ML072850111
| ML072850111 | |
| Person / Time | |
|---|---|
| Site: | Oyster Creek |
| Issue date: | 01/31/2007 |
| From: | - No Known Affiliation |
| To: | Office of Nuclear Reactor Regulation, NRC/SECY |
| SECY RAS | |
| References | |
| 50-219-LR, Amergen-Intervenor-60, RAS 14378 | |
| Download: ML072850111 (10) | |
Text
-RA5 (lq3713 U.S. i LA E u3 tt, (of .. L -" A fit&IC~flseeg treo N ff Other OnfWItrAssDPaneJ Citizens Exhibit 60 Portions Deliberately Omitted Only Included 77-85 Other Relevant Pages Submitted by NRC Staff as Staff Exhibit 6 DOCKETED USNRC October 1, 2007 (10:45am)OFFICE OF SECRETARY RULEMAKINGS AND ADJUDICATIONS STAFF'inp (c(,e~ = 5e:cy- o F U.D.
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..itLens Exhibits 6.0 N -2._... J-1ici .v ..'. ' -OFFERED by: .,.:.5. Sandbed Region Minimum wimowU -In addition to the stress and stability analysis of the drywell shell using the average rneas-urements in the sandbed region (thicknesses described in Section 2.6, and analyses outlined in Sections 3 and 4), a minimum sandbed thickness study was also performed.
These analyses aim to establish the minimum uniform thickness in the sandbed region that maintains compliance with the ASME B&PV code. The minimum acceptable shell thickness established here is based on a buckling (stability) analysis for the refueling load case. The refueling load case appears to govern the potential for instability since a relatively low effective factor of safety was produced in the average UT measurement analysis at 2.15. For Service Level B (refueling condition), a factor of safety of 2.0 is required by ASME N-284.The previous GE analysis (GE, 1991b) assumed a uniform sandbed shell thickness of 0.736".Their analyses produced an applied compressive stress of 7.58 ksi in the sandbed region and an inelastic buckling stress of 21.30 ksi (per ASME N-284). This produces an effective factor of safety of 2.81. A subsequent calculation documented in a 1993 GPU Nuclear Calculation Sheet (GPU Nuclear, 1993) shows an applied compressive stress of 7.58 ksi in the sandbed region for a shell thickness of 0.736", but with a lower value for the inelastic buckling stress at 15.18 ksi.This produces an effective factor of safety of 2.0, or at the required ASME N-284 value.The inconsistency between the two calculations appears to stem from a difference in the applica-tion of the increased capacity reduction factor due to the tensile stresses in the circumferential (hoop) direction.
This issue was discussed in detail in the previous stability analysis section. Ar-ticle 1500 of ASME N-284 states clearly that an increased capacity reduction factor may be jus-tified if an internal pressure loading is present and causes tensile stresses in the circumferential direction.
This internal pressure aids in "smoothing" the initial imperfections and increased the buckling capacity under-compressive meridional stresses.
The lack of an internal pressure load for the refueling load case prevents the justified use of.an increased capacity reduction factor.As with the buckling calculations for the refueling load. case in the previous section, the mini-mum thickness study does not employ any increase in the capacity reduction factor.The shell thicknessesused in the minimum thickness study are summarized in Table 5-1 for re-gions outside of the sandbed region. The degraded thickness values for the majority of the dry-well are equivalent to the values Used in the average UT measurement analysis.
The only exception being the thickness assigned to the lower sphere above an elevation of 15'-6.8", or the center of the ventlines.
In this region of the lower sphere (see Figure 5-1), the thickness is set to 1.154", or the nominal as-built value. This remains consistent with inspections of the upper por-tions of the lower sphere. In the average UT measurement analysis, additional conservatism was introduced by degrading the entire lower sphere uniformly in each bay combination.
However, several confirmatory analyses performed during this studyshowed that the thickness assigned to the lower sphere above elevation 15'-6.8" has only a negligible effect since the buckling occurs in the sandbed below 15'-6.8".In the lower sphere below elevation 15'-6.8" (sandbed region), the drywell shell is set to a uni-form thickness.
This region is shown in Figure 5-1. While the same finite element mesh was used as for the average UT measurement analyses, the local thinned regions under the ventlines 77 for Bays 1 and 13 are uniformly thinned consistent with the surrounding shell. In addition, this study only examined the minimum thickness required in the sandbed region and not in the upper portions of the sphere or in the cylinder.Table 5-1. Main Drywell Shell Model Thicknesses Outside of Sandbed Region Original Degraded Original Degraded Section Thickness, Thickness, Section Thickness, Thickness, in in in in Head 1.1875 N/C Reinforcing Around Ventlines 2.875 2.618 Upper Cylinder 1.1875 N/C Lower Sphere (below Sandbed) 1.154 N/C Main Cylinder 0.640 0.585 Bottom Sphere 0.676 N/C Knuckle 2.5625 2.54 Middle Sphere Thickened 1.0625 0.9625 Upper Sphere 0.722 0.676 Reinforcing Around Hatch 2.625 2.525 Middle Sphere 0.770 0.670 Lower Sphere (above El. 15'-6.8")
1.154 N/C N/C -No Change' Elevation 15'-6.8" Lower:Sphere Below Elevation Thickneess
= Varied to .Establish MinirnI6ni m u Figure 5-1. Drywell Lower Sphere for Establishing a Minimum Thickness in the Sandbed Region (Ventlines and Hatch Removed for Clarity)78 The thickness values assigned to the sandbed region were varied from 0.800" up to 1.050" with a concentration of analyses performed between 0.800" and 0.860". In the previous buckling analysis section, a buckling analysis was also performed for the undegraded drywell containment which included a uniform thickness of 1.154" throughout the sandbed region. The results of each of these analyses are summarized in Figure 5-2. Here, the effective factor of safety is plot-ted against the associated shell thickness in the sandbed region. This study shows that a thick-ness of 0.844" is required in the sandbed region to produce an effective factor of safety equal to the ASME N-284 value of 2.0.Figure 5-2 also plots the datapoint established in the previous buckling analysis section using average UT measurement data. In that analysis, the bay combination that buckled first was set to a thickness of 0.842" and resulted in an effective factor of safety equal to 2.15. Although the thicknesses used in the minimum thickness analysis and the average UT measurement analysis are essential equivalent, there are several important factors that produce the difference in safety factors. First, the average UT measurement analysis included two locally thinned regions that, in general, cause lower effective factors of safety for buckling in the adjacent bays than without the locally thinned regions. However, the effect of the locally thinned is outweighed by the exis-tence of bay combinations with thickness far exceeding 0.842" (see Figure 2-32). For the aver-age UT measurement analyses, 5 out of the 10 bay combinations were assigned thicknesses near or above 0.9". The existence of thicker bays enables a redistribution of the compressive loads leading to buckling.
Therefore, the average UT measurement analysis produced an effective fac-tor of safety of 2.15 with a thickness of 0.842", while the minimum thickness study produced an effective factor of safety of 2.0 with a thickness of 0.844". In the minimum thickness study, the entire sandbed region was uniformly thinned which prevents any redistribution of the load through thicker shell regions. The effect of the locally thinned regions was not rigorously ex-plored in the average UT measurement analyses, but it is likely that the effective factor of safety of 2.15 would increase without the presence of the locally thinned region.4 F S 3 .5 *:: ....
- ,.g,.; ,.,. , .. :5 ; z*, S-0 3 0 2.5 T 2 2.* Minimum Thickness Study .: o,.Average UT Measurment Analysis ','... .....-....
.......-r:- : ..A.°j 0.6 0.7 0.8 0.9 1 1.1 1.2 Sandbed Shell Thickness, in Figure 5-2. Effective Factor of Safety Values Computed for Various Thicknesses in the Sandbed Region for the Refueling Load Combination 79 Figure 5-3 and Table 5-2 illustrate the buckling location and ASMIE N-284 calculations for the sandbed with a thickness of 0.844". The major displacements for the first buckling mode in the sandbed are located between the ventline in Bays I and 3.Figure 5-3. Buckling in the Sandbed Region with a Thickness of 0.844" for the Refueling Load Combination Table 5-2. Buckling Evaluation for the Refueling Load Case with a Thickness of 0.844" in the Sandbed Sphere Radius, in 420 Sphere Thickness, in 0.844 Material Yield Stress, ksi 38 Elastic Modulus, ksi 29500 Factor of Safety, FS 2 Applied Meridional Compressive Stress from Analysis, ac, ksi 4.78 Load Factor from Bucking Analysis, X 9.67 Theoretical Elastic Buckling Stress, Gie = Xac, ksi 46.19 Capacity Reduction Factor, a 0.207 Reduced Elastic Instability Stress, 0e = ayie, ksi 9.56 Yield Stress Ration, A = aelay 0.252 Plasticity Reduction Factor, Tn 1.0 Inelastic Instability Stress, a0 = 7lae, ksi 9.56 Allowable Compressive Stress, ay1 = acIFS, ksi 4.78 Applied Compressive Stress Percentage of Allowable, Uc/aa,* 100 100.0%Effective Factor of Safety, FSE = os/aC 2.00 80
- 6. Summary of Assumptions The study performed for this program required a number of assumptions.
A summary of the most significant assumptions is provided below." The Accident and Post-Accident load combinations are assumed to govern the stress analysis." The Refueling and Post-Accident load combinations are assumed to govern the buckling (stability) analysis." Information of.the loads applied to the finite element model was taken from the previous study by GE. These loads were not independently verified." The seismic loading was applied using static coefficients provided in the Final. Safety Analysis Report (FSAR, 2003). The static coefficients were applied using body forces in both the vertical and lateral directions.
The displacement time histories (ground motions)were not made available for this study. The body forces used for the seismic loads were increased in the post-accident load combination to account for the. mass of the water flooding the drywell.o The ventlines were modeled down to the intersection with the ventline header. Here, springs acting in the radial and vertical directions were added to approximate the compli-ance of the ventline header. The spring constants were based on a simple submodel analysis of the ventline header. Since the ventline is connected to the torus with a flexi-ble bellow, all interaction between the ventline and torus was neglected.
o The ventline jet deflector was modeled as a solid plate. In reality, the deflector has mul-tiple holes throughout the plate. The thickness of the solid plate in the current model was reduced to account for the holes.* In a number of cases, the exact location that a specific load acts upon the drywell shell was not known. The magnitude and elevation of these loads were provided in the GE re-port, but the azimuth locations remained unknown. In these cases (mainly in the case of the penetration loads), the loads were distributed along the entire circumference of the drywell as a surface traction.* The loads applied to the drywell shell were "smeared" along a region defined on the shell surface. Typically, the region of application was taken as the area where an item is actu-ally attached to the shell in the real structure.
As mentioned above, the penetration loads were smeared along the entire circumference since loads for individual penetrations were not provided.* The spacing of the upper and lower beam seats around the circumference is not constant, but the appropriate load distribution at each seat was not known. The loads for the upper 81 and lower beam seats were distributed equally at each point of attachment to the drywell shell." The concrete that fills the drywell shell interior from an elevation of 8'-I 1.25" to 10'-3", and the additional curbs, have not been accounted for in this model. The drywell shell is assumed encased in concrete below elevation 8'-] 1.25" (bottom of sandbed)." For the accident load combination, the internal 44 psi pressure and the thermal load of 292"F (starting at 70TF) were applied to the entire drywell shell down to an elevation of 8'-11.25".
The concrete within the interior of the drywell shell extends up to 10'-3" with curbs extending up to 12'-3". Since the bond between the steel shell and the concrete is not known, it was assumed that a gap could exist which would enable gas to pressurize and heat the shell down to 8'-11.25".
or the bottom of the sandbed region on the exterior of the shell. Even if no gap exists initially, it is likely that the initial pressurization (pres-sure << 44 psi) acting on the shell above elevation l0'-3" would cause a gap to open.This would allow heated gas to flow between the shell and concrete." The Personnel Lock & Equipment Hatch penetration geometry (extent modeled and thicknesses assigned) was approximated and the outer surface fixed against vertical dis-placement.
The coefficient of thermal expansion for the A-212-61T Grade B pressure vessel steel used for the drywell was assumed to be 6.5E-6oF1.
A number of assumptions were made to develop the thicknesses assigned to the model in its degraded state. Section 2.6 provides a detailed discussion of these items.o A very limited mesh convergence study was performed which led to the use of a 4" nomi-nal element size. It was assumed that this mesh size was acceptable even though all load combinations were not examined in the convergence study and no checks on buckling were performed using different mesh sizes. In addition, a 1" nominal mesh size was used in the two local regions under the ventlines in Bay 1 and 13. No checks were performed to assess the mesh size in these regions.o Several assumptions were made in developing the ASME stress limits. These are dis-cussed in Section 3..o ASME Code Case N-284 was used to assess the stability of the degraded drywell shell.It was assumed that since the refueling case does not include any internal pressure, that the increase in buckling capacity used by GE for cases with circumferential (hoop) ten-sion was not appropriate.
Since the post-accident load case includes internal pressure, an increase in the capacity was applied.82
- 7. Conclusions The structural integrity of the degraded Oyster Creek drywell shell has been analyzed in this study. The allowable stresses and the buckling stability were both examined in accordance with the ASME B&PV code. The ASME allowable stresses are met for all three load cases examined here given the modeling and loading procedures outlined in Section 2. The only potential excep-tion is for the primary plus secondary stresses located at the base of the sandbed region of the accident condition due to the thermal expansion of the shell. There are a number of modeling and loading assumptions in this region that may contribute to the stress magnitudes recorded in the current analysis.
In addition, the primary plus secondary stresses were compared to the al-lowables use in the previous GE analysis (GE, 1991a). The current code does not require an evaluation of the primary plus secondary stresses for Service Level C. However, these stresses were assessed in this report to be consistent with the previous evaluation by GE. The buckling evaluation perfonned here using ASME N-284 show that based on the loadings and the model described in Section 2 both the refueling and post-accident load combinations met buckling re-quirements with a one exception.
The buckling at the upper beam seat for the refueling load case with degradation does not meet the required factor of safety of 2. As described in Section 4, the potential constraint provided by the attached beam has not been included in this analysis.
Table 7-1 summarizes the major conclusions for this study and for the previous GE analyses.Table 7-1. Comparison of Conclusion Between GE Study (GE, 1991a and b) and the Current Study Current Study Conclusion GE Study Conclusion The ASME B&PV stress analysis of the de- The ASME B&PV stress analysis of the de-graded Oyster Creek drywell shows all values graded Oyster Creek drywell shows all values within code limits. The current study uses within code limits. The GE study assumed a average UT measurement data to assign conservative uniform thickness of 0.736" in the thicknesses in the sandbed region. (Note that sandbed region.some primary plus secondary stresses for the accident condition are of concern as dis-cussed in Section 3.)ASME B&PV Code Case N-284 stability analy- ASME B&PV Code Case N-284 stability analysis sis of the degraded Oyster Creek drywell of the degraded Oyster Creek drywell shows shows that acceptable factors of safety are that acceptable factors of safety are met. The met. The current study uses average UT GE study assumed a conservative uniform measurement data to assign thicknesses in thickness of 0.736" in the sandbed region.the sandbed region. (Note that the buckling at the upper beam seats produces an effec-tive factor of safety slightly less than 2 for the refueling load case, but this may be affected by the modeling of that specific detail.)The minimum uniform shell thickness re- The minimum uniform shell thickness required quired to meet the ASME N-284 buckling to meet the ASME N-284 buckling safety factor safety factor was determined to be 0.844" in was determined to be 0.736". This thickness the sandbed region. This thickness was es- was established using the refueling load case.tablished using the buckling analysis for the (The thickness of 0.736" was established in a refueling load case. calculation by GPU Nuclear, 1993. This calcula-tion included an increase in the capacity reduc-tion factor not used in the current study.)83 The assessments performed here employ a uniform thinning of the drywell shell over large sec-tions of the surface. The thicknesses assigned in each region were based on limited measure-ment data since a very small percentage of the shell has been examined.
In many cases, the raw data was not available.
This led to the use of averages provided by AmerGen throughout the relevant documentation.
84
- 8. References ABAQUS Analysis User's Manual, Version 6.5. 2004. Pawtucket, RI: Hibbit, Karlsson & Soren-son, Inc.AmerGen April 4, 2006, Letter to U.S. NRC, 2130-06-20284, NRC Docket No. 50-219.AmerGen April 7, 2006, Letter to U.S. NRC, 2130-06-20289, NRC Docket No. 50-219.ASME (American Society of Mechanical Engineers).
2004. Boiler and Pressure Vessel Code.Sections III and VIII. New York, NY: American Society of Mechanical Engineers.
Cherry, J.L., and Smith, J.A., Capacity of Steel and Concrete Containment Vessels with Corro-sion Damage, NUREG/CR-6707, SAND2000-1735, U.S. Nuclear Regulatory Commission, 2000.Chicago Bridge & Iron (CB&I, 1980) Structural Drawings of the Oyster Creek Drywell, Torus, and Ventlines, attached to AmerGen letter to NRC, July 7, 2006, NRC Docket No. 50-219.Final Safety Analysis Report (Updated), Oyster Creek Nuclear Generating Station (FSAR), AmerGen, Exelon, Revision 13, April 2003.GE Stress Analysis Report, "An ASME Section VIII Evaluation of the Oyster Creek Dry'Well for Without Sand Case -Part I -Stress Analysis", GE Index #9-3, DRF #00664, February 1991 a.GE Stability Analysis Report, "An ASME Section VIII Evaluation of the Oyster Creek Drywell for Without Sand Case -Part HI- Stability Analysis", GE Index #9-4, DRF #00664, February 1991 b.GPU Nuclear, Calculation Sheet, Oyster Creek Drywvell External UT Evaluation, C-1302-187-5320-024, 1993.Individual Plant Examination Submittal Report, Oyster Creek, (IPE), June 1992.Johnson, B,G. editor, Guide to Stability of Design Criteria for Metal Structures, Third Edition, John Wiley & Sons, 1976.Spencer, B.W., Petti, J.P., and Kunsman, D.M., Risk-Informed Assessment of Degraded Con-tainment Vessels, NUREG/CR-6920, SAND2006-3772P, U.S. Nuclear Regulatory Commission, 2006.Standard Review Plan for the Review of Safety Analysis Reports for Nuclear Power Plants (SRP), NUJREG-0800, U.S. Nuclear Regulatory Commission, 1996.85