not necessarily unsafe when SM values are lower than 1.0.
However, it does indicate that the strength of the containment sphere is lower than the code requirements. The containment was previously tested under 53.4 psig.7 This test pressure is higher than the peak pressure used in this study. Therefore, we considered the SM values of 0.94 and 0.84 acceptable even though they are lower than 1.0 Figures 10 and 15 indicate that the containment sphere is under very high compressive hoop stress in the sand filled transition zone due to the postulated MSLB thermal condition. The peak hoop stress is 27,500 psi at the fixed location. It maintains a rather hi'gh stress level throughout the sand filled zone and drops off sharply above grade level.
The buckling of the shell under this high compressive hoop stress is a very complicated problem, and there is no simple method that can predict accurately the critical buckling stress.
Nevertheless, two extreme cases of critical buckling loads were calculated. One approach was based on ASME code buckling stress for a uniform external pressure on a spherical shell.
This buckling stress was calculated at 1100 psi.
Another approach was based on the local circumferential buckling of a uniformly heated thin cylindrical shell with a rigid bulkhead. 12 The predicted buckling stress for this case was 262,800 psi, which is well beyond the elastic range. These two cases do not reflect the true condition at the sand zone but rather comprise upper and lower bounds.
The pressure load applied at the same time as the thermal load provides some stabilizing effect, in addition to that provided by the rigid restraint at the base of the sand filled zone. Unlike some of the boiling water reactor (BWR) drywells, which also have a sand filled transition zone at the base, the 15
AFT FOR COMMET San Onofre Unit 1 containment sphere does not have concrete backing inside in
. this region. The top of the concrete base of the reactor building stops at the bottom elevation of the sand filled cavity.
Inward buckling of the steel shell is possible under the high thermal condition caused during a DBA.
However, further detailed study is needed to assess the safety margin of the containment sphere in this area.
16
- 5.
SUMMARY
AND CONCLUSIONS The containment sphere of the San Onofre Nuclear Generation Station Unit 1 was analyzed for the combined effect of dead load, SSE loads, and the pressure and thermal loads generated by either a LOCA or an MSLB.
The combined stresses were evaluated according to the 1980 edition of the ASME Boiler and Pressure Vessel Code,Section III Subsection NE for Class MC components. All the calculated stress intensities are within those allowed by the code, except for the case of general primary membrane stress (Pm) under level A service limits. The SMs for this case are 0.94 and 0.84 for metal strength established by Bechtel and specified by the code, respectively.
Since the containment sphere was previously tested under that particular load combination (dead load plus pressure load) for an even higher pressure (53.4 psig), the containment is considered acceptable for that condition.
The compressive hoop stress of the sphere in the sand filled transition zone is rather high (27,500 psi) due to the thermal condition induced by a postulated MSLB case. Since there is no simple method that can accurately predict the critical buckling stress in that region,-further detailed study is needed to reach a final conclusion about possible shell buckling in the circumferential direction.
17
Table 3a. Maximum general primary membrane stress (P m) for the level A service limits. Measurements were taken in psi a few ft above grade.
Meridian Hoop Shear Load (aS)
(
Dead, D
-400 488 0
Pressure, Pa 21,924 22,493 0
Sum, D + Pa 21,524 22,981 0
Maximum stress intensity = 22,981 Table 3b. Assessment based on the stress intensity of Table 3a.
Assessment Bechtel ASME Allowable stress intensity limit (psi)a.
21,600 19,250 Safety margin 0.94 0.84 1.0 x Smc' 18
Table 4a. Maximum local membrane stress (PL) for the level A service limits. Measurements were taken in psi at grade elevation.
Meridian Hoop Shear Load (a )
( )T)
Dead, D
-403
-24 0
Pressure, Pa 21,924 10,750 0
Sum, D + Pa 21,521 10,726 0
Maximum stress intensity 21,521 Table 4b. Assessment based on the stress intensity of Table 4a.
Assessment Bechtel ASME Allowable stress intensity limit (psi)a 32,400 28,875 Safety margin 1.50 1.34 a 1.5 x S.
mc
- a 19
Table Sa. Maximum primary general or local membrane stress plus primary bending stress (PL
+ b) for the level A service limits. Measurements were taken in psi a few ft above grade.
Meridian Hoop Shear Load (aS)
(
Dead plus pressure, O + Pa 26,779 22,780 0
Maximum stress intensity = 26,779 Table 5b.
Assessment based on the stress 4
intensity of Table 5a.
Assessment Bechtel ASME Allowable stress intensity limit (psi) a 32,400 28,875 Safety margin 1.21 1.08 a 1.5 x Smc,
- a 20
Table 6a.
Maximum primary general or local membrane stress
- plus secondary stress intensity (PL b + Q) for the level A service limits. Measurements were taken in psi at grade elevation.
Meridian Hoop Shear Load (as)
( )
()
Dead plus pressure plus
- thermal, D + P a+
T 54,170
-1,980 0
Maximum stress intensity = 56,150 Table 6b.
Assessment based on the stress intensity of Table 6a.
Assessment Bechtel ASME Allowable stress intensity limit (psi)a 76,400b 68,100 Safety margin 1.36 1.21 a 3.0 x S ml b No S was given in Refs. 2 and 7, since the 1974 edition of the ASME code was used in Bechtel's study. The Sml for material property established by Bechtel was assumed to be S ml Smc X
S mc Bechtel.
21
Table 7a.
Maximum general primary membrane stress (P ) for the level C service limits.
Measurements were taken in psi a few ft above grade.
Meridian Hoop Shear Load (aS)
(a )
(1)
Dead, D
-400 489 0
Pressure, Pa 21,924 22,483 0
SSE, E' 6,685 2,034 1,849 Sum, D + Pa + E' 28,209 25,006 1,849 Maximum stress intensity = 29,050 O
Table 7b.
Assessment based on the stress intensity of Table 7a.
Assessment Bechtel ASME Allowable stress intensity limit (psi)a 38,000 34,000 Safety margin 1.31 1.17 a The larger of 1.25 Smc and S.
- a 22
Table 8a. Maximum local membrane stress (PfL for the level C service limits. Measurements were taken in psi at grade elevation.
Meridian Hoop Shear Load (as)
(a )
()
Dead, 0
-406 41 0
Pressure, Pa 21,924 12,082 0
SSE, E' 6,685 2,034 1,849 Sum, D + Pa + E' 28,203 14,157 1,849 Maximum stress intensity = 28,440 Table 8b. Assessment based on the stress intensity of Table 8a.
Assessment Bechtel ASME Allowable stress intensity limit (psi)a 57,000 51,000 Safety margin 2.00 1.79 a The larger of 1.8 S and 1.5 S.
mc y
23
Table 9a.
Maximum primary general or local membrane stress plus 0
primary bending stress (PL b) for the level C service limits. Measurements were taken in psi at grade elevation.
Meridian Hoop Shear Load (aS)
(oe Dead plus pressure, D + Pa 31,758 13,680 0
SSE load, E' 6,685 2,034 1,849 Sum, D + Pa + E' 38,443 15,714 1,849 Maximum stress intensity = 38,592 Table 9b.
Assessment based on the stress intensity of Table 9a.
Assessment Bechtel ASME Allowable stress intensity limit (Psi)a 57,000 51,000 Safety margin 1.48 1.32 a The larger of 1.8 Smc and 1.5 S y 24
1 in.
53 @ 20 = 1060 angle (degrees),
AIB 2 20 160!,4 6 @ lo =.60 A = 30.94970 B = 25.3769 0 !
54 60 12 @ 0.2814 = 3.376902 Grade-el. 20 ft Fixed at el. 14 ft!
12 @ 0.4644= 5.57280
.F 65 66 6-7 Tb) 68 69 70 71 74 El. 20 ft 72' 71; 8
- Fi
- 7. Finite-element model of the containment structure showing (a) the overall containment structure and (b) detail of the sand filled transition zone.
25 F
.84-
Mse N s '
s MsM N9, aGO' M
MeM Fig. 8. Stresses or force/moment resultants of a shell.
26
<Cirvor
-f.
O RAFT FCR COMENT 500 (a)
Dead Load 4 0 0 300 N6 200 100-0 10 20 30 40-50 60 70 80 90 100 110 120 Angle (degrees)
-100
-200-I
-3001
--400 100- (b)
Dead Load Ms 80.
)
60 40 Me 20-
-C o
0 3
0-3m 2030 40 50 60 70 80 90 100 11 0 120 Angle (degrees)
-201
-40 Fig. 9.
Results of the analyses using dead load and the higher estimate of the static sand stiffness.
(a) shows force resultants and (b) shows moment resultants.
27
20000 !(a)
Uniform Temperature 16000-1 12000 8000]
Z 4 000 -u 0
10 20 30 40 50 60 70 80 90 100 110 1
S 5
-4000-Angle (degrees)
C
-8000 U-12000-N6
-16000
-20000
-24000
-28000 3200 -
Uniform Temperature 2800 -Ms 2400 2
2000 1600 1200 Cz M
800 400 A
10 20 30 40 50 60 70 80 90 100 i
1
-400 Angle (degrees)
-800
-1200
-1600 Fig. 10.
Results of the analyses using a peak metal temperature of 268 OF and the higher estimate of the static sand stiffness.
(a) shows force resultants and (b) shows moment resultants.
28
500 (a)
Pressure (1 psi) 400 N
c 300 No 200 100 CD 0
0~
0 10 20 30 40 50 60 70 80 90 100 110 120 Angle_ (degrees)
-100
-200 40- (b)
Pressure (1 psi) 36 32-Ms 28 24]
201 161 12 8MO us LL 41 S0 E
0 10 20 30 40 50 60 70 80 90 100 110 120 2
Angle (degrees)
-81
-12
-16 -24
. Fig. 11.
Results of the analyses using an internal pressure of 1.0 psi and the higher estimate of the static sand stiffness.
(a) shows force resultants and (b) shows moment resultants.
29
(a)
ERX AI FOR -C@MRE NT 32000 Dead + Pressure Loads 28000 24000 2
20000 16000 12000 E6 8000 cE C
a 4000 -
- u.
76 0
0
'1 20 0
4 60 70 80 90 100 110 120 C -4000-Angle_(degrees)
-8000
-12000-Normal stress at mid-surface
-16000 Extreme fibre stresses
-20000 S
32000D (b) ead+ Pressure Loads 28000-24000.
20000 16000 12000 8000
_4000 0
0 20 30 40 50' 60 70 80 90 100 110 120
-4000
- 0+ 0 Angle (degrees) 0 -80001
-12000 j
Normal stress at mid-surface
-16000 Extreme fibre stresses
-20000 O
Fig.
- 12.
Results of the analyses using dead load plus peak pressure load and the higher estimate of the static sand stiffness. (a) shows meridional normal stresses and (b) shows hoop normal stresses.
30
_R
-AFT FOR COMMENT (a) 50000-(
Dead + Thermal + Pressure Loads 5000 A40000 30000 o-
-0 afl U10000-0 10 20 30 40 50 60 70 80 90 100 110 120 Angle (degrees)
-10000
-lNorm tress at mid-surface
-200007-ExtreIme fibre stresses---
(b) 50000-Dead + Thermal + PressureLoads 40000 30000 20000 E(b 10000-Da T
0 0
20 30 40 50 610 70 80 90 100 110 120 Angle (degrees)
-10000
- 200001
-Normal stress at mid-surface
-30000-Extreme fibre stresses Fig. 13.
Results of the analyses using dead load plus thermal load plus pressure load and the higher estimate of the static sand stiffness.
(a) shows meridional normal stresses and (b) shows hoop normal stresses.
31
500(a)
RAFT FOR CTMMENT Dead Load 400-,
300 N
2 200 10 0,
C 0
10 20
.3'0 4'0.-.-.* 50.60 70 80 90o 100.
1-10 120
,--Angle (degrees) 0
- a.
1_0
-2.00 N S
-300
-400 70 - (b)
Dead Load 60 M
s 50-240 30 20 10-Mb 1
10 0
10 20 3
40 50 60 70 80 90 100 110 12
]1 Angle (degrees)
-20
-30.
-40 Fig. 14.
Results of the analyses using dead load and the lower estimate of the dynamic sand stiffness. (a) shows force resultants and (b) shows moment resultants.
32
1 60Q0- o Uniform Temperature 12000-kM, 8000-<
40001 U
4ooo u
o 1t 2
N 0
10 20 30 40 50 60 70 80 90 100 110120 4000 Angle (degrees) 0
- u. -8000
-12000 N6
-16000
-20000
-24000]
728000 2 4 0 0 J (b)
MS 2000 Uniform Temperature
- 1600
'T 1200 800-T 400-U 00 0 0 20 30 40 50 60 70 80 90 100 110 120 0 10 20Angle (degrees)
\\
-400
-800
-1 200
-1600 Fig. 15.
Results of the analyses using a peak metal temperature of 268 OF W
and the lower estimate of the dynamic sand stiffness. (a) shows force resultants and (b) shows moment resultants.
33
500- (a)
Pressure (1 psi)
NS 40Q 300-N 200 100 Cz JL 0
0 10 20 30 40 50 60 70 80 90 100 110 120 T
Angle (degrees)
-100
-200 (b)
Pressure (1 psi)
) 28 24 20
.S -16 12 MO 4-( U a
0 10 20 30 40 50 60 70 80 90 100 110 120 0
2
-4 AngIe (degrees)
-8
-12
-16 Fig. 16. Results of the analyses using an internal pressure of 1.0 psi and the lower estimate of the dynamic sand stiffness.
(a) shows force resultants and (b) shows moment resultants.
34
AFT F~i l
C EMMET 28000 Dead + Pressure Loads 24000
/
20000 16000 12000 8000 4000-L o
0 0
10 20 30 40 50..
60 70 80 90 100 110 120
-4000-Angle (degrees)
-8000 Normal stress at mid-surface 12000-Extreme fibre stresses
-16000 28000 -()
Dead + Pressure Loads 24000 20000 16000 12000
-8000 -
-C 4
0-CZ.
~ oo U,02.
_0 4
5 0
70 8
0 10 1
0 o
0 10 20 3
40 50 0
70 0
o io
-o c -4000 Angle (degrees)
-8000 Normal stress at mid-surface
........ Extreme fibre stresses
-16000 Fig. 17.
Results of the analyses using dead load plus peak pressure load and the lower estimate of the dynamic sand stiffness. (a) shows meridional normal stresses and (b) shows hoop normal stresses.
35
~~V
'Ynlagl'ol REFERENCES
- 1. Seismic Reevaluation and Modification, San Onofre Nuclear Generation Station Unit 1, Southern California Edison Company and San Diego Gas &
Electric Company, Rosemead, CA, Nuclear Regulatory Commission Docket 50-206 (April 1977).
- 2. A. Sanders, Bechtel Power Corporation, Norwalk, CA, letter to R. C.
Blaschke (August 7, 1981).
- 3. D. M. Crutchfield, Nuclear Regulatory Commission, Washington, DC, letter to R. Dietch (January 12, 1982).
- 4. Final Engineering Report and Safety Analysis, San Onofre Nuclear Generating Station Unit 1, Southern California Edison Company and San Diego Gas & Electric Company, Rosemead, CA, Nuclear Regulatory Commission Docket 50-206 (year).
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- 6. D. B. Vassallo, Nuclear Regulatory Commission, Washington, DC, letter to C. Eicheldinger (March 12, 1975).
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- 8. K. P. Baskin, Southern California Edison Company, Rosemead, CA, letter to D. M. Crutchfield (March 6, 1981).
- 9. J. M. Geets, MARVEL--A Digital Computer Code for Transient Analysis of a Multiloop PWR System, Westinghouse, WCAP-7907 (June 1972).
- 10.
Standard Review Plan, Nuclear Regulatory Commission, Washington, DC, NUREG-0800, revision 1 (July 1981).
- 11.
A. Sanders, Bechtel Power Corporation, Norwalk, CA, letter to R. C.
Blaschke, BPC/SCE-81-149 (June 9, 1981).
- 12.
- 0. J. Johns, Local Circumferential Buckling of Thin Circular Cylindrical Shells, National Aeronautics and Space Administration, Washington, DC, TN D-1510 (1962).
O CDJ/mg 36