ML17296A205
| ML17296A205 | |
| Person / Time | |
|---|---|
| Site: | Monticello  |
| Issue date: | 10/18/2017 |
| From: | O'Connor T J Northern States Power Company, Minnesota, Xcel Energy |
| To: | Document Control Desk, Office of Nuclear Material Safety and Safeguards, Office of Nuclear Reactor Regulation |
| References | |
| CAC L25058, L-MT-17-053 | |
| Download: ML17296A205 (461) | |
Text
Document Control Desk Page 4 Enclosure 10 provides Applied Analysis Corp. Calculation MNGP-018, "Accident Dose Assessment for MNGP DSCs 11-15". This calculation determines the offsite dose assuming a non-mechanistic release from the DSC closure welds. Enclosure 11 provides Jensen Hughes Report 016045-RPT
-01, "Risk Assessment of MNGP DSCs 11-15 Welds Using N UREG-1864 1 Methodology".
This report compares the calculated risk of the alternative of leaving these casks, as-is, in their current stored location versus the alternative of transferring these casks back into the reactor building for inspection and then returning them to their storage locations.
Enclosure 12 contains an affidavit executed by AREVA. As the owner of the proprietary information submitted in Enclosure 9, AREVA certifies that the enclosed proprietary information has been handled and classified as proprietary, is customarily held in confidence, and has previously been withheld from public disclosure.
AREVA requests that the enclosed proprietary information be withheld from public disclosure in accordance with 10 CFR 2.390. NSPM requests the NRC grant the requested exemption by October 31, 2018, to support restoration of compliance with 10 CFR 72 and also to meet the requirements of the Confirmatory Order issued in Reference 1 . If there are any questions or if additional information is required, please contact Mr. Shane Jurek at (612) 330-5788.
Summary of Commitments r makes no ew commitments and no revisions to existing commitments.
J. 0 Con or Vice President and Chief Nuclear Officer ern States Power Company-Minnesota Enclosures (12) cc: Administrator, Region Ill, USNRC Rob Kuntz, Project Manager, Monticello Nuclear Generating Plant, USNRC Christian Jacobs, Project Manager, Spent Fuel Storage and Transportation, USNRC Resident Inspector, Monticello Nuclear Generating Plant, USNRC 1 NUREG-1864, "A Pilot Probabilistic Risk Assessment of a Dry Cask Storage System At a Nuclear Power Plant"
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Page 1 of 30F0306-01R2
Quality Program Type:
Nuclear Commercial Monticello ISFSI - DSC 11 through 16 Exemption Request 1005, Release 48, Amendment 6 Xcel Energy
Monticello Nuclear Generating Station Development of an Analysis Based Stress Allowable Reduction Factor (SARF)
- Dry Shielded Canister (DSC) Top Closure Weldments 0 1 - 30 A A-2 B B-7 Initial Issue Richard Bax 10/23/14 James W. Axline 10/23/14 Preparer:
Wilson Wong 10/23/14 Checkers: J. Wu 10/23/14 C. Fourcade 10/23/14 File No.:
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1.0OBJECTIVE .................................................................................................................. 42.0TECHNICAL APPROACH .......................................................................................... 52.1Finite Element Model and Flaw Simulation ...................................................... 53.0ASSUMPTIONS / DESIGN INPUTS ........................................................................... 64.0CALCULATIONS ......................................................................................................... 74.1Pressure Loading ............................................................................................... 74.2Side Drop Loading ............................................................................................. 85.0RESULTS OF ANALYSIS ...........................................................................................
86.0CONCLUSION
S AND DISCUSSION ....................................................................... 1
07.0REFERENCES
............................................................................................................ 12APPENDIX A ANSYS INPUT FILES ................................................................................ A-1APPENDIX B SI REPORT 1301415.405, REVISION 0, "EXPECTATIONS FOR FIELD CLOSURE WELDS ON THE AREVA-TN NUHOMS 61BTH TYPE 1 & 2 TRANSPORT ABLE CANISTER FOR BWR DRY FUEL STORAGE," .............................................................................. B-1
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Table 1: OTCP Stress Reduction Factor Results - Pressure Loading .................................... 13Table 2: OTCP Stress Reduction Factor Results - Side Drop Loading ................................. 14Table 3: ITCP Stress Reduction Factor Results - Pressure Loading ...................................... 15Table 4: ITCP Stress Reduction Factor Results - Side Drop Loading ................................... 16Table 5: OTCP and ITCP Deflection Load Cases - Pressure Load Case .............................. 17
Figure 1. Finite Element Model and OTCP and ITCP Details ............................................... 18Figure 2. OTCP Postulated Flaw Configuration - Radial #1 ................................................. 19Figure 3. OTCP Postulated Flaw Configuration - Radial #2 ................................................. 20Figure 4. OTCP Postulated Flaw Configuration - Laminar ................................................... 21Figure 5. OTCP Postulated Flaw Configuration - Circumferential #1 .................................. 22Figure 6. OTCP Postulated Flaw Configuration - Circumferential #2 .................................. 23Figure 7. OTCP Postulated Flaw Configuration - Circumferential #3 .................................. 24Figure 8. OTCP Postulated Flaw Configuration - Circumferential #4 .................................. 25Figure 9. ITCP Postulated Flaw Configuration - Circumferential ......................................... 26Figure 10. OTCP Pressure Load Case - Displaced Shape (Exaggerated) .............................. 27Figure 11. ITCP Pressure Load Case - Displaced Shape (Exaggerated) ............................... 28Figure 12. Side Drop Model ...................................................................................................
29Figure 13. OTCP and ITCP Stress Path Definitions ............................................................... 30 File No.:
Revision: 0 Page 4 of 30F0306-01R2 The objective of this calcul ation is to develop a quan titative basis for a stress allowable reduction factor (SARF) to address weld quality in the inner top cover plate (ITCP) and outer top cove r plate (OTCP) weldments of the NUHOMS dry shielded canister (DSC
) system. This workscope is in support of the USNRC CofC Exemption submittal for DSC's 11 through 16, currently at the Monticello Nuclear Generating Plant (MNGP).
Weld quality is described as a global effect, for which a factor is used to reduce the stress allowables to account for potentially less than sound weldments. Th e SARF has historically be en tied to the level of non-destructive examination (NDE) performed on the weld ment. That is to say, th e greater the degree of NDE performed (such as volumetric) the greater the SARF (less reduction in stress allowable).
The ASME Code [5, NG-3352] contains values for SARF for a range of NDE. Specifically, a VT only scope of NDE would state an SARF of 0.35 for a partial penetration weldment. However, it should be clearly noted that the ASME Code table for SARF's has no limitations/definitions/requirements on the weld size, the weld/base metal ma terials, the welding configuration, the welding position, and most importantly, the welding process. In addition, as this table is from NG, the level and comprehensiveness of the design analysis is less than that for an NB-type component, such as the DSC. The 0.35 SARF is a
conservative factor that addresses all types of welding. In the case of the DSC weldments, these are specific joint geometries, with high quality materials, favorable welding positions, and again, most importantly, a high purity welding processes (GTAW),
and therefore, strict adherence to the 0.35 SARF number for a VT only NDE examination weldment is not warranted.
The intent of this calculation, for this exemption request only, is to evaluate a series of postulated weld flaws and determine, for each configuration, the effect on the unflawed stress results. The effect of the stress results will be comparative, performed by comparing the analysis results of the flawed configuration to those from the same geomet ry, but in an unflawed configuration.
The determination of the impact on stress results will be performed by finite element analysis (FEA) in which selected elements of the ITCP and OTCP weldments will be "removed" to represent "flawed/suspect" weld quality.
Various distributions of flaw size (length and depth) and frequency (spacing), will be examined.
The intent of this calculation is to analytically determ ine the type of flaw distri bution that would justify a specific SARF. A separate work scope has been performed to evaluate, for the specific DSC weldments (DSC's 11 through 16), what are the e xpected type and density of flaw distributions. It is the overall intent for this project workscope that it can be sh own that the type of flaw distribution, which would support an acceptable SARF, will be of significantly greater magnitude than those populations that would be expected for the type of welding used for the DSC weldments.
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Revision: 0 Page 5 of 30F0306-01R2 The determination of the impact of weld quality on stress results (SARF) will be performed by the finite element methods. Both the flawed and unflawed geom etry of the top end of the DSC will be modeled. To represent the presence of postulated flaws, selected elements within the model will be removed and analyses performed using representative load cases. By comparing the results from the unflawed and flawed FE models for these loads cases, a ratio, or stress allowable reduction factor can be determined. A range of flaws will be analyzed to develop a range of SARF values corresponding to the range of flaw populations.
Typical types of flaws will be considered, and a range of distributions of flaw size (length and depth) and frequency (spacing), will be examined.
Three types of flaws will be addressed.
Radial: a postulated flaw oriented in a plane radial to the DSC longitudinal axis and spanning the weldment from cover plate to shell. Circumferential: a planar flaw oriented in a plane parallel to the DSC axis and oriented circumferentially around the DSC. Laminar: a planar flaw in a plane perpendicular to the longitudinal axis of the DSC and spanning the weldment from cover plate to shell.
In the determination of what flaw types to analyze in the OTCP and the ITCP, the size/volume of the weldment was considered. The OTCP weldment is both large in size and volume absolutely, and also relative, to the weldment volume of the ITCP. Therefore, all three types of flaws are evaluated for the OTCP. The ITCP weldment, due to its reduced weldment size, is evaluated using a single flaw of
significant cross-section, which represents elements of all three types. Figures showing these flaw types, location, and orientation are shown in Figures 2 through 9.
A single finite element model (FEM) is developed using the ANSYS finite element analysis software [2]. The model represents a 180° sector of the upper end of the DSC. The model includes the outer top cover plate and weldment, the inner top cover plate and weldment, and a portion of the DSC shell.
The FEM utilizes the ANSYS 3-D structural element (SOLID45). The unflawed model contains all portions of the two weldments.
The modeling of the postulated flaw is done by "killing" the selected elements that represent the flaw size and location, using the EKILL command in ANSYS. This command deactivates the element such that it contributes near zero stiffness to the overall stiffness matrix.
The result is a redistribution of loading and stresses around "killed" elements.
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Revision: 0 Page 6 of 30F0306-01R2 The ANSYS model of the top end geometry is shown in Figure 1 which illustrates the full model and then localized sections through the OTCP and ITCP.
The top end geometry of the DSC is defined in Reference 3. The OTCP, ITCP, and DSC shell dimensions, as well as the materials, are provided in Reference 3. A number of assumptions were made during development of the finite element model, which are listed as follows:
The model consists of a half-symmetric portion of the inner top cover plate (ITCP), outer top cover plate (OTCP), and the top 20 inches of the outer DSC cylinde
- r. The 20 inches equates to greater than 4.0Rt, thus avoiding any end affects at the free end constraint. The model is constructed of approximately 840,000 SOLID45 elements to ensure adequate mesh refinement for the ITCP and OTCP welds in the circumferential direction. The OTCP is modeled with the top surface set 1/8 of an inch below the end of the DSC. The J-groove weld preparation is as shown in Reference 3. The weldment is shown flush with the surface of the OTCP and not set below, as is allowed by the Reference 3 field assembly drawing. The modeled set back weldment is considered acceptable as this is a comparative analysis and the same geometries are used in bot h the flawed and unflawed condition. The ITCP is modeled as a flat plate and the closure weldment is modeled flush with the top surface of the ITCP. The DSC shell, the OTCP, the ITCP, and the OTCP and ITCP weldments are modeled as SA-240, Type 304 stainless steel. Material properties are taken from Reference 4. Standard room temperature material properties for T ype 304 stainless steel are used: Young's Modulus = 28.30E6, Density = 0.283 lbs/in 3, and Poisson's Ratio of 0.3. The analysis is performed at 70°F. This temperature is selected as this is a comparative analysis and both the unflawed and flawed runs utilize the same temperature. The bottom edge of the outer cylinder is fixed in the axial and circumferential directions, and symmetry boundary constraints are placed on the symmetry plane. For the side drop runs, the outer cylinder is released in the circumferen tial direction and is s upported at the point of "impact" via radial displacement couples to a support block with reduced stiffness properties. The analyses are all treated as elastic. The localized effects of the vent and siphon block and the ITCP weldment are not modeled. This is acceptable as the weldment connection to the V/
S block (1/4" groove) is similar to the majority portions of the ITCP weldment, and the intent is to determine the effects of global weld quality, not localized stress concentrations. The effect of stress discontinuity at the V/S block will be addressed by the design analysis which models this explicitly, and then uses the SARF to further modify the stress allowables. The siphon/vent port cover plates are not modeled as the nominal stresses (primarily due to pressure) are sufficiently low to accommodate extremely low SARF's. Assuming a 3/16" closure groove weld [3] on a nominal 2 inch diameter cover plate results in a weld shear stress of File No.:
Revision: 0 Page 7 of 30F0306-01R2 less than 500 psi. Thus even a worst case SARF of 0.10, would be acceptable given the nominal weld filler metal shear stress allowable of 0.6 Sm [5, NB-32 27.2] = 0.6 * ~16 ksi = ~9.6 ksi. Dimensions for the components are taken as the nominal. This is acceptable as this is a comparative analysis. The evaluated paths for which the stress results are extracted and used for comparison (flawed vs unflawed) are shown in Figure 13.
The determination of the SARF, as a function of weld quality (number and dens ity of postulated flaws), is performed using two load cases. The pressure load is the primary normal and off normal load for these weldments and consists of internal pressure app lied to the inner top cover and outer top cover. The specific definition and modeling details are desc ribed below for the pressure load case.
The drop load cases consist of a can ister end drop, a canister corner dr op, and a canister side drop. For this comparative analysis the canister side drop load case is utilized as it best re presents the behavior of the drop event (an event that is germane to the MNGP ISFSI DSC hardware configuration) and is a more easily evaluated/modeled condition. Th e side drop load case develops lo calized stresses along a line of contact similar to the corner drop. The specific deta ils for the side drop load case are described below.
The pressure loading consists of a nominal 100 psig internal pressure ap plied to the top cover plates. For evaluation of the ITCP (the nominal pressure boundary) weldment quality, the pressure is applied to the inside surface of the ITCP and the DSC shell, an d the contacting surfaces be tween the ITCP and OTCP are bonded with sliding capability using ANSYS contact elements to allow for load transfer from the ITCP to the OTCP. For the ITCP pressure analysis, CONTA174 and TARGE170 contact elements were used to prevent the ITCP from pe netrating the OTCP. In these cases the OTCP acts as a non-pressure retaining structural support for the ITCP. Figure 11 shows the displaced shape for the ITCP pressure load case.
For evaluation of the OTCP weldment quality, the pressure is applied only to the inside surface of the OTCP and the inside surface of the DSC. The ITCP and the weldment to the shell are both contained within this model and are not modeled as containing flaws, nor are they loaded by pressure. The intent of applying the pressure loading to the OTCP alone is to maximize the response of the OTCP-to-DSC shell weldment, as a result of postulated flaws with in the weld. Applying the pressure to the ITCP, which in turn will load the OTCP, will diminish the response of the OTCP, as there exists supplemental stiffness from the ITCP. Figure 10 shows the displaced shape for the OTCP pressure load case.
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The side drop loading case is evaluated as a static 75G load case in which the FEM of the DSC shell is oriented with the symmetry plane in the direction of the drop. For the side drop analysis, the same contact element types (CONTA174 and TARGE170) were used to prevent the ITCP from penetrating the DSC outer cylinder. These are not used for the OTCP weld prep-to-DSC shell potential contact region, as the area of potential contact is small relative to the OTCP weld size. To simulate the support of the transfer cask, the lower 20° of the DSC model is supported by a material which represents the stiffness of the transfer cask given that there is a difference in diameter between the
DSC and the transfer cask. In the transfer condition, the DSC is suppor ted within the Transfer Cask on thin guide rails, and the use of a le ssor stiffness support in the lower 20° degree region is representative. Again this is a comparative analysis and the intent is to show the effect of weld quality in the weldments in the most highly stressed area of contact, which is at bottom dead center. Radial displacement couples between the DSC and support block ar e used. Figure 12 shows the geometry of this load case.
The determination of the SARF for a given postulate d flaw population is performed by extracting the stress results from the unflawed geometry, and the flawed geometry for the specific load case. These stresses are extracted and linearize d along identical paths to capture the change in stresses due to the missing/flawed elements.
The comparison to determine the change in stress re sults, as a result of the postulated flaw population, typically compares the linearized membrane (P m) and membrane plus bending (P m + Pb) stress intensities for a path adjacent to the postulated flaw and at other regular spacings between the postulated flaws.
These discrete ratios are then combined to produce a weighted SARF for the weld flaw pattern.
Figure 13 shows the path locations an d orientations for the three types of flaws for which stresses are extracted.
In general the comparison of stress results is done by comparing linearized membrane (P m) and membrane plus bending (P m + Pb) stress intensities. However, in the case of the side drop event for the radial and laminar flaws, the high compressive stresses in all three principal stresses make the use of stress intensity not representative. In these cases, where all three principal stresses are compressive, and the resultant stress intensity is of lesser magnitude than the principal stresses, the resulting SARF's are unrealistic. In these cases the greater stress values of the three principal stresses are combined by SRSS and compared for the flawed and unflawed configuration.
An initial set of postulated flaw populations for the radial, circumferential and laminar flaw were developed and analyzed. Subsequent to initial runs, additional flaw populati ons for the radial and circumferential flaw cases were run. The specific geometry of th e flaw populations are shown in Tables 1 through 4, along with the resulting SARF's.
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Revision: 0 Page 9 of 30F0306-01R2 It should be noted that the intent of the calculation is to show a flaw population that is severe and thus demonstrate that large flaw populati ons (size, length, and density) can be tolerated, as the calculated SARF is acceptable. In the selection of the flaw population parameters, the depth of the flaws is typically set as a through-wall flaw
. Obviously, such a flaw would ha ve been unacceptable, and would have been identified by leak test examination. Howe ver, the intent of this calculation is to address structural capacity of the weldment, not confinement.
1 Thus the use of the through-wall flaw allows for a conservative determ ination of the SARF.
Table 1 documents the calculated SARF's for the OTCP weldment subjected to pressure loading. Table 2 documents the calculated SARF's for the OTCP weldment subjected to the side drop loading.
Table 3 documents the calculated SARF's for the ITCP weldment subjected to pressure loading. Table 4 documents the calculated SARF's for the ITCP weldment subjected to the side drop loading.
Table 5 presents the axial deflection at the centerline of the OTCP for the various flaw configurations analyzed for the pressure load case. The intent is to show that, as expected, the stiffness of the combined OTCP and ITCP is greater (less deflection) than the OTCP alone. This is the reason that the pressure
loading was applied to the OTCP alone, so as to maximize the deflection of the OTCP, and therefore challenge to the OTCP weldment. A review of the table shows that the cha nge in deflection of the OTCP as a result of the introduction of postulated flaws, in either the OTCP or ITCP weldment, is relatively low (< 15% in the worst case). Thus the evaluation of flaws does not require the explicit evaluation of concurrent flaws in the OTCP and ITCP, as their respons es (unflawed/flawed) are basically similar, and this is a comparative evaluation.
In addition, a comparisons of the deflections of th e OTCP in the unflawed and postulated flawed cases shows that for the less severe, but still significant flaw populations (Radial 2, Laminar, Circ 3, and Circ 4), the change in response (OTCP deflection) is small, typically 1%
or less. It can therefore be presumed that a mix of flaw types would produce similar results as that for a single flaw type, e.g. a mix of radial, laminar, and circumferential flaws would have similar results as that for the bounding single flaw type. The worst case SARF for the selected flaw types will be utilized, thus any substitution of lesser SARF flaws (e.g. laminar) for greater SARF flaws (Circ) would be bounded.
Finally, the postulated 50% circumferential flaw for Circ 4 is positioned in the upper half of the weldment. The change in SARF values (Tables 3 a nd 4) between the Circ 3 and Circ 4 cases is an increase of ~4% for the pressure case, and ~14%
for the side drop case.
A 50% through-wall flaw, located in the lower portion of the weldment, would have an SARF no worse than the Circ 3 case, and the Circ 3 case SARF, for both pres sure and side drop, is greater than 0.80. The placement of the 50%
through-wall flaw in the lower half of the weldment would thus not change the results to a point where the Circ 3 case would not be bounding.
1 The results demonstrate that the remaining ligaments of the DSC weldments have sufficient structural capacity, even with very severe and conservative penalties (postulated flaws) for nonconforming PT examinations, to perform their design function of restraining the OTCP and ITCP's, and additionally maintaining the confinement function during all service level load cases.
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Revision: 0 Page 10 of 30F0306-01R2 The OTCP and ITCP weldments are made using both materials and processes, and in conditions which would result in high quality (very small flaw distribution). Specifically it is a stainless steel weldment made with argon cover gas in a flat position using a machine GTAW process. As such, concerns over weld porosity are minimized and the machine welding process will produce a very uniform and consistent weldment. Report 1301415.405 [1, See Appendix B] details the expected flaw distribution for this type of weldment.
A review of Tables 1 through 4 documents the calcula ted SARF for the selected flaw populations. The question of which flaw population to consider representative or typical, or bounding is based not on these analytical results but on the separate Refere nce 1 report. This report is based on the actual elements of the OTCP and ITCP welding, and considers industry experience and ISFSI Vendor experience [1, See Appendix B].
Reference 1 states in the conclusion that:
Comparing this to the an alyzed flaw populations:
OTCP: Both the radial and laminar flaws are not representative of the circumferentially oriented flaw described above. However, in both cases
, the postulated flaws for these types are full thickness and full width, and thus would be considered more severe than a 1/8" thick, 25% total weld length flaw, with a width of one weld bead. As an example, the laminar flaw is the full width of the weld, and covers 72% of the circumferential arc.
The radial Configuration 2 flaw (more limiting), shown in Figure 3, is a full height (through-wall) flaw, spanning the full weldment width, and occurring less than 2" apart.
The circumferential flaw, Configuration 3, show n in Figure 7, is a full height (through-wall) flaw, 1" long and occurring every 5". The 1" in 5" spacing is a 20% occurrence of postulated flaws, which although less than 25%, is tempered by the fact that th e analyzed flaw is full height, not the expected one bead thickness dimensi on (~ 1/8") described above. With this consideration, the Configuration 3 circumferential flaw bounds the "conservatively assumed"
flaw stated in Reference 1.
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Revision: 0 Page 11 of 30F0306-01R2 ITCP: The 360 degree embedded flaw postulated and evaluated (Figure 9), is much more adverse than the expected flaw of Reference 1 described above.
In both the OTCP and the ITCP weldments, the weld is a multi-layer weldment, and both received multi-level VT and PT examinations. Although the PT cannot be credited, the VT can be assumed to have seen large surface breaking flaws. As a further argument that the postulated and analyzed flaws are bounding for flaws that would have not have been identified by the VT exams, the likelihood that
multiple through-layer thickness flaws of the postulated percentage of arc length (e.g. the Circ 3 case flaw covers 20% of the total arc length) would occur in every layer, and would also line up with flaws below and above to create a through-wall combined flaw, and not be detected by the multiple VT's, is highly unlikely and not realistic.
Again the use of through-wall flaws is done to evaluate the structural integrity of the weldments. The validation of confinement of the weldments was separately confirmed by su ccessful leak testing.
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Revision: 0 Page 12 of 30F0306-01R2 1.SI Report No. 1301415.405, Revision 0, "Expecta tions for Field Closure Welds on the AREVA-TN NUHOMS 61BTH Type1 & 2 Trans portable Canister for BWR Dry Fuel Storage," October 2014, SI File No. 1301415.405. [Appendix B]
2.ANSYS Mechanical APDL and PrepPost, Rel ease 14.5 (w/ Service Pack 1), ANSYS, Inc., September 2012.
3.AREVA Design Drawings for the 61BTH, Type 1 and 2, NUH61BTH-3000, Rev 1, "NUHOMS 61BTH Type 1 DSC Main Assembly," and NUH61BTH-4008, Rev 1, NUHOMS 61BTH Type 1 & 2 Transportable Canister for BWR Fuel Field Welding, PROPRIETARY SI File No. 1301415.201P.
4.ASME Boiler and Pressure Vessel Code,Section II, Part D, Materi al Properties, 2004 Edition.
5.ASME Boiler and Pressure Code,Section III, Division 1, Rules for Construction of Nuclear Facility Components, 2004 Edition.
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Pm Pm+Pb(I) Pm+Pb(O) Pm Pm+Pb(I) Pm+Pb(O) Pm Pm+Pb(I) Pm+Pb(O) Pm Pm+Pb(I) Pm+Pb(O) Pm Pm+Pb(I) Pm+Pb(O) Pm Pm+Pb(I) Pm+Pb(O) Pm Pm+Pb(I) Pm+Pb(O) Average 0.908 0.762 0.900 0.955 0.879 0.973 0.911 0.911 0.950 0.515 0.534 0.436 0.759 0.771 0.703 0.924 0.920 0.888 0.940 0.956 0
.919 MIN Through Wall Flaw Through Wall Flaw Through Wall Flaw Through Wall Flaw Through Wall Flaw Through Wall Flaw 50% Part Through Wall Flaw Pattern Arc Spacing (in) 0.864 Pattern Arc Spacing (in) 1.734 Pattern Arc Spacing (in) 5.760 Pattern Arc Spacing (in) 5.184 Pattern Arc Spacing (in) 5.184 Pattern Arc Spacing (in) 5.184 Pattern Arc Spacing (in) 5.184 Flaw Width (in) 0.144 Flaw Width (in) 0.144 Flaw Arc Length (in) 4.176 Flaw Arc Length (in) 3.600 Flaw Arc Length (in) 2.016 Flaw Arc Length (in) 1.012 Flaw Arc Length (in) 1.012 Un-Flawed Arc Spacing (in) 0.720 Un-Flawed Arc Spacing (in) 1.590 Un-Flawed Arc Spacing (in) 1.584 Un-Flawed Arc Spacing (in) 1.584 Un-Flawed Arc Spacing (in) 3.168 Un-Flawed Arc Spacing (in) 4.172 Un-Flawed Arc Spacing (in) 4.172
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Revision: 0 Page 14 of 30 F0306-01R2 Pm Pm+Pb(I) Pm+Pb(O) Pm Pm+Pb(I) Pm+Pb(O) Pm Pm+Pb(I) Pm+Pb(O) Pm Pm+Pb(I) Pm+Pb(O) Pm Pm+Pb(I) Pm+Pb(O) Pm Pm+Pb(I) Pm+Pb(O) Average 0.976 0.921 0.912 0.882 0.957 1.000 0.542 0.606 0.762 0.720 0.756 0.903 0.846 0.861 0.972 0.979 0.974 0.974 MIN Through Wall Flaw Through Wall Flaw Through Wall Flaw Through Wall Flaw Through Wall Flaw 50% Part Through Wall Flaw Pattern Arc Spacing (in) 0.864 Pattern Arc Spacing (in) 5.760 Pattern Arc Spacing (in) 5.184 Pattern Arc Spacing (in) 5.184 Pattern Arc Spacing (in) 5.184 Pattern Arc Spacing (in) 5.184 Flaw Width (in) 0.144 Flaw Arc Length (in) 4.176 Flaw Arc Length (in) 3.600 Flaw Arc Length (in) 2.016 Flaw Arc Length (in) 1.012 Flaw Arc Length (in) 1.012 Un-Flawed Arc Spacing (in) 0.720 Un-Flawed Arc Spacing (in) 1.584 Un-Flawed Arc Spacing (in) 1.584 Un-Flawed Arc Spacing (in) 3.168 Un-Flawed Arc Spacing (in) 4.172 Un-Flawed Arc Spacing (in) 4.172
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Revision: 0 Page 15 of 30 F0306-01R2 Pm Pm+Pb(I) Pm+Pb(O) 0.964 1.000 0.954 Flaw Cross Section Area 0.006 in2 Pattern Arc Spacing (in) 5.184 Flaw Arc Length (in) 2.590 Flaw Arc Spacing (in) 2.590 File No.:
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Pm Pm+Pb(I) Pm+Pb(O) 1.000 0.931 1.000 Flaw Cross Section Area 0.006 in2 Pattern Arc Spacing (in) 5.184 Flaw Arc Length (in) 2.590 Flaw Arc Spacing (in) 2.590 File No.:
Revision: 0 Page 17 of 30 F0306-01R2 OTCP Radial 1 0.9089 0.921 1.3% Radial 2 0.9089 0.9149 0.7% Laminar 0.9089 0.918 1.0% Circ 1 0.9089 1.0391 14.3% Circ 2 0.9089 0.9507 4.6% Circ 3 0.9089 0.9208 1.3% Circ 4 0.9089 0.9169 0.9% ITCP Circ 0.629 0.6314 0.4%
Note: 1) The deflection value was taken at the center top of each plate.
l)swotut&llnlsgrtty Assoclat9s, Inc.* Inner Top Cover Plate (ITCP) Outer Top Cover Plate (OTCP) Figure 1. Finite Element Model and OTCP and ITCP Details File No.: 1301415.301 Revision:
0 STORED IN HARD COPY Page 18 of30 F0306-0IR2
Integrity Associates, Inc.* File No.: 1301415.301 Revision: 0 ___ Postulated Radial Flaw Figure 2. OTCP Postulated Flaw Configuration-Radial #1 STORE IN HARD COP Page 19 of30 F0306-0IR2
()stnrahnl Jntqrtty Associat9s, Inc.* OTCP Postulated Flaw Configuration-Radial #2 File No.: 1301415.301 Revision:
0 STOREDI HA COP Page 20 of30 F0306-0IR2 eStluclznlllllegrlly Assoclatss, Inc.* ----Postulated Laminar Flaw Pressure: l.a;ilinar Flaw Figure 4. OTCP Postulated Flaw Configuration-Laminar File No.: 1301415.301 Revision:
0 STORE I HAR 0 Page 21 of30 F0306-0IR2
--Postulated Circ Flaw Pressure:Circl I Figur*e 5. OTCP Postulated Flaw Configuration
-Circumferential
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Inc.* Postulated Circ Flaw Figure 6. OTCP Postulated Flaw Configuration
-Circumferential
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,__ __ Postulated Circ Flaw Figure 7. OTCP Postulated Flaw Configuration-Circumferential
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Associalls.
Inc.* sn=-t Sil*l 'TI!£-1 m; -. 90930'1 Pressure:No Crack Figure 10. OTCP Pressure Load Case-Displaced Shape (Exaggerated)
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(;Stnlclulal
/nl.sfrlty Associates, Inc.* Drop Direction Figure 12. Side Drop Model File No.: 1301415.301 Revision:
0 A Page 29 of30 F0306-0IR2 lJstntotuTBIIRIIgrtty Associams, Inc!' OTCP Radial Flaw Stress Path OTCP Circumferential Flaw Stress Path OTCP Laminar Flaw Stress Path JTCP Stress Path _j Figure 13. OTCP and ITCP Stress Path Definitions File No.: 1301415.301 Revision:
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Revision: 0 Page A-2 of A-2F0306-01R2 Base_Model.INP ANSYS input file to construct the 3-dimensional model. C*_$_%.INP ANSYS input file to perform OTCP flawed stress analyses * = 1-4 (Case Number)
$ = Side, Pressure (Loading)
% = Radial, Circ, Lam (Flaw Direction)
I1_$_CIRC.INP ANSYS input file to perform ITCP flawed stress analyses
$ = Side, Pressure (Loading) Pressure.INP ANSYS input file to perform OTCP non-flawed pressure stress analyses.
Side.INP ANSYS input file to perform OTCP non-flawed side drop stress analyses. I1_Pressure.INP ANSYS input file to perform ITCP non-flawed pressure stress analyses. I1_Side.INP ANSYS input file to perform ITCP non-flawed side drop stress analyses. Genstress.mac Macro to perform linearized path stress extraction. Lin_out.mac Macro to perform linearized path stress extraction using the native ANSYS PRSECT command.
GETPATH.TXT Path listing for stress extraction. Data.xlsm Excel file to compile stresses and compute ratios.
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Associat9s, Inc.* e Structural Integrity Associates, Inc.* October 23, 2014 Report No. 1301415405.RO Quality Program: Nuclear D Commercial 1v1r. James F. Becka Xcel Energy Project Supervisor-2013 DFS Loading Campaign Monticello Nuclear Generating Plant 2807 W. Country Road 75 Monticello, MN 55362 11515 Vmslory Dnve SLJie 125 Hmlersvtlle, NC 20078 Prone Fax www.strtdintccm rsmith@slrt.etirtcom
Subject:
Expectations for Field Closure Welds on the AREVA-TN NUHOMS 61BTH Type 1 & 2 Transportable Canister for BWR Dty Fuel Storage
References:
- 1. Xcel Energy Contract No. 1005, Release 48, Amendment
- 6. 2. SI Report 1301415402 RO, "Review ofTRIVIS INC Welding Procedures used for Field Welds on the Transnuclear NUHOMS 61BTH Type 1 & 2 Transportable Canister for BWR Fuel", January 30, 2014 3. SI Report 1301415403 R2, "Assessment ofMonticello Spent Fuel Canister Closure Plate Welds based on Welding Video Records",
May 2014 4. "E-mail train on Questions Regarding Postulated DCS Welding Flaw Distrubutions.pdf, from Peter Quinlan to Dick Smith, October 10, 2014, SI File No. 1301415.205.
- 5. Repair Rates in Welded Construction
-An Analysis of Industry Trends, TWI, Cambridge!UK, Welding and Cutting, November 2012, SI File No. 1301415.204.
Dear 1v1r. Becka:
Details of the machine gas tungsten arc welding ( G TAW) field closure welds used on the NUHOMS 61BTH transportable dry shielded canisters (DSC) located at Xcel Energy's Monticello Nuclear Generating Plant (tvfNGP) have been reviewed in an attempt to perform a qualitative assessment of the likelihood that the welds might contain unacceptable defects.
It is known that the required NDE acceptance testing was not performed according to approved procedures.
Sequential dye penetrant (PT) examinations were required on the inner top cover plate weld -first after the weld root and hot pass( es) were completed and again, after the final weld layer was completed.
This is a relatively small weld (3/16 inch partial penetration weld) and it was not required to perform an intermediate inspection.
The second weld is a 1/2 inch partial penetration weld that requires a root, intermediate, and final PT inspection due to the -----------------
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Revision: 0 Page B-3 of B-7F0306-01R2 eSintolu1'8/lllllfrlty Associat9s, Inc.* James F. Becka ReportNo. 1301415.405.RO October 23, 2014 Page 2 of6 larger size. The problem identified was that the dwell times used for both dye penetrant and developer were less than required by procedure.
The PT tests were performed, but procedures were not followed. This point is being emphasized because large open defects are seen very quickly with PT testing and likely would have been identified even though the dwell times were too short to meet procedure. Smaller tight defects might have been missed as the dye requires sufficient dwell time to wick and then be pulled out via the developer. This statement is in no way intended to justify the failure to follow approved PT procedures, but rather to apply perspective from a qualitative sense. There are a number of reasons to believe that the field closure welds in their current condition do not contain large discontinuities that could challenge the effectiveness of the closure welds to meet their intended design fimction.
It is the purpose of this review, performed in accordanc e with Reference 1, to identify valid reasons to support this conclusion.
A qualitative justificati on is provided that is outlined in the listing below: Reasons to expect the subject spent fuel canister welds are free from large discontinuities: 1. Use of qualified and proven welding procedures and techniques
. [Reference 2] 2. Use of a machine GTAWprocess. [Reference 2] 3. Application of a proven and robust welding system designed specifically to support these types of field welds in these specific types of canisters. [Reference 5] 4. Use of ductile and easily weldable base materials (SA-240 Type 304 stainless steel). [Reference 2] 5. Use of solid wire filler metal designed for welding these base materials and formulated to eliminate hot cracking and other types ofmicrofissures (SFA 5.9 ER308 austenitic stainless steel flller metal and welding grade gases for shielding the weld puddle. [Reference 2] 6. Canisters are oriented in the vertical position during welding such that the weld is performed in the flat welding position (the most forgiving welding orientation
). [References 2,3 and 4] 7. Weld roots are typically about 118 inch or slightly thicker which is good practice for GTA W machine welds. [Reference 4] 8. Weld layers are thin (between 1116 inch and 1/8 inch) requiring multiple layers (and multiple weld passes) to assist with developing weld deposit consistenc y via remelting. Layers become thinner as the groove is filled because the width is greater.
rReference 4] 9. AREVA-TN's historical record with these welds is excellent having a significant history of welds made with this system and these welding procedures that shows 1% repairs rates. [Reference 4] The welding procedures and welder control documentation were reviewed in detail and specifics of that review are reported in Reference 2. The review concluded that " ... the procedures the GTAW welds in the subject spent fuel canisters can reasonably be expected to be of good quality and free of injurious defects.
The expectation was based on the characteristic s of the G TAW weld, the excellent controls outlined for the welding program, and the fact that the welds and base materials are austenitic stainless steel. Also the welding consumables are compatible with the strnctural materials used in the design .... " [Reference 2] e Stntcturallntegrlly Associates, Inc.*
l)swctutWIInlefrlty Associates, Inc.* James F. Becka Report No. 1301415.405
.RO October 23,2014 Page 3 of6 The welding application
_ itself is performed entirely in the flat position-a welding position that eliminates any complications related to welding out of position or having to negotiate restricted access. The reason for this viewpoint is that out of position welds have to compete against the forces of gravity and the joint design provides adequate access for arc manipulation The result of a welding in the flat position is that defects are less likely to be introduced than might be expected with other weld orientations or restrictions.
The spent fuel canister welding system is robust and is proven The welding head is mounted on a non-metallic shielding material weighing over 1500 lbs and is shown in Figure 1 below. Figure 1 Photo of the robust welding head that ls positioned on the dry storage cask as shown In Figure 2. The welding torch Is visible In the photo just behind the rope. (Photo provided by AREVA-TN)
Associates, Inc.* File No.: 1301415.301 Revision:
0 A Page B-4 of B-7 0 F0306-0IR2
(;swcMIIIIIegrtty Associates, Inc!' James F. Becka Report No. 1301415.405
.RO October 23, 2014 Page 4of6 Figure 2 Welding system positioned on the storage for \Vel ding (Photo provided by AREVA-TN)
The' entire welding system rotates similar to a "lazy-susan
" and the welding torch is manipulated in and out as required for proper positioning.
There are other torch adjustments such as tilt, lead, height, etc. Leading and trailing cameras are mounted to provide video of the front and rear of the torch and weld puddle. Welding videos have been reviewed
[Reference 3] in an attempt to assess whether or not weld quality could be assessed One objective of the video review was to look for key discontimrities such as porosity and evidence for any lack of fusion The conclusions from the video review were that circumstances were observed at various times during welding that might support the generation of defects such as oxide buildup, weld root b:um-tbru.,
localized contamination on the surface, weld deposit surface irregularities, and tungsten drift requiring realignment.
However, nothing could confirm either the generation of defects or the lack of defects.
Since each weld is a unique entity one must rely on tendencies or trends if post weld inspections are not available.
There were also observations of good welding practices as well as.those events stated above. These included root repair, periodic adjustment of tungsten positioning, tungsten electrode replacement, electrode steering as needed, etc. Most of the videos were very similar (all having the same types of observations at about the same frequency)
. Canister No. 16 iuso had the same types of observations but the frequency appeared to be aboUt twice any of the others. This was a judgment call by the reviewers and not quantitative
. It was carefully pointed out that even so, there was no evidence to indic-ate that any specific discontinuities were generated-only that welding conditions were observed that sometimes lead to the various types of discontinuities
. In addition, since these welds use multiple weld beads to complete the weld, there is the opportunity to "heal" conditions created by welding over them. Historical Perspective AREVA-TN was asked to describe their historical perspeCtive on the welding of the canisters with this system. It is recognized that all of the canisters were not welded by AREV A-TN but File No.: 1301415.301 Revision:
0 STOR I A p Page B-5 ofB-7 F0306-0IR2 File No.:
Revision: 0 Page B-6 of B-7F0306-01R2 eSintolu1'8/lllllfrlty Associat9s, Inc.* James F. Becka ReportNo. 1301415.405.RO October 23, 2014 Page 5 of6 might include a contractor or the utility themselve
- s. However the same welding system likely would have been used (often rented from AREVATN).
AREV A-TN noted that typical discontinuities might include local porosity (rare), occasional tungsten inclusions, usually resulting from torch tip contact with the solidifying weld puddle, lack of fusion or overlap.
Regarding the potential for any linear indications (holidays or breaks), cracking typically does not occur with austenitic stainless welds. Maximum size of indications typically would be less than 1" to 2". Irregularities at starts and stops can occur, and rollover has been seen in some cases. AREVA-TN also was asked for their historical experience regarding canister closure weld acceptance rates (i.e. first time PT rate). The response indicated that a best estimate would be less than 1 UN SAT PT per 10 canisters, with an average of 10 PT examinations per canister (includes root and final layer on inner top cover, vent port cover, siphon port cover and test port, with root, mid and final layer on outer top cover for certain DSC models). Therefore, the historical experience suggests a rate of about 1% UN SAT PTs for field closure welds. Further, the recent field experience as the welding process matured produced no weld repairs at all -on SO+ canisters the findings were 1 PT indication from starts and stops was found to hold developer, but light grinding was performed to smooth the surface and eliminated the indication.
Thus, these minor indications required no weld repairs.
AREVA-TN was also asked regarding how many stainless steel canisters have been loaded and closed by welding to date. The estimate was for approximately 750 loaded/closed NUHOMS canisters, with closure performed by AREVA-TN, end user or other contractor.
This represents an extensive sampling that indicated an indication rate ofless than 1% and that rate appeared to significantly improve over the last 50 that have been welded. There were no applicable mockups that had been used to examine for discontinuities or defects, so that information was unavailable. The historical evidence seems to paint a favorable picture lending a degree of comfort that the canisters in question at MNGP are not likely to have indications of a significant size. Finally, literature was examined to find information regarding generation of defects in stainless steel weldments
. The best paper found is indicated in Reference 5. This paper written by The Welding Institute in Cambridge, UK was published in Welding and Cutting, November 2012. The paper titled "Repair Rates in Welded Construction-An Analysis of Industry Trends" provided good insight.
More than 800 professionals were contacted with about 10% responding. There were different kinds of responses such as % of welds requiring repair or % weld lengths requiring repair being the most prevalent. The following applicable conclusions were noted. GT A W stainless steel welds returned under 2% repair rates. The impact of different welding factors were parsed and suggested the following impacts:
root repairs at 22.5%, fill layers 7.5%, joint type 15%, access limitations 26%, and other welding factors 11% . Most of these are not present in the canister welds as pointed out previously.
It appears that the ARE VA-TN canister weld repair experienc es are slightly lower, but nevertheless are considered consistent with industrial expectations for a variety of manufactur ed and installed components. Since all welding is in the flat position using a proven welding system, the 1% defect rate appears to be reasonable. In addition it was pointed out that experience with the past 50 canisters has been even better. e Stntcturallntegrlly Associates, Inc.*
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Revision: 0 Page B-7 of B-7F0306-01R2 l)swalurallllllgrtty Associatfls, Inc.* James F. Becka Report No. 1301415.405
.RO Conclusions October 23, 2014 Page 6 of6 Based on the sum of the infonnation reviewed, it can be said that the likelihood for the occurrence of large defects is not supported by historical evidence.
While there remains the potential for long lack of fusion defects either interbead or sidewall, the thin multilayer design and potential for subsequen t bead healing by remelting would significantly limit the thickness dimension of any long defect. In fact, the most likely lack of fusion indication(s) would be intcrmiuent in nature and not expected to have a through-thickness dimension greater than one weld bead. While a quantitative estimate of a limiting naw size cannot be produced, the qualitative likelihood that large defects would not be present is assuring. It is suggested a bounding subsurface defect condition is conservatively represented as an intennittent lack of fusion (LOF) defect evenly distributed along the canister weld. Further, the total length for LOF is conservatively estimated at 25% of the canister cover plate weld circumference.
The estimated through thickness dimension is 1/8 inch. because this dimension represents a maximum weld bead thickness.
One eighth inch is considered to be a conservative assumption. because it is recognized that most weld beads will be thinner especially as the weld cavity begins to fill. No credit is being taken for remelting even though remelting is normally associated with multipass welding."
Very truly yours, Richard E. Smith. PhD. FA WS Senior Associate res
Report No.: 700388.401 Revision:
1 Project No.: 1700388.00 August 2017 Purchase Order:
67027
WelderIDFillerHeatDSCNo.WeldNo.1111821314355 737880736908Spool527221Spool111xxxx 114xxxxxxxx 121xxxx124 xxx 131xxx 134xxxxxxxx141xxx 144xxxxx 151xxx x154xxxxxxx161xxx x164xx x
Photo of welding arc about 10 inches further along the wdd length displaying non-unifOffil filli.ng of the trough under the eleotrode an:. Notice the molten filler metal dropped into the trough \vifuout fusion_ Fusion of bead #4 tbe shell sidewall is seen along the top of the bead, but less so wtth the previously deposited bead #3 and in the bottom of the trough where a drop of molten metal is seen in the trough but not fused. This conrution can. insulate the sidewall of wdd bead #3 and make jt difficult to fuse the trough interbead sidewalL This condition promotes "interbead LOF'. Distance stated above is baserl on the timestamp v.'ith an assumed travel speed of3-i:nch tntin. Innnediatd y foUowing the photo above shows a srufa.ce ,condition
. \vhere the electrode position for the bead #3 deposit was adjusted abruptly towards the shell wall. This action changed s:ur:face topography on which bead #4 was deposited (see bottom of photo). 'This \Velding feature i.s seen as contributing to the intermittent distribution of i:nterbea.d LOF defects along the olosure weld. S} Structural Integrity Associates, Inc.
Report No.: 1700388.401.R1 Page C-1 Report No.: 1700388.401.R1 Page C-2Weld3/16"highX0.31"wideWeldLength(OTCP66.25xPI=208")&(ITCP=197")SiphonVentBlock80.5"to91.5"locationofWeldstarts)WeldLocationVIDFileNameDateTimeStartTimeEndDuration(min)Length(in)TS(in/min)compLayerNo.TungstenBiasITCPtoShell0110/16/20139:24:1910:29:4265.51973.01(Root)LidITCPtoShell02and0310/16/201311:14:4812:32:08721972.72ShellOTCPtoShell0310/17/20139:39:4010:50:18702083.01(Root)LidOTCPtoShell0510/17/201311:50:3212:56:10662083.22ShellOTCPtoShell0610/17/201313:01:5714:10:25682083.13LidOTCPtoShell0910/17/201314:47:3015:46:40592083.54ShellOTCPtoShell1110/17/201315:52:1116:49:4157.52083.65LidDSC16(VIDsInnerandOuter16)1700388MonticelloSpentFuelDryStorageCask(DSC)VIDRecordsITCPisInnerTopCoverPlate0.75"ThickOTCPOuterTopCoverPlate1.25"ThickWeld5/8"highX1.0"wideComments6minlosttoadjustmentsofwirefeedon02series.
Report No.: 1700388.401.R1 Page C-3WeldLocationVIDFileNameDateTimeStartTimeEndDuration(min)Length(in)TS(in/min)compLayerNo.TungstenBiasOTCPtoShell039/14/201316:28:3117:36:46692083.01(Root)LidbiasatrootOTCPtoShell069/14/201318:56:0919:03:287213.02stripLidOTCPtoShell079/14/201319:07:2119:34:5327.582.53.02stripLidtoShellOTCPtoShell089/14/201320:11:3120:12:37133.02aLidOTCPtoShell099/14/201320:17:1420:17:410.51.53.02bLidOTCPtoShell109/14/201320:19:3820:41:3022663.02cLidOTCPtoShell119/14/201320:44:0321:00:18441323.02dLidCompletedbead.OTCPtoShell159/14/201322:30:0023:07:57381143.03ShellOTCPtoShell169/14/201323:12:0723:32:3720.561.53.03LidOTCPtoShell179/14/201323:37:4223:45:227213.03LidOTCPtoShell009/16/20132:02:392:21:4419573.03stripShellWelddepositionappearedtoimproveOTCPtoShell029/16/20132:31:333:05:0433.5100.53.03stripLidWelddepositionappearedtoimproveOTCPtoShell039/16/20134:16:545:25:07682043.04ShellWirefeeddraggingandarcadjustmentOTCPtoShell049/16/20135:46:186:52:07661983.05LidWelddepositionappearedtoimproveDSC12(VIDs1212and464956_12)CommentsTopofRootSurfacewasgroundsmoothatleastinplaces.Blowthroughofroot06_00009WeldapparentlyweldedatlowerpowerandstartedweavingacrossgrooveGrindingthenmovedarctolidsideforshortrepairofdeposit.Wireentryatwrongplaceandhadtoshutdownquickly.Oscillatingatplatewall(0.15or0.20)blowholeatendonrootseepicTapestoppedforunknownreasonStoppedweldforpossiblewireguideorweldheadadjustmentStoppedweldforpossiblewireguideorweldheadadjustment Report No.: 1700388.401.R1 Page C-4TS Durat (in/ Weld VID File Length min Layer Tungsten Date Time Start Time End ion Comments Location Name (min) (in) ) No. Bias com p DSC-13 (VID 464956_13) -, I Small oscillation in root pass at start, soon OTCP to Shell 01 09/24/2013 14:37:49 14:46:29 8.5 25.5 1 Lid stopped osc. Gap widene-d 8 min into root (picture capture-d for record. Blow out at 14:46:29 and stoppe*d to repair. , Second try to repair. Manually adjuste-d TS to add more wire and procee-de-d using oscillation to push OTCP to Shell 03 09/24/2013 14:53:01 15:13:50 21 63 1 Lid filler metal from lid to gap. Photo record in file. Oscillation stoppe-d as groove tightene-d about the 15 :00:37 mark. Lava flow can be seen f rom mirror reflection position on molten material. , Restarte-d to repair blowout. Oscillation continue-d OTCP to Shell 04 09/24/2013 15:15:21 15:31:51 16.5 49.5 1 Lid with bias on lid side. Another blowout at 15:31:51. Appears to be burning through weld land on the lid side. , OTCP to Shell 05 09/24/2013 15:37:08 15:37:43 0.5 L5 1 Lid Restarte-d to repair blowout but burne-d through at the same location being repaire-d. , Restarte-d to repair blowout. Oscillation continue-d OTCP to Shell 07 09/24/2013 15:59:18 16:02:38 3 9 1 Lid with bias on lid side. Another blowout occurre-d at 16:02:38 , Restarted to repair blowout. Oscillation continue-d OTCP to Shell 09 09/24/2013 16:23:49 16:46:43 23 69 1 Lid with bias on lid side. Steppe*d back about 12 inches and reinitiate
<l arc with the same technique. , OTCP to Shell H 09/24/2013 16:52:50 17:16:45 23 69 1 strip Shell Applie-d a wider oscillation but shifte-d to center on the shell side. OTCP to Shell " 13 09/24/2013 17:36:25 18:05:14 28.5 85.5 2 Shell No asci II atio n , OTCP to Shell 14 09/24/2013 18:09:41 18:38:39 29 87 2 strip Lid Covering the 2-3 overlap. Appears to be clean. Having intermittent camera issues. OTCP to Shell " 16 09/24/2013 19:48:18 20:33:33 45 135 t 2 Shell Surface grinding prior to welding. , 09/24/2013 OTCP to Shell 18 20:49:39 21:35:53 50.5 151.5 3 Lid Begin new layer. No oscillation e Structural Integrity Associates
, Inc.
Report No.: 1700388.401.R1 Page C-5WeldLocationVIDFileNameDateTimeStartTimeEndDuration(min)Length(in)TS(in/min)compLayerNo.TungstenBiasOTCPtoShell2009/24/201321:58:1122:42:08441324ShellOTCPtoShell2109/24/201322:53:1222:53:35N/AStripfillOTCPtoShell2209/24/201322:58:4423:46:54471414ShellOTCPtoShell2509/25/20130:31:191:22:48511534ShellOTCPtoShell2809/25/20131:34:232:38:40641925LidITCPtoShell0109/23/201312:08:3013:18:38701972.81LidITCPtoShell0209/23/201314:36:1315:42:09661972.92ShellCommentsDSC13(VID464956_13)cont.Initiallycenteredoncrownofrootpassbutweldmetalflowwasseentolapoverthetoponthelid.Initiallytheflowwasminimalonshellsideandtungstenrepositionedtoshellside.Lavaflowcoverssidetoside.Photocapturedshowslavaflowwellaheadoftungstenonlidside.GoodbitofdriftingofthetungstenBeginnewlayer.NooscillationWeldransmoothlyWeldappearedtohavebeenground,weldsteppedbacktocorrectfillinpatter,thentheweldappearedtorunsmoothlyWeldransmoothlyRansmoothlyBeginnewlayer.Nooscillation Report No.: 1700388.401.R1 Page C-6WeldLocationVIDFileNameDateTimeStartTimeEndDuration(min)Length(in)TS(in/min)compLayerNo.TungstenBiasOTCPtoShell0210/2/201312:49:4914:00:32712082.91LidOTCPtoShell0510/02/201315:43:3616:37:19542aCenterOTCPtoShell0610/02/201316:39:1016:54:291520832bCenterOTCPtoShell0910/02/201317:11:1817:54:1443129est.3ShellOTCPtoShell1110/02/201318:06:5019:09:51632083.34aLidOTCPtoShell1310/02/201320:11:2520:54:1343129est.5aShellOTCPtoShell1510/02/201321:01:0621:48:3947.5142.5est.5bLidDSC14(VID464956_14)CommentsCenteredonTackthenbiastolid.Wireappearstohaveagoodbitofcastcausingittowanderfromsidetoside.Nostopsforentirerun.Beginoscillatingfromcentertowardsshell.Wireentryadjustementsrequired.Maybeleavingintermittentdefectsontheshellsidewallatlayer2level.Willbedifficulttomeltinonthefourthbead.OscillationdugintotheshellsidewallatframeO5_00119.Photorecordmade.Repositionedtungstenandcompletedbead.Smoothrunonlayer3.Weldpassfilledinthearcgougeontheshellsidewallfromthepreviouspass.(likelypartialcircumferencetoaddressanincompletelyfilledlocation)Thereareintermittentdeeppockets(limitedlengths)thatmayresultinintermittentshortdefectsbetweenbeads3and4shortlyafterstartofthe4thbead.SimilartotheDSC16conditions.SurfacegroundinsomeplacestobeginSteppedovertotheLidsidetofillinincompleted4thbead.TheentireDSC14seemedtostripbeadstoattempttocorrectincompletefillingofthegrooveweldingsequence.
Report No.: 1700388.401.R1 Page C-7WeldLocationVIDFileNameDateTimeStartTimeEndDuration(min)Length(in)TS(in/min)compLayerNo.TungstenBiasOTCPtoShell0010/11/201303:04:483:48:374413231CentertoLidOTCPtoShell0110/11/201304:10:0904:35:54262083.01CentertoLidOTCPtoShell0210/11/201306:04:0907:13:27692083.02ShellOTCPtoShell0310/11/2013NoWeldingOTCPtoShell0410/11/201307:30:3908:37:26672083.13LidOTCPtoShell0510/11/2013NoWeldingOTCPtoShell0610/11/2013NoWeldingOTCPtoShell0710/11/2013NoWeldingDSC15(VID1515)CommentsWirefeedoscillatingbackandforth.TungstenmovedfurtherfromthecentertotheLidsideafterthetackispassed.Appearstobeagoodbitoflavaflowdownintothegapbuttherootisflowingsmoothlyat03:09:2013.AdjustedarcpositiontowardsLidinprocessafterithaddriftedmoretocenterat03:12:00am.Runningverysmoothly.Gapwideningabout03:33:20am.ContinuestorunsmoothlyatVID55.Thereisalargeamountoflavaflowseeninthegaparea.Largeanddeepseparationbeingweldedovertomakeuniformtieinwithmoltenmaterialusedforremelting.Wirefeedentrypointmovedwayofftotheshellsidewallandmayhavechilledthepuddlesuchthatitcreatedfusiondefectsat3:48andarcwasstopped.Photoisprovidedshowingthisissue.Continuationofrootpassafterrepairingblowthrough.Wirefeedentrypointwascorrectedandapparentlythelackofcoolingderrivedfromthewireentrycausedtheblowthrough.Arcextinguishednormallyaccordingtothedownslopeprogrammed.Weldingsmoothlywithwirefeedinitiallywellcenteredonthetungstentip.MovedtotheLidsideattheveryend.ManipulatingtheTungstentobeginwelding.Sawnoevidenceofsurfacegrindingorcleaning.Tungstenappearstobewellcenteredonthetroughformedfrombeadwandthelidsidewall.Wirefeedisatthetorchtipasitshouldbe.Appearstobebreakingdownthematerialproperly.Somewanderingofthetungstenbutnotsevere.Appearedtoberelatedtotheovalityoftheshellrelativetothelidposition.
Report No.: 1700388.401.R1 Page C-8
WeldLocationVIDFileNameDateTimeStartTimeEndDuration(min)Length(in)TS(in/min)compLayerNo.TungstenBiasOTCPtoShell0810/11/201309:42:2809:57:26154ShellOTCPtoShell0910/11/201310:13:3111:07:34542083.04ShellOTCPtoShell1010/11/2013NoWeldingOTCPtoShell1110/11/201311:12:2912:21:09692083.05LidOTCPtoShell1210/11/201312:30:5213:19:50492084.26ShellStartsincenterofbeadthenmovestothesideastheweldpassesthepriorstoppingpoint.Beadsranprettymuchsteadilywithoutapparentincidents.Tieintotheshellsidewallappearedtobeuniformandsuccessful.Tungstenpositionedatlidsidewallwherepriorbeadtruncated.LavaflowiswellaheadoftungstenmakingaUshapearoundthewireentry.Notsurewhatthatisdoing.Flowingovertopoflidstartingat11:20andfinishingat12:21.Completedwithoutincident.Fillinginfromshellsideandflowingtoapproximatelythecrownoftheweld.Mayhaveincreasedthetravelspeedsincetheweldpoolseemedsmallerandflowonlywenttotheapexoftheweldcrown.Builduptoshellwallbutflowingallthewayacrosstothelidinsomeinstances.Beadrunningwellbutwirefeedappearedtostopandmayhaverequiredchangingwire.CommentsDSC15(VID1515)cont.
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Calculation 11042-0204 A Form 3.2-1 No.: Calculation Cover Sheet Revision 3 J.\REVA Revision 10 No.: Page 1 of 10 PROJECT NAME: OCR NO (if applicable)
- NUHOMS 61 BTH Type 1 11042-022 Rev.O DSCs for Monticello Nuclear Generating Plan PROJECT NO: 11042 CLIENT: Xcel Energy CALCULATION TITLE: Allowable Flaw Size Evaluation in the Inner Top Cover Plate Closure Weld for DSC #16 SUMMARY DESCRIPTION:
- 1) Calculation Summary An allowable flaw size of 0.15 inch is calculated for a 0.25 inch Inner Top Cover Plate (ITCP) weld in DSC # 16. Limit load analysis per ASME Code,Section XI, Appendix Cis used to determine the allowable flaw size. 2) Storage Media Location Rev.O: Secure network drive initially, then redundant tape backup. Rev.1, 2 and 3-No additional software files. If original issue, is licensing review per TIP 3.5 required?
Yes D No D (explain below) Licensing Review N/A Software Utilized (subject to test requirements of TIP 3.3): Software Software Log ANSYS Version:
Revision:
14.0.3 Not applicable when Rev. 0 was issued. Calculation is complete:
Originator Name and Signature:
Veeresh Sayagavi
!W b 5.?9.1 0 I 13.52.13 i
--04'00' Date: 09/10/2015 Calculation has been checked for consistency, completeness and correctness:
HAROON Raheel jltJ 2015.09.10 15:07:27
-04'00' Date: 9/10/2015 Checker Name and Signature:
Raheel Haroon Calculation is approved for use: SHIH Digitally signed by SHIH Yueh-9/10/2015 Kan DN: o=oAREVA GROUP, 2.5.4.45=5A3923106548495977F Yueh-Kan s.cn=sHIHYueh-Kan Project Engineer Name and Signature:
. Date: 2015.09.10 16:20:11
- <WOO' Date:
A Calculation No. 11042-0204 Revision No. 3 AREVA Calculation Page 2 of 10 REVISION SUMMARY Affected Affected Rev. Description Pages Disks 0 Initial issue All All 1 Excel Energy comments incorporated.
1-11,13 None and 14 Correct editorial error on Table 4 and Table 6. The nodal 2 force reported should be "lbs" instead of "kips". Add 1 ,2, 13 and None clarification about nodal force and force/in in Table 4 14 through Table 7 per OCR 11042-020, Rev. 0. Files related Revised such that this calculation only evaluates the As indicated to Outer Top 3 critical flaw size in the Inner Top Cover Plate (ITCP) and by the Cover Plate increased the weld size to 0.25 inch from 3/16 inch revision Weld (OTCP) evaluated earlier.
bars. are removed.
A Calculation No. 11042-0204 Revision No. 3 AREVA Calculation Page 3 of 10 TABLE OF CONTENTS Page 1.0 PURPOSE ..........................................................................................................................................
4 2.0 CONSERVATISM I ASSUMPTIONS
...................................................................................................
4 3.0 DESIGN INPUT/DATA
........................................................................................................................
4 3.1 Bounding Load Combinations
....................................................................................................
4 4.0 METHODOLOGY
................................................................................................................................
4 4.1 Allowable Flaw Evaluation
.........................................................................................................
4 4.2 Limit Load Analysis
....................................................................................................................
5
5.0 REFERENCES
...................................................................................................................................
6 6.0 NOMENCLATURE
..............................................................................................................................
6 7.0 COMPUTATIONS
.............................................................................................................
- .................
7 7.1 Allowable Flaw Size Evaluation
..................................................................................................
7 7.1.1 Weld Post-Processing and Stress Calculation
................................................................
7 7.1.2 Determination of Allowable Weld Flaw Size ...................................................................
7 8.0 RESULTS ...........................................................................................................................................
8
9.0 CONCLUSION
S
..................................................................................................................................
8 10.0 LISTING OF FILES .............................................................................................................................
8 LIST OF TABLES Page Table 1 Safety Factors for Circumferential Flaw (Ref. [5.1]) .........................................................................
9 Table 2 Weld Stress Results of Inner Top Cover Plate Welds for Individual Loads ......................................
9 Table 3 Load Combination Weld Membrane Stress ( o-'m) Result for Inner Top Cover Plate Weld ...............
9 LIST OF FIGURES Page Figure 1 Subsurface Crack Model for ITCP Welds .....................................................................................
10 Figure 2 Surface Crack Model for ITCP Welds .......................................
- ...................................................
10 A Calculation No. 11042-0204 Revision No. 3 AREVA Calculation Page 4 of 10 1.0 PURPOSE The calculation calculates the NUHOMS 61 BTH Type 1 DSC allowable flaw size for increased Inner Top Cover Plate (ITCP) closure weld size of 0.25 inch. 2.0 CONSERVATISM I ASSUMPTIONS
- 1. The weld allowable flaw size is based on radial tensile membrane force acting on the weld; however it is conservatively evaluated based on SRSS method excluding the compressive stresses in the weld. 2. ASME Code,Section XI, Appendix C Limit Load evaluation uses only primary stresses.
Residual stress being a secondary stress are not considered.
3.0 DESIGN INPUT/DATA Per Ref. [5.9], the distance between the weld root and crown at the canister wall ranges from 0.25 inches to 0.4 inches for ITCP lid weld. Thus, the ITCP weld size is modified to 0.25 inch in lieu of 3/16 inches per design. 3.1 Bounding Load Combinations All bounding normal, off-normal and accident load combinations are taken from Ref. [5.2]. 4.0 METHODOLOGY 4.1 Allowable Flaw Evaluation The allowable flaw evaluation is based on flaw evaluation methodology per ASME Code,Section XI, Appendix C Ref. [5.1]. Although the affected component is not subject to in-service inspection activities, the methodology of Section XI is deemed appropriate for this application.
Determination of the allowable surface and sub-surface flaw depth is accomplished by means of the methodology, outlined below. Figure 1 shows the possible circumferential flaw for ITCP Welds. It is stipulated that the allowable flaw configuration is a circumferential weld flaw exposed to the tensile component radial stress. Conservatively the weld flaw is evaluated for all the component stresses except the compressive stresses onto the weld. Safety factors used to determine the allowable flaw size are taken from Appendix C, Section C-2621 of Ref. [5.1]. All bounding normal, off-normal and accident load combinations are taken from Ref ([5.2], Table 52). The following are basic steps that are performed in order to determine the allowable flaw depth: 1) Identification of bounding load and load combinations analyzed in Ref. [5.2]. 2) Calculate the resultant force acting on the weld ignoring the compressive load. Evaluate membrane stresses occurring at the ITCP weld. 3) Determine limiting membrane stresses in the ITCP weld for all load combinations.
- 4) Multiply limiting stresses with safety factors SFm for the corresponding Service Levels (Ref. [5.1]) as presented in Table 1.
A Calculation No. 11042-0204 Revision No. 3 AREVA Calculation Page 5 of 10 5) Since ITCP weld is GTAW (Non-Flux weld), thus according to ASME Code Sec XI, Division 1, Fig C-421 0-1, Ref. [5.1] maximum allowable flaw depth is estimated using Limit Load criteria.
4.2 Limit Load Analysis The relation between the allowable membrane stress and flaw depth at incipient stress is taken from Ref. [5.7], Table 12.28, which is given as ' 3o-1(1-a)2 O'"m= .*................................................................*......
(1) A-+ A-2+9(1-a)2 where: o-'m = The allowable membrane stress, which is the applied membrane stress times the different service factors, SFm determined from Appendix C, Section C-2621 of Ref. [5.1]. o-1 = the flow stress, defined as crt = (Sy + Su)/2, where Sy and Su are yield and ultimate strengths, respectively.
a= ac and A,= o-h = 0, for no bending stress on the weld. tW O'"m a = half crack length for center cracked plate, = crack depth for single edge cracked plate t = half plate thickness for center cracked plate, = plate thickness for single edge cracked plate for a 360° circumferential flaw, c = w, hence equation (1) reduces to o-'m = 0'" I ( 1-; ) ..................................................
- .. * .. * ......................................
(2) Using equation (2) the allowable flaw depth (a) is obtained as .........................................................................................
(3) Equation (3) can-be applied for both surface and subsurface crack (center-cracked plate and single edge crack plate model), respectively.
A Calculation No. 11042-0204 Revision No. 3 AREV.A Calculation Page 6 of 10
5.0 REFERENCES
5.1 ASME Boiler and Pressure Vessel Code,Section XI, Division 1, Appendix C, 2004 edition through 2006 Addenda.
5.2 TN Calculation NUH61 BTH-0200, Rev.O, "NUHOMS-61 BTH Type 1 Dry Shielded Canister Shell Assembly Structural Analysis".
5.3 ANSYS Computer Code and User's Manual, Release 14 (used only for post processing results).
5.41SG-15, Rev. 0, "Materials Evaluation".
5.5 ASME Boiler and Pressure Vessel Code, Division 1, Subsection NG, 1998 edition through 2000 Addenda.
5.6 TN Calculation No. NUH61 BTH-0403, Rev. 2, "NUHOMS-61 BTH DSC Thermal Evaluation for Storage and Transfer Conditions".
5.7 T.L. Anderson, "Fracture Mechanics, Fundamentals and Applications",
Second Edition.
5.8 TN Engineering Evaluation No. 11 042-EE-001, "Monticello Nuclear Generating Plant: Engineering Evaluation of Spent Fuel Storage Canisters with Nonconforming Closure Welds". 5.9 Design Input Document Dl-11042-02 Rev.O, AREVA Document Number 180-9236022-000, NDE Services Final Report, Monticello, DSC-16, Phased Array UT Examination Results of the Inner and Outer Top Cover Lid Welds. 6.0 NOMENCLATURE ITCP: Inner Top Cover Plate DSC: Dry Shielded Canister DWH: Horizontal Dead Weight PI : Internal Pressure Fweld: Resultant weld load (excluding compressive load) R : Radius of the ITCP weld T weld : Weld size Weld Stress: The weld stress for the ITCP. a'm: Weld membrane stress at limit load for ITCP. SRSS: Square root of sum of squares.
A Calculation No. 11042-0204 Revision No. 3 AREVA Calculation Page 7 of 10 7.0 COMPUTATIONS 7.1 Allowable Flaw Size Evaluation 7 .1.1 Weld Post-Processing and Stress Calculation All the controlling load combinations for ITCP weld are listed in Ref. [5.2, Table 54]. It is evident from these results that the critical cases are 75g side drop and 25g corner drop load cases. Weld nodal forces for ITCP weld nodes are post-processed using ANSYS. The compressive radial forces on the welds would have no impact on the allowable weld flaw evaulation.
Thus, these forces are excluded from the weld flaw evaluation.
The weld membrane stress ( CJ'm) at limit load for ITCP is calculated using SRSS method, for a 0.25 inch ITCP weld, while excluding the compressive loads onto the weld. Weld membrane stress for individual load cases and the bounding load combinations are listed in Table 2 and Table 3 respectively for ITCP. The top cover plate welds are evaluated assuming the shear load on the top cover plate welds due to a 25g corner drop. The outer top cover plate of the DSC is assumed to be unsupported by the cask in the axial direction.
ITCP welds resist the load such that the stress can be calculated based on the total weld area of both ITPC and OTCP welds. The allowable is based on a maximum temperature of not more than 300°F for any transfer condition (Ref. [5.6]). For the corner drop the total shear load on the welds is 9,437 lb/in (Ref. [5.2], Section 1 0.2). The load shared by the ITCP weld (1/4") and OTCP weld (0.50") are calculated below. = ( 114 ) X 9,437 = 3,146 fb fin (1/ 4 + 0.50) These shear loads are used in calculating the load combinations for the 25g corner drop. Weld membrane stress ( CJ'm) for individual load cases are calculated for ITCP. 7.1.2 Determination of Allowable Weld Flaw Size Table 3 lists the bounding load combinations to specify limiting depth of weld flaw for ITCP weld. The yield strength (cry) and ultimate tensile strength (cru) for SA-240 Type 304 at 300 °F are 22.4 ksi and 66.2 ksi (Ref. [5.2]). So flow stress (crt) as per Section 4.3 is crt= (22.4+66.2)/2
= 44.3 ksi The allowable flaw depths, calculated by means of the methodology described in Section 4.0. Note that in the case of subsurface flaws, the 't' and 'a' in equation (1) are half-width and half-crack depth, respectively, whereas for surface flaws 't' and 'a' respectively represent the weld thickness and the crack depth.
A Calculation No. 11042-0204 Revision No. 3 AREVA Calculation Page 8 of 10 ITCP Allowable Weld Flaw The weld membrane stress ( o-'m) are listed in Table 3. The bounding weld membrane stress is 17.08 ksi. The allowable flaw size for a 360° weld flaw is calculated below. = (44.3 -17.08) 0.25/44.3
= 0.15" (Using a single edge cracked plate model) For center crack plate model (used for a subsurface flaw), the half-crack length a, is 0.15/2 =0.075".
The total allowable crack length is 2*a = 0.15". 8.0 RESULTS The ITPC closure welds for individual and combination load cases are listed in Table 2 and Table 3 respectively.
The allowable flaw for surface (crack depth =a) and subsurface (half-crack length =a, total crack length =2a) flaws for ITCP is 0.15 inch and 360° along the circumference.
9.0 CONCLUSION
S The evaluations performed in this calculation indicate that the minimum allowable flaw size for the ITCP is 0.15" for a full 360° weld flaw. 10.0 LISTING OF FILES Below is the listing of all files used in the ANSYS for Finite Element Analysis.
All the nodal forces have been extracted using ANSYS Release 14.0.3 Ref. [5.3]. Load Case No. 4 4 File Name QT61 BIP.db and .rst T61 BSD.db and .rst Weld_forces_QT61 BIP, inp and .out, W9PFQK-L.err, QT61 BIP _weld_20psi_ITCP.txt Weld_forces_
T61 BSD, inp and . out, WT35B3-E.err, T61 BSD_weld_ITCP.txt Date Time 12/09/1999 12:40a 05/27/2000 2:40a 05/22/2014 14:14:34 05/22/2014 14:09:35 Description 20 psi internal pressure evaluation, Ref. [5.2, Table 22]. These files are not part of the archived files. 75g side drop acceleration, Ref. [5.2, Table 22]. These files are not part of the archived files. Post processing files for 20psi internal pressure.
Post processing files for 75g side drop . Note: For the above listed files, date is reported by the OS on the report issue date and time, these values may be changed by windows depending on time of the year (e.g., daylight savings time) and time zones A Calculation No. 11042-0204 Revision No. 3 AREVA Calculation Page 9 of 10 Table 1 Safety Factors for Circumferential Flaw (Ref. [5.1]) Circumferential Flaws Service Membrane Stress Level SFm A 2.7 B 2.4 c* 1.8 D 1.3 Table 2 Weld Stress Results of Inner Top Cover Plate Welds for Individual Loads Load F'm F'm Force (2) CY' (3) Description Load Case m Step Nodal Force (lbs/in)
(lbs)<1> (ksi) 4 20 psi internal pressure on inner pressure Pl(20) 277 139 0.56 boundary 4 75g side drop acceleration Side Drop 5495 2761 11.05 Notes <1> The .db and .rst files are taken from Ref. [5.2] and are listed in Section 9.0. <2> The element size of ANSYS elements is 1.99 inch Ref. [5.2]. Hence, the F'm Force= F'm Nodal force /1.99 <3> The weld throat size in., hence the CY'm = F'm Force /(1/4) Table 3 Load Combination Weld Membrane Stress ( CY'"') Result for Inner Top Cover Plate Weld Load Service Stress Safety Factor CY' m Case Level Category Loads CY'm (ksi) SFm (ksi) x SFm TR-9 D p Pl(20) + 25g Corner Drop {ll 13.14 1.3 17.08 TR-10 D p Pl(20) + 75g Side Drop 11.61 1.3 15.09 Notes <1> The corner drop load combination is calculated by adding 3,146 lbs/in of shear load to the individual loads for Pl(20) load case obtained from Ref. [5.2].
A Calculation No. 11042-0204 Revision No. 3 .AREVA Calculation Page 10 of 10 t t t t t
! Figure 1 Subsurface Crack Model for ITCP Welds Figure 2 Surface Crack Model for ITCP Welds
3/21/16HAROON Raheel 2016.03.21
17:41:22 -04'00' t\ Calculation No. 11042-0205 Form 3.2-1 Revision No. 3 Calculation Cover Sheet AREVA TIP 3.2 (Revision
- 10) Page 1 of 90 OCR NO (if applicable):
PROJECT NAME: NUHOMS 61BTH Type 1 DSCs for 11042-025 Revision 0 Monticelto Nuclear Generating Plant PROJECT NO: 11 042 CLIENT: Xcel Energy CALCULATION TITLE: 61 BTH ITCP and OTCP Closure Weld Flaw Evaluation SUMMARY DESCRIPTION:
- 1) Calculation Summary This calculation qualifies Monticello DSC-16, a 61BTH Type 1 DSC, for all design basis loads in consideration of observed flaws in the Inner Top Cover Plate (ITCP) and Outer Top Cover Plate (OTCP) closure welds. 2) Storage Media Location
-Coldstor
-/areva_tn
/11 042/11042-0205-000
-Coldstor-/areva_tn
/11 042/11042-0205
-002 If original issue, is licensing review per TIP 3.5 required?
N/A Yes D No D (explain below) Licensing Review No.: Software utilized (subject to test requirements of TIP 3.3): Software Version: Software Log ANSYS 14.0 Revision R-31 Calculation is complete Date: Originator Name and Signature: Jeff Pieper Calculation has been checked for consistency, completeness
, and correctness Date: Checker Name and Signature
- Gabriel Lomas CJJJ.L) 3/Jo//6 Calculation is approved for use Date: Project Engineer Name and Signature
CalculationCalculation No.11042-0205Revision No.
3Page2of 90REVISION SUMMARYRev.DescriptionAffectedPagesAffectedData0Initial IssueAllAll1Revised per DCR11042-022 Revision 0.Made editorial clarifications, updated information from revised Reference calculations, removed extraneous sensitivity analyses.1-10, 13, 14, 17, 18,21, 22, 24-34, 36, 45, 46, 48-50, 59-77Removed extraneous data.2Revised per DCR 11042-023 Revision 0.Added elastic-plastic analyses in Appendix A.
Additional discussion and clarifications added throughout.1-4, 6-10, 13-17, 19-22, 24, 27-29, 31, 33-35, 44, 60, 62-63, 68-78, Appendix A (79-87)Added files in Appendix A 3Revised per DCR 11042-025Revision 0to address RAI-2-1. Clarified limit load vs elastic plastic results.Added Table 7. Added additional results plots in Appendix A.1-7, 13, 15, 21, 22, 29, 35, 36, 45, 46, 81, 82, 90None CalculationCalculation No.11042-0205Revision No.
3Page3of 90TABLE OF CONTENTSPage1.0PURPOSE................................................................................................................................................72.0ASSUMPTIONS.......................................................................................................................................83.0DESIGN INPUT/DATA.............................................................................................................................93.1DSC Geometry................................................................................................................................93.2Flaw Details and Geometry..........................................................................................................103.2.1Outer Top Cover Plate......................................................................................................103.2.2Inner Top Cover Plate......................................................................................................113.3Material Properties........................................................................................................................133.4Design Criteria..............................................................................................................................134.0METHODOLOGY...................................................................................................................................144.1Analysis Method and Acceptance Criteria...................................................................................144.2Load Cases...................................................................................................................................174.3FEA Model Details........................................................................................................................224.3.1Axisymmetric Case #1......................................................................................................244.3.2Axisymmetric Case #2......................................................................................................244.3.3Axisymmetric Case #0......................................................................................................244.3.4Half-Symmetry (3D) Case #1...........................................................................................254.3.5Half-Symmetry (3D) Case #0...........................................................................................274.4Limit Load Solution Details...........................................................................................................2
75.0REFERENCES
.......................................................................................................................................286.0ANALYSIS..............................................................................................................................................296.1Axisymmetric Analyses for Internal Pressure...............................................................................296.1.1Axisymmetric Case #1 -Initial Mesh Model....................................................................296.1.2Axisymmetric Case #1 -Refined Mesh Models..............................................................296.1.3Axisymmetric Case #2......................................................................................................316.1.4Axisymmetric Case #0......................................................................................................316.2Half Symmetry Analyses for Internal Pressure (Benchmark Cases)...........................................326.3Half Symmetry Analyses for Side Drop Loading..........................................................................336.3.1Half-Symmetry Case #1...................................................................................................336.3.2Half-Symmetry Case #0...................................................................................................346.4Evaluation of the 25g Corner Drop...............................................................................................347.0DISCUSSION AND CONCLUSIONS.....................................................................................................
358.0LISTING OF COMPUTER FILES..........................................................................................................379.0TABLESAND FIGURES........................................................................................................................3910.0APPENDIX A -ELASTIC-PLASTIC ANALYSES...................................................................................81 CalculationCalculation No.11042-0205Revision No.
3Page4of 90LIST OF TABLESPageTable 1 -Summary of Design Basis Load Combinations for the 61BTH DSC [Ref. 5.8]..............................39Table 2 -Internal Pressure in the 61BTH Type 1 DSC..................................................................................42Table 3 -Maximum Temperatures in the 61BTH Type 1 DSC Shell.............................................................42Table 4 -Properties of SA-240 Type 304. [Ref. 5.11]....................................................................................43Table 5 -Properties of SA-36. [Ref. 5.11]......................................................................................................44Table 6 -Summary of Load Cases, Mesh Refinement Results, and NB-3228.1 Limit Load Analysis Results.....................................................................................................................................45Table 7 -Evaluation of Peak Strain Values at Specified Loads and at 1.5x Specified Loads from Elastic-Plastic Analyses.......................................................................................................................46Table A-1-Summary of Elastic Plastic Analysis Results..............................................................................83 CalculationCalculation No.11042-0205Revision No.
3Page5of 90LIST OF FIGURESPageFigure 1 -Sketch of the 61BTH DSC Top End and Transfer Cask from Reference 5.1...............................47Figure 2 -Details of the 61BTH Top End Component Interfaces..................................................................48Figure 3 -ITCP and OTCP Closure Weld Details from Reference 5.5..........................................................49Figure 4 -DSC Top End Detailed Dimensions...............................................................................................50Figure 5 -OTCP Flaws -Raw Data from Reference 5.1...............................................................................51Figure 6-OTCP Flaws -Main Flaw Group Reduced and Bounded.............................................................51Figure 7 -OTCP Flaws -Bounding Set #1 for ANSYS Collapse Analysis...................................................52Figure 8 -OTCP Flaws -Bounding Set #2 for ANSYS Collapse Analysis....................................................52Figure 9 -ITCP Flaws -Raw Data from Reference 5.1.................................................................................53Figure 10 -ITCP Flaws -Bounding Flaw Set for ANSYS Collapse Analysis................................................53Figure 11 -Overview of the Axisymmetric Model...........................................................................................54Figure 12 -Mesh Details Near the Lid Regions of the Axisymmetric Model.................................................54Figure 13 -Mesh Details at the Welds for Axisymmetric Case #1.................................................................5 5Figure 14 -Flaw Locations for Axisymmetric Case #1...................................................................................55Figure 15 -Refined Mesh (Weld Region) for Axisymmetric Case #1............................................................56Figure 16 -Refined Mesh (Weld and Lid Interior Region) for Axisymmetric Case #1...................................56Figure 17 -Mesh Details at the Welds for Axisymmetric Case #2.................................................................5 7Figure 18 -Flaw Locations for Axisymmetric Case #2...................................................................................57Figure 19 -Overview of the Half-Symmetry Model........................................................................................58Figure 20 -Detail Views and Mesh Plots of the Half Symmetry Model.........................................................59Figure 21 -Isometric Views of Half-Symmetry Model....................................................................................60Figure 22 -Isometric Views of Half-Symmetry Model (Refined Circumferential Mesh)................................61Figure 23 -Results for Axisymmetric Case #1 -Initial Mesh -Service Level A/B.......................................62Figure 24 -Deflection History of the Center of the OTCP for the Axisymmetric Case #1 Initial Mesh.........63Figure 25 -Results for Axisymmetric Case #1 -Refined Mesh in Weld Region -Service Level A/B.........64Figure 26 -Results for Axisymmetric Case #1 -Refined Mesh in Weld and Lid Interior Region -Service Level A/B..................................................................................................................................65Figure 27 -Deflection History of the Center of the OTCP for the Axisymmetric Case #1 Refined Mesh.....66Figure 28 -Comparison of Maximum Displacement Histories for Axisymmetric Model Sensitivity Studies.67Figure 29 -Comparison of Maximum Displacement Histories for Axisymmetric Model with Lid Contact Defined using Nodal DOF Couples vs. Contact Elements......................................................68Figure 30 -Comparison of Maximum Displacement Histories for Axisymmetric Model With and Without Pressure Loading Applied to the ITCP Weld Root Flaw Faces..............................................69Figure 31 -Results for Axisymmetric Case #2 -Refined Mesh in Weld and Lid Interior Region -Service Level A/B..................................................................................................................................70Figure 32 -Results for Axisymmetric Case #0 -Refined Mesh in Weld and Lid Interior Region -Service Level A/B..................................................................................................................................71Figure 33 -Comparison of Maximum Center-of-Lid Displacement Histories for the Various Flaw Models..72Figure 34 -Results for Half-Symmetry Case #1 Internal Pressure Loading Benchmark Analysis -Service Level A/B..................................................................................................................................73Figure 35 -
Benchmark of the Half Symmetry model with the Axisymmetric Analysis..................................74Figure 36 -Equivalent Stress and Plastic Strain Plots from the Half-Symmetry #1 Side Drop Analysis......75Figure 37 -Additional Results Plots from the Half-Symmetry #1 Side Drop Analysis...................................76Figure 38 -Equivalent Stress and Plastic Strain Plots from the Half-Symmetry #1 Side Drop Analysis with Off-Normal Internal Pressure...................................................................................................
77Figure 39 -Equivalent Stress and Plastic Strain Plots from the Half-Symmetry #1 Side Drop Analysis with Refined Circumferential Mesh.................................................................................................78 CalculationCalculation No.11042-0205Revision No.
3Page6of 90Figure 40 -Equivalent Stress and Plastic Strain Plots from the Half-Symmetry #0 (No Flaws) Side Drop Analysis....................................................................................................................................79Figure 41 -Comparison of Maximum Displacement Histories for the Various Half-Symmetry Analyses.....80Figure A-1-Ramberg-Osgood Derived Stress Strain Curve for SA-240 Type 304 at 500 oF......................84Figure A-2-Ramberg-Osgood Stress Strain Curves for SA-240 Type 304 from ANSYS Model at Various Temperatures..........................................................................................................................85Figure A-3-Service Level A Internal Pressure -Equivalent Plastic Strain at 32 psi....................................86Figure A-4-Service Level D Internal Pressure -Equivalent Plastic Strain at 65 psi....................................87Figure A-5-Service Level D Side Drop -Equivalent Plastic Strain at 75g...................................................88 Figure A-6-Service Level D Internal Pressure-Equivalent Plastic Strain at 100 psi..................................90Figure A-7-Service Level D Side Drop -Equivalent Plastic Strain at 112.5g..............................................90 CalculationCalculation No.11042-0205Revision No.
3Page7of 901.0PURPOSEThe purpose of this calculation is to evaluate DSC-16 at the Monticello Nuclear Generating Plant (MNGP)per ASME Section III criteriain consideration of flaws observed in the Inner and Outer Top Cover Plate (ITCP and OTCP) closure welds. The flaws are documented in the Reference 5.1Phased Array Ultrasonic Testing (PAUT)inspection report. The canister is a 61BTH Type 1 design. The ASME Section III Subsection NB Code limits on primary stress are evaluated using the limit load analysis criteria prescribed in the Code[Ref. 5.7].Additional elastic-plastic analyses are performed to document the actual predicted strains in the welds and to demonstrate adequate margin against plastic collapse.The body of this calculation is predominately concerned with the limit load analyses, including several finite element model mesh sensitivity analyses. The limit load analyses demonstrate satisfaction of the ASME NB limits on primary stress. The elastic-plastic analyses are performed in Appendix Aand the results are summarized in Section 7.0andTable 7. The elastic plastic analyses demonstrate adequate margin against the material ductility limits and against plastic collapse.
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3Page8of 902.0ASSUMPTIONS1.The ITCP weld to the siphon and vent blockand the welds of the siphon and vent port cover platesareinaccessible for PAUTinspection. Approximately 11" are obscured due to the locationof the siphon and vent block. Whereas the main circumferential lid-to-shell welds are made with an automated welding machine, some manual welding was performed around the siphon/vent block and ports.As discussed in Section 3.4, a strength reduction factor of 0.8 is considered for both the ITCP and OTCP welds. This factor accounts for the siphon and vent block welds and uncertainties in the UT technique.Note that the bounding flaws evaluated in this analysis are treated as full circumferential flaws. In other words, it is not assumed that the siphon and vent block is free of flaws, but rather contains the same bounding flaws as the examined welds. The geometry of the siphon and ventblock is not assumed in this analysis. It is assumed that the stresses in the circular configuration bound the stresses that would be computed for a configuration that explicitly includes the siphon and vent block.2.The longitudinal seams in the canister shell caused attenuation in the PAUTenergy beam at locations 24.3" to 24.8" and 129.5" to 130" [Ref. 5.1]that can potentially diminish the effectiveness of the examination in thesehalf inchareas. These regions are considered limited examination zones. It is assumed that the flaws observed outside of these regions are representative, and that no larger or more bounding flaws exist in the regionsbehind the canister seam welds.The use of the 0.8 weld strength reduction factor discussed above in Assumption 1 accounts for any uncertainty in this region.3.[Not used]4.The flaws are considered to be planar cracks lying on circumferential planes, parallel with the longitudinal axis of the cask. (I.e. the crack tips are pointed in the axial directions of the cask). This is a conservative flaw orientation since the welds primarily resist normal stresses in the plane of the lids due to plate bending caused by DSC internal pressure. Also, during the side drop loading, normal stresses in the plane of the lids resist the ovalizing mode of shell deformation.This flaw orientation is also conservative for through-thickness shear stresses in the lid weldssince it maximizes the reduction in available shear area.(A flaw of equal length, but placed at an angle, would result in less reduction of the weld throat thickness).5.Many of the flaws identified in the Reference 5.1PAUT examination report lie in very similar locations within the weld cross section. As discussed in detail in Section 3.2, flaws that lie in similar radial and axial positions within the weld are considered bounded by a representative "group flaw." The locations and sizes of the "group flaws" are chosen conservatively to ensure they are bounding of the individual flaws.6.The analysis is based on the nominal dimensions of the components as shown in the design drawings [References 5.3and 5.4] including theas-fabricatedradial gap between the outer diameter of the lids and the inner diameter of the DSC shell. Although weld shrinkage will close this gapduring closure operations, the resulting compressive load path between the lids and shell is conservatively ignored. Further discussion is provided in Section 4.3.
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3Page9of 903.0DESIGN INPUT/DATA3.1DSC GeometryThe 61BTH Type 1 DSC geometry is detailed in the Reference 5.3and 5.4drawings. The Reference 5.5drawing shows the details for the final ITCP and OTCP closure field welds. Sketches from Reference 5.1and details from References 5.3and5.4are shown in Figure 1through Figure 4.Thematerialfor all structural components(DSC Shell, OTCP, and ITCP) isSA-240 Type 304 stainless steel.The shield plug material is SA-36 carbon steel.The DSC shell is 0.5" nominal thickness.The ITCP is 0.75" nominal thickness. Per the Reference 5.5drawing, it is welded to the DSC shell and vent/siphon block with a 3/16" groove weld.However, the ITCP lid groove (weld prep) is 0.25"minimum, and it was confirmed that the weld is also 0.25"[Ref. 5.1].The OTCP is 1.25" nominal thickness. It is welded to the DSC shell with a 1/2" groove weld.The ITCP and OTCP closure welds(with the exception of the ITCP welds around the vent/siphon blockand the welds of the vent and siphon port cover plates)are made using the GTAW process with an automated welder. This is a non-flux type of weld.The vent/siphon block and the vent and siphon port cover plate welds are performed manually, also using a non-flux process.
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3Page10of 903.2Flaw Details and GeometryVarious setsof bounding flaws arechosen for the detailed analyses based on theflaw dimensions in Reference 5.1and the discussion below.Note that flaws are identified in thiscalculation using the numerical flaw listings in the Reference 5.1inspection report.3.2.1Outer Top Cover Plate3.2.1.1Case 1Figure 5shows all of the OTCP weld flaws from Reference 5.1plotted on an outline of the DSC geometry. Figure 6shows a similar plot but with the main cluster of flaws bounded by a box, and showing a representative "group flaw" for this region. The longest flaw within the group region is 31.7" long and the tallest flaw is 0.14" high. Therefore, the bounding flaw for this region is taken as a full circumferential flaw, 0.14" in height. Note that all flaws in thegroup region were reviewed to ensure that no two flaws in close circumferential proximity, considered as being joined, would produce a taller flaw. For example, OTCP Flaw #9 and OTCP Flaw #10 are within 0.17" of each other in the circumferential direction, but their combined height is only 0.47-0.38=0.09". Therefore these flaws, considered combined, are bounded by the 0.14" high group flaw.The radial and axial positions of the bounding flaw were chosen to be at the center of the group region. This radial position is within the critical failure planeof the weld(i.e. a plane containing the minimum weld throat thickness of 0.5").Figure 6also shows additional information about the flaws outside of the group region. OTCP Flaw #2 is intermittent around the entire circumference of the DSC. Therefore this flaw, at 0.12" in height, is considered a full circumferential flaw. Since OTCP Flaw #14 is in close proximity to Flaw #2, it is conservatively considered joined toOTCPFlaw #2, and the combined flaw height is considered to be present around the entire circumference. The combined flaw height is determined based on the geometry to be 0.195".As seen in Figure 6,OTCP Flaw #20 is remote from the group region and from OTCP Flaw #s 2 and 14. OTCP Flaw #20 is only 0.32" in length, and only 0.07" in height. This flaw is separated from OTCP Flaw #19 by 0.36" in the circumferential direction and by 0.19"in the axial direction. It is separated from OTCP Flaw #21 by 1.66" in the circumferential direction and by 0.23" in the axial direction. Since extension of the flaws under the postulated loading is negligible (since only one cycle of the critical loads is applied)this flaw willnot join with the adjacent flaws. Additionally, since OTCP Flaw #20 is much smaller than the critical surface flaw size of 0.29" from Reference5.17, it is not considered explicitly in the FEA analyses and is considered bounded by the other modeled flaws which are very conservative.Similarly, OTCP Flaw #3 is remote from all flaws with the exception of OTCP Flaw #2. However, OTCP Flaw #3 is very small, only 0.18" long and 0.09" tall. Inspection of the PAUTplots (see Page 22 of Reference 5.1) also shows that OTCP Flaw #2, which is considered as fully continuous in this analysis, is actually very intermittent at the circumferential position of OTCP Flaw #3. Furthermore, OTCP Flaw#3 is much smaller than the critical subsurface flaw size of 0.29" from Reference5.17. Therefore, it is not considered explicitly in the FEA analyses and is considered bounded by the other modeled flaws which are very conservative.Figure 7shows the first bounding flaw setconsidered for the OTCP in the ANSYS collapse analyses.
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3Page11of 903.2.1.2Case 2The discussion above and the flaw locations shown in Figure 5through Figure 7are based primarily on the tabulated flaw data from Reference 5.1. Since OTCP Flaw #2 is intermittent around the circumference of the weld, a closer inspection of the PAUTscan images is performed, and an additional flaw set for the OTCP is created. In this additional case, the location of OTCP Flaw #2 is based on the PAUTscan image of the flaw at the circumferential position of OTCP Flaw #14, which is the only additional flaw that could be considered to interact with OTCP Flaw #2. Based on the PAUTscanimages, the flaws are located as seen in Figure 8.In this case the height of both Flaw #2 and Flaw #7 are estimated based on the PAUT scan images and are conservatively larger than the flaw heights tabulated in Reference 5.1.3.2.2Inner Top Cover PlateFigure9shows all of the ITCP weld flaws from Reference 5.1plotted on an outline of the DSC geometry. All but two of the flaws are clustered in the region of the weld root at the inner surface of the DSC shell. Figure 10showsthebounding flaw set considered for the ITCP in the ANSYS collapse analyses. Both the representative group flaw and ITCP Flaw #7 are considered to be full circumferential flaws. ITCP Flaw #11 is remote from all other flaws (in the circumferential direction) and istherefore considered bounded by the representative group flaw. The representative group flaw for the ITCP is conservatively placed at the tension side of the weld when resisting internal pressure.All of the ITCP flaws documented in Reference 5.1were reviewed to ensure that no two (or more) flaws,which are in close proximity to each other,could be considered as combined and therefore creating a more critical flaw. The following cases are considered in particular:ITCP Flaw #2 and Flaw #3 are within 0.12" from each other in the circumferential position, but their maximum combined height (1.58-1.49 = 0.09") is bounded by the group flaw height of 0.09". ITCP Flaw #5 and Flaw #8 partially overlaps with Flaw #6 in the circumferential direction and would have a combined height of 0.15". However, Flaw #5 (0.15" in length) and Flaw #8 (0.14" inlength) are extremely small. Due to their overlap in the circumferential direction, their combined length would be only 0.16", and therefore wouldnot affect the global or local stability of the weld. This very short region with a potential 0.15" high flaw is bounded by the full-circumferentialrepresentation of the modeled flaws. ITCP Flaw #10 is within 0.04" of Flaw #12 in the circumferential direction. The individual flaws are 0.05" tall and 0.04"tall, respectively, and 0.49" long and 0.18" long, respectively.They are also separated in the axial direction by 0.09". Postulating a flaw from the bottom of Flaw #12 to the top of Flaw #10 would imply a height of 0.18". However, the combined-height region would be over a very short length and would not affect the global or local stability of the weld. Therefore this postulated combined flaw is considered bounded by the full-circumferential representation of the modeled flaws. It is noted thatbased on Figure 9and Figure 10,ITCP Flaw #7 appears to be in the base metal of the inner top cover plate. It is likely that the flaw is actually at the fusion / heat affected region between the weld metal CalculationCalculation No.11042-0205Revision No.
3Page12of 90and the base metal. The ANSYS models used in this calculation place the flaw at 0.81" inward from the outer surface of the DSC shell whereas the tabulated data in Reference 5.1places the flaw at 0.80" from the outer surface. The 0.01" discrepancy is considered negligible. The exact locationof the flaw is not considered critical in light of the significant margin that is available (See Section 7.0) and the generally very conservative idealization of the flaws (i.e. full circumferential).
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3Page13of 903.3Material PropertiesThe material properties for the DSC structure are taken from Reference 5.11.The properties of the two materials of construction, SA-240 Type304 stainless steel and SA-36 carbon steel, are provided in Table 4andTable 5, respectively. The weld metal is considered to be composed of the same properties as the base metal, as the welds are made with the non-flux GTAW method[Reference5.14]using bare metal ER308(stainless) filler material.The tensile strength of the ER308 electrode (80 ksi at room temperature[Ref. 5.15]) is slightly greaterthan the type 304 base metal (75 ksi at room temperature[Ref. 5.16]). The yield stress valueof the weld metal is assumed to be equal to or greater than the base metal.Therefore, the treatment of the weld metal as being identical to the base metal is appropriate for the Section III limit loadanalyses and the elastic-plastic analyses performed in this calculation. Temperatures used for material properties are discussed in Section 4.2and are shown in Table 3.Poisson's ratio for all modeled parts is taken as 0.29.Weight density for SA-240 Type 304 is taken as 0.285 lb/in 3.Weight density for SA-36 is taken as 0.284 lb/in 3.3.4Design CriteriaAll of the applicable design basis loading conditions are considered in accordance with the requirements of ASME Section III Subsection NB [Ref. 5.7]. Section 4.1details the methods used to perform the code [Ref. 5.7]qualifications. Section 4.2details the selection of the bounding load cases. The mockup used in the PAUT process development contained weld manufacturing flaws intentionally distributed in locations that would be expected with the weld process used for the DSC lid closure welds.
Approximately 30% of those flaws were placed at the weld root and 27% were placed near the weld toe to demonstratethat they could be reliably detected in the presence of typical geometric responses from those regions. The flaws include incomplete root penetration, lack of fusion, and tungsten inclusions. AREVA document 54-PQ-114-001 [Ref. 5.19], Section 8.0, provides images of the UT responses for these flaws and demonstrates that the PAUT process can effectively detect these flaws. Furthermore, the qualification performed on the blind mockup provides objective evidence that detection of flaws in these regions of the weld is not a problem. The blind mockup used for qualification contained a similar percentage/number and distribution of flaws as the development mockup. Although the flaw information for the blind mockup cannot be disclosed in order to preserve the security of the mockupfor future qualifications, EPRI and NRC personnel present at the demonstrationhave reviewed that information. In addition, uncertainties in the PAUT examination are accounted for by using a 0.8 reduction factor on the limit loadand a 0.8 reduction factor on the material ductility for the elastic-plastic analyses. This factor, which is in agreement with ISG-15
[Ref. 5.20], conservatively accounts for any additional limitations in the efficacy of the PAUT examinations and also accounts for the inaccessible area around the vent and siphon block as well as the geometric reflectors at the root and near the toe of the weld.
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3Page14of 904.0METHODOLOGY4.1Analysis Method and Acceptance CriteriaThe 61BTH DSC including the ITCP and OTCP welds are designed and analyzed per ASME Section III Subsection NB (the Code) [Ref.5.7] in the Reference 5.2calculation. The presence of the ITCP and OTCP weld flaws will cause high local stresses and complex stress fields that will render an elastic analysis (such as those performed in Reference 5.2)very difficult.Therefore, the flaws are explicitly included in the finite element models as "design features", and the applicable ASME code[Ref. 5.7]stress limits are evaluated as described below.Primary Stress LimitsIn order to satisfy the primary stress limits of Reference 5.7paragraphs NB-3221.1, NB-3221.2, and NB-3221.3, a Limit Analysis will be performed per Paragraph NB-3228.1. The acceptance criterionis that the specified loadings not exceed two-thirds of the lower bound collapse load, as determined using an ideally plastic (non-strain hardening) material model, with the yield stress set at a value of 1.5*S m.This criterionis used for evaluation of the Service Level A and B load cases discussed in Section 4.2.Note that Service Level C acceptance criteria aregenerally 20% greater than Service Level A criteria, per Paragraph NB-3224 of Reference 5.7. This information is used in the discussion in Section 4.2to eliminate some non-critical load cases.FortheServiceLevel D loadings (accident level internal pressure and side drop),the rules of ASME Section III Appendix F Paragraph F-1341.3 [Ref.5.9]areused, whichindicate that the loads "shall not exceed 90% of the limit analysis collapse load using a yield stress which is the lesser of 2.3S mand 0.7Su."This criterionis used for evaluation of the Service Level D load cases discussed in Section 4.2.An additional increase factor of 1/0.8=1.25 is applied to the required limit load collapse pressure in order to account for the weld strength reduction factor of 0.8 to account for UT sensitivity and inaccessible weld regions discussed in Section 3.4.Typically, the weld strength reduction factor is applied to the weld allowable stress during qualification. In the case of limit-load analysis, reduction of the material yield stress is applicable. The reduction in yield stress would have a direct, 1:1correlation to the calculated lower bound collapse pressure due to the perfectly-plastic (i.e. non-strain hardening) material model. In this analysis, rather than decrease the material yield stress the required calculated collapse pressure is increased by the factor of 1.25.Note that the Service Level D criterionis essentially 2.1 timesgreaterthan the Service Level A/B criterion,as calculatedbelow. This information is used in the discussion in Section 4.2to eliminate some non-critical load cases.At a temperature of 500 ºF,the limit load yield stress for SA-240 Type 304 for Service Levels A/B and D are 26.3 ksi and 40.3 ksi, respectively. The code[Ref. 5.7]required factors against the lower bound collapse load as determined by the limit load analyses for Service Levels A/B and D are 1.5 and 1.11, respectively.The ratio of the acceptance criteria is therefore:.(i.e. the Service level A/B criteria are 2.1 times as severe)
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3Page15of 90Limit Load Analysis BackgroundASME Section III Subsection NB provides only a basic description of the Limit Load analysis technique. A more thorough description is provided in ASME Section VIII Division 2Paragraph 5.2.3 [Ref.5.18]:
Separately, in order to address questions on thepotential for material rupture due topotentiallyhigh plastic strains, supplemental elastic-plastic analyses are performed in Section 10.0(Appendix A).Material Ductility LimitsIn order to show adequate margin against material failure at regions of high localized plastic strain, elastic-plastic analyses are performed in Appendix A. The peak strain values are compared against the material minimumspecified elongation limit reduced by the weld uncertainty factor of 0.8 discussed in Section 3.4. Primary Plus Secondary Stress LimitsThe Code[Ref. 5.7]also prescribeslimitson primary plus secondarystressesfor Service Levels A and B[Ref. 5.7Paragraph NB-3222.2].Secondary stresses may be developed in the DSC due to differential thermal expansion of the interconnected parts and thermal gradients within the structure. The code stress limit for primary plus secondary stress (calculated on an elastic basis) is 3S
- m. However, as shown in Ref. 5.7Figure NB-3222-1, rules for exceeding the 3S mlimit are provided in Paragraph NB-3228.5, which states that "the 3Smlimit - may be exceeded provided that the requirements of (a) through (f) below are met."Requirement (a) states that "the range of primary plus secondary membrane plus bending stress intensity, excluding thermal bending stresses, shall be 3Sm." This provision is related to the potential for "plastic strain concentrations" occurring in "localized areas of the structure", and the potential for these concentrations to affect the "fatigue behavior, ratcheting behavior, or buckling behavior of the structure" [Ref. 5.7Paragraph NB-3228.1]. Requirements (b) through (d) are also limitationsrelated to fatigue and thermalstress ratchet. As detailed in Section 10.5 of Reference 5.2, the DSC is exempt from fatigue analysis requirements since all of the criteria in NB-3222.4 of Reference 5.7are satisfied.Similarly, since the DSC CalculationCalculation No.11042-0205Revision No.
3Page16of 90thermal loads are not cyclic in nature (other than small daily and seasonal fluctuations), thermal stress ratchet is not a concern.Therefore, the 3S mlimit as it relates to fatigue is not applicable. Requirement (e) requires that the component temperature be less than 800 ºF for austenitic stainless steel.
The maximum DSC shell temperature (entire shell including the lid region) is 611 ºF (See Table 3).Therefore this requirement is satisfied.Requirement (f) states that the material must have a specified yield stress to ultimate stress ratio ofless than 0.8. For the 61BTH DSC which used SA-240 Type 304 steel, the ratio is 30/75 = 0.4. Therefore this requirement is satisfied.Based on the discussion above (primarily the fact that cyclic conditions are not a design factor for the DSC), there isno need to consider limits on primary plus secondary stresses. Therefore, thermal stresses are not included in this analysis.Special Stress LimitsIn addition to the primary and primary plus secondary stress limitsthe Code[Ref. 5.7]also imposes Special Stress Limits as detailed in paragraph NB-3227. Theapplicablespecial stress limits are discussed belowin relation to the DSC top end cover plate welds.Bearing Loads: There are no significant bearing loads affecting the ITCP and OTCP closure weldsduring Service Level A, B, or C loading.During the Service Level D side drop event, bearing stress exists at the contact surface between the DSC and Transfer Cask. However, as noted in ASME Section III Appendix F
[Ref. 5.9] paragraph F-1331.3, bearing stress need only be evaluated for pinned and/or bolted joints.
Therefore this special stress limit is not applicable to this evaluation.Pure Shear: Although the ITCP and OTCP closure welds are loaded in shear by internal pressure loading, the stress state is not pure shear due to the additional bending stresses. Paragraph NB-3227.2 of Reference 5.7clarifies that this stress limit is applicable to"for example, keys, shear rings, screw threads."
Therefore this special stress limit is not applicable to this evaluation. Progressive Distortion of Nonintegral Connections: The ITCP and OTCP closure welds are integral and therefore not nonintegral connections. Furthermore, there are no sources of significant cyclic loading that would cause progressive distortion of the DSC. Therefore this special stress limit is not applicable to this evaluation. Triaxial Stress: The purpose of the code[Ref. 5.7]limit on triaxial stress is to provide protection against failure due to uniform triaxial tension [Ref. 5.13Chapter 4.5]. Internal pressure in the DSC and bending of the cover plates may cause tension in the weld in the radial and circumferential directions, but there is no source for tension in the axial direction. Therefore failure due to hydrostatic tension in the weld metal is not credible. Therefore this special stress limit is not applicable to this evaluation. Fracture and Flaw ExtensionAlthough linear-type flaws have been identified in the structure, the critical failure mode of the welds is plastic collapse. Under one-time loading, elastic and plastic crack extension are not a concern for the very tough type 304 stainless steel materials of the DSC shell, OTCP, and ITCP. This conclusion is supported by ASME Section XI Article C-4000 "Determination of Failure Model" [Ref. 5.10] which states that for austenitic CalculationCalculation No.11042-0205Revision No.
3Page17of 90wrought material and non-flux welds, "plastic collapse is the controlling failure mode." Note that the 61BTH Type 1 DSC OTCP and ITCP closure welds are made with the GTAW method[Reference 5.14]which is a non-flux type of weld.Additionally, there is no source for fatigue flaw extension. The only cyclic loads on the DSC are minor daily and seasonal temperature fluctuations. Therefore, cyclic fatigue growth of the flaws in not a credible phenomenon. Combined withthe discussions above, the limit load analysis of the DSC top cover plates and closure welds is sufficient to satisfy all of the applicable stress criteria oftheCode [Ref. 5.7].Residual StressResidual stress due to welding is a secondary stress and therefore is not considered in the limit load analyses performed in this calculation, as the Section III Code[Ref. 5.7]does not require it in the limit load analysis.4.2Load CasesTable 1lists the design basis load combinations for the 61BTH DSC. This calculation is concerned with all load cases beginning with the inner top cover plate weld, identified as Load FL-6 in Table 1.The loading conditions of interest in this evaluation are internaland externalpressure and inertial loads due to handling, transfer, seismic, and accidental drop conditions.
As discussed in Section 4.1, secondary (thermal)loading is not considered.
Note that the discussions below, and the analyses performed in this calculation, are based on the conservative design values for internal pressure loading, rather than the actual calculated values of internal pressure. Table 2summarizes the conservative design values as well as the actual calculated values. Temperatures used for the material properties for each Service Level condition are listed in Table 3anddiscussed further in the paragraphs below.Service Level AThe bounding Service Level A load combination for the DSC top end cover plates and welds is load case TR-5 which combines the hot ambient condition with internal pressure and 1g axial inertial loading. The other directions of inertial loading are not considered critical since their effects are not directly additive to the internal pressure loading, and furthermore they are bounded by the 75g side drop load discussed further below.The 1g axial load will cause the DSCpayloadweight (fuel, basket, holddown ring, shield plug)to bear against the ITCP. Thetotal maximum payload weightis 75,811 lbs conservatively including the weights of the ITCP and OTCP [Ref. 5.2Section 10.2]. The equivalent uniform pressure applied to the top-end components is therefore:
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3Page18of 90Where the inner diameter of the DSC shell is 66.25 inches.Therefore, the bounding Service Level A case is a uniform 10psi internal pressure(for a Type 1 DSC)plus an additional 22.0psi acting on the shield plug in the outward axial direction of the DSC Shell.Conservatively, this analysis considers the combined 10+22=32psi load as a uniform internal pressure in the DSC Shell.This is very conservative since the fuel pressure load which is applied to the inner surface of the shield plug would in reality be distributed to the perimeter of the ITCP as a line load by the significant stiffness of the 7-inch thick shield plug. In other words, the approach used in this calculation maximizes the bending loads on the cover plates and therefore maximizes the loading on the closure welds.Note that the cases with external pressure loading are discussed below.
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3Page19of 90Service Level BThe bounding Service Level B load combination for the DSC topend cover plates and welds is the combination of the hot ambient condition withthe off-normal internal pressure of 20 psi(LD-6). All of the other Service Level B conditions, such as ram push/pull loads, do not affect thetop end components. Therefore, the bounding Service Level B case is a uniform 20 psi internal pressure. Since the pressure loading is smaller (20 psifor SL B versus 32 psi for SL A as described above), the temperature used for SL A (500 oF) bounds the maximum SL B temperatures (416 oF),and since the same limit load acceptance criterion is used for Service Levels A and B, this case is bounded by Service Level A.Service Level CThe bounding Service Level C load combination for the DSC top end cover plates and welds is one that bounds HSM-4 and HSM-8which combines the hot ambient condition, normal internal pressure(20 psi), and seismic loading. However, the seismic loads are bounded by the handling loads [Ref. 5.2Section 7.8] discussed above for Service Level A.In addition, the acceptance criteria for Level C limit load analysis is greater than Service Levels A and B.Therefore, all Service Level C conditions are bounded by the Service Level A case described above.Note that the other Service Level C cases (such as LD-7 and UL-7) are for accident condition DSC ram push/pull loads. These loads do not affect the DSC top end components. Therefore they are not applicable to this analysis. Note that cases with external pressure loading are discussed below.
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3Page20of 90Service Level DThree load combinations are found to be critical for Service Level D loading of the DSC top end components, namely:accident level internal pressurecorner dropside dropThe first load combination isbounded byHSM-5 or HSM-6 which consist of 65 psi internal pressure due to HSM blocked vent thermal conditions. This load is not combined with any other load that affects the top-end components. Therefore, the first bounding Service Level D load case considered inthis analysis is 65 psi internal pressure.Note that in this condition the maximum DSC shell temperature is 611 oF and 625 oFis conservatively used in this analysis (See Table 3).The other Service Level D conditions consist of the drop eventsandaccident-level seismic loading.The accident seismic loads are bounded by the handling loads [Ref. 5.2Section 7.8] discussed above for Service Level A.The end-drop load is not a credible event[see footnote 12 to Table 1] but was usedin the original calculation [Ref. 5.2]to bound the corner drop event.However, that analysis produced negligible load in the top cover plate welds due to the idealized boundary conditions. As a result of an RAI by the NRC, the corner drop is considered using an alternate idealization that maximizes the load in the top cover plate welds. In this case, the25-gcorner drop load has an axial component that may be considered to load the top end cover plates with the inertia of the fuel, shield plug, hold-down ring, ITCP and OTCP. This case is evaluated in Section 6.4.The 75g side drop load TR-10 is considered a critical load case and is evaluated in detail. Note that this load case represents 75x more load than the Service Level A 1g inertial loads. As discussed in Section 4.1, the Service Level D acceptance criterionis only 2.1 times less stringent than the Service Level A/B criterion.
Therefore, evaluation of the 75g side drop case using the Service Level D criterionis bounding of the Service Level A transverse inertial loading. (Also, as discussed in Section 4.3, the boundary conditions used for the 75g side drop analysis are conservative and representative of the boundary conditions encountered for theService Level A inertial loads and seismic loading.)The 75g side drop case also includes the off-normal internal pressure of 20 psi, as shown in Table 1.Note that the side drop event TR-10 occurs during transfer operations which result in a maximum DSC shell temperature of 500 oF as shown in Table 3. The higher Service Level D temperature of 625 oF discussed above occurs only during DSC storage in the HSM, and therefore is not combined with the side drop loading.
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3Page21of 90External Pressure LoadingExternal pressure is present on the DSC in load cases DD-2(vacuum drying, Service Level A) and HSM-9/10 (flood load, Service Level C).(The load cases with hydrostatic external pressure are due to the cask annulus being filled with water while the cask and DSC are in the vertical position. In this case the pressure load varies from zero at the top of the DSC to a maximum value at the bottom of the cask. Since the external pressure near the cover plates is essentially zero, these cases are not critical and are not considered further in this calculation.)In loadcase DD-2, the external pressure is 14.7 psi (full vacuum). This pressure is bounded by the Service Level B off-normal pressure (20 psi) and therefore primary stresses in the cover plates and welds are bounded by the internal pressure load cases. Stability concerns of the DSC shell are not affected by the presence of weld flawssince they are at the end of the cask, remote from the locations at which buckling would occur.Additionally, the external pressure is resisted directly by the shield plug and the shield plug support ring, rather than by the OTCP and ITCP welds. Therefore external pressure load case DD-2 is not critical and is not considered further in this analysis. In load case HSM-9/10, the flood load is due to a 50-foot static head of water, which is equivalent to 22 psi external pressure [Ref. 5.2Section 7.9]. This pressure is bounded by the 32 psi internal pressure considered for Service Level A discussed above.Therefore the flood load case HSM-9/10 is bounded by the other internal pressure load cases. SummaryThe bounding load cases considered for the limit load collapse analyses are therefore:(See Table 3for temperature references)(See Section 4.1for explanation of the 1.5 and 1.11 factors for Service levels A/B and D, respectively, and also for the 0.8 factor which accounts for limitations in the weld examinationand inaccessible weld regions, as discussed in Section 2.0, Assumption No. 1. Service Level A/B:32psi Uniform Internal Pressure, Properties at 500 ºF(Accounts for internal pressure + inertial load of DSC contents onto Lid)Limit load collapse pressure required tosatisfy criteria: 1.5*32/0.8 = 60 psiService Level D-1:65 psi Uniform Internal Pressure, Properties at 625ºFLimit load collapse pressure required to satisfy criteria: 1.11*65/0.8 = 90.2psiService Level D-2:75g Side Drop Acceleration plus 20 psi Uniform Internal Pressure, Properties at 500ºF.Limit load collapse acceleration required to satisfy criteria: 1.11*75/0.8 = 104gFor the elastic-plastic analyses performed in Appendix A, the same boundingload cases described above are performed in order to predict plastic strains for comparison to the material strainlimitsand to demonstrate adequate margin against collapse.
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3Page22of 904.3FEA Model DetailsSeveral finite element models of the top half of the 61BTH DSC areconstructed in ANSYS based on the Reference 5.3,5.4,5.5drawings. The models fall into two basic categories: axisymmetric(2D)and half-symmetric(3D).The axisymmetric models use ANSYS plane element type PLANE182, a 4-node axisymmetric plane element with non-linear capabilities. Each node has 2 degrees of freedom (translation in the X (radial)andY (axial)directions).The default element options are used in the analysis. Sensitivity studies were performed to ensure that there were no adverse effects on the results due to the potential shear locking of the elements.
(Sensitivity runs used KEYOPTION 1=3 to invoke the simplified enhanced strain formulation to relieve shear locking.)Additional discussion of the sensitivity analyses is provided in Section 6.0.Contact between the ITCP and OTCP is simulated using nodal coupling in the Y (axial) direction. (See Section 6.1.2for a sensitivity study using contact elements at this interface.)No contact is defined between the opposing faces of the weld flaws. In other words, whereas compressive loading normal to the plane ofthe flaw may in reality be transmitted via compression through the crack face surfaces, this load path is ignored. This is conservative, and considered necessary since it is difficult (or impossible) to deduce from the PAUT data what separation may exist between the two faces of the flaws.Also, no contact is consideredbetween the DSC shell inner diameter and the ITCPand OTCP outer diameters.As seen in Figure 4, the fabricated dimensions of the lids and shell result in small radial gaps between the outer diameterof the lids and the inner surface of the shell. During the welding process, these gaps close, but since a small remaining gap cannot be ruled out, this analysis conservatively assumes that the as-fabricated gap exists, as shown in Figure 4.Even if the lids deflect in the analysis such that the gaps would close, the resulting contact/compressive load path is conservatively neglected.This is conservative since it forces all loads in the lid to travelthrough the weld, rather than through compression between the lids and shell. Figure 11and Figure 12show images of the axisymmetric model.Loading and boundary conditionsare discussed in the following sections.These sections are focused on the limit load analyses. See Appendix A for discussion of the elastic-plastic analyses.
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3Page23of 90The 3D, half-symmetric model uses ANSYS solid element type SOLID185, an 8-node brick (or 6-node prism) element with non-linear capabilities. Each node has 3 degrees of freedom (translation in the X, Y, and Zdirections).The default element options are used in the analysis. Sensitivity studies were performed to ensure that the mesh was adequate. Additional discussion of the sensitivity analyses is provided in Section 6.0.Contact in the half-symmetry model isdefined using ANSYS element types CONT173 and TARGE170. Contact is defined between the following interfaces:OTCP to ITCPITCP to Shield PlugShield Plug outer diameter to DSC ShellShield Plugbottom surface to Support RingSupport Ring to DSC ShellThe default contact parameters are used, although the contact stiffness is reduced in some cases to aid in convergence.Due to the large contact areas and since the contact areas are generally remote from the critical stress regions, the contact stiffness is not considered a critical parameter.The default contact parameters include:[Reference 5.6]Penetration tolerance factor: Default value = 0.1. This parameter controls the acceptable level of penetration of the contact node into the target surface, based on the depth of the element underlying the target element.Pinball region scale factor: Default Value= 1.0. This parameter controls the extents of the region around each contact node that is checked for contact with target segments. The default volume is a sphere of radius 4*depth of the underlying element.KeyOption 2: Contact algorithm: Default = Augmented Lagrangian. The contact method is an iterative penalty method where the contact pressure is augmented during the equilibrium iterations so that the final penetration is within the acceptable tolerance. KeyOption 4: Location of contact detection point: Default = On Gauss Point. Other options include using the nodal points, normal to either the contact surface or the target surface. The default option is suggested for general cases. Other features and controls of the CONTA173 elements are related to advanced features (bonded contact, cohesion, etc.) and initial penetration and gap controls which are not utilized in this analysis.Figure 19through Figure 21show images of the half-symmetry model.Loading and boundary conditions are discussed in the following sections.Table 6shows a summary of the ANSYS models and analyses which are performed. Further details on the various ANSYS models are provided below.
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3Page24of 904.3.1Axisymmetric Case #1The first case considered is a combination of OTCP Flaw Set #1 and the ITCP bounding flaw setdiscussed in Sections 3.2.1.1and3.2.2, respectively. The meshand flawdetails for this case, called Axisymmetric Case #1, are shown in Figure 13andFigure 14.The mesh shown in these figures was created based on a basic goal of having at least 4 elements across the thickness of the net sections of the weld, as reduced by the flaws. In order to investigate the effects of mesh density, a refined mesh was created for this case, as shown in Figure 15.Since the sensitivity model shown in Figure 15only refined the weld region an additional model was created as shown in Figure 16to ensure a sufficient mesh in the lid interior region.This model, and all of the other axisymmetric models discussed below, areused for analysis of uniform internal pressure loading. The model is constrained in the radial direction at the axis of symmetryand in the axial direction at the bottom cut of the DSC shell near the mid-length of the cask (remote from the top end components of interest.)The pressure loading is applied to the internal pressure boundary (bottom surface of ITCP, surface of ITCP weld to Shell, and Shell inner surface). (See Section 6.1.2for a sensitivity analysis where internal pressure is included on the ITCP weld root flawinternal surfaces.)4.3.2Axisymmetric Case #2The second case considered is a combination of OTCP Flaw Set #2 and the ITCP bounding flaw setdiscussed in Sections 3.2.1.2and3.2.2, respectively. The mesh and flaw details for this case, called Axisymmetric Case #2, are shown in Figure 17andFigure 18for the refined mesh.Based on the results of the Axisymmetric Case #1 (See Section 6.1.2), the initial mesh level described above for Case #1 would be sufficient. However, since the run times remained reasonable, arefined mesh model (weld and lid interior regions) was generated and is used for Case #2.4.3.3Axisymmetric Case #0In order to study the effect of the flaws, a 3 rdcase is considered in which the flaws are removedand the as-designed collapse load is determined. Only therefined meshmodel(weldand lid interior regions)isconsidered.The mesh is identical to Figure 16but the coincident nodes along the crack faces are merged.
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3Page25of 904.3.4Half-Symmetry (3D) Case#1The 3D model is based on the Axisymmetric Case #1. (Analysis results showed that there was negligibledifference in the results from Axisymmetric Case #1 and Case #2. The total projected cross-sectional area of the flaws in Case #1 is greater than Case #2. Therefore, Case #1 is considered critical for the side drop loading). The same flaw pattern is modeled, but the initial mesh is slightly less refined in order to obtain reasonable run times. Mesh sensitivity studies are described below. The half-symmetry model isused for internal pressure loading (as a benchmark case to study the effects of mesh refinement) and also for side-drop loading.The shield plug support ring is connected to the DSC shell at the two corners using nodal DOF couples to represent the fillet welds used to join the two parts. In order to improve the numerical stability of the ANSYS model, soft springs (COMBIN14) elements are used to connect the shield plug to the support ring. The springs have a stiffness of 1 lb/in. The low stiffness combined with the very small relative deflections between these parts results in negligible internal force in the springs. The forcesin the springs at the final converged solution are reviewed to confirm that the spring forces are small.In all load cases, symmetry conditions are applied to the cut face of the model. Axial constraints are applied at the bottom cut of the DSC shell near the mid-length of the cask (remote from the top end components of interest.) For the internal pressure load case, the model was further reduced to a 90-degree model and symmetry constraints were placed on both cut faces of the model.The purpose of thiscalculation is to evaluate the effects of the closure weld flaws and qualify the welds and any other components affected by the welds. All other aspects of the DSC (such as the shell remote from the welds) are not in the scope of this calculation. The modeling approach (loads and boundary conditions) for the side drop event are considered in light of this purpose and are described in the following paragraphs.For the side drop cases, the OD of the canister shellis constrained in the vertical (drop) direction for a small sector (approximately 1.5"inches or 2.8degrees) of assumed contact. In reality the DSC is supported inside the Transfer Cask (TC) during this event. Therefore the true boundary condition would either be a line of contact along aTC rail (which is 3" wide) or a line of contact at areas remote from the rails. As deformations increase, the area of contact would also increase. As discussed below in Section 4.4, deflections are over-estimated in a limit load analysis. Therefore, the area of contact with the TC rail or inner surface is assumed to be constant. This conservatively neglects the increase in contact area that would occur during the drop deformations. Additionally, this boundary condition is representative of the DSC storage condition inside the HSM, where the DSC rests on the 3-inch wide steel rails.As discussed in the Reference 5.2calculation, the DSC payload (basket and fuel) are located approximately 21.5 inches away from the ITCP and are therefore considered to have no effect on the DSC lid components.
The effect of the basket and fuel loading on the DSC shell is considered in the basket design-basis calculation for side-drop loading. The basket hold-down ring is a grid-type structure that does not represent significant weight and is of sufficient strength and stiffness to be self-supporting during the side drop and not significantlyaffect the DSC shell and adjacent regions. Therefore, as in the Reference 5.2calculation, the DSC payload is not considered as affecting the top-end components and the weight is applied as a pressure along a strip of elementsat the impact region, beginning approximately 23" below the ITCP. Since the loads CalculationCalculation No.11042-0205Revision No.
3Page26of 90are essentially applied directly over the supported (impact) region of the DSC shell, they have no appreciable effect on the shell deformations. Images of the Half-Symmetry model are shown in Figure 19to Figure 21.In order to study the adequacy of the mesh for the half-symmetry model, an internal pressure load case was performed and compared to the results of the axisymmetric case refined mesh. This study confirms the adequacy of the mesh in the cross-section of the 3D model. In order to evaluate the mesh in the circumferential direction, a model was created with a refined mesh in the regions of the model showing large plastic strains(the impact region) and locations where tensile stress is expected in the weld (at the 90-degree location where the lid resists ovalization of the DSC shell). This model is shown Figure 22.
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3Page27of 904.3.5Half-Symmetry (3D) Case #0In order to study the effect of the flaws, an additional case is considered in which the flaws are removed from the model and the as-designed side drop limit load capacity is determined. 4.4Limit Load Solution DetailsAs discussed in Section 4.1, this calculation is based on predicting the lower-bound collapse loads of the DSCbased on limit load analysis. All materialsare modeled as elastic-perfectly plastic 1, with yield stress values based on the limit load analysis requirements of the ASME code[Ref. 5.7].Table 3lists the temperatures used for each load case, and the values of the material properties are shown in Table 4andTable 5.The prescribed loads are applied to the model, and then are increased linearly until the solution fails toconverge. The analyses use small deflection theory (NLGEOM,OFF). This is conservativesince deflections are unrealistically high in a limit load analysis due to the lower-bound non-strain-hardening material properties that are used. If large deflections were to be considered, the beneficial effects of OTCP and ITCP membrane action and of increased contact areas would be over-estimated, resulting in non-conservative effects. This was verified with a sensitivity study using NLGEOM,ON, which resulted in much higher collapse pressures. This confirmed that using NLGEOM,OFF is appropriate, and conservative.In addition, Paragraph 5.2.3.1 of Reference 5.18states that small displacement theory is to be used in a limit load analysis.
1"Elastic-perfectly plastic is standard mechanics of materials term that describes an idealized material that behaves in a linear-elastic manner up to the yield point, and thereafter is perfectly-plastic, i.e. non-strain hardening.
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05.0REFERENCES
5.1.AREVA Document No.180-9236022-000. NDE Services Final Report. "Monticello, DSC-16, Phased Array UT Examination Results of the Inner and Outer Top Cover Lid Welds." Revision0.5.2.AREVA (Transnuclear) Calculation No. NUH61BTH-0200 Revision 0. "NUHOMS-61BTH Type1 Dry Shielded Canister Shell Assembly Structural Analysis."5.3.AREVA (Transnuclear) Drawing No. NUH61BTH-3000 Revision 8. "NUHOMS 61BTH Type 1 DSC Main Assembly."5.4.AREVA (Transnuclear) Drawing No. NUH61BTH-3001 Revision 4. "NUHOMS 61BTH Type 1 DSC Shell Assembly."5.5.AREVA (Transnuclear) Drawing No. NUH61BTH-4008 Revision 1. "NUHOMS 61BTH Type 1&2 Transportable Canister for BWR Fuel Field Welding."5.6.ANSYS Version 14.0. ANSYS Inc.(Including the ANSYS Mechanical APDL Documentation).5.7.ASME Boiler and Pressure Vessel Code,Section III Subsection NB. 1998 Edition with Addenda through 2000.5.8.AREVA (Transnuclear) Document Number NUH-003 Revision 14. "Updated Final Safety Analysis Report for the Standardized NUHOMS Horizontal Modular Storage System for Irradiated Nuclear Fuel."5.9.ASME Boiler and Pressure Vessel Code,Section III Appendices. 1998Edition with Addenda through 2000.5.10.ASME Boiler and Pressure Vessel Code,Section XI. Rules for Inservice Inspection of Nuclear Power Plant Components. 1998 Edition with Addenda through 2000.5.11.AREVA (Transnuclear) Document No. NUH61BTH1-0101 Revision 2."Design Criteria Specification for the NUHOMS-61BTH Transportable Storage Canister."5.12.AREVA Calculation No. 11042-0204Revision 3. "Allowable Flaw Size Evaluation in the Inner Top Cover Plate Closure Weld for DSC #16"5.13.Chattopadhyay, Somnath. "Pressure Vessels Design and Practice." CRC Press. 2004.5.14.TriVis Incorporated Welding Procedure Specification No. SS-8-M-TN Revision 10.5.15.ASME Boiler and Pressure Vessel Code,Section II, Part C. "Specifications for Welding Rods, Electrodes, and Filler Metals." 1998 Edition with Addenda through 2000.5.16.ASME Boiler and Pressure Vessel Code,Section II, Part D. "Properties." 1998 Edition with Addenda through 2000.5.17.AREVA (Transnuclear) Calculation No. NUH61BTH-0253 Revision 0. "NUHOMS 61BTH Type1 DSC Shell Assembly Outer Top Cover Plate Critical Flaw Size of Weld."5.18.ASME Boiler and Pressure Vessel Code,Section VIII Division 2. 2010.
5.19.AREVA Document No. 54-PQ-114-001 Revision 0. "Phased Array Ultrasonic Examination of Dry Storage Canister Lid Welds."5.20.NRC Spent Fuel Project Office Interim Staff Guidance -15. Materials Evaluation. (ISG-15).1/10/2001.
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3Page29of 906.0ANALYSISTable 6shows a summary of the results of all of the limit load analyses performed for this calculationandincludes a comparison of the results with the acceptance criteria. Each limit load analysiscase is discussed in more detail below.6.1Axisymmetric Analyses for Internal Pressure6.1.1Axisymmetric Case #1-Initial MeshModelTwo analyses are performed with the Axisymmetric Case #1 initial-mesh model described in Section 4.3.1:one case using the Service Level A/B material properties and one case using the Service Level D material properties. The collapse pressureswere determined to be 95.9psi for Service Level A/B and 136.6psi for Service Level D. Figure 23showsvarious plots of the plastic strain in the initial-mesh modelfor Service Level A/B at various locations and levels of loading. These strain plots arealsorepresentative of the behavior of the Service Level D analysis.Figure 24shows the deflection history at the center of the lid, and indicates the expected plastic instability that occurs as the limit load is approached. Note that both the strains and displacements presented in these figures have no physical meaning and the displacement plots show only the loading (pressure) at which the solution fails to converge. Since the initial mesh contains several element divisions at each critical cross-section, it is not expected that element shear locking (due to the default fully-integratedelements) will be significant. To confirm this, a test case was done using the Service Level A/B modelbut with the Simplified Enhanced Strain element formulation (KEYOP 1=3). The collapse pressure was found to be 96.1 psi, which is essentially identical to the initial results. 6.1.2Axisymmetric Case #1 -Refined Mesh ModelsAdditional analyses are performed using the Service Level A/B material properties with the refined mesh models described in Section 4.3.1.Figure 25and Figure 26show the plastic strain results for the refined mesh at the weld regionand the refined mesh at the weld and lid interior regions, respectively. The collapse pressures were found to be 94.8 psi and 93.8 psi, respectively, for these models. The OTCP deflection histories are shown in Figure 27.Note that both the strains and displacements presented in these figures have no physical meaning and the displacement plots show only the loading (pressure) at which the solution fails to converge.Figure 28shows a comparison of the maximum displacement history curves for the various Axisymmetric Case #1 models, done as part of the mesh sensitivity study. As seen in the figure, the results match very well. The results of the refined mesh models deviate at most (95.9-93.8)/93.8 = 2.2% from the initial mesh results. This is very close agreement particularly due to the non-linearnature of the analysis. Therefore, the initial mesh is considered sufficient. However since the analysis run times for the axisymmetric cases are reasonable even for the refined mesh model, the remaining axisymmetric cases use a refined mesh.The Axisymmetric Case #1 with refined weld andlids for Service Level D criteria reported a collapse pressure of 132.6 psi.Note that the nodal coupling in the axial direction between the ITCP and OTCP is a valid method to model the contact between the plates since the internal pressure loading ensures that the ITCP lid will bear against CalculationCalculation No.11042-0205Revision No.
3Page30of 90the OTCP, and since the nodes that are coupled remain coincident throughout the analysis, with only very minor differences in radial position occurring at the later load steps.In order to confirm the behavior of the nodal coupling, the Axisymmetric Case #1 model with refined welds and lids was modified to include contact between the ITCP and OTCP. The model replaces the nodal coupling with CONTA171 and TARGE169 elements, using the default element parameters. Figure 29shows a comparison between the model using DOF couples and the model using contact elements. As seen in the figure, the results are very similar, with the DOF-couple-model showing slightly more conservative results. Therefore, the nodal coupling is acceptable and is used in all other axisymmetric models.Note that in all of the FEA models, the internal pressure loading was not applied to the faces of the ITCP weld root flaw that is exposed to the internal region of the cask. Pressure loading on this crack face is negligible since the flaw is only 0.09" high, and in reality the ITCP flaws are generally very short (i.e. not full-circumferential flaws). In order to support this conclusion, a sensitivity analysis is performed where the pressure loading is applied to the ITCP weld root crack faces. The results, shown inFigure 30, confirm that pressure loading on the faces of this flaw are negligible.
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3Page31of 906.1.3Axisymmetric Case #2Two analyses are performed with the Axisymmetric Case #2 refined-mesh model describedin Section 4.3.2:one case using the Service Level A/B material properties and one case using the Service Level D material properties. The collapse pressureswere determined to be 93.7psi for Service Level A/B and 132.9psi for Service Level D. Figure 31showsvarious plots of the plastic strain for Service Level A/B at various locations and levels of loading. These strain plots are also representative of the behavior of the Service Level D analysis. Note that both the strains and displacements presented in these figures have no physical meaning and the displacement plots show only the loading (pressure) at which the solution fails to converge.6.1.4Axisymmetric Case #0One analysisisperformedwith the Axisymmetric Case #0 refined-mesh model described in Section 4.3.3 using the Service Level A/B material properties.The collapse pressures were determined to be 94.5psi for Service Level A/B.Figure 32showsvarious plots of the plastic strain at various locations and levels of loading. Note that both the strains and displacements presented in these figures have no physical meaning and the displacement plots show only the loading (pressure) at which the solution fails to converge.Figure 33shows a comparison of the maximum center-of-lid displacement history for all three axisymmetric cases. As seen in the figure, there is essential no difference between Axisymmetric Case #0, Case #1 and Case #2. The Case #1 and Case #2 analyses show slightly larger deflections early in the analysis due to the slightly reduced rotational fixity of the welds. However, the final collapse pressure are within(94.5-93.7)/93.7=0.9% of each other.This supports a suppositionthat theobservedflaws have negligible impact on the governing failure mode of the top end closure platesand welds. Again, displacements from the limit load analyses have no physical meaning, other than to show the onset of non-convergence of the FE model.
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3Page32of 906.2Half Symmetry Analyses for Internal Pressure (Benchmark Cases)The model described in Section 4.3.4is used for an internal pressure collapse analysis in order to benchmark the model against the axisymmetric cases. The collapse pressure was calculated to be approximately 97 psi.(The run was terminated at 95 psi and the final collapse pressure was estimated to avoid excessive computer run time). Figure 34shows various plots of the plastic strain at various locations and levels of loading.A comparison of the half-symmetry case to the refined-mesh axisymmetric case is shown inFigure 35.As seen in the figure, the half-symmetry case closely matches the behavior of the refined mesh axisymmetric modelalthough the results indicate a slightly greater collapse pressure.
Therefore, the half-symmetry model is considered sufficiently accurate for this analysis. As shown by the results, and as discussed in Section 7.0, there is significant safety margin available such that further mesh refinement of the half-symmetry model is not warranted. However, the effects of circumferential mesh density for the half-symmetry model can be seen in Section 6.3.1.
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3Page33of 906.3Half Symmetry Analyses for Side Drop Loading6.3.1Half-Symmetry Case #1The model described in Section 4.3.4is used to perform twoside-drop limit load analysis. One case includes side-drop acceleration loading only, while the second case includes the DSC off-normal internalpressure of 20 psi. For this later case, the 20 psi internal pressure is applied simultaneously with a 75g acceleration, and then both the pressure and the acceleration are increased linearly until the collapse g-load is obtained. (For example, for collapse occurring at 181g, the internal pressure at collapse is 20*181/75 = 48.3 psi.)Note that the side drop loading is combined with the design-basis off-normal internal pressure of 20 psi, as opposed to the internal pressure value of 32 psi used for the SL A/B cases which was the sum of the 10 psi normal pressure and an additional 22 psi to account for inertial handling/seismic loads. See Section 4.2The collapse g-load for side-drop-only loading was found to be approximately 181g. The collapse g-load when internal pressure loading was included was found to be greater than 181g. This later run terminated at 181g, but based on the collapse behavior(seeFigure 41) it is expected that smaller time steps would allow the solution to continue to larger loads.Various images of the stress and strain in the side drop analyses are shown in Figure 36to Figure 38.Note that both the strains and displacements presented in these figures have no physical meaning and the displacement plots show only the loading (pressure) at which the solution fails to converge.The Half-Symmetry Case #1 model with refined mesh in the circumferential direction was used to evaluate the side drop load case (without internal pressure).This analysis was performed up until a load of 185 g's, at which time the analysis was terminated manually to avoid large file sizes and excessive run time. As seen inFigure 41, this model showed a greater resistance to the side-drop loading, and would eventually result in collapse g-loads in excess of 185 g if smaller timesteps and longer run times were provided. Images of the stress and strain from this analysis are shown inFigure 39.Note that both the strains and displacements presented in these figures have no physical meaning and the displacement plots show only the loading (pressure) at which the solution fails to converge.This analysis confirms that the mesh used in the other half-symmetry cases is adequate, and conservative.
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3Page34of 906.3.2Half-Symmetry Case #0One side drop analysis is performed with the Half-Symmetry Case #0 model (no flaws) described in Section 4.3.5. Based on the results discussed above, only the case without internal pressure loading was considered. This analysis resulted in a collapseload of 189g. Stress and strain plots from this analysis are shown in Figure 40.As noted previously, these strain plots have no physical meaning. As shown inFigure 41, the collapse behavior was nearly identical to the case with weld flaws, indicating that the flaws had negligible effect on the results. 6.4Evaluation of the 25g Corner DropReference 5.2Section 10.2 evaluated the OTCP weld to resist a 25g inertial load on the entire DSC contents and neglecting the strength of the ITCP weld. Furthermore, a conservative stress was assumed in the weld due to internal pressure. The Reference 5.2calculation is revisedbelow to account for the strength of both weldsand include a reduction in the weld thickness due to the observed flaws. The total weld thickness is taken as the combined weld throats from the ITCP and OTCP minus the height of the flaws present in the welds. (See Reference 5.2Section 10.2 for the basis of the following values and calculations.)Note that the allowable weld stress noted below includes a joint efficiency factor of 0.7as described in Reference 5.2Section 6.2. This reduction factor conservatively bounds the reduction factor of 0.8 discussed in Section 3.4.Therefore, no further reduction factor is applied, and the calculation below is conservative.
- Note: the reduction of the weld to account for the flaws is based on the maximum flaw heights in any one plane through each of the welds. This is taken as 0.23" for the OTCP weld and 0.11" for the ITCP weld.)Therefore, the top end closure welds, with the observed flaws,are OK for the Service Level D corner drop event.
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3Page35of 907.0DISCUSSION AND CONCLUSIONSThis calculation qualifies the as-welded DSC-16 canister using a combination of limit load analyses and elastic-plastic analyses. The limit load analyses are used to show that the DSC satisfies the primary stress limits of ASME Section III Subsection NB. The elastic-plastic analyses are used to show that the actual predicted strain values are below the material ductility limits and that adequate design margin above and beyond the specified loading exists. Both the limit load and elastic-plastic analyses account for any remaining uncertainty in the weld (e.g. non-inspected weld regions and PAUT technique limitations) by includingan uncertainty factor of 0.8 which is described in detail in Section 3.4. Limit Load Analyses:The lower bound collapse pressure for Service Level A/B criteria was found to be 93.7psi which is greater thanthe required pressure of 1.5x32/0.8=60psi(where 1.5 is the code-required[Ref. 5.7]factor on the 32 psi design pressure loading and 0.8 is the weld strength reduction factor-see Section 4.2). Therefore the Service Level A/B criteria is satisfied.The lower bound collapse pressure for Service Level D criteria was found to be 132.6 psi which is greater than the required pressure of 1.11x65/0.8=90.2psi (where 1.11 is the code-required[Ref. 5.7]factor on the 65 psi design pressure loading and 0.8 is the weld strength reduction factor -see Section 4.2). Therefore the Service Level D criteria for internal pressure is satisfied.As noted in Section 6.1.4and as shown inFigure 33, there is essentially no difference in the collapse pressure and extremely little difference in the overall collapse behaviorof the DSC subjected to internal pressure loading with and without flaws in the weld. The lower bound collapse acceleration for side drop (Service Level D) loading was found to be 181gwhich is greater thanthe required load of 1.11x75/0.8=104g. Therefore the Service Level D criteria for side drop loading is satisfied.As noted in Section 6.3.2and as shown inFigure 41, there is essentially no difference in the collapse load and behavior between the as-designed DSC and the DSC with closure weld flaws. Elastic-Plastic Analyses:Table 7lists the peak strains predicted by the elastic-plastic analyses for the bounding Service Level D load cases performed in Appendix A. As shown in the table, the peak strain values remainbelow the material ductility limitsat the specified loading conditions, and also at 1.5x the specified loads. The ductility limit conservatively includes a reduction factor of 0.8 to account for weld uncertainties as discussed in Section 3.4. The Reference 5.12and 5.17calculationsdocument the ITCP and OTCP closure weld critical flaw sizes,respectively,based on the maximum radial stresses in the welds. The guidance and safety factors of Reference 5.10are used in the critical flaw size analysis. The critical flaw sizes are determined to be 0.19 and0.29 inches for surfaceand subsurface flaws, respectively,in the OTCP weld and 0.15 inches for surface and subsurface flaws in the ITCP weld.The largest singleOTCPflaw size documented in Reference 5.1is 0.14 inches. As discussed in Section 3.2a very conservative maximum combined flaw height of 0.195 inches is assumed inthis analysis. The largest single ITCP flaw size documented in Reference 5.1is 0.11 inches. Therefore, the observed flaws actually are smaller thanthe critical flaw size limits and therefore it is not surprising that the flaws are shown to have little effect on the capacity of the structure. This analysis CalculationCalculation No.11042-0205Revision No.
3Page36of 90shows that the quantity and close proximity of some of the flaws also has no significant adverse effects on the structural capacity of the DSC.Even though all observed flaws in the ITCP and OTCP welds are included in the analysis models using conservative representations, an additional weld strength reduction factor of 0.8 is considered by increasing the limit load acceptance criteria by a factor of 1/0.8=1.25 timesand by reducing the elastic-plastic strain limit by a factor of 0.8. The 0.8factor, which is the same magnitude reduction factor as in ISG-15 [Ref. 5.20], conservatively accountsfor any additional limitations in the efficacy of the PAUT examinations and also accounts for the inaccessible area around the vent and siphon block as well as the geometric reflectors at the root and near the toe of the weld.Therefore it is concluded that Monticello DSC-16, remains in compliance withthe ASME Section III Subsection NB [Ref. 5.7]stress limitsand has adequate design margin above and beyond the specified loadingswith the presence of the ITCP and OTCP closure weld flaws as documented in Reference 5.1.
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3Page37of 908.0LISTING OF COMPUTER FILESAnalyses performed on Computer HEA-0213AusingANSYS Version 14.0 [Ref.5.6].FileDate & Time listing is as displayed by the Windows 7 Operating System -Differences may occur due local time zone and daylight savings settings.Analysis CaseFile NameDate & TimeAxisymmetric 1Initial MeshInternal PressureSL A/B61BTH_WeldFlaw_1F_AX_2_DETACH.db4/7/2015 10:45 AM61BTH_WeldFlaw_1F_AX_2_DETACH.rst4/7/2015 10:45 AM61BTH_WeldFlaw_1F_AX_2_DETACH.mntr4/7/2015 10:45 AMSOLUTION_AXISYMM_IP_LimitLoad.INP4/7/2015 10:20 AMAxisymmetric 1Refined Weld MeshInternal PressureSL A/B61BTH_WeldFlaw_1F_AX_2_DETACH.db4/7/215 11:59 AM61BTH_WeldFlaw_1F_AX_2_DETACH.rst4/7/215 11:58 AM61BTH_WeldFlaw_1F_AX_2_DETACH.mntr4/7/215 11:59 AMSOLUTION_AXISYMM_IP_LimitLoad.INP4/7/2015 11:55 AMAxisymmetric 1Refined Weld and Lid MeshInternal PressureSL A/B61BTH_WeldFlaw_1F_AX_2_DETACH.db4/21/2015 9:04 AM61BTH_WeldFlaw_1F_AX_2_DETACH.rst4/21/2015 9:03 AM61BTH_WeldFlaw_1F_AX_2_DETACH.mntr4/21/2015 9:04 AMSOLUTION_AXISYMM_IP_LimitLoad.INP4/7/2015 11:55 AMAxisymmetric 1Initial MeshInternal PressureSL D61BTH_WeldFlaw_1F_AX_2_DETACH.db4/20/2015 10:43 AM61BTH_WeldFlaw_1F_AX_2_DETACH.rst4/20/201511:09 AM61BTH_WeldFlaw_1F_AX_2_DETACH.mntr4/20/2015 11:09 AMSOLUTION_AXISYMM_IP_LimitLoad_SLD.INP4/7/2015 11:20 AMAxisymmetric 1Refined Weld and Lid MeshInternal PressureSL D61BTH_WeldFlaw_1F_AX_2_DETACH.db4/30/2015 8:12 AM61BTH_WeldFlaw_1F_AX_2_DETACH.rst4/30/2015 8:12 AM61BTH_WeldFlaw_1F_AX_2_DETACH.mntr4/30/2015 8:12 AMSOLUTION_AXISYMM_IP_LimitLoad_SLD.INP4/7/2015 12:02 PMAxisymmetric 2Refined Weld and Lid MeshInternal PressureSL A/B61BTH_WeldFlaw_2G_AX_2.db4/21/2015 3:06 PM61BTH_WeldFlaw_2G_AX_2.rst4/21/2015 2:58 PM61BTH_WeldFlaw_2G_AX_2.mntr4/21/2015 3:06 PMSOLUTION_AXISYMM_IP_LimitLoad.INP4/7/2015 11:55 AMAxisymmetric 2Refined Weld and Lid MeshInternal PressureSL D61BTH_WeldFlaw_2G_AX_2.db4/21/2015 3:10 PM61BTH_WeldFlaw_2G_AX_2.rst4/21/2015 3:10 PM61BTH_WeldFlaw_2G_AX_2.mntr4/21/2015 3:10 PMSOLUTION_AXISYMM_IP_LimitLoad_SLD.INP P4/7/2015 12:02 PMAxisymmetric 0Refined Weld and Lid MeshInternal PressureSL A/B61BTH_WeldFlaw_1F_AX_2_DETACH.db4/21/2015 10:39 AM61BTH_WeldFlaw_1F_AX_2_DETACH.rst4/21/2015 10:31 AM61BTH_WeldFlaw_1F_AX_2_DETACH.mntr4/21/2015 10:39 AMSOLUTION_AXISYMM_IP_LimitLoad.INP4/15/2015 11:07 AMAxisymmetric 1Initial Mesh with Keyoption 1=3Internal PressureSL A/B61BTH_WeldFlaw_1F_AX_2_DETACH.db4/17/2015 5:40 PM61BTH_WeldFlaw_1F_AX_2_DETACH.rst4/17/2015 5:40 PM61BTH_WeldFlaw_1F_AX_2_DETACH.mntr4/17/2015 5:40 PMSOLUTION_AXISYMM_IP_LimitLoad.INP4/16/2015 12:27 PM CalculationCalculation No.11042-0205Revision No.
3Page38of 90Analysis CaseFile NameDate & TimeAxisymmetric 1Refined Weld and Lid MeshInternal Pressure,SL A/BITCP/OTCP couples replaced with Contact61BTH_WeldFlaw_1F_AX_2_DETACH.db5/19/2015 8:45AM61BTH_WeldFlaw_1F_AX_2_DETACH.rst5/19/2015 8:02 AM61BTH_WeldFlaw_1F_AX_2_DETACH.mntr5/19/2015 8:45 AMSOLUTION_AXISYMM_IP_LimitLoad.INP5/18/2015 5:03PMAxisymmetric 1Refined Weld and Lid MeshInternal Pressure,SL A/BWith Pressure on ITCP Weld Root Flaw Surfaces61BTH_WeldFlaw_1F_AX_2_DETACH.db5/18/2015 2:42 PM61BTH_WeldFlaw_1F_AX_2_DETACH.rst5/18/2015 1:37 PM61BTH_WeldFlaw_1F_AX_2_DETACH.mntr5/18/2015 2:42 PMSOLUTION_AXISYMM_IP_LimitLoad.INP5/18/2015 1:25 PMHalf Symmetry 1Initial MeshInternal PressureSL A/B61BTH_WeldFlaw_1GC.db4/29/2015 2:10 PM61BTH_WeldFlaw_1GC.rst4/29/2015 4:52 PM61BTH_WeldFlaw_1GC.mntr4/29/2015 4:52 PMSOLUTION_HALFSYM_LimitLoad.INP4/29/2015 2:10 PMHalf Symmetry 1Initial MeshSide DropSL D61BTH_WeldFlaw_1GC.db4/30/2015 8:20 AM61BTH_WeldFlaw_1GC.rst4/30/2015 3:35 PM61BTH_WeldFlaw_1GC.mntr4/30/2015 3:35 PMSOLUTION_HALFSYM_SD.INP4/30/2015 8:21 AMHalf Symmetry 1Initial MeshSide Drop + Internal PressureSL D61BTH_WeldFlaw_1GC.db5/1/2015 6:58 PM61BTH_WeldFlaw_1GC.rst5/1/2015 4:29 PM61BTH_WeldFlaw_1GC.mntr5/1/2015 4:13 PMSOLUTION_HALFSYM_SD.INP4/30/2015 10:26 PMHalf Symmetry 1Refined Circumferential MeshSide DropSL D61BTH_WeldFlaw_1GD_Refined.db5/6/2015 1:54 PM61BTH_WeldFlaw_1GD_Refined.rst5/6/2015 11:48 AM61BTH_WeldFlaw_1GD_Refined.mntr5/6/2015 11:47 AMSOLUTION_HALFSYM_SD.INP5/5/2015 9:01 PMHalf Symmetry 0Initial MeshSide DropSLD61BTH_WeldFlaw_1GC.db5/2/2015 6:53 AM61BTH_WeldFlaw_1GC.rst5/2/2015 6:53AM61BTH_WeldFlaw_1GC.mntr5/2/2015 3:37AMSOLUTION_HALFSYM_SD.INP4/30/2015 8:21 AM CalculationCalculation No.11042-0205Revision No.
3Page39of 909.0TABLESAND FIGURESTable 1-Summary of Design Basis Load Combinations for the 61BTH DSC [Ref. 5.8]
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3Page40of 90Table 1 (Continued) -Summary of Design Basis Load Combinations for the 61BTH DSC [Ref. 5.8]
A AREVA HSl\I LOADING Hoauontal OW nmuiO\Y lnttru:ll E.xterul Thmnal Handling Other Loads Stt'lice DSC Fuel DSC fut1 Pnssnn"' Condition Loads Lenl LD-1 Noiill3!
Loadmg-Cold Cask X --10/15 ps1g -OOF Cask 80Klp -A LD-2 Normal Loading -Hoi Cask X --10115 psig -1000 F Cask -SO Kip -A LD-3 Cask X --10115 psig --+80Kip -A w/shade'" LD-4 Loading -Cold Cask X --20psig -O"FCask -80Klp fF B LD-5 Off-Normal Loading -Hot Cask X --20psig -!OO*F CasktJI -80Klp fF B LD-6 Cask X --10psig --117* F wtshade<'> +SO Kip fF B LD-7 Accident Loading Cask X --20psig -117* F Kip FF OD HSM RAG£ Horizontal OW YerticaJDW loternal External Thennal Handling Other Senice DSC Fuel DSC Fuel Pre.ssurel'l Condition Loads Loads Len I HSM-1 Off-Normal HSM X --15psig --4Q*FHSM --B HSM-1 StoraJte HSM X --15psll!:
-O*FHSM --A HSM-3 Off-Normal HSM X --15 psig -117*FHSM --B HSM-4 Off-Normal Temp.-Failed Fuel HSM X ---20psig -117*FHSM -FF c HSM-5 Blocl:ed Vent Storage HSM X --65/120psig -117* F --D HSM-6 B.V. +Failed Fuel Storage HSM X --651120psig
-HSMIB\PX41 -FF D 117* F HSM-7 EarthquakeLrod.ing_
-Cold HSM X --10115 psig -OOF -EQ CJI)IIJ/ HSM-8 Earthquake Load.ing_ -Hot HSM X .. -10115 psig -!OO*FHSM -EQ CJI)IIll HSM-9 Flood Load (50' H20) -Cold HSM X --10115 psig psig O*FHSM -c EISM-10 Flood Load (50' H20)-Hot HSM X --10115 ps1g l1 ps1g IOO*FHSM -Flooci<ll c HS:O.I t:l\1.0.\Dll\G Horizontal DW Yertical DW Internal ExtHn<>l Thermal Handling Other DSC Fuel DSC Fuel Condition Loads Loads Len I lJL.I :4ormal Unloading-Cold HSM X --10115 psag -
-60 Kip -A UL-2 Unloading -Hot HSM X ---10115 p>ig -lOOOFHSM 60Kip -A liL-3 HSM X --10115 psig -117*F -60 Kip -A wlshade UL-4 Off-Normal Uuloadm.g
-Cold HSM X --10psig -OOFHSM +60Kip FF B UL-5 Off-Norma l Unloadm.g-Hot HSM X --20pstg -!OO*FHSM -60 Klp FF B UL-6 HSM X --20psig -117*F -60Kip FF B UL-7 Off. Norm. Unloading
-FF!Hot<W) HSM X --20psig -wlsbade ... so Kip FF c tOO* F HSM UL-8 Accident Unloading
-FF/Hot"*"' HSM X --65 1120' 'J psig -tOO*FHSM Kip FF D RF-1 DSC Reflood --Cask X 20 psi!!, (max) Hydrostatic 1200 F Cask --D CalculationCalculation No.11042-0205Revision No.
3Page41of 90Table 1 (Concluded) -Summary of Design Basis Load Combinations for the 61BTH DSC [Ref. 5.8](Notes for the preceding portions of Table 1)
A AREVA L 25g and 75g drop acceleration includes gravity effects.
Therefore, it is not necessary to add an additional LOg load. 2. For D events, only maximum temperature case is considered.
(Thermal stresses are not limited for level D events and maximum temperatures give minimum allowables). 3. Flood load is an external pressure equi\*alent to 50 feet of water. 4. BV = HSM vents are blocked. 5. At. temperatur e over l00°F a sunshade is required over the Transfer Cask. Temperamr es for these cases are enveloped by the 100°F (without sunshade) case. 6. As described in Section T.4, pressure assumes release of the fuel cover gas and 30% of the ftssion gas. Since unloading requires the HSM door to be removed, the pressure and temperantre s are based on the normal (unblocked vent) condition
. Pressure is applied to the confinement botmdary. 7. As described in Section T.4. this pressure assumes release of the fuel cover gas and 30% of the fission gas. Although unloading requires the HSM door to be removed.
the pressure and temperatures are based on the blocked vent condition.
Pressure is applied to the shell. irmer bonom and irmer and outer top cover plates. 8. Not usecl 9. Unless noted otherwise
, pressure is applied to the confinement boundary. 10 psig and 65 psig are applicable to Type 1 DSC. while 15 psig and 120 psig are applicable to Type 2 DSC. 10. Fuel deck seismic loads are assumed en\*eloped by handling
- 11. Load Cases UL-7 and UL-8 envelop loading cases where the due to insertion loading of 80 kips are added to stresses due to internal pressure (in reality. the insertion force is opposed by internal pressure). 12. The 60g top end drop and bottom end drop are not credible therefore these drop analyses are not required. However, consideration of 60g end drop aud 75g side drop conservatively em*elops the effect of25g comer drop. 13. Consen*atively based on normal operating pressure times 1.5 to cover future 1 OCFR Part 11 requiremet1ts. 14. A 25g comer drop analysis (30° from horizontal
) of61BTH DSC without support from the TC is to be documented. 15. Service Levcl C is for the standard seismic event and Service Llfl*el D is for the high seismic CalculationCalculation No.11042-0205Revision No.
3Page42of 90Table 2-Internal Pressure in the 61BTH Type 1 DSCDesign ConditionMaximum Calculated Pressure[psi]Design Pressure used in Ref. 5.2and This Calculation[psi]ReferenceNormal7.310Ref. 5.8Table T.4-16Off-Normal10.920Ref. 5.8Table T.4-20Accident56.165Ref. 5.8Table T.4-24Table 3-Maximum Temperaturesin the 61BTH Type 1 DSC ShellDesign ConditionMaximum Calculated Temperature[ºF]Design Temperature used in This Calculation[ºF]ReferenceNormalStorage374500Ref. 5.8Table T.4-13Transfer439500Off-NormalStorage399500Ref. 5.8Table T.4-18Transfer416500AccidentStorage611625Ref. 5.8Table T.4-22Transfer467500 CalculationCalculation No.11042-0205Revision No.
3Page43of 90Table 4-Properties of SA-240 Type 304. [Ref. 5.11]
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3Page44of 90Table 5-Properties of SA-36. [Ref. 5.11]
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3Page45of 90Table 6-Summary of Load Cases, Mesh Refinement Results, and NB-3228.1 Limit Load AnalysisResults CalculationCalculation No.11042-0205Revision No.
3Page46of 90Table 7-Evaluation of Peak Strain Values at Specified Loads and at 1.5x Specified Loads from Elastic-Plastic Analyses.
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3Page47of 90Figure 1-Sketch of the 61BTH DSC Top End and Transfer Caskfrom Reference 5.1
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3Page54of 90Figure 11-Overview of the Axisymmetric ModelFigure 12-Mesh Details Near the Lid Regions of the AxisymmetricModel(Small differences in the mesh exist amongst the sub-models)
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3Page55of 90Figure 13-Mesh Details at the Welds for Axisymmetric Case #1Figure 14-Flaw Locations for Axisymmetric Case #1 CalculationCalculation No.11042-0205Revision No.
3Page56of 90Figure 15-Refined Mesh(Weld Region)for Axisymmetric Case #1Figure 16-Refined Mesh (Weld and Lid Interior Region) for Axisymmetric Case #1 CalculationCalculation No.11042-0205Revision No.
3Page57of 90Figure 17-Mesh Details at the Welds for Axisymmetric Case #2Figure 18-Flaw Locations for Axisymmetric Case #2 CalculationCalculation No.11042-0205Revision No.
3Page58of 90Figure 19-Overview of the Half-Symmetry Model CalculationCalculation No.11042-0205Revision No.
3Page59of 90(a) Lid Region Solid View(b) Lid Region Mesh(c) Weld Region Solid View with Flaws Visible(d) Weld Region MeshFigure 20-Detail Views and Mesh Plots of the Half Symmetry Model CalculationCalculation No.11042-0205Revision No.
3Page60of 90Figure 21-Isometric Views of Half-Symmetry Model A AREVA ANSYS 14.0 MAY 4 2015 10:04:39 ELEMENTS Powe rGraphics EFACET=1 TYPE NUM XV =-.373471 YV =-.679348 zv =.631669 *DIST=35.7666 *XF =12.2625 *YF =29.0871 *ZF =167.331 A-ZS=41.1572 Z-BUFFER ANSYS 14.0 MAY 4 2015 10:04:54 ELEMENTS PowerGraphics EFACET=1 TYPE NUM XV =-.373471 YV =-.679348 zv =.631669 *DIST=4.86713 *XF =44.6517 *YF =26.3642 *ZF =167.034 A-ZS=41.1572 Z-BUFFER CalculationCalculation No.11042-0205Revision No.
3Page61of 90Figure 22-Isometric Views of Half-Symmetry Model (Refined Circumferential Mesh)
A AREVA ANSYS 14o0 MAY 4 2015 10:22:15 ELEMENTS PoW'erGraphics TYPE NUM XV YV zv *XF *llo4989 *YF *ZF Z-BUFFER ANSYS 14o0 MAY 4 2015 10:22:36 t>LF.MEN'l'S Pow'"rGraphics I':F'ACE'r*l TYPE NUM XV -o 290468 YV *-o 652082 zv o700299 o 164 25 *XF 43o 55 'YF *ZF 165 o 967 Z-BUFFER CalculationCalculation No.11042-0205Revision No.
3Page62of 90(a) Equivalent Plastic Strain in Weld Region [in/in] at 20 psi Internal Pressure(b) Equivalent Plastic Strain in Weld Region [in/in] at 65 psi Internal Pressure(c) Equivalent Plastic Strain in Weld Region [in/in] at 95.9 psi Internal Pressure(d) EQV Plastic Strain in the Cover Plates at 95.9 psiFigure 23-Results for Axisymmetric Case #1 -Initial Mesh -Service Level A/B(Note that the magnitude of the strains and deflections has no true physical meaning due to the nature of limit load analysis)
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3Page63of 90(a) Service Level A/B Material Properties(b) Service Level D Material PropertiesFigure 24-Deflection History of the Center of the OTCP for the Axisymmetric Case #1 Initial Mesh(Maximum deflection occurs at the center point of the lids, in the outward axial direction)(Note that the magnitude of the deflections has no true physical meaning due to the nature of limit load analysis)
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3Page64of 90(a) Equivalent Plastic Strain in Weld Region [in/in] at 20 psi Internal Pressure(b) Equivalent Plastic Strain in Weld Region [in/in] at 65 psi Internal Pressure(c) Equivalent Plastic Strain in Weld Region [in/in] at 94.8 psi Internal Pressure(d) EQV Plastic Strain in the Cover Plates at 94.8 psiFigure 25-Results for Axisymmetric Case #1 -Refined Mesh in Weld Region -Service Level A/B(Note that the magnitude of thestrains anddeflections has no true physical meaning due to the nature of limit load analysis)
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3Page65of 90(a) Equivalent Plastic Strain in Weld Region [in/in] at 20 psi Internal Pressure(b) Equivalent Plastic Strain in Weld Region [in/in] at 65 psi Internal Pressure(c) Equivalent Plastic Strain in Weld Region [in/in] at 93.7 psi Internal Pressure(d) EQV Plastic Strain in the Cover Plates at 93.7psiFigure 26-Results for Axisymmetric Case #1 -Refined Mesh in Weld and Lid Interior Region -Service Level A/B(Note (c) and (d) are plotted one timestep before the collapse pressure)(Note that the magnitude of thestrains anddeflections has no true physical meaning due to the nature of limit load analysis)
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3Page66of 90(a) Service Level A/B Material PropertiesRefined Mesh at Weld Region Only(b) Service Level A/B Material PropertiesRefined Mesh at the Weld and Lid Interior RegionsFigure 27-Deflection History of the Center of the OTCP for the Axisymmetric Case #1 Refined Mesh(Maximum deflection occurs at the center point of the lids, in the outward axial direction)(Note that the magnitude of the deflections has no true physical meaning due to the nature of limit load analysis)
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3Page67of 90Figure 28-Comparison of Maximum Displacement Histories for Axisymmetric Model Sensitivity Studies(Maximum deflection occurs at the center point of the lids, in the outward axial direction)(Note that the magnitude of the deflections has no true physical meaning due to the nature of limit load analysis)(Service Level A/B material Properties)
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3Page68of 90Figure 29-Comparison of Maximum Displacement Histories for Axisymmetric Model with Lid Contact Defined using Nodal DOF Couples vs. Contact Elements(Maximum deflection occurs at the center point of the lids, in the outward axial direction)(Note that the magnitude of the deflections has no true physical meaning due to the nature of limit load analysis)(Service Level A/B material properties)
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3Page69of 90Figure 30-Comparison of Maximum Displacement Histories for Axisymmetric Model With and Without Pressure Loading Applied to the ITCP Weld Root Flaw Faces(Maximum deflection occurs at the center point of the lids, in the outward axial direction)(Note that the magnitude of the deflections has no true physical meaning due to the nature of limit load analysis)(Service Level A/B material properties)
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3Page70of 90(a) Equivalent Plastic Strain in Weld Region [in/in] at20 psi Internal Pressure(b) Equivalent Plastic Strain in Weld Region [in/in] at 65 psi Internal Pressure(c) Equivalent Plastic Strain in Weld Region [in/in] at 93.6psi Internal Pressure(d) EQV Plastic Strain in the Cover Plates at 93.6psiFigure 31-Results for Axisymmetric Case #2 -Refined Mesh in Weld and Lid Interior Region -Service Level A/B(Note (c) and (d) are plotted one timestep before the collapse pressure)(Note that the magnitude of thestrains anddeflections has no true physical meaning due to the nature of limit load analysis)
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3Page71of 90(a) Equivalent Plastic Strain in Weld Region [in/in] at 20 psi Internal Pressure(b) Equivalent Plastic Strain in Weld Region [in/in] at 65 psi Internal Pressure(c) Equivalent Plastic Strain in Weld Region [in/in] at 94.0 psi Internal Pressure(d) EQV Plastic Strain in the Cover Plates at 94.0psiFigure 32-Results for Axisymmetric Case #0 -Refined Mesh in Weld and Lid Interior Region -Service Level A/B(Note (c) and (d) are plotted one timestep before the collapse pressure)(Note that the magnitude of thestrains anddeflections has no true physical meaning due to the nature of limit load analysis)
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3Page72of 90Figure 33-Comparison of Maximum Center-of-Lid Displacement Histories for the Various Flaw Models(Service Level A/B material properties)(Note that the magnitude of the deflections has no true physical meaning due to the nature of limit load analysis)
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3Page73of 90(a) Equivalent Plastic Strain in Weld Region [in/in] at 20 psi Internal Pressure(b) Equivalent Plastic Strain in Weld Region [in/in] at 65 psi Internal Pressure(c) Equivalent Plastic Strain in Weld Region [in/in] at 95psi Internal Pressure(d) EQV Plastic Strain in the Cover Plates at 95psiFigure 34-Results for Half-Symmetry Case #1 Internal Pressure Loading Benchmark Analysis -Service Level A/B(Note that the magnitude of thestrains anddeflections has no true physical meaning due to the nature of limit load analysis)
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3Page74of 90Figure 35-Benchmark ofthe Half Symmetrymodel with theAxisymmetric Analysis(Service Level A/B material properties)(Note that the magnitude of the deflectionshas no true physical meaning due to the nature of limit load analysis)
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3Page75of 90(a) Equivalent (von Mises) Stress[psi] at 75g(b) Equivalent Plastic Strain in Weld Region [in/in] at 75g.(c) Equivalent (von Mises) Stress [psi] at 181g(d) Equivalent Plastic Strain in Weld Region [in/in] at 181g.Figure 36-Equivalent Stress and Plastic Strain Plots from the Half-Symmetry #1 Side Drop Analysis(Note that the magnitude of thestrains anddeflections has no true physical meaning due to the nature of limit load analysis)
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3Page76of 90(a) Deformed Shape Plot -Axial View -Exaggerated Scale(b) Deformed Shape Plot of DSC Shell -Axial View -Exaggerated ScaleFigure 37-Additional Results Plots fromthe Half-Symmetry #1 Side Drop Analysis(Note that the magnitude of the deflections has no true physical meaning due to the nature of limit load analysis)
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3Page77of 90(a) Equivalent (von Mises) Stress [psi] at 75g(b) Equivalent Plastic Strain in Weld Region[in/in] at 75g.(c) Equivalent (von Mises) Stress [psi] at 181g(d) Equivalent Plastic Strain in Weld Region [in/in] at 181g.Figure 38-Equivalent Stress and Plastic Strain Plots from the Half-Symmetry #1 Side Drop Analysis with Off-Normal Internal Pressure(Note that the magnitude of thestrains anddeflections has no true physical meaning due to the nature of limit load analysis)
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3Page78of 90(a) Equivalent (von Mises) Stress [psi] at 75g(b) Equivalent Plastic Strain in Weld Region [in/in] at 75g.(c) Equivalent (von Mises) Stress [psi] at 185g(d) Equivalent Plastic Strain in Weld Region [in/in] at 185g.Figure 39-Equivalent Stress and Plastic Strain Plots from the Half-Symmetry #1 SideDrop Analysis with Refined Circumferential Mesh(Note that the magnitude of thestrains anddeflections has no true physical meaning due to the nature of limit load analysis)
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3Page79of 90(a) Equivalent (von Mises) Stress [psi] at 75g(b) Equivalent Plastic Strain in Weld Region [in/in] at 75g.(c) Equivalent (von Mises) Stress [psi] at 189g(d) Equivalent Plastic Strain in Weld Region [in/in] at 189g.Figure 40-Equivalent Stress and Plastic Strain Plots from the Half-Symmetry #0 (No Flaws) Side Drop Analysis(Note that the magnitude of thestrains anddeflections has no true physical meaning due to the nature of limit load analysis)
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3Page80of 90Figure 41-Comparison of Maximum Displacement Histories for the Various Half-Symmetry Analyses(Service Level D material properties)(Note that the magnitude of thestrains anddeflections has no true physical meaning due to the nature of limit load analysis)
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3Page81of 9010.0APPENDIX A -ELASTIC-PLASTIC ANALYSESPurposeThe purpose of this appendix is to document elastic-plastic analyses of DSC-16. The models listed below are used as a basis for the analyses. Axisymmetric 1 with Refined Welds and Lid MeshHalf Symmetry 1 with Initial MeshThese models produced the bounding results using the limit load methodology. The models are updated to include the elastic-plastic material properties described below. In addition, these new runs consider the effects of large-deformations (NLGEOM,ON). The intent of these analyses is to provide a more realistic prediction of the actual material strains that would occur under the design basis loading, as opposed to the over-estimated strains and deformations which result from the limit-load analysis methodology.Material PropertiesThe elastic-plastic behavior of SA-240 Type 304stainlesssteel is idealized using Ramberg-Osgood stress-strain curve constants calculated using the equations in Appendix B of Reference A1 2.The constants are calculated using the ASME code[Ref.5.16]specified minimumyield and ultimate strength values at the applicable temperatures. In order to incorporate the curves into the ANSYS analysis, the initial slope of the curves must match the defined elastic modulus. Therefore the first data point in the curves is defined at the (strain,stress) data point (S y/E,Sy).The material behavior is based on true stress and true strain, since the ANSYS analysis accounts for changes in geometry (e.g. necking).The following equations from Reference A1 were used to develop the curves:
2The equations in Reference A1 to develop the full-range true stress-strain curve are based on curve fits of tensile test data. The resulting curve is not indicative of a specific failure type or analysis approach. Rather, it is a method to develop a full-range stress-strain curve of a material using a limited set of data (i.e. minimum specified yield and ultimate strengths.)
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3Page82of 90The value of e uis taken as 0.35, which is assumed to be etot-0.05, where etotis taken as the minimum specified elongation of the material (40%), per Reference A2. The relationship between true and engineering stress and strain is per the following equations:Figure A-1 shows both the true and engineering stress strain curves based on the Ramberg-Osgood equations. The curves at various temperatures as coded into the ANSYS analysis are shown in Figure A-2.The SA-36 shield plugs use a bi-linear stress strain curve with a tangent modulus of 1% of the initial elastic modulus.This results in a less stiff representation of the shield plug, which will result in conservatively greater strains in the DSC.Load CasesAnalyses are performed for the following load cases:1.Internal pressure loading (32 psi) for Service Level A/B.2.Internal pressure loading (65 psi) for Service Level D.3.Side drop Loading (75g) for Service Level D.As discussed in Section 4.2, these three load cases bound all of the design loading conditions for the DSC OTCP and ITCP welds.Resultsand ConclusionPlots of the equivalent plastic strain for the three analyses are shown in Figure A-3through A-5. The results are summarized in Table A-1.As shown by the results, the strain levels remain well below the minimum specified elongation limits of Type 304 steel and Type 308 weld electrodes [Ref. A2 and A3]. Therefore, material rupture will not occur at the design conditions. The maximum strains at loads up to 1.5x the specified loading are also extracted. These results are shown in Table 7, which also includes a comparison of the peak strain values to the ductility limit of the material reduced by the weld uncertainty factor of 0.8 discussed in Section 3.4. See Section 7.0 for further discussion and conclusions.
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3Page83of 90ReferencesA1.EPRI NP-5531. Evaluation of High-Energy Pipe Rupture Experiments. January 1988.A2.ASME Section II Part A. Ferrous Material Specifications. 1998 Edition with Addenda through 2000.
A3.ASMESection II Part C. Specifications for Welding Rods, Electrodes, and Filler Metals. 1998 Edition with Addenda through 2000.Computer FilesAnalyses performed on Computer HEA-0213A using ANSYS Version 14.0 [Ref. 5.6]File date & time listing is as displayed by the Windows 7 Operating System -Differences may occur due to local time zone and daylight savings settings.Analysis CaseFile NameDate & TimeElastic-PlasticAxisymmetric 1Refined Lids and WeldsInternal PressureSL A/B61BTH_WeldFlaw_1F_AX_2_DETACH.db11/29/2015 8:41 AM61BTH_WeldFlaw_1F_AX_2_DETACH.rst11/29/2015 8:18 AMSOLUTION_AXISYMM_IP_500F.INP11/27/2015 4:03PMElastic-PlasticAxisymmetric 1Refined Lids and WeldsInternal PressureSL D61BTH_WeldFlaw_1F_AX_2_DETACH.db11/27/2015 3:36 PM61BTH_WeldFlaw_1F_AX_2_DETACH.rst11/27/2015 3:33 PMSOLUTION_AXISYMM_IP_625F.INP11/19/2015 11:46 AMElastic-PlasticHalf-Symmetry 1Initial MeshSide DropSL D61BTH_WeldFlaw_1GC.db11/29/2015 8:16AM61BTH_WeldFlaw_1GC.rst11/27/2015 6:26PMSOLUTION_HALFSYM_SD.INP11/20/2015 10:08AMStress-Strain Curve DevelopmentStress-Strain.xls11/30/2015 10:59 AM CalculationCalculation No.11042-0205Revision No.
3Page84of 90Table A-1-Summary of Elastic-Plastic Analysis Results.Analysis CaseResultValue[in/in]Internal PressureService Level AAxisymmetric(Note 1)Equivalent Plastic Strain at 32 psi Internal Pressure(Note 1)0.0183(1.83%)Internal PressureService Level DAxisymmetricEquivalent Plastic Strain at 65 psi Internal Pressure0.0597(5.97%)Side DropService Level DHalf-SymmetryEquivalent Plastic Strain at 75g Acceleration0.0609(6.09%)Note 1: The 32 psi internal pressure is bounding for Service Levels A and B and includes designinternal pressure of 10 psi plusan additional 22 psi to account for inertial loading of the DSC contents onto the lid. See Section 4.2 for details.
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3Page85of 90Figure A-1-Ramberg-Osgood Derived Stress Strain Curve for SA-240 Type 304 at 500 oF.
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3Page86of 90Figure A-2-Ramberg-Osgood Stress Strain Curvesfor SA-240 Type 304from ANSYS Modelat Various Temperatures.Upper image shows full range of curves(strain axis from 0 to 40%), lower image shows detail of the initial yield point (strain axis from 0 to 0.2%)
A AREVA SIG SIG 900 800 700 zoo 100 (XlO.Ul) 2500 2250 2000 KUIH for 1 .08 .16 .24 .32 .4 .04 .12 .2 .28 .36 EI?S KINH Table For Matcr1al 1 L..-------------------
(*10"-3) .4 .2 .8 l EPS 1.6 1.4 LA ANSVS 14.0 NOV 27 2015 15:25:00 Tabla Data Tl* 500.00 T2* WO.OO n-10o.oo zv -1 OI$7=.75 XF 5 YF *.5 ZF "7.-BUFFER ANSYS 14.0 NOV 27 2015 15:26:05 Table Data Tl= soo. 00 T: 600.00 *r-::-100. oo zv *1 OT$'1'=.75 XE' VE' =.5 ZF =.5 Z-BUFFER CalculationCalculation No.11042-0205Revision No.
3Page87of 90Figure A-3-Service Level A Internal Pressure -Equivalent Plastic Strain at 32 psi*Note*Note: The 32 psi internal pressure is bounding for Service Levels A and B and includes design internal pressure of 10 psi plus an additional 22 psi to account for inertial loading of the DSC contents onto the lid. See Section 4.2 for details.
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3Page88of 90Figure A-4-Service Level D Internal Pressure -Equivalent Plastic Strain at 65 psi CalculationCalculation No.11042-0205Revision No.
3Page89of 90Figure A-5-Service Level D Side Drop -Equivalent Plastic Strain at 75g.Upper image shows all DSC components, lower image is without shell to allow viewofthe weld surface. The peak strain of 6.09% occurred on the surface of the shell. Therefore, when the shell was removed for the lower image, the peak strain reported reduced to 5.49%.
A AREVA ANSYS 14.0 NOV 29 2015 08:15:52 NODAL SOLUTION SUB EPPLEQV (AVG) PowerGraphics DMX SMX 0 -.006768 c::J . 013536 CJ .020304 CJ .027072 CJ .03384 CJ .040608 CJ .047376 CJ .054144 -.060912 ANSYS 14.0 NOV 29 2015 08:16:20 NODAL SOLUTION SUB EPPLEQV (AVG) PowerGraphics DMX SMX 0 -.006095 liliill CJ D c:::J CJ CJ CJ -.01219 .018284 .024379 .030474 .036569 .042664 .048759 .054853 CalculationCalculation No.11042-0205Revision No.
3Page90of 90Figure A-6-Service Level D Internal Pressure-Equivalent Plastic Strain at 100 psi.Figure A-7-Service Level D Side Drop -Equivalent Plastic Strain at 112.5g.
RaheelHaroon 8/9/2017 CalculationCalculation No.11042-0207Revision No.
0Page2of 34REVISION SUMMARYRev.DescriptionAffectedPagesAffectedFiles0Initial issue AllAll CalculationCalculation No.11042-0207Revision No.
0Page3of 34TABLE OF CONTENTS Page1.0 PURPOSE...........................................................................................................................................5 2.0 ASSUMPTIONS..................................................................................................................................5 3.0 DESIGN INPUT/DATA........................................................................................................................5 3.1 Flaw Details and Geometry........................................................................................................5 3.2 Material Properties.....................................................................................................................5 3.3 Design Criteria...........................................................................................................................5 4.0 METHODOLOGY................................................................................................................................6 4.1 Analysis Method and Acceptance Criteria..................................................................................6 4.2 FEA Model Details......................................................................................................................6 4.3 Limit Load Solution Details.........................................................................................................6 4.4 Elastic Plastic Solution Details....................................................................................................6 4.5 Load Cases................................................................................................................................6
5.0 REFERENCES
....................................................................................................................................8 6.0 ANALYSIS AND RESULTS.................................................................................................................8 6.1 LIMIT LOAD ANALYSIS............................................................................................................
.8 6.1.1 2D-Axisymmetric Analyses for Internal Pressure............................................................8 6.1.2 3D-Half Symmetric Analyses for Side Drop Loading.......................................................8 6.2 ELASTIC-PLASTIC ANALYSIS..................................................................................................8 6.2.1 2D-Axisymmetric Analyses for Internal Pressure............................................................8 6.2.2 3D-Half Symmetric Analyses for Side Drop Loading.......................................................9 7.0 DISCUSSION AND CONCLUSIONS...................................................................................................9 8.0 LISTING OF COMPUTER FILES......................................................................................................
10 APPENDIX A..............................................................................................................................................31 LIST OF TABLESPageTable 1 -Internal Pressure in the 61BTH Type 1 DSC (Ref. [5.3]).............................................................11 Table 2 -Maximum Temperatures in the 61BTH Type 1 DSC Shell (Ref. [5.3]).........................................11 Table 3 -Properties of SA-240 Type 304. Ref. [5.3]...................................................................................12 Table 4 -Properties of SA-36. Ref. [5.3]...................................................................................................13 Table 5 -Summary of Limit Load Analysis for the maximum weld flaws.....................................................14 Table 6 -Summary of Peak Strain Values for Elastic-Plastic Analyses for the maximum weld flaws..........14 Table 7 -Summary of Elastic-Plastic Analysis Results for the maximum weld flaws..................................15
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0Page4of 34LIST OF FIGURESPageFigure 1 -Weld Flaws in Original Model (Ref. [5.3])...................................................................................16 Figure 2 -Maximum Weld Flaws based on the allowed design limits.........................................................16 Figure 3 -Overview of the 2D-Axisymmetric Model....................................................................................17 Figure 4 -Mesh Details at the Welds for 2D-Axisymmetric Model..............................................................17 Figure 5 -Flaw Locations for 2D-Axisymmetric Model...............................................................................18 Figure 6 -Overview of the 3D-Half-Symmetric Model................................................................................19 Figure 7 -Detail Views and Mesh Plots of the 3D-Half Symmetric Mo del...................................................20 Figure 8 -Isometric Views of 3D-Half-Symmetric Model............................................................................21 Figure 9 -Results of Limit Load for 2D-Axisymmetric Model -Service Level A/B......................................22 Figure 10 -Results of Limit Load for 2D-Axisymmetric Model -Service Level D.......................................23 Figure 11 -Deflection at the Center of the OTCP for the 2D-Axisymmetric Model for Limit Load...............24 Figure 12 -Equivalent Plastic Strain at 32 psi for 2D-Axisymmetric Elastic Plastic Analysis -SL A/B Internal Pressure..............................................................................................................................25 Figure 13 -Equivalent Plastic Strain at 65 psi for 2D-Axisymmetric Elastic Plastic Analysis -SL D Internal Pressure..............................................................................................................................26 Figure 14 -Equivalent Plastic Strain at 100 psi for 2D-Axisymmetric Elastic Plastic Analysis -SL D Internal Pressure..............................................................................................................................27 Figure 15 -Equivalent Plastic Strain Plots for 3D-Half-Symmetric Limit Load Analysis -SL D Side Drop with Off-Normal Internal Pressure...............................................................................................28 Figure 16 -Equivalent Plastic Strain at 75g for 3D-Half-Symmetric Elastic-Plastic Analysis -SL D Side Drop 29 Figure 17 -Equivalent Plastic Strain at 112.5g for 3D-Half-Symmetric Elastic-Plastic Analysis -SL D Side Drop....................................................................................................................................30
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0Page5of 341.0PURPOSEThe purpose of this calculation is to evaluate NUHOMS 61BTH Type 1 (DSCs 11-15)at the Monticello Nuclear Generating Plant (MNGP) per ASME Section III criteria withthemaximum flaws in the Inner and Outer Top Cover Plates(ITCP and OTCP) closure weldsbased on the evaluation performed in the reference calculation [5.3].2.0ASSUMPTIONS1.Assumptions 1 through 6 of Ref.[5.3]are applicable tothis calculation.2.The flaws (at the same locations as Ref. [5.3]) are allowed to be increased until the design limits criteria are reached.3.The DSC design in this calculation is typical of MNGP DSCs 11-16, and the modeled baseline flaws are representative of those indications identified by Phased Array Ultrasonic examination (PAUT) of DSC 16 (performed in 2015).3.0DESIGN INPUT/DATA3.1Flaw Details and GeometryTwo casesof flawsare describedand analyzed in Ref.[5.3]. The ITCP weld flaw is the same for both cases,and OTCP increased weld flaw covers both sets (case #1 & case #2 weld flaws). TheresultsofLimit load for both cases are very similar.Figure 1showsOTCP & ITCP flaws in the reference model (Flaw case#1 and Flaw case#2) andFigure 2shows maximized OTCP & ITCP flaws evaluated in this calculation.3.2Material Properties Thematerial properties for the DSC structure are identicalto Ref.[5.3].They are duplicated here in Table 3andTable 4.3.3Design CriteriaAll of the applicable design bases loading conditions areconsidered in accordance with the requirements of ASME Section III Subsection NB Ref. [5.2]. Section 4.1details the methods used to perform the code Ref. [5.2]qualifications.The uncertainties in the PAUT examination are accounted for by using a 0.8 reduction factor on the limit load. This factor is in agreement with ISG-15, conservatively accounts for any additional limitations in the PAUT examinations. This weld uncertainty factor of 0.8 is applied to the minimum of the ASME specified minimum elongation of SA-240 304 (40%) and E308-XX (35%). Therefore strain limit is taken as 0.8*35=28% Ref.[5.3].
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0Page6of 344.0METHODOLOGY4.1Analysis Method and Acceptance CriteriaThe analysis methods, finite element models details and acceptance criteria are the same as discussed tin Ref. [5.3]. The ITCP and OTCP weld flaws are maximized and analyzed per Limit load and Elastic Plastic analyses.Initial ANSYS finite element iterations were performed by increasing all the four flaws by a very small length resulting in a negligible increase in plastic strain. In the second step very large flaws where considered (leaving only one element of the model connected at each flaw) resulting inexcessive strain for the elastic-plastic side drop analysis (Section 4.4).Similarly, few more iterations were performed such that the weld flaw reaches close to acceptable strain limit for the elastic-plastic side drop analysis.Only the final flaw configuration (see Figure 2) is presented in the document.4.2FEA Model DetailsFinite element models of the top half of the 61BTH DSC are used based on Ref.[5.3]. The models fall into two basic categories: axisymmetric (2D) and half-symmetric (3D).The original evaluation in Ref. [5.3] uses ANSYS 14.0. The evaluation in this calculation uses ANSYS 17.1 Ref. [5.1]. APPENDIX Aperforms the sensitivity analysis between the 2 ANSYS versions. As discussed in APPENDIX A, the default ANSYS 17.1 contacts stiffness's for the 3D-Half-Symmetric model were modified to match the default ANSYS 14.0 stiffness's.The models were modified to increase the weld flaws as described in Section 4.1.AxisymmetricModel (2D)An axisymmetric model is used as described in Section 4.3.1 of Ref.[5.3]. Figure3to Figure 5show images of the axisymmetric modelwith maximum flaws.Half-Symmetric Model (3D)A half-symmetric model is used as described in Section 4.3.4 of Ref.[5.3]. Figure 6toFigure 8show images of the half-symmetricmodelwith maximum flaws.4.3Limit Load Solution DetailsLimit load solution details are the same as detailed in Section 4.4 of Ref.[5.3].4.4ElasticPlasticSolution DetailsElastic Plasticsolution details are the same as detailed in Appendix-A of Ref.[5.3].4.5Load CasesThe analyses performed in this calculation, are based on the conservative design values for internal pressure loading, rather than the actual calculated values of internal pressure. Table 1summarizes the conservative design values as well asthe actual calculated values which are taken from Ref. [5.3].
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0Page7of 34Temperatures used for the material properties for each Service Level condition are listed in Table 2.Four 2D-Axisymmetric analyses for bounding Service Level (SL) A/B and D, and two 3D-Half-Symmetric analyses for bounding SL D are performed in this calculation.
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0Page8of 3
45.0REFERENCES
5.1.ANSYS Version 17.1. ANSYS Inc. (Including the ANSYS Mechanical APDL Documentation).5.2.ASME Boiler and Pressure Vessel Code,Section III Subsection NB. 1998 Edition with Addenda through 2000.5.3.AREVA Document No. 11042-0205 Revision 3. "61BTH ITCP and OTCP Closure Weld Flaw Evaluation"5.4.ASME Section II Part A. Ferrous Material Specifications. 1998 Edition with Addenda through 2000.5.5.ASME Section II Part C. Specifications for Welding Rods, Electrodes, and Filler Metals1998Editionwith Addenda through 20006.0ANALYSISAND RESULTS6.1LIMIT LOAD ANALYSIS6.1.12D-Axisymmetric Analyses for Internal PressureTwo analyses are performed with the 2D-Axisymmetric model: one case forService Level A/B and the othercase forService Level D. The collapse pressures were determined to be 86.3psi for Service Level A/B and 122.2 psi for Service Level D.Figure 9shows various plots of the plastic strainfor Service Level A/B atvarious locations and levels of loading. Figure 10shows various plots of the plastic strainfor Service Level D.These strain plots are also representative of the behavior of the Service Level D analysis. Figure 11showsthe deflection history at the center of the lid, and indicates the expected plastic instability that occurs as the limit load is approached. Note that both the strains and displacements presented in these figures show only the load (pressure) at which the solution fails to converge.6.1.23D-Half SymmetricAnalyses for Side Drop LoadingThe 3D-half-symmetricmodel described in Section4.2is used to perform the side-drop limit load analysis.
The case includestheside-drop acceleration loading of 75g as well as the off-normal internal pressure of 20 psi. The collapse g-load for side-drop loading was found to be approximately 179.5g.Plotsof the plastic strainsin the side drop analyses are shown inFigure 15.The resultsfor Limit load analysisare summarized in Table 5.6.2ELASTIC-PLASTIC ANALYSIS6.2.12D-Axisymmetric Analyses for Internal PressureTwo analyses are performed with the 2D-Axisymmetric model:one case for Service Level A/B and the other case for Service Level D. The Equivalent Plastic Strain wasdetermined to be 3.1%forService Level A/B pressure and7.4%for Service Level Dpressure.Figure 12showsplot of the plastic strain for Service Level A/B.Figure 13shows plot of the plastic strain for Service Level D.The results for elastic-plastic analyses are CalculationCalculation No.11042-0207Revision No.
0Page9of 34summarized inTable 7.As shown by the results, the strain levels remain well below the minimum specified elongation limits of Type 304 steel and Type 308 weld electrodes Ref.[5.4]and Ref. [5.5]. Therefore,material rupture will not occur at the design conditions.The maximum strains at loads up to 1.5x the specified loading are also extracted. These results are shown in Table 6, which also includes a comparison of the peak strainvalues to the ductility limit of the material reduced by the weld uncertainty factor of 0.8 discussed in Section 3.4of Ref.[5.3].6.2.23D-Half SymmetricAnalyses for Side Drop LoadingThe 3D-half-symmetric model described in Section 4.2is used to perform the SL D side-drop limit load analysis. Thecase includes the 75g side-drop acceleration loading only. The maximum strains at loads up to 1.5x the specified loading (112.5g) are also extracted and compared with the material strain limit. The equivalent plastic strainwas determined to be 11.1% for75g and 23.0% for112.5gpresentedinTable 6.Figure 16and Figure 17showthe corresponding plastic strain plots.The results for elastic-plastic analyses are summarized inTable 7.7.0DISCUSSION AND CONCLUSIONSThis calculation qualifies the NUHOMS 61BTH Type 1 (DSCs 11-15)at the Monticello Nuclear Generating Plantwith maximum weld flaw using a combination of limit load analyses and elastic-plastic analyses. The limit load analyses are used to show that the DSC satisfies the primary stress limits of ASME Section III Subsection NB. The elastic-plasticanalyses are used to show that the actual predicted strain values are below the material ductility limits. Both the limit load and elastic-plastic analyses account for any remaining uncertainty in the weld (e.g. non-inspected weld regions and PAUT technique limitations) by including an uncertainty factor of 0.8 which is described in detail in Section 3.4 of Ref.[5.3].For both OTCP and ITCP, all weld flaws were maximized such that the weld flaw reaches close to acceptable design limits. The maximum modeled weld flaws for OTCP to DSC shell weld are 0.43" and 0.42" in length, which represents about 85% through-wall of the 0.5-inch minimum weld throat. The maximum modeled full-circumferential weld flaws for ITCP to DSC shell weld are 0.16"
- cos(45°)=0.11" and 0.14" in length, which represents respectively 58% and 74% through-wall of the 0.19-inch minimum weld throat as shown in Figure 2.All four assumed flaws represent defects spreading over more than one weld bead. These flaws were located based on DSC #16 PAUT results and are considered representative locations for DSC's # 11 to 15.
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0Page10of 348.0LISTING OF COMPUTER FILESFinite Element Analyseswereperformed using ANSYS Version 17.1Ref. [5.1].All analyses were performed on HPC v2 Linux platform.Load CaseAnalysis TypeFile NameDescription Date / Time(1)Internal Pressure 2D-Axisymmetric modelLimit load analysis SL-A/B61BTH_WeldFlaw_1F_AX_2_DETACH.dbReference .db file for Axisymmetric SL-A/B Limit load analysisNote (2)AXISYMM_IP_LimitLoad.ext.ext = .inp, .err, .mntr, .out, .db, .rstLimit load analysis files for SL-A/B06/20/201711:33:31Limit load analysis SL-D61BTH_WeldFlaw_1F_AX_2_DETACH.dbReference .db file forAxisymmetric SL-D Limit load analysis Note (2)AXISYMM_IP_LimitLoad_SLD.ext.ext = .inp, .err, .mntr, .out, .db, .rstLimit load analysis files for SL-D06/20/201712:29:27Elastic-plastic analysis SL-A/B61BTH_WeldFlaw_1F_AX_2_DETACH.dbReference .db file for Axisymmetric SL-A/B Elastic-plastic analysisNote (2)AXISYMM_IP_500F.ext.ext = .inp, .err, .mntr, .out, .db, .rstElastic-plastic analysis files for SL-A/B06/20/201712:34:31Elastic-plastic analysis SL-D61BTH_WeldFlaw_1F_AX_2_DETACH.dbReference .db file for Axisymmetric SL-A/B Elastic-plastic analysisNote (2)AXISYMM_IP_625F.ext.ext = .inp, .err, .mntr, .out, .db, .rstElastic-plastic analysis files for SL-D06/20/201712:39:21Side Drop 3D-Half-Symmetric modelLimit loadanalysis SL-D61BTH_WeldFlaw_1GC.dbReference .db file for half-symmetric limit load analysis Note (2)LIMIT_HALFSYM.ext.ext = .inp, .err, .mntr, .out, .db, .rstunmerge.mac, unmerge2.macLimit load SL D analysis files.06/20/201711:39:15Elastic-plastic analysis SL-D61BTH_WeldFlaw_1GC.dbReference .db file for half-symmetric elastic-plastic analysisNote (2)STRAIN_HALFSYM.ext.ext = .inp, .err, .mntr, .out, .db, .rst unmerge.mac, unmerge2.macElastic-plastic SL D analysis files.06/19/201722:48:11Notes:(1)The date & time (EST) for the main runs are from the listing at the end of output file.
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0Page11of 34(2)ANSYS FE models are taken from Section 8.0 of Ref. [5.3].Table 1-Internal Pressure inthe 61BTH Type 1 DSC(Ref.[5.3])Design ConditionMaximum Calculated Pressure[psi]Design Pressure used in this Calculation[psi]Normal7.310Off-Normal10.920Accident56.165Table 2-Maximum Temperatures in the 61BTH Type 1 DSC Shell(Ref.[5.3])Design ConditionMaximum Calculated Temperature[ºF]Design Temperature used in This Calculation[ºF]NormalStorage374500Transfer439500Off-NormalStorage399500Transfer416500AccidentStorage611625Transfer467500 CalculationCalculation No.11042-0207Revision No.
0Page12of 34Table 3-Properties of SA-240 Type 304. Ref. [5.3]Temp [oF] E Modulus of Elasticity
[ksi] Sm Allowable Stress Intensity
[ksi] Sy Yield Stress
[ksi] Su Ultimate Tensile Strength
[ksi] Yield Stress for SL A/B Limit Load Analysis
[ksi] Yield Stress for SL D Limit Load Analysis
[ksi] 70 28,300 20.0 30.0 75.0 30.0 46.0 100 28,138 20.0 30.0 75.0 30.0 46.0 200 27,600 20.0 25.0 71.0 30.0 46.0 300 27,000 20.0 22.4 66.2 30.0 46.0 400 26,500 18.7 20.7 64.0 28.1 43.0 500 25,800 17.5 19.4 63.4 26.3 40.3 600 25,300 16.4 18.4 63.4 24.6 37.7 625 25,175 16.3 18.2 63.4 24.5 37.5 700 24,800 16.0 17.6 63.4 24.0 36.8 CalculationCalculation No.11042-0207Revision No.
0Page13of 34Table 4-Properties of SA-36. Ref. [5.3]Temp [oF] E Modulus of Elasticity
[ksi] Sm Allowable Stress Intensity
[ksi] Sy Yield Stress
[ksi] Su Ultimate Tensile Strength
[ksi] Yield Stress for SL A/B Limit Load Analysis
[ksi] Yield Stress for SL D Limit Load Analysis
[ksi] 70 29,500 19.3 36.0 58.0 29.0 40.6 100 29,338 19.3 36.0 58.0 29.0 40.6 200 28,800 19.3 33.0 58.0 29.0 40.6 300 28,300 19.3 31.8 58.0 29.0 40.6 400 27,700 19.3 30.8 58.0 29.0 40.6 500 27,300 19.3 29.3 58.0 29.0 40.6 600 26,700 17.7 27.6 58.0 26.6 40.6 625(1) 26,400 17.6 27.2 58.0 26.4 40.4 700 25,500 17.3 25.8 58.0 26.0 39.8 Note:(1)All values are interpolated from the 600 oF and 700 oF values.
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0Page14of 34Table 5-Summary of LimitLoad Analysisfor the maximum weld flawsSl. No.NameLoadingTemp.F]Analysis CriteriaDesign Pressure(psi)Requirement of pressure to Safety Limit load Criteria(psi)Limit LoadCollapse Pressure(psi)12D-AxisymmetricInternal pressure500SL A/B326086.3 22D-AxisymmetricInternal pressure625SL D6590.2122.2Sl. No.NameLoadingTemp.F]Analysis CriteriaDesign G-load(g)Required G-load to Satisfy Limit load Criteria(g)Limit LoadCollapse G-Load(g)33D-Half-symmetricSide dropwith 20psi off-normal IP500SL D75104179.5 (1)Note: (1)To be compared with 188.5g with the original Case #1 weld flaws of Ref. [5.3], see APPENDIX ATable 6-Summary ofPeak Strain Values for Elastic-Plastic Analyses for the maximum weld flawsLoad Case Specific loading Internal Pressure (psi) Peak Equivalent Plastic Strain Material Strain Limit(1) at 65 psi internal Pressure at 100 psi internal Pressure 2D-Axisymmetric Internal Pressure Service Level D 65 7.4% 13.6% 28% Load Case Specific loading Side Drop G-Load (g) Peak Equivalent Plastic Strain Material Strain Limit(1) at 75g loading at 112.5g loading 3D-Half-symmetric Side Drop Service Level D 75 11.1% 23.0% 28% Note:
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0Page15of 34(1)Theweld uncertaintyfactor of0.8 (See Section 3.4 of Ref. [5.3]) is applied to the minimum of the ASME specified minimum elongation of SA-240 304 (40%) and E308-XX (35%). Therefore strain limit is taken as 0.8*35=28%-See Section 3.3.Table 7-Summary of Elastic-Plastic Analysis Resultsfor the maximum weld flawsAnalysis CaseResultPlastic StrainInternal PressureService Level A2D-Axisymmetric (1)Equivalent Plastic Strain at 32 psi Internal Pressure(1)3.1%Internal PressureService Level D2D-AxisymmetricEquivalent Plastic Strain at 65 psi Internal Pressure7.4%Side DropService Level D3D-Half-SymmetryEquivalent Plastic Strain at 75g Acceleration11.1%Note:(1)The 32 psi internal pressure is bounding for Service Levels A and B and includes design internal pressure of 10 psi plus an additional 22 psi to account for inertial loading of the DSC contents onto the lid.
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0Page16of 34Flaw Case #1Flaw Case #2Figure 1-Weld Flaws in OriginalModel(Ref. [5.3])Figure 2-Maximum Weld Flaws based on the allowed design limits CalculationCalculation No.11042-0207Revision No.
0Page17of 34Figure 3-Overview of the 2D-Axisymmetric ModelFigure 4-Mesh Details at the Welds for 2D-Axisymmetric Model CalculationCalculation No.11042-0207Revision No.
0Page18of 34Figure 5-Flaw Locations for 2D-Axisymmetric Model CalculationCalculation No.11042-0207Revision No.
0Page19of 34Figure 6-Overview of the 3D-Half-Symmetric Model CalculationCalculation No.11042-0207Revision No.
0Page20of 34(a) Lid Region Solid View(b) Lid Region Mesh(c) Weld Region Solid View with Flaws Visible(d) Weld Region MeshFigure 7-Detail Views and Mesh Plots of the 3D-Half SymmetricModel CalculationCalculation No.11042-0207Revision No.
0Page21of 34Figure 8-Isometric Views of 3D-Half-SymmetricModel CalculationCalculation No.11042-0207Revision No.
0Page22of 34(a) Equivalent Plastic Strain in Weld Region [in/in] at 20 psi Internal Pressure(b) Equivalent Plastic Strain in Weld Region [in/in] at 65 psi Internal Pressure(c) Equivalent Plastic Strain in Weld Region [in/in] at 86.3psi Internal Pressure(d) EQV Plastic Strain in the Cover Plates at 86.3 psiFigure 9-Resultsof Limit Loadfor 2D-Axisymmetric Model-Service Level A/B CalculationCalculation No.11042-0207Revision No.
0Page23of 34(a) Equivalent Plastic Strain in Weld Region [in/in] at 65psi Internal Pressure(b) Equivalent Plastic Strain in Weld Region [in/in] at 100psi Internal Pressure(c) Equivalent Plastic Strain in Weld Region [in/in] at 122.2psi Internal Pressure(d) EQV Plastic Strain in the Cover Plates at 122.2 psiFigure 10-Resultsof Limit Loadfor 2D-Axisymmetric Model-Service Level D CalculationCalculation No.11042-0207Revision No.
0Page24of 34(a) Service Level A/B(b) Service Level DFigure 11-Deflection atthe Centerof the OTCP for the 2D-Axisymmetric Modelfor Limit Load(Maximum deflection occurs at the center point of the lids, in the outward axial direction)
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0Page25of 34Figure 12-Equivalent Plastic Strain at 32 psi for 2D-Axisymmetric Elastic Plastic Analysis-SLA/BInternal Pressure CalculationCalculation No.11042-0207Revision No.
0Page26of 34Figure 13-Equivalent Plastic Strain at 65 psi for2D-Axisymmetric Elastic Plastic Analysis -SLD Internal Pressure CalculationCalculation No.11042-0207Revision No.
0Page27of 34Figure 14-Equivalent Plastic Strain at 100 psi for 2D-Axisymmetric Elastic Plastic Analysis -SL D Internal Pressure CalculationCalculation No.11042-0207Revision No.
0Page28of 34(a) Equivalent Plastic Strain in Weld Region [in/in] at 75g.(b) Equivalent Plastic Strain in Weld Region [in/in] at 179.5g.Figure 15-Equivalent Plastic Strain Plots for3D-Half-SymmetricLimit Load Analysis -SL D Side Drop with Off-Normal Internal Pressure CalculationCalculation No.11042-0207Revision No.
0Page29of 34Figure 16-Equivalent Plastic Strain at 75g for 3D-Half-Symmetric Elastic-Plastic Analysis -SLD Side Drop CalculationCalculation No.11042-0207Revision No.
0Page30of 34Figure 17-Equivalent Plastic Strain at 112.5g for 3D-Half-Symmetric Elastic-Plastic Analysis -SLD Side Drop CalculationCalculation No.11042-0207Revision No.
0Page31of 34APPENDIX ASensitivity Study ofANSYS Release 14.0and 17.1ANSYS computer program Release 14.0 has been used in stress calculation in Ref. [5.3]. ANSYS Release 17.1 is used in this calculation. Release 17.1 Ref. [5.1]was installed in accordance with QAP and TIP3.3 requirements and is verified against empirical Data. The purpose of Appendix Ais to determine the effect of using different releases of ANSYS on the same FE model. The following bounding 3D-half-symmetric load cases from the main part of this document are considered for the sensitivity analysis:1)Elastic-Plastic analysis: Side drop 75g and 112.5g2)Limit Load analysis: Side drop with off-normal internal pressureA.1Elastic-Plastic sensitivity analysisRef. [5.3] Elastic-Plastic analysis on the 3D-half-symmetricFE modeluses ANSYS 14.0 and provides a peak equivalent plastic strain of 6.09% for 75g and 12.6% for 112.5g (Line 1 of Table A-1). The same ANSYS FEmodel was resumed in ANSYS 17.1 Ref. [5.1] and analyzed without any modification. The results for ANSYS 17.1 peak equivalent plastic strainare found to be 5.60% and 11.76% for 75g and 112.5g respectively (Line 2 of Table A-1). The default surface-to-surface contact stiffness's between the two releases are differentandare found to be higher in ANSYS 17.1 resulting in lower equivalent plastic strains. Therefore the contact stiffness's were reduced by a 4.2873factor to match the default surface-to-surface contact stiffness's of ANSYS 14.0. As the contact stiffness coefficient FKN used in ANSYS 14.0is 0.1, the newcontact stiffness coefficient in ANSYS 17.1 is0.1 / 4.2873=0.02332.Once this modification implemented, ANSYS 17.1 provides exactly the same results (Line 3 of Table A-1) as ANSYS 14.0.Table A-1: Comparison ANSYS 14.0vs 17.1-3D-half-symmetric Model -Elastic Plastic analysis Peak Equivalent Plastic Strain Sl. No. Side Drop at 75g at 112.5g 1 ANSYS 14.0 6.09% Table 7 of [5.3
] 12.6% Table 7 of [5.3
] 2 ANSYS 17.1 5.60% 11.76% 3 ANSYS 17.1 modified 6.09% 12.59% A.2Limit Load sensitivity analysisRef. [5.3] Limit Load analysis on the 3D-half-symmetricFE modeluses ANSYS 14.0 and provides a limit load of 180.6g (Line 1 of Table A-2). The same ANSYS FE model was resumed in ANSYS 17.1 Ref. [5.1] and CalculationCalculation No.11042-0207Revision No.
0Page32of 34analyzed without any modification. The result for ANSYS 17.1 limit load is found to be 188.52g (Line 2 of Table A-2). The same contact stiffness's modification described in Section A.1 was implemented for the Limit Load case. However, the limit load stayed identical (188.56g, Line 3 of Table A-2) to the unmodified ANSYS
17.1 result.Table A-2: Comparison ANSYS 14.0vs 17.1-3D-half-symmetric Model -Limit Load analysisSl. No. Side Drop Limit Load Collapse G-Load (g) Loading Temp [°F] Design G-load (g) Required G-load to Satisfy Limit load Criteria (g) 1 ANSYS 14.0 180.6 Table 6 of [5.3] Side drop with off-normal IP 500 75 104 2 ANSYS 17.1 188.52 3 ANSYS 17.1 modified 188.56 Although the ANSYS 17.1 runs converge up to 188.5g instead of 180.6g for ANSYS 14.0, Figure A-1 clearly shows thatthe results (here the maximum displacement in the model) are identical up to the point where ANSYS 14.0 stop converging. The limit load for the Case #1 weld flaws is thus considered to be 188.5g in this calculation and is the reference for comparison with the increased flaws calculation results presented in Table 5.A.3ConclusionBased on the sensitivity evaluations performed in Appendix A, it is concluded that the results are independent of the ANSYS release for the 3D-Half-Symmetric model.
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0Page33of 3400.10.2 0.3 0.4 0.5 0.6 0.7 0.8 0.910102030405060708090100110120130140150160170180190Maximum Displacement (in)Acceleration (g)ANSYS 14.0ANSYS 17.1Figure A-1: Comparison ANSYS 14.0 vs 17.1-Limit Load analysis CalculationCalculation No.11042-0207Revision No.
0Page34of 34A.4Listing of computer filesFinite Element Analyseswereperformed using ANSYS Version 17.1 Ref. [5.1]. All analyses were performed on HPC v2 Linux platform.Load CaseAnalysis TypeFile NameDescription Date / Time(1)Side Drop Half-Symmetric modelInput Identical to Ref [5.3]Limit load analysis SL-D61BTH_WeldFlaw_1GC.dbReference .db file for half-symmetric limit load analysis Note (2)SOLUTION_HALFSYM_SD.INPSOLUTION_HALFSYM_SD.out3D_WeldFlaw.ext .ext = .mntr, .db, .rstLimit load analysis files05/25/201721:46:23Elastic-plastic analysis SL-D61BTH_WeldFlaw_1GC.dbReference .db file for half-symmetric elastic-plastic analysisNote (2)STRAIN_HALFSYM.ext.ext = .inp, .err, .mntr, .out, .db, .rstElastic-plastic analysis files06/07/201716:18:0261BTH_WELDFLAW_MATERIALS_ElasticPlastic_RamOsTrue.INPSide Drop Half-Symmetric modelInput Modified(See Section A-1)Limit load analysis SL-D61BTH_WeldFlaw_1GC.dbReference .db file for half-symmetric limit load analysis Note (2)SOLUTION_HALFSYM_SD.ext .ext=.INP, .out, .err3D_WeldFlaw.ext .ext = .mntr, .db, .rstLimit load analysis files05/28/201705:12:40Elastic-plastic analysis SL-D61BTH_WeldFlaw_1GC.dbReference .db file for half-symmetric elastic-plastic analysisNote (2)SOLUTION_HALFSYM_SD.ext .ext=.INP, .out, .err3D_WeldFlaw.ext .ext = .mntr, .db, .rstElastic-plastic analysis files05/27/201716:39:1161BTH_WELDFLAW_MATERIALS_ElasticPlastic_RamOsTrue.INPNotes:(1)The date & time (EST) for the main runs are from the listing at the end of output file.
(2)ANSYS FE models are taken from Section 8.0 of Ref. [5.3].
08/09/2017RaheelHaroon 8/9/2017 CalculationCalculation No.11042-0208Revision No.
0Page2of 23Rev.DescriptionAffectedPagesAffectedFiles0Initial issue AllAll CalculationCalculation No.11042-0208Revision No.
0Page3of 23TABLE OF CONTENTS Page1.0 PURPOSE...........................................................................................................................................5 2.0 ASSUMPTIONS..................................................................................................................................5 3.0 DESIGN INPUT/DATA........................................................................................................................5 3.1 Flaws Details and Geometry.....................................................................................................
.5 3.2 Material Properties.....................................................................................................................5 3.3 Design Criteria...........................................................................................................................5 4.0 METHODOLOGY................................................................................................................................5 4.1 Analysis Method and Acceptance Criteria..................................................................................5 4.2 FEA Model Details......................................................................................................................6 4.3 Limit Load Solution Details.........................................................................................................6 4.4 Elastic Plastic Solution Details....................................................................................................6 4.5 Load Cases................................................................................................................................6
5.0 REFERENCES
....................................................................................................................................7 6.0 ANALYSIS AND RESULTS.................................................................................................................7 6.1 LIMIT LOAD ANALYSIS............................................................................................................
.7 6.1.1 2D-Axisymmetric Analyses for Internal Pressure............................................................7 6.1.2 3D-Half Symmetric Analyses for Side Drop Loading.......................................................7 6.2 ELASTIC-PLASTIC ANALYSIS..................................................................................................8 6.2.1 2D-Axisymmetric Analyses for Internal Pressure............................................................8 6.2.2 3D-Half Symmetric Analyses for Side Drop Loading.......................................................8 7.0 DISCUSSION AND CONCLUSIONS...................................................................................................8 8.0 LISTING OF COMPUTER FILES........................................................................................................9 LIST OF TABLESPageTable 1 -Internal Pressure in the 61BTH Type 1 DSC...............................................................................10 Table 2 -Maximum Temperatures in the 61BTH Type 1 DSC Shell [5.5]...................................................10 Table 3 -Properties of SA-240 Type 304. Ref. [5.4]...................................................................................11 Table 4 -Properties of SA-36. Ref.[5.4]...................................................................................................12 Table 5 -Summary of Limit Load Analysis for the maximum weld flaws.....................................................13 Table 6 -Summary of Peak Strain Values for Elastic-Plastic Analyses for the maximum weld flaws..........13 Table 7 -Summary of Elastic-Plastic Analysis Results for the maximum weld flaws..................................14
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0Page4of 23LIST OF FIGURESPageFigure 1 -Results of Limit Load for 2D-Axisymmetric Model -Service Level A/B......................................15 Figure 2 -Results of Limit Load for 2D-Axisymmetric Model -Service Level D.........................................16 Figure 3 -Deflection at the Center of the OTCP for the 2D-Axisymmetric Model for Limit Load.................17 Figure 4 -Equivalent Plastic Strain at 29.3 psi for 2D-Axisymmetric Elastic Plastic Analysis -SL A/B Internal Pressure..............................................................................................................................18 Figure 5 -Equivalent Plastic Strain at 45.9 psi for 2D-Axisymmetric Elastic Plastic Analysis -SL D Internal Pressure..............................................................................................................................19 Figure 6 -Equivalent Plastic Strain at 69 psi for 2D-Axisymmetric Elastic Plastic Analysis -SL D Internal Pressure..............................................................................................................................20 Figure 7 -Equivalent Plastic Strain Plots for 3D-Half-Symmetric Limit Load Analysis -SL D Side Drop with Off-Normal Internal Pressure...............................................................................................21 Figure 8 -Equivalent Plastic Strain at 75g for 3D-Half-Symmetric Elastic-Plastic Analysis -SL D Side Drop............................................................................................................................................22 Figure 9 -Equivalent Plastic Strain at 112.5g for 3D-Half-Symmetric Elastic-Plastic Analysis -SL D Side Drop....................................................................................................................................23
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0Page5of 231.0PURPOSEThe purpose of this calculation is to evaluatethe margins forthe NUHOMS 61BTH Type 1 DSCsat the Monticello Nuclear Generating Plant (MNGP) per ASME Section III criteria withthemaximumpostulated flaws in the Inner and Outer Top Cover Plates(ITCP and OTCP) closure weldsbased on the evaluation performed in the reference calculation [5.4].The as-loaded site specific bounding temperatures and pressures used in this calculation are provided in Ref. [5.5].2.0ASSUMPTIONS1.Assumptions 1 through 6 of Ref.[5.3]are applicable tothis calculation.3.0DESIGN INPUT/DATA3.1FlawsDetails and GeometryTheflawsdetailsare identicalto the maximum weld flaws evaluated in Ref.[5.4]. The geometry ofthe DSC structureis identical to the geometry used in References [5.3] and [5.4].3.2Material Properties Thematerial properties for the DSC structure are identicalto the material properties of References [5.3] and
[5.4].They are duplicated here in Table 3andTable 4for convenience.3.3Design CriteriaAll of the applicable design bases loading conditions areconsidered in accordance with the requirements of ASME Section III Subsection NB Ref. [5.2]. Section 4.1details the methods used to perform the code Ref. [5.2]qualifications.4.0METHODOLOGY4.1Analysis Method and Acceptance CriteriaThe analysis methods, finite element model details and acceptance criteria are the same as discussed in Ref. [5.3]. The ITCP and OTCPmaximumweld flaws as evaluated in Ref. [5.4]are analyzedand margins are evaluatedforLimit load and Elastic Plastic analyses.The as-loaded site specific bounding temperatures and pressures used in this calculation are provided in Ref. [5.5].
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0Page6of 234.2FEA Model DetailsFinite element model details ofthe DSC structure are identicalto the ones described in Ref.[5.4].AxisymmetricModel (2D)A2D-axisymmetric model is used as described in Section 4.2 of Ref.[5.4].Half-Symmetric Model (3D)A3D-half-symmetric model is used as described in Section 4.2 of Ref.[5.4].4.3Limit Load Solution DetailsLimit load solution details are the same as detailed in Section 4.4 of Ref.[5.3].4.4Elastic PlasticSolution DetailsElastic Plastic solution details are the same as detailed in Appendix-A of Ref.[5.3].4.5Load CasesThe analyses performed in this calculation are usingvalues of peak accident internal pressurecalculated using the bounding value for actual canister heat load.Table 1summarizes the actual values which are taken fromTable 2 ofRef.[5.3].Peak Accident internal pressure is taken as 45.91 psi from Table 7-5 of Ref.[5.5].Temperatures used for the material properties for each Service Level condition are listed in Table 2.The Maximum Service Level (SL) D temperature of the DSC shell is taken as 370
°F as per Table 7-2 of Ref.[5.5]for a blocked vent accident.The same table also gives the maximum DSC shell SL-B temperature as 237°F, before the blocked vent accident.Four 2D-Axisymmetric analyses for bounding Service Level (SL) A/B and D, and two 3D-Half-Symmetric analyses for bounding SL D are performed in this calculation.
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0Page7of 2
35.0REFERENCES
5.1.ANSYS Version 17.1. ANSYS Inc. (Including the ANSYS Mechanical APDL Documentation).5.2.ASME Boiler and Pressure Vessel Code,Section III Subsection NB. 1998 Edition with Addenda through2000.5.3.AREVA Document No. 11042-0205 Revision 3. "61BTH ITCP and OTCP Closure Weld Flaw Evaluation"5.4.AREVA Document No. 11042-0207Revision 0. "NUHOMS 61BTH Type 1 DSC ITCP and OTCP Maximum Weld Flaw Evaluation"5.5.AREVA Document No. 11042-0400Revision 0."Site-Specific Thermal Evaluation of 61BTH Type 1 DSCsstoredin HSM-H at Monticello Nuclear Generating Plant"5.6.ASME Section II Part A. Ferrous Material Specifications. 1998 Edition with Addenda through 2000.5.7.ASME Section II Part C. Specifications for Welding Rods, Electrodes, and Filler Metals1998Editionwith Addenda through 20006.0ANALYSISAND RESULTS6.1LIMIT LOAD ANALYSIS6.1.12D-Axisymmetric Analyses for Internal PressureTwo analyses are performed with the 2D-Axisymmetric model: one case forService Level A/B and the othercase forService Level D. The collapse pressures were determined to be 98.4psi for Service Level A/B and 144.1psi for Service Level D.Figure 1shows various plots of the plastic strainfor Service Level A/B at various locations and levels of loading. Figure 2shows various plots of the plastic strainfor Service Level D.These strain plots are also representative of the behavior of the Service Level D analysis. Figure 3showsthe deflection history at the center of the lid, and indicates the expected plastic instability that occurs as the limit load is approached. Note that both the strains and displacements presented in these figures show only the load (pressure) at which the solution fails to converge.6.1.23D-Half SymmetricAnalyses for Side Drop LoadingThe 3D-half-symmetricmodel described in Section4.2is used to perform the side-drop limit load analysis.
The case includestheside-drop acceleration loading of 75g as well as the off-normal internal pressure of 10.9psi. The collapse g-load for side-drop loading was found to be approximately 204g.Plotsof the plastic strainsin the side drop analyses are shown inFigure 7.The resultsfor Limit load analysisare summarized in Table 5.
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0Page8of 236.2ELASTIC-PLASTIC ANALYSIS6.2.12D-Axisymmetric Analyses for Internal PressureTwo analyses are performed with the 2D-Axisymmetric model:one case for Service Level A/B and the other case for Service Level D. The Equivalent Plastic Strain wasdetermined to be 2.7%forService Level A/B pressure and4.4%for Service Level Dpressure.Figure 4showsaplot of the plastic strain for Service Level A/B.Figure 5shows aplot of the plastic strain for Service Level D.The results for elastic-plastic analyses are summarized inTable 7.As shown by the results, the strain levels remain well below the minimum specified elongation limits(28%)of Type 304 steel and Type 308 weld electrodes Ref.[5.6]andRef. [5.7]. Therefore, material rupture will not occur at the as loadedconditions.The maximum strains at loads up to 1.5x the specified loading are also extracted. These results are shown in Table 6, which also includes a comparison of the peak strain values to the ductility limit of the material reduced by the uncertainty factor of 0.8due to PAUT examinationdiscussed in Section 3.4of Ref.[5.3].6.2.23D-Half SymmetricAnalyses for Side Drop LoadingThe 3D-half-symmetric model described in Section 4.2is used to perform the SL D side-drop limit load analysis. Thecase includes the 75g side-drop acceleration loading only. The maximum strains at loads up to 1.5x the specified loading (112.5g) are also extracted and compared with the material strain limit.The equivalent plastic strainwas determined to be 9.8% for 75g and 19.0%for112.5g.Figure 8and Figure 9showthe corresponding plastic strain plots.The results for elastic-plastic analyses are summarized inTable 7.7.0DISCUSSION AND CONCLUSIONSLimit Load Analyses:The lower bound collapse pressure for Service Level A/B criteria was found to be 98.4psi which is greater thanthe limiting pressure of 60 psi(Table 5).Therefore the Service Level A/B criterion is satisfied.The lower bound collapse pressure for Service Level D criteria was found to be 144.1psi which is greater than the limiting pressure of 90.2 psi(Table 5). The lower bound collapse G-Load for Service Level D side drop criteria was found to be 204 gwhich is greater than the limiting G-Load of 104 g(Table 5).Therefore the Service Level D criterion is satisfied.Elastic-Plastic Analyses:Table 6lists the peak strains predicted by the elastic-plasticanalyses for the bounding Service Level D event. As shown in the table, the peak strain values remain below the material ductility limits(28%)at the specified loading conditions, and also at 1.5x the specified loads, with a minimum margin of safety of 1.86. Therefore the elastic plastic analysescriteria are satisfied.
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0Page9of 238.0LISTING OF COMPUTER FILESFinite Element Analyseswereperformed using ANSYS Version 17.1Ref. [5.1].All analyses were performed on HPC v2 Linux platform.Load CaseAnalysis TypeFile NameDescription Date / Time(1)Internal Pressure 2D-Axisymmetric modelLimit load analysis SL-A/BAXISYMM_IP_LimitLoad.dbReference .db file for Axisymmetric SL-A/B Limit load analysis Note (2)AXISYMM_IP_LimitLoad-237.ext
.ext = .inp, .err, .mntr, .out, .db, .rstLimit load analysis files for SL-A/B06/27/2017 16:12:31Limit load analysis SL-DAXISYMM_IP_LimitLoad_SLD.dbReference .db file for Axisymmetric SL-D Limit load analysis Note (2)AXISYMM_IP_LimitLoad_SLD-370.ext
.ext = .inp, .err, .mntr, .out, .db, .rstLimit load analysis files for SL-D06/27/2017 13:00:25Elastic-plastic analysis SL-A/BAXISYMM_IP_500F.dbReference .db file for Axisymmetric SL-A/B Elastic-plastic analysis Note (2)AXISYMM_IP-237F.ext
.ext = .inp, .err, .mntr, .out, .db, .rst11042-0208_Material_Properties_Macro.INPElastic-plastic analysis files for SL-A/B06/27/2017 16:03:51Elastic-plastic analysis SL-DAXISYMM_IP_625F.dbReference .db file for Axisymmetric SL-D Elastic-plastic analysis Note (2)AXISYMM_IP-370F_SLD.ext
.ext = .inp, .err, .mntr, .out, .db, .rst 11042-0208_Material_Properties_Macro.INPElastic-plastic analysis files for SL-D06/27/2017 16:05:25Side Drop 3D-Half-Symmetric modelLimit load analysis SL-DLIMIT_HALFSYM.dbReference .dbfile for half-symmetric limit load analysis Note (2)LIMIT_HALFSYM-237.ext.ext = .inp, .err, .mntr, .out, .db, .rstLimit load SL-Danalysis files.06/29/2017 03:44:12Elastic-plastic analysis SL-DSTRAIN_HALFSYM.dbReference .dbfile for half-symmetric elastic-plastic analysisNote (2)STRAIN_HALFSYM-237.ext
.ext = .inp, .err, .mntr, .out, .db, .rst 11042-0208_Material_Properties_Macro.INPElastic-plastic SL-Danalysis files.06/28/2017 10:47:28Excel File11042_0208_Elastic-Plastic_Stress-Strain.xlsSS304true strain / stress temperature dependent curves evaluation6/29/201713:15:19Notes:(1)The date & time (EST) for the main runs are from the listing at the end of output file.
(2)ANSYS FE models are taken from Section 8.0 of Ref. [5.4].
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0Page10of 23Table 1-Internal Pressure in the 61BTH Type 1 DSCDesign ConditionMaximum Calculated Pressuresused in this Calculation[psi]Design Pressures [psi]Normal7.3 [5.3]10Off-Normal10.9 [5.3]20Accident45.9 [5.5]
65Table 2-Maximum Temperatures in the 61BTH Type 1 DSC Shell[5.5]Design ConditionMaximum as loaded calculatedTemperaturesused in This Calculation[ºF]Design Temperature [
°F]NormalStorage237500Transfer237500Off-NormalStorage237500Transfer237500AccidentStorage370625Transfer237500 CalculationCalculation No.11042-0208Revision No.
0Page11of 23Table 3-Properties of SA-240 Type 304. Ref. [5.4]Temp [oF] E Modulus of Elasticity
[ksi] Sm Allowable Stress Intensity
[ksi] Sy Yield Stress
[ksi] Su Ultimate Tensile Strength
[ksi] Yield Stress for SL A/B Limit Load Analysis
[ksi] Yield Stress for SL D Limit Load Analysis
[ksi] 70 28,300 20.0 30.0 75.0 30.0 46.0 100 28,138 20.0 30.0 75.0 30.0 46.0 200 27,600 20.0 25.0 71.0 30.0 46.0 300 27,000 20.0 22.4 66.2 30.0 46.0 400 26,500 18.7 20.7 64.0 28.1 43.0 500 25,800 17.5 19.4 63.4 26.3 40.3 600 25,300 16.4 18.4 63.4 24.6 37.7 700 24,800 16.0 17.6 63.4 24.0 36.8 CalculationCalculation No.11042-0208Revision No.
0Page12of 23Table 4-Properties of SA-36. Ref. [5.4]Temp [oF] E Modulus of Elasticity
[ksi] Sm Allowable Stress Intensity
[ksi] Sy Yield Stress
[ksi] Su Ultimate Tensile Strength
[ksi] Yield Stress for SL A/B Limit Load Analysis
[ksi] Yield Stress for SL D Limit Load Analysis
[ksi] 70 29,500 19.3 36.0 58.0 29.0 40.6 100 29,338 19.3 36.0 58.0 29.0 40.6 200 28,800 19.3 33.0 58.0 29.0 40.6 300 28,300 19.3 31.8 58.0 29.0 40.6 400 27,700 19.3 30.8 58.0 29.0 40.6 500 27,300 19.3 29.3 58.0 29.0 40.6 600 26,700 17.7 27.6 58.0 26.6 40.6 700 25,500 17.3 25.8 58.0 26.0 39.8 CalculationCalculation No.11042-0208Revision No.
0Page13of 23Table 5-Summary of Limit Load Analysisfor the maximum weld flaws Sl. No.NameLoadingTemp.Analysis CriteriaApplied Pressure(psi)Requirement of pressure to Safety Limit load Criteria(1)(psi)Limit Load Collapse Pressure(psi)Code Limit Load Criteria Satisfied?
12D-AxisymmetricInternal pressure237SL A/B29.36098.4Yes 22D-AxisymmetricInternal pressure370SL D45.990.2144.1Yes Sl. No.NameLoadingTemp.Analysis CriteriaDesign G-load(g)Required G-load to Satisfy Limit load Criteria(1)(g)Limit Load Collapse G-Load(g)Code Limit Load Criteria Satisfied?
33D-Half-symmetricSide drop with 10.9 psi off-normalIP237SL D75 104204.0YesNote: (1)See paragraph Limit Load Analyses, Section 7.0, Ref. [5.3]Table 6-Summary ofPeak Strain Values for Elastic-Plastic Analyses for the maximum weld flawsLoad Case Specific loading Internal Pressure (psi) Peak Equivalent Plastic Strain Material Strain Limit Margin of Safety at Specified Loading(2) at 45.9 psi internal Pressure at 69 psi internal Pressure(1) 2D-Axisymmetric Internal Pressure Service Level D 45.9 4.4% 7.1% 28% 5.36 Load Case Specific loading Side Drop G-Load (g) Peak Equivalent Plastic Strain Material Strain Limit Margin of Safety at Specified Loading(2) at 75g loading at 112.5g loading(1) 3D-Half-symmetric Side Drop Service Level D 75 9.8% 19.0% 28% 1.86 Note:(1)1.5x Specified Loads (2)Margin of Safety is calculated as (Strain Limit/Actual Strain)-1 CalculationCalculation No.11042-0208Revision No.
0Page14of 23Table 7-Summary of Elastic-Plastic Analysis Resultsfor the maximum weld flawsAnalysis CaseResultPlastic StrainInternal PressureService Level A/B2D-Axisymmetric (1)Equivalent Plastic Strain at 29.3psi Internal Pressure(1)2.7%Internal PressureService Level D2D-AxisymmetricEquivalent Plastic Strain at 45.9psi Internal Pressure4.4%Side DropService Level D3D-Half-SymmetryEquivalent Plastic Strain at 75g Acceleration9.8%Note:(1)The 29.3psi internal pressure is bounding for Service Levels A and B and includes calculated internal pressure of 7.3psi plus an additional 22 psi to account for inertial loading of the DSC contents onto the lid.
CalculationCalculation No.11042-0208Revision No.
0Page15of 23(a) Equivalent Plastic Strain in Weld Region [in/in] at 10.9psi Internal Pressure(b) Equivalent Plastic Strain in Weld Region [in/in] at 45.9psiInternal Pressure(c) Equivalent Plastic Strain in Weld Region [in/in] at 98.4psi Internal Pressure(d) EQV Plastic Strain in the Cover Plates at 98.4 psiFigure 1-Resultsof Limit Loadfor 2D-Axisymmetric Model-Service Level A/B CalculationCalculation No.11042-0208Revision No.
0Page16of 23(a) Equivalent Plastic Strain in Weld Region [in/in] at 45.9psi Internal Pressure(b) Equivalent Plastic Strain in Weld Region [in/in] at 100psi Internal Pressure(c) Equivalent Plastic Strain in Weld Region [in/in] at 144.1psi Internal Pressure(d) EQV Plastic Strain in the Cover Plates at 144.1 psiFigure 2-Resultsof Limit Loadfor 2D-Axisymmetric Model-Service Level D CalculationCalculation No.11042-0208Revision No.
0Page17of 23(a) Service Level A/B(b) Service Level DFigure 3-Deflection atthe Center of the OTCP for the 2D-Axisymmetric Modelfor Limit Load(Maximum deflection occurs at the center point of the lids, in the outward axial direction)
CalculationCalculation No.11042-0208Revision No.
0Page18of 23Figure 4-Equivalent Plastic Strain at 29.3psi for 2D-Axisymmetric Elastic Plastic Analysis-SLA/BInternal Pressure CalculationCalculation No.11042-0208Revision No.
0Page19of 23Figure 5-Equivalent Plastic Strain at 45.9psi for2D-Axisymmetric Elastic Plastic Analysis -SLD Internal Pressure CalculationCalculation No.11042-0208Revision No.
0Page20of 23Figure 6-Equivalent Plastic Strain at 69psi for 2D-Axisymmetric Elastic Plastic Analysis -SL D Internal Pressure CalculationCalculation No.11042-0208Revision No.
0Page21of 23(a) Equivalent Plastic Strain in Weld Region [in/in] at 75g.(b) Equivalent Plastic Strain in Weld Region [in/in] at 204g.Figure 7-Equivalent Plastic Strain Plots for 3D-Half-Symmetric Limit Load Analysis -SL D Side Drop with Off-Normal Internal Pressure CalculationCalculation No.11042-0208Revision No.
0Page22of 23Figure 8-Equivalent Plastic Strain at 75g for 3D-Half-Symmetric Elastic-Plastic Analysis -SLD Side Drop CalculationCalculation No.11042-0208Revision No.
0Page23of 23Figure 9-Equivalent Plastic Strain at 112.5g for 3D-Half-Symmetric Elastic-Plastic Analysis -SLD Side Drop
08/09/2017 8/9/2017RaheelHaroon CalculationCalculation No.11042-0209Revision No.
0Page2of 20REVISION SUMMARYRev.DescriptionAffectedPagesAffectedFiles0Initial issue AllAll CalculationCalculation No.11042-0209Revision No.
0Page3of 20TABLE OF CONTENTS Page1.0 PURPOSE...........................................................................................................................................5 2.0 ASSUMPTIONS..................................................................................................................................5 3.0 DESIGN INPUT/DATA........................................................................................................................5 3.1 Flaws Details and Geometry.....................................................................................................
.5 3.2 Material Properties.....................................................................................................................5 3.3 Design Criteria...........................................................................................................................5 4.0 METHODOLOGY................................................................................................................................5 4.1 Analysis Method and Acceptance Criteria..................................................................................5 4.2 g-load Evaluation........................................................................................................................6 4.2.1 Bounding Static Deceleration for End Drop:....................................................................7 4.2.2 Bounding Static Deceleration for Side Dr op:...................................................................8 4.2.3 g-load Evaluation results................................................................................................8 4.3 FEA Model Details......................................................................................................................9 4.4 Limit Load Solution Details.........................................................................................................9 4.5 Elastic Plastic Solution Details....................................................................................................9 4.6 Load Cases................................................................................................................................9
5.0 REFERENCES
..................................................................................................................................10 6.0 ANALYSIS AND RESULTS...............................................................................................................10 6.1 LIMIT LOAD ANALYSIS...........................................................................................................1 0 6.2 ELASTIC-PLASTIC ANALYSIS................................................................................................10 7.0 DISCUSSION AND CONCLUSIONS.................................................................................................11 8.0 LISTING OF COMPUTER FILES......................................................................................................
11 9.0 APPENDIX A -CALCULATION FOR FACTOR G............................................................................19 LIST OF TABLESPageTable 1 -Parameters for g-load Evaluation from Ref. [5.9]...........................................................................6 Table 2 -Internal Pressure in the 61BTH Type 1 DSC [5.5].......................................................................
12 Table 3 -Maximum Temperatures in the 61BTH Type 1 DSC Shell [5.6]...................................................12 Table 4 -Properties of SA-240Type 304. Ref. [5.4]...................................................................................13 Table 5 -Properties of SA-36. Ref. [5.4]...................................................................................................14 Table 6 -Summary of Limit Load Analysis for the maximum weld flaws [5.5].............................................15 Table 7 -Summary of Peak Strain Values for Elastic-Plastic Analyses for the maximum weld flaws..........15
CalculationCalculation No.11042-0209Revision No.
0Page4of 20LIST OF FIGURESPageFigure 1 -Equivalent Plastic Strain Plots for 3D-Half-Symmetric Limit Load Analysis -SL D Side Drop with Off-Normal Internal Pressure (Ref. [5.5]).............................................................................16 Figure 2 -Equivalent Plastic Strain at 52.5g for 3D-Half-Symmetric Elastic-Plastic Analysis -SL D Side Drop...........................................................................................................................................17 Figure 3 -Equivalent Plastic Strain at 79g for 3D-Half-Symmetric Elastic-Plastic Analysis -SL D Side Drop............................................................................................................................................18
CalculationCalculation No.11042-0209Revision No.
0Page5of 201.0PURPOSEThe purpose of this calculation is to evaluatethe margins forthe NUHOMS 61BTH Type 1 DSCsat the Monticello Nuclear Generating Plant (MNGP) per ASME Section III criteria withthemaximum postulated flaws in the Inner and Outer Top Cover Plates(ITCP and OTCP) closure weldsfor Limit load and Elastic-Plastic analysesbased on theside dropload casesperformed in the reference calculation [5.5].The as-loaded site specific bounding temperatures and pressures used in this calculation are providedin Ref. [5.6].The site specific side drop load is used to better represent the actual Independent Spent Fuel Storage Installation's (ISFSI)approach slabat MNGP, instead of using thebounding 75g design side drop load.2.0ASSUMPTIONS1.Assumptions 1 through 6 of Ref.[5.3]are applicable tothis calculation.3.0DESIGN INPUT/DATA3.1FlawsDetails and GeometryTheflawsdetailsare identicalto the maximum weld flaws evaluated in Ref.[5.4] and[5.5]. The geometry ofthe DSC structureis identical tothe geometry used in Ref.[5.3],[5.4]and [5.5].3.2Material Properties Thematerial properties for the DSC structure are identicalto the material properties of Ref.[5.3],[5.4]and [5.5].They are duplicated here in Table 4andTable 5for convenience.3.3Design CriteriaAll of the applicable design bases loading conditions areconsidered in accordance with the requirements of ASME Section III Subsection NB Ref. [5.2]. Section 4.1details the methods used to perform the code Ref. [5.2]qualifications.4.0METHODOLOGY4.1Analysis Method and Acceptance CriteriaThe analysis methods, finite element model details and acceptance criteria are the same as discussed in Ref. [5.3]. The ITCP and OTCPmaximumweld flaws as evaluated in Ref. [5.4]are analyzedand margins are evaluatedforLimit load and Elastic Plastic analyses.The as loaded site specific bounding temperatures and pressures used in this calculation are provided in Ref. [5.6].The side drop and end drop design loads are set to 75g. These g-loads represent the ultimate capacity of the concrete slab, independently of the drop height(usually assumed at 80"), calculated with the Target hardness model used in Ref. [5.7] and validated in Ref. [5.8]. This ultimate capacity includes a 1.5 Dynamic Load Factor(DLF) and is assessed for the main ISFSI pad which is usually only present under the HSM's location. Therefore, severalmodifications of the drop g-load can be done by taking into account:
CalculationCalculation No.11042-0209Revision No.
0Page6of 201.Alow drop height. For MNGP, the drop height could be reduced to 64.5", but this method requiresa full dynamic finite element analysisof the Transfer Cask (TC) drop on the pad.2.The actual MNGP ISFSI pad design instead of the generic 36" thick pad design leading to the 75g load3.The actual Approach slab parameters, instead of the ISFSI pad.In this calculation, only modification3is done, using the slabs Target hardness model to derive a g-load reduction factor between the 30" MNGP ISFSI pad and the 15" Approach slab, for both side and enddrop loads.Conservatively the 75g drop load derived for the generic 36" thick pad is considered also for the MNGP ISFSI 30" thick pad. The evaluation is done using the characteristic of the NUHOMS TC OS197.4.2g-load EvaluationThe Target hardness model methodology of g-load evaluation is presented and validated in Ref. [5.8]. It was previouslyused for the NUHOMS OS197 transfer cask inRef. [5.7]. Although the NRC questioned the validity of the Target hardness modelmethodology (Page 3-19 of Ref. [5.11]), the 75g bounding drop load is accepted by the NRC. The methodology is not used here to evaluate a specific g-load value, but rather to find a ratio to evaluate the site specific g-load compared to the generic design 75g drop load.Table 1-Parameters for g-load Evaluation from Ref. [5.9]Sl. No.Parameters30" ISFSI Pad15" Approach Slab 128 day compressive strength of concrete, f'c (1)4000 psi4690 psi2Yield strength of reinforcement, fy (1)60 ksi72.3 ksi3Soil subgrade stiffness, k50pci100pci 4Elastic modulus of concrete, Ec3.834E6psi4.152E6psi5Poisson ratio of soil,S0.330.336Poisson ratio of concrete,C0.170.177Impulse duration0.016sec0.016sec 8Width of TC contact area (side drop), b20 in.20 in.
9Concrete Pad thickness, h30in.15 in.
Note:(1)Parameters conservativelytaken as designedfor 30" ISFSI pad and maximum measured for 15" Approach slabThe correlation has been established between "limiting static deceleration" of the cask and "Target Hardness"(Page 2-1 of Ref. [5.8])
G34533.5ln(S) for 120,000 S14.7106 CalculationCalculation No.11042-0209Revision No.
0Page7of 20G8811.5 ln(S) for 13,300 S120,000G15.353.85 ln(S) for S 13,300where:G=limiting equivalent static deceleration as a multiplier on gravityS=target hardness number (non-dimensional)4.2.1Bounding Static Deceleration for End Drop:The bounding static deceleration values, based on the Ref. [5.8], are function of a "target hardness" parameter, S, given empirically as:
r)cose-(1WfM2rAK r-3'cuwhere:Mu= ultimate moment capacity of the slab (lb-in),W= weight of Transfer Cask (lb),186175 lb for TC OS197 [5.7]
'= ultimate strength of concrete (psi), A= cask footprint area=
2(in2),S=Poisson's ratio of soilC= Poisson's ratio of concrete k= soil subgrade stiffness, r= cask radius (in),39.56" for TC OS197 [5.7]h= concrete pad thickness (in),= 1.15 (for a circle per Page 2-3 of Ref. [5.8])
414,23112concrete slab rigidity (lb-in 2),21foundation modulus, CalculationCalculation No.11042-0209Revision No.
0Page8of 20,12soil's elastic modulus,,4)22(2effective bearing area of the concrete slab/soil interface,,ultimate moment capacity of the slab (lb-in)with= area of steel reinforcement= effective concrete cover of the reinforcement4.2.2Bounding Static Deceleration for SideDrop:The bounding static deceleration values, based on the Ref. [5.8], are function of a "target hardness" parameter, S, given empirically as:3'cuSWfMELb2where:L= Length of cask (in), 207 in for TC OS197 Ref. [5.10]= 1.41 (for a rectangle per Page 2-3 of Ref. [5.8])
4143L121)2)(2((All other parameters are identical to the End drop parameters.4.2.3g-load Evaluation resultsThe above expressionsof S depict the relationship between the g-load for end and side dropsand the above input parameters. The parameter S is directly proportional to the concrete strength, soil modulus of elasticity, concrete padthickness & inversely proportional to the weight of cask.Effectsof concrete strength,concrete pad thickness and soil subgrade stiffness arehigher as compared to the other parameters. Table 1shows that compressive strength of concrete is increased and concrete padthickness is decreased.The dynamic amplification factor is the same, so it does not affect the evaluation here.The ratio new Gto old Gfor sidedrop is conservatively taken as 0.7as per parameterscalculated inSection 9.0. 75g load is taken for the Side drop in Ref. [5.5]. The new site specific g-load is=0.7X 75 = 52.5g.
CalculationCalculation No.11042-0209Revision No.
0Page9of 204.3FEA Model DetailsFinite element models detail ofthe DSC structure are identicalto the ones described in Ref.[5.4].4.4Limit Load Solution DetailsLimit load solution details are the same as detailed in Section 4.4 of Ref.[5.3].4.5Elastic PlasticSolution DetailsElastic Plastic solution details are the same as detailed in Appendix-A of Ref.[5.3].4.6Load CasesTable 2summarizes the actual internal pressure values which are taken fromTable 2 ofRef.[5.3].Temperatures used for the material properties for each Service Level condition are listed in Table 3.The Maximum Service Level (SL) D temperature of the DSC shell is taken as 370
°F as per Table 7-2 of Ref.[5.6]for a blocked vent accident.The same table also gives the maximum DSC shell SL-B temperature as 237°F, before the blocked vent accident.Two 3D-Half-Symmetric analyses for bounding SL D are performed.
CalculationCalculation No.11042-0209Revision No.
0Page10of 2
05.0REFERENCES
5.1.ANSYS Version 17.1. ANSYS Inc. (Including the ANSYS Mechanical APDL Documentation).5.2.ASME Boiler and Pressure Vessel Code,Section III Subsection NB. 1998 Edition with Addenda through 2000.5.3.AREVA Document No. 11042-0205 Revision 3. "61BTH ITCP and OTCP Closure Weld Flaw Evaluation"5.4.AREVA Document No. 11042-0207Revision 0. "NUHOMS 61BTH Type 1 DSC ITCP and OTCP Maximum Weld Flaw Evaluation"5.5.AREVA Document No. 11042-0208Revision 0. "Site Specific NUHOMS 61BTH Type 1 DSC ITCP and OTCP margin evaluation for Maximum Weld Flaw"5.6.AREVA Document No. 11042-0400Revision 0. "Site-Specific Thermal Evaluation of 61BTH Type 1 DSCsstoredin HSM-H at Monticello Nuclear Generating Plant"5.7.AREVA Document No. NUH-04.0110Revision 0. "NUHOMS ISFSICask Drop Acceleration"5.8.Anatech Report TR-108760prepared for EPRI, "Validation of EPRI Methodology of Analysis of Spent-Fuel Cask Drop and Tipover Events",August 1997.5.9.AREVA Document DI-11042-04 Revision 0, Xcel Energy "DIT No. 60115-002-Transmittal of MNGP Design Documents", 07/06/20175.10.AREVA Drawing DWG-NUH-06-8003, Revision 11. "NUHOMS
-OS197-1 Outer Transfer Cask Main Assembly"5.11.US NRC, SER 1004, December 1994, "SER of SAR for the Standardized NUHOMS Horizontal Modular Storage System for Irradiated Nuclear Fuel"6.0ANALYSISAND RESULTS6.1LIMIT LOAD ANALYSIS3D-Half SymmetricAnalyses for Side Drop LoadingLimit load analysis result for 3D half Symmetric model is presented in Section6.1.2 of Ref. [5.5]. The collapse g-load for side-drop loading was found to be approximately 204g.Plotsof the plastic strainsin the side drop analyses are shown inFigure 1.The resultsfor Limit load analysisare summarized in Table 6.6.2ELASTIC-PLASTIC ANALYSIS3D-Half SymmetricAnalyses for Side Drop LoadingThe 3D-half-symmetric model described in Section 4.3is used to perform the SL D side-drop limit load analysis. Thecase includes the 52.5gside-drop acceleration loading only. The maximum strains at loads up to 1.5x the specified loading (79g) are also extracted and compared with the material strain limit.The equivalent plastic strainwas determined to be 5.8% for 52.5g and 10.6% for 79g.Figure 2andFigure 3show the corresponding plastic strain plots.The results for elastic-plastic analyses are summarized inTable
- 7.
CalculationCalculation No.11042-0209Revision No.
0Page11of 207.0DISCUSSION AND CONCLUSIONSThe analyses are done for site specific side drop g-load based on the Target hardness model which is used to derive the ultimate reinforced concrete slab capacity on which the TC could drop. This bounding g-load is found to be 52.5g for the MNGP site as compared to the 75g design load.Limit Load Analyses:The lower bound collapse G-Load for Service Level D side drop criteria was found to be 204 gwhich is greater thanthe limiting G-Load of 104g(Table 6).Therefore the Service Level D criterion is satisfied.Elastic-Plastic Analyses:Table 7lists the peak strains predicted by the elastic-plastic analyses for the bounding Service Level D. As shown in the table, the peak strain values remain below the material ductility limits atthe specified loading conditions with a minimum margin of safety of 3.83.Therefore the elastic plastic analysescriteria are satisfied.8.0LISTING OF COMPUTER FILESFinite Element Analyseswereperformed using ANSYS Version 17.1Ref. [5.1].All analyses were performed on HPC v2 Linux platform.Load CaseAnalysis TypeFile NameDescription Date / Time(1)Side Drop 3D-Half-Symmetric modelLimit load analysis SL-DLIMIT_HALFSYM.dbReference .db file for half-symmetric limit load analysisRef.[5.5]LIMIT_HALFSYM-237.ext.ext = .inp, .err, .mntr, .out, .db, .rstLimit loadSL-Danalysis files.Elastic-plastic analysis SL-DSTRAIN_HALFSYM-237.dbReference .db file for half-symmetric elastic-plastic analysisNote (2)STRAIN_HALFSYM-52.5g.ext
.ext = .inp, .err, .mntr, .out, .db, .rstElastic-plastic SL-D analysis files.07/14/2017 20:13:57G-Factor_calculation.xlsxExcel file to calculate G-Factor07/17/2017 12:10Notes:(1)The date & time (EST) for the main runs are from the listing at the end of output file.
(2)ANSYS FE models are taken from Section 8.0 of Ref. [5.5].
CalculationCalculation No.11042-0209Revision No.
0Page12of 20Table 2-Internal Pressure in the 61BTH Type 1 DSC[5.5]Design ConditionMaximum Calculated Pressuresused in this Calculation[psi]Design Pressures [psi]Normal7.3 10Off-Normal10.9 20Accident45.9 65Table 3-Maximum Temperatures in the 61BTH Type 1 DSC Shell[5.6]Design ConditionMaximum as loaded calculatedTemperaturesused inThis Calculation[ºF]Design Temperature [
°F]NormalStorage237500Transfer237500Off-NormalStorage237500Transfer237500AccidentStorage370625Transfer237500 CalculationCalculation No.11042-0209Revision No.
0Page13of 20Table 4-Properties of SA-240 Type 304. Ref. [5.4]Temp [oF] E Modulus of Elasticity
[ksi] Sm Allowable Stress Intensity
[ksi] Sy Yield Stress
[ksi] Su Ultimate Tensile Strength
[ksi] Yield Stress for SL A/B Limit Load Analysis
[ksi] Yield Stress for SL D Limit Load Analysis
[ksi] 70 28,300 20.0 30.0 75.0 30.0 46.0 100 28,138 20.0 30.0 75.0 30.0 46.0 200 27,600 20.0 25.0 71.0 30.0 46.0 300 27,000 20.0 22.4 66.2 30.0 46.0 400 26,500 18.7 20.7 64.0 28.1 43.0 500 25,800 17.5 19.4 63.4 26.3 40.3 600 25,300 16.4 18.4 63.4 24.6 37.7 700 24,800 16.0 17.6 63.4 24.0 36.8 CalculationCalculation No.11042-0209Revision No.
0Page14of 20Table 5-Properties of SA-36. Ref. [5.4]Temp [oF] E Modulus of Elasticity
[ksi] Sm Allowable Stress Intensity
[ksi] Sy Yield Stress
[ksi] Su Ultimate Tensile Strength
[ksi] Yield Stress for SL A/B Limit Load Analysis
[ksi] Yield Stress for SL D Limit Load Analysis
[ksi] 70 29,500 19.3 36.0 58.0 29.0 40.6 100 29,338 19.3 36.0 58.0 29.0 40.6 200 28,800 19.3 33.0 58.0 29.0 40.6 300 28,300 19.3 31.8 58.0 29.0 40.6 400 27,700 19.3 30.8 58.0 29.0 40.6 500 27,300 19.3 29.3 58.0 29.0 40.6 600 26,700 17.7 27.6 58.0 26.6 40.6 700 25,500 17.3 25.8 58.0 26.0 39.8 CalculationCalculation No.11042-0209Revision No.
0Page15of 20Table 6-Summary of Limit Load Analysisfor the maximum weld flaws[5.5]
Sl. No.NameLoadingTemp.Analysis CriteriaSite SpecificG-load(g)Required G-load to Satisfy Limit load Criteria(1)(g)Limit Load Collapse G-Load(g)Code Limit Load Criteria Satisfied?
13D-Half-symmetricSide drop with 10.9 psi off-normal IP237SL D52.5104.0204.0 YesNote: (1)See paragraph Limit Load Analyses, Section 7.0, Ref. [5.3]Table 7-Summary ofPeak Strain Values for Elastic-Plastic Analyses for the maximum weld flawsLoad Case Specific loading Side Drop G-Load (g) Peak Equivalent Plastic Strain Material Strain Limit Margin of Safety at Specified Loading(2) at 52.5g loading at 79g loading(1) 3D-Half-symmetric Side Drop Service Level D 52.5 5.77% 10.6% 28% 3.83 Note:(1)1.5x Specified Loads (2)Margin of Safety is calculated as (Strain Limit/Actual Strain)-1 CalculationCalculation No.11042-0209Revision No.
0Page16of 20(a) Equivalent Plastic Strain in Weld Region [in/in] at 75g.(b) Equivalent Plastic Strain in Weld Region [in/in] at 204g.Figure 1-Equivalent Plastic Strain Plots for 3D-Half-Symmetric Limit Load Analysis -SL D Side Drop with Off-Normal Internal Pressure(Ref. [5.5])
CalculationCalculation No.11042-0209Revision No.
0Page17of 20Figure 2-Equivalent Plastic Strain at 52.5gfor 3D-Half-Symmetric Elastic-Plastic Analysis -SLD Side Drop CalculationCalculation No.11042-0209Revision No.
0Page18of 20Figure 3-Equivalent Plastic Strain at 79g for 3D-Half-Symmetric Elastic-Plastic Analysis -SL D Side Drop CalculationCalculation No.11042-0209Revision No.
0Page19of 209.0APPENDIX A -CALCULATION FOR FACTORG30" ISFSI Pad End drop Side drop r 39.56 L 207 W 186175 186175 b 20 A 4916.57 1.15 1.41 h 30.00 30.0 As 1.27 1.27 ceff 3.0 3.0 k50 50 fc' 4000 4000 fy 60000 60000 S0.33 0.33 C0.17 0.17 As 1.520E+04 2.136E+04 Ic 4.658E+05 Ec 3.83E+06 3.83E+06 Mu 1.824E+06 1.824E+06 Es 4777 4.618E+03 K 1.684E+04 1.915E-02 5.043E-03 Dc 8.884E+09 8.884E+09 S 1.123E+04 8.575E+03 Gold 20.6 19.5 DLF 1.5 1.5 CalculationCalculation No.11042-0209Revision No.
0Page20of 2015" Approach Slab End drop Side drop r 39.56 L 207 W 186175 186175 b 20 A 4916.57 1.15 1.41 h 15.00 15.0 Ast 0.44 0.44 ceff 3.0 3.0 k100 100 fc' 4690 4690 fy 72300 72300 S0.33 0.33 C0.17 0.17 As 9.352E+03 1.185E+04 Ic 5.822E+04 Ec 4.152E+06 4.152E+06 Mu 2.875E+05 2.875E+05 Es 7.493E+03 6.880E+03 K 2.642E+04 3.533E-02 9.184E-03 Dc 1.202E+09 1.202E+09 S 2.243E+03 1.296E+03 Gnew 14.4 12.2 DLF 1.5 1.5 For End drop: G new/Gold= 14.4/20.6 = 0.70For Side drop: G new/Gold=12.2/19.5 = 0.63The designed g-load is 75g, the maximum ratio of G new/Goldis0.70. So the new g-load is =0.7X75 = 52.5g.
APPLIED ANALYSIS CORP. aac Calculation for: XCEL ENERGY INC MONTICELLO NUCLEAR GENERA TlNG PLANT MNGP-018 Accident Dose Assessment for MNGP DSCs 11 -15 Assumptions Requiring Later Verification:
D No 00 Yes Assumption
- 3. 7 Nuclear Quality Status: Nuclear Quality Method of Verification:
0 Design Review D Alternate Calculation APPROVAL Revision:_O_
Date: 08/16/2017 Reviewed By:_J_. M __ . __ d __ O __________
_ Date: 08/16/2017 Approval By: J. M. Cajigas Date: 08/16/2017 For signatures see electronic file: MNGP-018 RO.pdf CALCULATION NO. MNGP-018 PAGE 2 of 38 REV 0 PAGE REV. ATT. NO. REV. ATT. NO. REV. 1-38 0 Att. 1 0 Att. 2 0 Att. 3 0 Att. 4 0 Att. 5 0 Att. 6 0 Att. 7 0 Att. 8 0 Att. 9 0 Att. 10 0 Att. 11 0 Att. 12 0 Att. 13 0 Att. 14 0 Att. 15 0 Att. 16 0 Att. 17 0
CALCULATION NO. MNGP-018 PAGE 3 of 38 REV 0 TABLE OF CONTENTS DESCRIPTION PAGE List of Effective Pages & Attachments ........................................................................................................2
Table of Contents .........................................................................................................................................3
Revision History ..........................................................................................................................................6
Definitions
................................................................................................
....................................................7
Computer Data .............................................................................................................................................8
1.0 Purpose .............................................................................................................................................9
2.0 List of References ..........................................................................................................................10
3.0 Assumptions ...................................................................................................................................12
4.0 Design Input ...................................................................................................................................13
5.0 Methodology ..................................................................................................................................17
6.0 Results ............................................................................................................................................18 6.1 ORIGEN-ARP DSC Source Term ...........................................................................................18 6.2 RADTRAD DSC Dose Calculation .........................................................................................22 6.3 ...........................................................................................................31 6.4 DSC Leakage Rate ...................................................................................................................32
7.0 Conclusions ....................................................................................................................................36
LIST OF FIGURES FIGURE # TITLE PAGE None
CALCULATION NO. MNGP-018 PAGE 4 of 38 REV 0 LIST OF TABLES TABLE # TITLE PAGE 4.1 Reference 2.8 Table 3.2.1 - Key Inputs
.........................................................................................13 6.1 DSC Activity ..................................................................................................................................21 6.2 Inhalation DCFs (Sv/Bq) - FGR-11 Table 2.1 ...............................................................................26 6.3 Submersion DCFs (Sv/sec per Bq/m3) - FGR-12 Table III-1 .......................................................27 7.1 Organ Dose ....................................................................................................................................36 7.1a - Additional Hole Sizes ....................................................36 7.2 Organ Dose with Realistic Dispersion Factor ................................................................................37 7.2a Organ Dose with Realistic Dispersion Factor - Additional Hole Sizes ........................................37 LIST OF ATTACHMENTS ATT. # DESCRIPTION
- 1. SQAP ORIGEN
-ARP Documentatio n, Electronic Files:
Sample Problem:
AAC-arp.arp, AAC-arp.inp, AAC-arp.out, AAC-arp.F71 Check Problem:
AAC-arp - SQAP Check.arp, AAC
-arp - SQAP Check.inp, AAC-arp - SQAP Check.out, AAC-arp - SQAP Check.F71 AAC-arp vs AAC-arp - SQAP Check Output Comparison and Computer Sample Problem ORIGEN-ARP Validation Sheet SQAP - MNGP-018 R0.pdf 2. SQAP RADTRAD Documentation, Electronic Files: Test13b.psf, Test13b.o0, Test14b.psf, Test14b.o0, Test15.psf, Test15.o0, Test16.psf, and Test16.o0, BWR_I131.nif, BWR_DBA.rft, Fgr11&12.inp, Computer Sample Problem RADTRAD Validation Sheet SQAP - MNGP-018 R0.pdf 3. ORIGEN-ARP Input/Output, Electronic Files:
MNGP-017 R0 Att. 3: Files MNGP EPU - GE14 37 GWD Fuel.inp/arp/out/f71 MNGP-017 R0 Att. 3: Files MNGP EPU - GE14 37 GWD Fuel - Match.inp/arp/out/f71 Windiff Comparison File - MNGP-017 R0 Att. 3 - MNGP-017 R0 Att. 3 - Match - Output Comparison.txt
- 4. ORIGEN-ARP Input/Output, Electronic Files:
- 0 Cutoff.inp/arp/out/f71 Attachments 5-9 are electronic files:
- 5. Excel Spreadsheet, "MNGP
- DSC Source Term to RADTRAD"
- 6. RADTRAD Source Term File
- "DSC Source Term.NIF"
- 7. RADTRAD Release and Timing File
- "DSC Source Term.RFT"
- 8. RADTRAD Input File
- "DSC Source Term.INP"
- 9. RADTRAD Input/Output Files - DSC Source Term.psf/o0
- 10. Electronic File, "PNL 10268 Figure 9 Data Bias Bins Depiction"
- 11. Excel Spreadsheet
, Electronic File Data Analysis, "PNL 10286, Figure 9, X
-Q Biases" 12. Excel Spreadsheet, Electronic File, "DSCs 11-15 Burnup Averages" CALCULATION NO. MNGP-018 PAGE 5 of 38 REV 0 13. ORIGEN-ARP Input/Output, Electronic Files:
- 0 Cutoff - 41 GWD.inp/arp/out/f71 Attachments 14-16 are electronic files:
- 14. RADTRAD Input/Output Files - DSC Source Term
- EPRI Area 1.psf/o0 15. RADTRAD Input/Output Files - DSC Source Term
- EPRI Area 2.psf/o0 16. RADTRAD Input/Output Files - DSC Source Term
- EPRI Area 3.psf/o0 17. Design Verification Comment Sheet, Electronic File: DVCS - MNGP-018 R0
CALCULATION NO. MNGP-018 PAGE 6 of 38 REV 0 0 08/16/17 Initial Issue
CALCULATION NO. MNGP-018 PAGE 7 of 38 REV 0 AAC Applied Analysis Corp. AST Alternative Source Term BWR Boiling Water Reactor CFR Code of Federal Regulations Ci/MTU Curies per Metric Ton Uranium Ci/MWth Curies per Megawatt Thermal DCF Dose Conversion Factor DI Design Input DSC Dry Shielded Canister EAB Exclusion Area Boundary (also referred to as Owner Controlled Area (OCA) Boundary) EPU Extended Power Uprate GUI Graphical User Interface GWD Gigawatt Days ISFSI Independent Spent Fuel Storage Installation MNGP Monticello Nuclear Generating Plant MTU Metric Tons Uranium MWD Megawatt Days MWth Megawatts Thermal PU Power Uprate Thermal Power RG Regulatory Guide SQAP Software Quality Assurance Program SVVR Software Validation and Verification Report USNRC United States Nuclear Regulatory Commission
CALCULATION NO. MNGP-018 PAGE 8 of 38 REV 0 The ORIGEN-ARP Version 5.1.01 computer program included with the SCALE Version 6.0 package (Reference 2.1) was used in accordance with the AAC SQAP Revision 0 (Reference 2.2). The originator has executed the sample problems per S VVR Section 6.4.1 and verified proper installation and execution (see Attachment 1).
The RADTRAD Version 3.03 program (Reference 2.3) was used in accordance with the AAC SQAP Revision 0 for RADTRAD Version 3.03 (Reference 2.4). The originator has executed the sample problems per SQAP Section 6.1.1 and verified proper installation and execution (see Attachment 2).
CALCULATION NO. MNGP-018 PAGE 9 of 38 REV 0 MNGP is planning an Exemption Request for five (5) NUHOMS Dry Shielded Canisters (DSCs) that were placed in service at the MNGP Independent Spent Fuel Storage Installation (ISFSI) with noncompliant dye penetrant examinations (PTs). In support of this exemption request, an offsite radiological dose consequence analysis from an accidental release from an affected DSC is provided
.
In accordance with Reference 2.5, the MNGP DSCs will not release radioactive effluents under any accidental circumstances. However, in order to support a justification of the above described Exemption Request, a radiation dose consequence analysis will be performed to determine the radiological dose consequences from a postulated accidental release from a single DSC at the MNGP ISFSI pad location to the site boundary (EAB).
As noted in calculation Section 7, the computed accident dose results are considered to be cons ervative dispersion value and the use of a 100% occupancy factor by the public at the nearest plant EAB. Conservatisms that are not directly scalable include the impact from the failed fuel percentage, the calculated natural deposition coefficient and the consideration of DSC leakage at the maximum critical flux rate for the entire 30 day postulated accident duration.
Accident dose acceptance criteria per 10CFR72.106 (Reference 2.31) and NUREG-1567 (Reference 2.12) is shown in Tables 7.1 and 7.2.
CALCULATION NO. MNGP-018 PAGE 10 of 38 REV 0 2.1 ORNL/TM-2005/39, Version 6, Vol. I, Section D1, "ORIGEN-ARP: Automatic Rapid Processing For Spent Fuel Depletion, Decay, and Source Term Analysis".
2.2 AAC Software Quality Assurance Plan SQAP
-ORIGEN-ARP-R0, "Software Quality Assurance Plan, ORIGEN
-ARP", Revision 0.
2.3 NUREG/CR-6604, "RADTRAD: A Simplified Model for RADionuclide Transport and Removal and Dose Estimation", December 1997, NUREG/CR-6604 (SAND98-0272/1), Supplement 1, "RADTRAD: A Simplified Model for RADionuclide Transport and Removal and Dose Determination", June 8, 1999, and NUREG/CR-6604, Supplement 2, "RADTRAD: A Simplified Model for RADionuclide Transport and Removal and Dose Determination", October 2002. 2.4 Applied Analysis Corp. Software Quality Assurance Plan, Revision 0, "SQAP - RADTRAD - R0". 2.5 Monticello Nuclear Generating Plant, Independent Spent Fuel Storage Installation (ISFSI), 10CFR72.212(b)(5)(iii) Radiological Evaluation.
2.6 MNGP Calculation 16-090, "MNGP EPU - Core Inventory With GE14 Fuel With 37 GWD/MTU Exposure", Revision 0.
2.7 EPU Correspondence Number: MNGP-375-GE, dated 11/15/07, "MNGP EPU Radiological Design Basis Update - AST DIR R2", Attachment EPU-DIT-0198. 2.8 MNGP Calculation 11-245, "Task Report T0802 Core Source Term (GE -
NE-0000-0064-6767-TR-RO)", Revision 0.
2.9 DIT-AAC-001 (EC-18508), "CORE RELOAD MOD FOR CYCLE 29 OPERATION", 12/14/2016.
2.10 Design Information Transmittal (DIT) No. 48 (EC-18508), 6/06/2013, "
Subject:
Design Inputs for MNGP AST Calculations (CRDA and LOCA)". 2.11 Design Information Transmittal (DIT) 67492-001 (Mod/Tracking No. 18624), "Spent Fuel Loading of Dry Shielded Canisters 6A-10B", 5/19/2017.
2.12 USNRC NUREG-1567, "Standard Review Plan for Spent Fuel Dry Storage Facilities," U.S. Nuclear Regulatory Commission, February 2000.
2.13 USNRC NUREG-1864, "A Pilot Probabilistic Risk Assessment Of a Dry Cask Storage System At a Nuclear Power Plant
", Published March 2007.
2.14 USNRC NUREG-1536, "Standard Review Plan for Spent Fuel Dry Storage Systems at a General License Facility," Revision 1, U.S. Nuclear Regulatory Commission, July 2010.
2.15 "Table of Isotopes", Eighth Edition, Richard B. Firestone, V. S. Shirley.
2.16 USNRC Regulatory Guide 1.109, Calculation of Annual Doses to Man from Routine Releases of Reactor Effluents for the Purpose of Evaluating Compliance With 10CFR50 Appendix I," Revision 1, October 1977.
CALCULATION NO. MNGP-018 PAGE 11 of 38 REV 0 2.17 USNRC Regulatory Guide 1.145, Atmospheric Dispersion Models For Potential Accident Consequence Assessments At Nuclear Power Plants," Revision 1, November 1982. 2.18 Federal Guidance Report (FGR) 11, "Limiting Values of Radionuclide Intake and Air Concentration and Dose Conversion Factors," 1989.
2.19 Federal Guidance Report (FGR) 12, "External Exposure to Radionuclides in Air, Water, and Soil," 1993.
2.20 Monticello Nuclear Generating Plant, Unit No. 1 Facility Operating License, Amendment No. 147 dated July 5, 2006.
2.21 USNRC Document SMSAB 03, "Best
-Estimate Offsite Dose from Dry Storage Cask Leakage," Prepared by Jason H. Shaperow, June 2000.
2.22 USNRC Regulatory Guide 1.183, Alternative Radiological Source Terms For Evaluating Design Basis Accidents At Nuclear Power Reactors," July 2000. 2.23 Transnuclear, "Final Safety Analysis Report, Standardized NUHOMS Horizontal Modular Storage System for Irradiated Nuclear Fuel", NUH-003, Revision 8, Appendices K & L.
2.24 PNL 10286, "Atmospheric Dispersion Estimates In The Vicinity Of Buildings", By J. V Ramsdell, Jr and C. J Fosmire, dated January 1995.
2.25 American National Standards
- ANSI N14.5
-1997, "For Radioactive Materials
- Leakage Tests on Packages for Shipment". 2.26 USNRC 10 CFR Part 71, "Packaging And Transportation Of Radioactive Material". 2.27 "Introduction to Unsteady Thermofluid Mechanics", F. J. Moody, 1990.
2.28 "Basic Principles and Calculations in Chemical Engineering", 2nd Edition, D. M. Himmelblau.
2.29 Design Information Transmittal (DIT) 67492-002 (Mod/Tracking No. 18624), "Spent Fuel Loading of Dry Shielded Canisters 6A-10B", 6/9/2017.
2.30 Transnuclear, "Final Safety Analysis Report, Standardized NUHOMS Horizontal Modular Storage System for Irradiated Nuclear Fuel", NUH-003 Revision 11, Appendix T.
2.31 USNRC 10 CFR Part 72, "Licensing Requirements For The Independent Storage Of Spent Nuclear Fuel, High
-Level Radioactive Waste, And Reactor
-Related Greater Than Class C Waste". 2.32 Holtec International .Topical Safety Analysis Report for the Holtec International Storage, Transport, and Repository Cask System (HI-Star 100 Cask System). NRC Docket No. 721008, Holtec Report HI-941184, Rev. 8. August 1998.
2.33 Design Information Transmittal (DIT) 67492-004 (Mod/Tracking No. 18624), "Spent Fuel Loading of Dry Shielded Canisters 6A-10B", 7/13/2017. 2.34 AREVA Calculation 11042-0400 R0, "NUHOMS° 61BTH Type 1 DSCs for Monticello Nuclear Generating Plant".
CALCULATION NO. MNGP-018 PAGE 12 of 38 REV 0 3.1 ORIGEN-ARP calculations are performed on a 1 MTU reference basis.
3.2 For DSC accident, fraction of fuel rods assumed to fail is 1.0 (100% of available fuel rods), Reference 2.11, Item A.9.
3.3 For DSC accident, fraction of gases in fuel rods assumed available for release is 1.0 (100%).
3.4 For the DSC accident, a public individual is assumed present at the nearest EAB boundary for the accident duration, i.e., 100% Occupancy Factor.
3.5 -1, use of the Reference 2.21 "best estimate" (50 th percentile) settling velocity of 0.00082 m/sec is used. Use of the 50 th percentile or "best estimate" is deemed as the most appropriate methodology for the evaluation of a "realistic" accident dose.
3.6 Calculations to be performed at 102% of the re-rated power level (1775 MWth) (Also referred to as Power Uprate, PU) which was the highest power limit when the subject fuel assemblies in DSCs 11 - 15 were irradiated., see Section 4.9.
3.7 MNGP site specific preliminary AREVA Calculation 11042-0400 R0 (Reference 2.34) based on actual DSC 11-15 heat loads predicts a peak accident DSC pressure of 45.91 psig (use 46 psig) and a peak accident temperature of 404.75 °F (use 405 °F).
3.8 No additional assumptions used.
CALCULATION NO. MNGP-018 PAGE 13 of 38 REV 0 ORIGEN-ARP DSC Source Term:
Design Input items 4.1 - 4.4 are initially pulled from calculation 16-090 R0, Reference 2.6 (with the exception of the reference numbers and wording) and modified as described below. This information is provided herein for completeness as the ORIGEN-ARP model used to determine the source term within the MNGP DSC is the exact same model as developed in 16-090 R0 (See Reference 2.11, Item A.1) with the exception of post-operation decay time and average fuel exposure (See Sections 4.3, 4.5 and Section 6.1 discussions). 4.1 The baseline DSC ORIGEN-ARP source term analysis is performed at 102% of the EPU power level (2044 MWth, Reference 2.7 Item B.1). Final ORIGEN-ARP results are scaled in Attachment 5 to PU (re-rated) power, see Section 4.9. Note that this is an appropriate conservative analytical approach for this evaluation because the actual fuel loaded in DSCs 11
- 15 pre-date MNGP EPU. The fuel stored in DSCs 11 -15 have an exposure based on plant operation at PU (1775 MWth) and 41 GWD/MTU or less
.
4.2 The MNGP EPU core is based on GE14 fuel. The design parameters of that fuel are identified per Table 3.2.1 of Reference 2.8 which is repeated below (Also see Reference 2.6). Item Parameter Value Units 1 Fuel type GE14 N/A 2 Fuel bundle mass (Uranium) 182 Kg 3 Fuel bundle average enrichment 4.6 % 4 MNGP specific fuel bundle thermal power density based on the 2044 MWth and 484 fuel bundles per core data from above (2044 MWth / 484 fuel bundles) 4.223 MWth 5 Core average end of cycle exposure GWD/MTU 6 Maximum discharge bundle exposure 58 GWD/MTU 4.3 In the Reference 2.
9 and Reference 2.6 ORIGEN-ARP model, the core average end of cycle exposure considered was 37 GWD/MTU. As noted in Item 4.1 and Reference 2.33 Item A.1, the DSCs 11-15 canister average fuel assembly exposure is based on PU plant operation (1775 MWth) at a maximum of 41 GWD/MTU. See the Attachment 12 spreadsheet which calculates the average exposure of the 61 fuel assemblies within each of DSCs 11-15 based on the Cask Loading Reports supplied via Reference 2.11, Item A.1.
4.4 Number of fuel assemblies (bundles) in MNGP GE14 core = 484 (Reference 2.10, Attachment 1, Item 3.c)
CALCULATION NO. MNGP-018 PAGE 14 of 38 REV 0 4.5 Decay time of fuel within DSC is 15.53 years (Reference 2.11, Item A.2). From inspection of Cask Loading Reports (CLRs) the lowest value of Cooling Years is 10.06, but that is just one assembly (an outlier). Since the source term is a summation of all assemblies in a DSC, it is mathematically appropriate to use an average across the entire canister (CLRs report Cooling Years ranging up to 20.30 years). However, to keep the calculation simple, it is appropriate and conservative to use the second-lowest cooling
-time increment (11.53 years associated with a discharged batch of fuel) and add the 4 years that have transpired from the data date on the CLRs (5/16/2017 minus 5/16/2013).
4.6 MNGP DSC is a NUHOMS 61BTH Type 1 Model containing 61 fuel assemblies (Reference 2.11, Item A.3)
4.7 Source term isotopes of concern determined using ORIGEN-ARP to be consistent with the guidance of NUREG-1567 (Reference 2.12) Section 9 and Table 9
.2 and Reference 2.14 Table 5.2 specifications (See Reference 2.11, Items A.4 and A.5). 1. Crud as Co-60 determined as per DI Item 4.8 (See Reference 2.12, Table 9.2 footnote "#",
Co-60 activity specified as 1254 µCi/cm 2 and Reference 2.14, Table 5.2
). 2. Iodines 3. Fission products which are greater than 0.1% of total design basis activity
- 4. Actinides which are greater than 0.01% of total design basis activity 4.8 DSC Co-60 activity per fuel assembly calculated as follows.
Fuel Rod Activity = 1254 µCi/cm 2 per DI Item 4.7 Maximum Crud Reduction Factor = 2 per NUREG 1864 Table D.1, axial distribution (Reference 2.13)
Fuel Rods per Assembly = 92 (Reference 2.11, Item A.6) Fuel Rod Length = 150 inches (Reference 2.11, Item A.7) Fuel Rod Outside Diameter = 0.483 inch (Reference 2.11, Item A.8)
Fuel Assembly Surface (cm 2 * (0.483
- 2.54) = 135,096 Co-60 half-life = 5.2714 years (Reference 2.15, Table 1, page 277)
Co-60 Activity (Ci) = ((1254 / 2) *135096) / 1E+06 * (exp (-ln(2)*15.53/5.2714)) = 10.99 4.9 The Power Uprate (PU) power level is 102% if 1775 MWth = 1810.5 MWth (References 2.20 and 2.33, Item A.1). The ORIGEN
-ARP source term (Attachment 13) will be scaled to this power level in Attachment 5
.
CALCULATION NO. MNGP-018 PAGE 15 of 38 REV 0 RADTRAD DSC Dose Calculation:
4.10 Isotopic fuel rod activity released from fuel rods to DSC volume (See Reference 2.11, Item A.10).
NUREG-1567 Table 9.2:
Gases Fraction = 0.3 Volatiles Fraction = 2.0E-04 Per (per footnote, Cs-134, 135 & 137; Ru-103 & 106, Sr-89 & 90) Fuel fines Fraction = 3.0E-05 Crud (Co-60) fraction = 1.0
Co-60 activity release fraction is reduced to 0.015 based on further justification per Reference 2.13, Section D.2.4.2 and Table D.7
. 4.11 Adult Breathing Rate = 2.50E-04 m 3/sec per Reference 2.11, Item A15 and Reference 2.16.
4.12 Distance to nearest EAB location from SW corner of HSM-6A = 245 m (Reference 2.29). Also given distance to nearest EAB location from SW corner of 30 HSM Array = 235 m (Reference
2.11, Item A.13).
4.13 2, page 9-15 based on wind stability Class F and a wind speed of 1 m/sec)
Nearest EAB Distance = 235 m minimum (See DI Item 4.12 above) Wind Stability Class = F Wind Speed = 1 m/sec
Per Reference 2.17 Section 1.3.1.a: Min Vertical Plane X
-Sectional Area = 0 m (conservative value)
Note that the scale of Figures 1, 2 & 3 (log-log scale) are too large to distinguish between 235 and 245 m. y, m = 10 (Interpolated per Reference 2.17, Figure 1) z, m = 5 (Interpolated per Reference 2.17, Figure 2) M = 4 (Per Reference 2.17, Figure 3) y = M
- y = 4
- 10 = 40
Equation 1 = 1 / (y
- z + 0/2)) = 6.37E-03 sec/m3 Equation 2 y
- z )) = 2.12E-03 sec/m3 Equation 3 y
- z )) = 1.59E-03 sec/m 3
6.37E-03 sec/m3 value from above and Eqn. 3 = 1.59E-03 sec/m3 See Section 6.3 for discussion of modification
CALCULATION NO. MNGP-018 PAGE 16 of 38 REV 0 4.14 Inhalation Dose Conversion Factors per Reference 2.18 and Reference 2.11, Item A.16 with additional guidance per Reference 2.3.
4.15 Submersion Dose Conversion Factors per Reference 2.19 and Reference 2.11, Item A.17 with additional guidance per Reference 2.3.
4.16 DSC free volume = 365000 in 3 = 211.2 ft 3, Reference 2.11, Item A.21.
4.17 DSC channel opening dimension are 5.8 inch by 5.8 inch, Reference 2.29.
4.18 DSC channel length = 164 inches per Reference 2.
- 29.
4.19 DSC cavity free volumes
- Table T.4-29 per Reference 2.11, Item A.21 & Reference 2.30. Cavity volume = 618,766 in 3 Basket volume = 108,888 in 3 Fuel volume = 141,947 in 3 Bounding free volume = 365,000 in 3
4.20 DSC accident pressure - 46 psig per Reference 2.
33, Item A.22 and Assumption 3.7 (requiring later verification)
4.21 DSC accident temperature
- 405 °F per Reference 2.33, Item A.23 and Assumption 3.7 (requiring later verification)
4.22 DSC gas mass from Table T.4-24 of Reference 2.30: DSC Cavity He Fill Gas = 192.9 g-mole Fuel Rod He Gas = 83 g
-mole Fission Product Gases = 369.7 g-mole
CALCULATION NO. MNGP-018 PAGE 17 of 38 REV 0 5.1 ORIGEN-ARP DSC Source Term:
The first step in this assessment is to re-run the Reference 2.6 Attachment 3 ORIGEN-ARP model and confirm a match.
The second step in this assessment is to modify the existing ORIGEN-ARP model from Attachment 3 of Reference 2.6 to reflect the additional decay time consistent with the fuel assemblies within the subject MNGP DSC and to subsequently run ORIGEN-ARP with an average fuel exposure of 37 GWD/MTU. This step is also used as a confirmation, this time to confirm that the isotopic activity at a decay time of 180 days matches the existing ORIGEN
-ARP model from Attachment 3
.
The third step repeats step two with the additional modification of an average fuel exposure of 41 GWD/MTU.
The fourth step is to take the step 3 ORIGEN-ARP output and manipulate this data to: 1. Determine fission product isotopes that meet the 0.1% of total design basis activity criterion
- 2. Determine actinide isotopes that meet the 0.01% of total design basis activity criterion
- 3. Convert the step 3
& 4 isotopic data (units of Ci/MTU) to Ci activity in a single DSC
This fourth step is accomplished using an Excel spreadsheet (Attachment 5)
. 5.2 RADTRAD DSC Dose Calculation:
The organ dose calculations are performed using the RADTRAD computer code (Reference 2.3) and the Attachment 5 Excel spreadsheet. A simple two volume (DSC & Environment), one pathway (DSC to Environment) and one dose location (EAB) model is developed. Default RADTRAD input files are modified as outlined below:
- 1. Source Term File (*.NIF) - 14 of the Table 6-1 isotopes exist in the default BWR LOCA "NIF" file. For those 14 isotopes, the Table 6-1 source is input appropriately.
An additional 7 isotopes (plus Sm-151) from Table 6.1 are added to the new "NIF" file by replacing default isotopes 21 - 28 which are not needed for this calculation. The source term for the remaining 39 isotopes (default 60 - 21) are zeroed.
- 2. Release Fraction & Timing File (*.RFT)
- The default BWR "RFT" is modified to instantaneously (0.01 Hours) release the Table 6.1 source term to the RADTRAD DSC volume in accordance with the defined "Fuel to DSC" release fractions given in DI Item 4.10. Release fractions for nuclide groups not represented by the 21 isotopes in Table 6.1 are zeroed.
- 3. Dose Conversion Factors, D CF (*.INP) - Inhalation DCFs are given per DI Item 4.14. Submersion DCFs are given per DI Item 4.15.
CALCULATION NO. MNGP-018 PAGE 18 of 38 REV 0 6.1 ORIGEN-ARP DSC Source Term:
The ORIGEN-ARP input deck is generated by importing the Reference 2.6, Attachment 3 ORIGEN-ARP model as described below. Changes to the Reference 2.6, Attachment 3 model to reflect the DI Item 4.3 specific DSCs 11-15 fuel assembly burnup and DI Item 4.5 DSC fuel assembly decay time of 15.53 years are shown in bold italics below.
Title: Fuel Type:
ORIGEN-ARP fuel type GE10x10-8 which represents GE 10x10 fuel (GE14) as per calculation Item 4.2 Uranium Mass: 1.0E6 grams (Assume 1.0 MTU as reference basis, Assumption 3.1) Fuel Enrichment: 4.6% per calculation Item 4.2 Fuel Burnup: 000 MWD/MTU ( GWD/MTU) per calculation Item 4.3 Irradiation Cycles: 3 (Using default number. The problem has 1 decay cycle following the third irradiation cycle and no intermediate decay cycles)
Cooling Time:
days (Assumed decay duration with default intervalsModerator Density: 0.7332 g/cc (Default value) Power History: 100% continuous operation (No intermediate decay cycles)
Average Power:
Data from calculation Item 4.2 and 4.4 1000kg/MTU
- 2044 MWth/ (484 bundles
- 182 kg/bundle) = 23.2 MWth/MTU Decay Cycle Options:
Access by selecting the "Apply" and "OK" tabs to exit the "Express" option. On leftmost column of ORIGEN-ARP screen select "Cases" tab. On the pop-up dialog box entitled "Case Data" under the "Select Existing Case" column select the "Decay
- 3 Cycle Down" case and Click "OK" at bottom of dialog box. On the lower horizontal tab bar select the "Options" tab. In the "Decay Output Options" dialog box, change "Table cutoff to 0.00%. Change "Results In" from grams to Curies. Check a ll selections under the "Tables" and "Edit By" tabs then click "OK" at dialog box bottom. Save and run case.
The input data was developed using the ORIGEN-ARP GUI. The problem was executed on a microcomputer operating under WindowsXP and the input/output files are saved as MNGP DSC Decay - 0 Cutoff - 41 GWD.* and may be found as Attachment 13 of this calculation.
CALCULATION NO. MNGP-018 PAGE 19 of 38 REV 0 In Attachment 5, Excel Spreadsheet entitled "MNGP
- DSC Source Term to RADTRAD", worksheet entitled "MNGP EPU - GE14 Fuel - 37 GWD", the Reference 2.6, Attachment 3 ORIGEN-ARP selected isotopic output data is copied and pasted for comparison purposes.
In Attachment 5, Excel Spreadsheet entitled "MNGP
- DSC Source Term to RADTRAD",
worksheet entitled "DSC Decay - 0 Cutoff - 37 GWD", the Attachment 4 ORIGEN-ARP selected isotopic output data from this calculation is summarized in columns A-L. The Attachment 4 ORIGEN-ARP model is identical to the Attachment 13 model described above except that the fuel burnup value is 37000 MWD/MTU. The sole purpose of Attachment 4 is to demonstrate that with the modification of decay timing that the output isotopic activity at 180 days replicates the 180 day decay data from the Attachment 3 base case. There is no additional use for the Attachment 4 output. Column Q is used to tabulate the Reference 2.6, Attachment 3 isotopic activity with 180 days decay in Ci/MTU. Column R is used to tabulate the worksheet isotopic activity with 180 days decay in Ci/MTU. Column S is used to calculate the Columns Q/R isotopic activity ratio. A value of 1.0 represents a match.
In Attachment 5, Excel Spreadsheet entitled "MNGP
- DSC Source Term to RADTRAD", worksheet entitled "DSC Decay - 0 Cutoff - 41 GWD", the Attachment 13 ORIGEN-ARP selected isotopic output data from this calculation is summarized in columns A
-L. Columns U & V are used to determine which isotopes meet the specifications of DI Item 4.7. As an actinide example, Cell U8 calculates whether the Cell L8 5668 day decay value of 3.122E-13 Ci/MTU exceeds the 0.01% threshold of total actinide activity given in Cell L150 of 5.610E+04 Ci/MTU.
As a fission product example, Cell U160 calculates whether the Cell L160 5668 day decay value of 2.864E+02 Ci/MTU exceeds the 0.1% threshold of total fission product activity given in Cell
L1424 of 3.246E+05 Ci/MTU. Isotopes that meet the criteria are highlighted in Column V.
In Attachment 5, Excel Spreadsheet entitled "MNGP
- DSC Source Term to RADTRAD",
worksheet entitled "Accident Source Term", the isotopes that meet the DI Item 4.7 criteria are tabulated in Column A. In column B these isotopes are characterized according to DI Item 4.10 categories. In Column C the isotopic Ci/MTU from the Worksheet entitled "DSC Decay - 0 Cutoff - 41 GWD" are copied. In Column D, the baseline core average power in MWth/MTU is entered (See Section 6.1, ORIGEN-ARP input description). In column E, 102% of PU is entered as the fuel contained in the DSC has an exposure consistent with the PU (See References 2.20, 2.33, Item A.1 and DI Item 4.9). In Column F, the average core isotopic Ci are calculated as the product of Columns C, D & E (For example, Am-241 - Cell F7 (1.550E+05 Ci) is the product of Cell C7 / Cell D7
- Cell E7).
CALCULATION NO. MNGP-018 PAGE 20 of 38 REV 0 In Column G, the number of fuel assemblies per MNGP core is entered as 484 (See Design Input Item 4.4). In Column H, the isotopic activity in Ci per fuel assembly is calculated (For example,
Am-241 - Cell H7 (3.202E+02 Ci) is the product of Cell F7 / Cell G7). In Column J, the number of fuel assemblies in the MNGP NUHOMS 61BTH Type 1 DSC is entered (See DI Item 4.6). In Column K, the isotopic activity per DSC is calculated (For example, Am-241 - Cell K7 (1.9533E+04 Ci) is the product of Cell H7
- Cell J7). The Column K results are subsequently used as input to the RADTRAD source term file given as Attachment 6. This work is provided below as Table 6.1.
CALCULATION NO. MNGP-018 PAGE 21 of 38 REV 0 DSCs 11-15 DSC Total EPU-41 GWD DSCs 11-15 DSCs 11-15 Fuel Assembly Average DSC Activity in Nuclide EPU-41 GWD Avg Power Power Level Avg Core Per Core Fuel Assembly Total Fuel All Fuel Rods Isotope Group (Curies/MTU)
(MWth/MTU)
(MWth) (Curies) (Curies) Assemblies (Curies) Am-241 Fuel Fines 1.986E+03 2.320E+01 1.8105E+03 1.550E+05 4.84E+02 3.202E+02 61 1.9533E+04 Am-242m Fuel Fines 8.342E+00 2.320E+01 1.8105E+03 6.510E+02 4.84E+02 1.345E+00 61 8.2047E+01 Am-242 Fuel Fines 8.304E+00 2.320E+01 1.8105E+03 6.480E+02 4.84E+02 1.339E+00 61 8.1674E+01 Am-243 Fuel Fines 1.737E+01 2.320E+01 1.8105E+03 1.356E+03 4.84E+02 2.801E+00 61 1.7084E+02 Ba-137m Fuel Fines 8.543E+04 2.320E+01 1.8105E+03 6.667E+06 4.84E+02 1.377E+04 61 8.4024E+05 Cm-242 Fuel Fines 6.868E+00 2.320E+01 1.8105E+03 5.360E+02 4.84E+02 1.107E+00 61 6.7550E+01 Cm-243 Fuel Fines 1.003E+01 2.320E+01 1.8105E+03 7.827E+02 4.84E+02 1.617E+00 61 9.8650E+01 Cm-244 Fuel Fines 9.512E+02 2.320E+01 1.8105E+03 7.423E+04 4.84E+02 1.534E+02 61 9.3555E+03 Co-60 Crud 1.099E+01 61 6.7047E+02 Cs-134 Volatiles 7.550E+02 2.320E+01 1.8105E+03 5.892E+04 4.84E+02 1.217E+02 61 7.4258E+03 Cs-137 Volatiles 9.047E+04 2.320E+01 1.8105E+03 7.060E+06 4.84E+02 1.459E+04 61 8.8981E+05 Eu-154 Fuel Fines 1.690E+03 2.320E+01 1.8105E+03 1.319E+05 4.84E+02 2.725E+02 61 1.6622E+04 Kr-85 Gaseous 4.557E+03 2.320E+01 1.8105E+03 3.556E+05 4.84E+02 7.348E+02 61 4.4820E+04 Np-239 Fuel Fines 1.737E+01 2.320E+01 1.8105E+03 1.356E+03 4.84E+02 2.801E+00 61 1.7084E+02 Pm-147 Fuel Fines 3.136E+03 2.320E+01 1.8105E+03 2.447E+05 4.84E+02 5.056E+02 61 3.0844E+04 Pu-238 Fuel Fines 2.532E+03 2.320E+01 1.8105E+03 1.976E+05 4.84E+02 4.083E+02 61 2.4903E+04 Pu-239 Fuel Fines 2.746E+02 2.320E+01 1.8105E+03 2.143E+04 4.84E+02 4.428E+01 61 2.7008E+03 Pu-240 Fuel Fines 4.956E+02 2.320E+01 1.8105E+03 3.868E+04 4.84E+02 7.991E+01 61 4.8745E+03 Pu-241 Fuel Fines 4.979E+04 2.320E+01 1.8105E+03 3.886E+06 4.84E+02 8.028E+03 61 4.8971E+05 Sr-90 Volatiles 6.868E+04 2.320E+01 1.8105E+03 5.360E+06 4.84E+02 1.107E+04 61 6.7550E+05 Y-90 Fuel Fines 6.870E+04 2.320E+01 1.8105E+03 5.361E+06 4.84E+02 1.108E+04 61 6.7570E+05 3.795E+05 3.7334E+06
CALCULATION NO. MNGP-018 PAGE 22 of 38 REV 0 6.2 RADTRAD DSC Dose Calculation:
RADTRAD "NIF" File:
The RADTRAD BWR default nuclide inventory as shown in Reference 2.3, Section 1.4.3.2 and Table 1.4.3.2-3 is modified for the purposes of this calculation as described below.
- 1. As noted in calculation Section 5.2, 14 of the isotopes from Table 6.1 are already part of the RADTRAD Table 1.4.3.2-3 nuclide inventory. The Ci value from Table 6.1 is input as line 6 for each of these isotopes.
- 2. The remaining 7 isotopes (plus Sm-151) are incorporated as nuclides 21 - 28 of the modified "NIF" file (replacing the RADTRAD default nuclides 21 - 28. For isotopes Ba-137m, Pm-147, Eu-154, Am-243 and Cm-243 nuclide parameters are as previously developed in Reference 2.21, Appendix A nuclides 31, 33, 34, 36 and 37 (except for decay times as noted). Sm-151 (previously developed for Attachment 4, but failed to satisfy the Design Input Item 4.7 criteria based on Attachment 13 results), Am-242m and Am-242 data is developed below.
Nuclide 021:
Ba-137m 6 1.5312000000E+02 Reference 2.15, Table 1, page 1242 0.1370E+03 8.4024E+05 Table 6.1 none 0.0000E+00 none 0.0000E+00 none 0.0000E+00 Nuclide 022:
Pm-147 9 8.2731542400E+07 Reference 2.15, Table 1, page 1431 0.1470E+03 3.0844E+04 Table 6.1 none 0.0000E+00 none 0.0000E+00 none 0.0000E+00 Nuclide 023:
Eu-154 9 2.7098884800E+08 Reference 2.15, Table 1, page 1630 0.1540E+03 1.6622E+04 Table 6.1 none 0.0000E+00 none 0.0000E+00 none 0.0000E+00 Nuclide 024:
CALCULATION NO. MNGP-018 PAGE 23 of 38 REV 0 Sm-151 9 Lanthanides Series - Reference 2.22, Table 5 2.8382400000E+09 Reference 2.15, Table 1, page 1532 0.1510E+03
0.0000E+00 Per Attachment 13, fails DI Item 4.7 criteria none 0.0000E+00 Reference 2.15, Table 1, page 1532 none 0.0000E+00 none 0.0000E+00 Nuclide 025:
Am-242m 9 Lanthanides Series - Reference 2.22, Table 5 4.4465760000E+09 Reference 2.15, Table 1, page 2807 0.2420E+03 8.2047E+01 Table 6.1 Am-242 1.0000E+00 Reference 2.15, Table 1, page 2807 none 0.0000E+00 none 0.0000E+00 Nuclide 026:
Am-242 9 Lanthanides Series - Reference 2.22, Table 5 5.7672000000E+04 Reference 2.15, Table 1, page 2807 0.2420E+03 8.1674E+01 Table 6.1 Cm-242 0.8270E+00 Reference 2.15, Table 1, page 2807 none 0.0000E+00 none 0.0000E+00 Nuclide 027:
Am-243 9 2.3242032000E+11 Reference 2.15, Table 1, page 2812 0.2430E+03 1.7084E+02 Table 6.1 none 0.0000E+00 none 0.0000E+00 none 0.0000E+00 Nuclide 028:
Cm-243 9 9.1769760000E+08 Reference 2.15, Table 1, page 2812 0.2430E+03 9.8650E+01 Table 6.1 none 0.0000E+00 none 0.0000E+00 none 0.0000E+00 CALCULATION NO. MNGP-018 PAGE 24 of 38 REV 0 3. Nuclides other than the 21 identified in Table 6.1 have their activity zeroed (line 6).
The RADTRAD "NIF" file for the MNGP DSC source term is given as Attachment 6.
RADTRAD "RFT" File: The RADTRAD BWR default release and timing file as shown in Reference 2.3, Section 1.4.3.1 and Table 1.4.3.1-3 is modified for the purposes of this calculation as described below.
- 1. Gap release timing is set to 0.01 hours1.157407e-5 days <br />2.777778e-4 hours <br />1.653439e-6 weeks <br />3.805e-7 months <br /> (instantaneous) consistent with Reference 2.21, Appendix A. All other times set to zero.
- 2. 6 RADTRAD nuclide groups are represented by the 21 isotopes in Table 6.1. The release fraction for these isotopes is defined per DI Item 4.10. All other RADTRAD nuclide groups release fractions are zeroed.
Note that Co-60 is part of the Noble Metals Group (or Ruthenium Series).
The RADTRAD "RFT" file for the MNGP DSC source term is given as Attachment 7.
CALCULATION NO. MNGP-018 PAGE 25 of 38 REV 0 RADTRAD "INP" File: The RADTRAD NUREG-1465 default conversion factors file as shown in Reference 2.3, Section 1.4.3.3 and Table 1.4.3.3-2 is modified for the purposes of this calculation as described below.
- 1. As noted in calculation Section 5.2, 14 of the isotopes from Table 6.1 are already part of the RADTRAD Table 1.4.3.3-2 nuclide inventory. Note that the values provided in RADTRAD Table 1.4.3.3-2 are consistent with Reference 2.18 and 2.19 (one exception is the Cs-137 Submersion DCF value where the Table 1.4.3.3-2 table list DCFs for the combined Cs-137 -
Ba-137m parent daughter. However, since the FGR 12 DCF value for Ba-137m is higher than the RADTRAD default file DCF for Cs-137, the FGR 12 DCF values are input for Ba-137m. 2. The remaining 7 isotopes (plus Sm-151) are incorporated as nuclides 21 - 28 of the modified
"INP" file (replacing the RADTRAD default nuclides 21 - 28. For isotopes Ba-137m, Pm-147, Eu-154, Sm-151 (DCFs zeroed), Am-242m, Am-242, Am-243 and Cm-243 DCFs are taken from References 2.18 (FGR 11, Table 2.1) and 2.19 (FGR 12, Table III-1). 3. All DCF values not used are zeroed. Thus DCF values remain only for the 9 organs listed for RADTRAD designated columns "Inhaled Chronic" (FGR 11 - Inhalation) and "Cloudshine" (FGR 12 - Submersion).
- 4. Inhalation DCFs are tabulated in Table 6.2 below. Submersion DCFs are tabulated in Table 6.3 below.
The base RADTRAD "INP" file for the MNGP DSC source term is given as Attachment 8.
CALCULATION NO. MNGP-018 PAGE 26 of 38 REV 0 Am-241 3.250E-05 2.670E-09 1.840E-05 1.740E-04 2.170E-03 1.600E-09 7.820E-05 1.200E-04 0.000E+00 Class W Am-242m 3.210E-05 1.380E-09 4.200E-06 1.690E-04 2.120E-03 5.640E-10 7.480E-05 1.150E-04 0.000E+00 Class W Am-242 1.940E-09 2.940E-12 5.200E-08 1.320E-08 1.650E-07 2.520E-12 8.540E-09 1.580E-08 0.000E+00 Class W Am-243 3.260E-05 1.520E-08 1.780E-05 1.730E-04 2.170E-03 8.290E-09 7.740E-05 1.190E-04 0.000E+00 Class W Ba-137m 0.000E+00 0.000E+00 0.000E+00 0.000E+00 0.000E+00 0.000E+00 0.000E+00 0.000E+00 0.000E+00 Cm-242 5.700E-07 9.440E-10 1.550E-05 3.900E-06 4.870E-05 9.410E-10 2.450E-06 4.670E-06 0.000E+00 Class W Cm-243 2.070E-05 6.290E-09 1.940E-05 1.180E-04 1.470E-03 3.830E-09 5.760E-05 8.300E-05 0.000E+00 Class W Cm-244 1.590E-05 1.040E-09 1.930E-05 9.380E-05 1.170E-03 1.010E-09 4.780E-05 6.700E-05 0.000E+00 Class W Co-60 4.760E-09 1.840E-08 3.450E-07 1.720E-08 1.350E-08 1.620E-08 3.600E-08 5.910E-08 0.000E+00 Class Y Cs-134 1.300E-08 1.080E-08 1.180E-08 1.180E-08 1.100E-08 1.110E-08 1.390E-08 1.250E-08 0.000E+00 Class D Cs-137 8.760E-09 7.840E-09 8.820E-09 8.300E-09 7.940E-09 7.930E-09 9.120E-09 8.630E-09 0.000E+00 Class D Eu-154 1.170E-08 1.550E-08 7.920E-08 1.060E-07 5.230E-07 7.140E-09 1.130E-07 7.730E-08 0.000E+00 Class W Kr-85 0.000E+00 0.000E+00 0.000E+00 0.000E+00 0.000E+00 0.000E+00 0.000E+00 0.000E+00 0.000E+00 Np-239 7.450E-11 1.630E-11 2.360E-09 2.080E-10 2.030E-09 7.620E-12 9.590E-10 6.780E-10 0.000E+00 Class W Pm-147 8.250E-15 3.600E-14 7.740E-08 1.610E-09 2.010E-08 1.980E-14 1.560E-09 1.060E-08 0.000E+00 Class Y Pu-238 1.040E-05 4.400E-10 3.200E-04 5.800E-05 7.250E-04 3.860E-10 2.740E-05 7.790E-05 0.000E+00 Class Y Pu-239 1.200E-05 3.990E-10 3.230E-04 6.570E-05 8.210E-04 3.750E-10 3.020E-05 8.330E-05 0.000E+00 Class Y Pu-240 1.200E-05 4.330E-10 3.230E-04 6.570E-05 8.210E-04 3.760E-10 3.020E-05 8.330E-05 0.000E+00 Class Y Pu-241 2.760E-07 2.140E-11 3.180E-06 1.430E-06 1.780E-05 9.150E-12 6.020E-07 1.340E-06 0.000E+00 Class Y Sm-151 0 0 0 0 0 0 0 0 0 Class W Sr-90 2.690E-10 2.690E-10 2.860E-06 3.280E-08 7.090E-08 2.690E-10 5.730E-09 3.510E-07 0.000E+00 Class Y Y-90 5.170E-13 5.170E-13 9.310E-09 1.520E-11 1.510E-11 5.170E-13 3.870E-09 2.280E-09 0.000E+00 Class Y CALCULATION NO. MNGP-018 PAGE 27 of 38 REV 0 Am-241 8.580E-16 1.070E-15 6.740E-16 5.210E-16 2.870E-15 7.830E-16 6.340E-16 8.180E-16 1.280E-15 Am-242m 3.800E-17 6.010E-17 1.720E-17 1.720E-17 7.940E-17 2.950E-17 1.940E-17 3.170E-17 1.360E-16 Am-242 6.090E-16 7.300E-16 5.510E-16 4.770E-16 1.880E-15 5.940E-16 5.180E-16 6.150E-16 8.200E-15 Am-243 2.190E-15 2.610E-15 1.920E-15 1.550E-15 7.470E-15 2.090E-15 1.790E-15 2.180E-15 2.750E-15 Ba-137m 2.820E-14 3.220E-14 2.800E-14 2.730E-14 4.630E-14 2.880E-14 2.680E-14 2.880E-14 3.730E-14 Cm-242 7.830E-18 1.480E-17 1.130E-18 1.890E-18 1.060E-17 4.910E-18 2.270E-18 5.690E-18 4.290E-17 Cm-243 5.770E-15 6.680E-15 5.500E-15 5.000E-15 1.500E-14 5.760E-15 5.190E-15 5.880E-15 9.790E-15 Cm-244 6.900E-18 1.330E-17 7.080E-19 1.460E-18 8.820E-18 4.190E-18 1.810E-18 4.910E-18 3.910E-17 Co-60 1.230E-13 1.390E-13 1.240E-13 1.230E-13 1.780E-13 1.270E-13 1.200E-13 1.260E-13 1.450E-13 Cs-134 7.400E-14 8.430E-14 7.370E-14 7.190E-14 1.200E-13 7.570E-14 7.060E-14 7.570E-14 9.450E-14 Cs-137 2.669E-14 3.047E-14 2.649E-14 2.583E-14 4.382E-14 2.725E-14 2.536E-14 2.725E-14 4.392E-14 Eu-154 6.000E-14 6.810E-14 5.990E-14 5.870E-14 9.430E-14 6.150E-14 5.750E-14 6.140E-14 8.290E-14 Kr-85 1.170E-16 1.340E-16 1.140E-16 1.090E-16 2.200E-16 1.180E-16 1.090E-16 1.190E-16 1.320E-14 Np-239 7.530E-15 8.730E-15 7.180E-15 6.500E-15 2.000E-14 7.520E-15 6.760E-15 7.690E-15 1.600E-14 Pm-147 7.480E-19 9.560E-19 5.450E-19 4.460E-19 2.180E-18 6.750E-19 5.260E-19 6.930E-19 8.110E-16 Pu-238 6.560E-18 1.270E-17 1.060E-18 1.680E-18 9.300E-18 4.010E-18 1.990E-18 4.880E-18 4.090E-17 Pu-239 4.840E-18 7.550E-18 2.650E-18 2.670E-18 9.470E-18 3.880E-18 2.860E-18 4.240E-18 1.860E-17 Pu-240 6.360E-18 1.230E-17 1.090E-18 1.650E-18 9.260E-18 3.920E-18 1.960E-18 4.750E-18 3.920E-17 Pu-241 7.190E-20 8.670E-20 6.480E-20 5.630E-20 2.190E-19 6.980E-20 6.090E-20 7.250E-20 1.170E-19 Sm-151 0 0 0 0 0 0 0 0 0 Sr-90 7.780E-18 9.490E-18 6.440E-18 5.440E-18 2.280E-17 7.330E-18 6.110E-18 7.530E-18 9.200E-15 Y-90 1.890E-16 2.200E-16 1.770E-16 1.620E-16 4.440E-16 1.870E-16 1.680E-16 1.900E-16 6.240E-14
CALCULATION NO. MNGP-018 PAGE 28 of 38 REV 0 RADTRAD Model:
A simple 2 volume, one pathway and one dose location RADTRAD model is developed to analyze the postulated DSC accident dose to the site boundary (EAB). The RADTRAD input is developed below.
RADTRAD Volume 1 represents the MNGP DSC: DSC free volume is modeled as 211.2 ft3 (DI Item 4.16) 100% of the Table 6.1 source term released to the D SC volume User defined natural deposition removal coefficient = 20 Hr See below No additional inputs Natural Deposition: The methodology utilized for the calculation of natural deposition of the aerosols released from the assumed damaged fuel rods to the DSC free volume is developed in Reference 2.21, page 7, paragraph underneath Table 8.
Per the above, the preferred method to determine the natural deposition removal coefficient is to divide the given settling velocity by the characteristic fall height within the DSC. For this analysis, the best
-estimate or 50th percentile settling velocity of 0.00082 m/sec is chosen for determination of the removal coefficient.
The fall height is calculated per the methodology described on page 10 of Reference 2.21 underneath Table 12. For a cask on its side, the fall height is modeled to be the free volume of the lodgment divided by the area of one side of the lodgment. For the NUHOMS 61BTH, each of the 61 horizontal channels houses 1 fuel assembly. Fall height calculation parameters are given below.
- 1. Per DI Item 4.17 the channel opening face dimensions are 5.8" x 5.8".
- 2. Per DI Item 4.18, the channel length is 164 inches.
- 3. Per DI Item 4.19, Reference 2.11, Item A.21 indicates that the channel free volume can be calculated using parameters in attached Table T.4-29.
618,766 in 3 (DSC Cavity Volume) - 108,888 in 3 (Basket Volume) = 509,878 in 3 509,878 in 3 - 141,947 in 3 (Fuel Volume) = 367,931 in 3 (free volume) 141,947 in 3 (Fuel Volume) / 61 (Fuel Assemblies) = 2,327 in 3 per channel Channel volume = 5.8 in
- 5.8 in
- 164 in = 5,516.96 in 3 Channel free volume = 5,516.96 in 3 - 2,327 in 3 = 3,189.96 in 3 Channel lodgment side area = 5.8 in
- 164 in = 951.2 in 2 Channel characteristic height = fall height = 3,189.96 in 3 / 951.2 in 2 = 3.3536 in = 0.0852 m
For conservatism, use wall height of 5.8 inch as fall height = 0.14732 m
CALCULATION NO. MNGP-018 PAGE 29 of 38 REV 0 Natural depoHr-1 -1 for 30 day duration. Note that in Reference 2.21, Section entitled "Effect of Aerosol Concentration," an evaluation of the effect of DSC aerosol concentration on the deposition rate is presented. That effect is not considered in the RADTRAD model. However, the Reference 2.21 evaluation demonstrates that the concentration effect is negligible. RADTRAD Volume 2 represents the Environment: No inputs RADTRAD Pathway 1 represents the leakage from the DSC to the Environment: A filter pathway is model ed with filter removal efficiencies set to zero Pathway flow rate is 3.333E-03 CFM (see Section 6.4) conservatively modeled for the 30 day accident duration based on continuous critical flux flow (not terminated when the DSC pressure reduces to 14.7 psia). No additional inputs RADTRAD Dose Location 1
- MNGP EAB: /Q value equals 1.59E-03 sec/m3 for 30 days per DI Item 4.13 Adult breathing rate value 2.50E-04 m 3/sec for 30 days per DI Item 4.11 No additional inputs RADTRAD Source Term:
User Inventory File "DSC Decay Source.NIF" (Attachment 6). This file is developed utilizing the released activity as described in Section 6.1 Modeled plant power set as 1 MWth, since the source is based on curies Model isotopic decay and daughter in-growth User Release Fraction and Timing File "DSC Decay Source.RFT" (Attachment 7)
User Inventory Input File "DSC Decay Source.INP" (Attachment 8) No additional inputs
The resulting RADTRAD input/output files (psf/o0) are given as Attachment 9.
From Attachment 9, the following 30 day organ doses are calculated directly via RADTRAD.
- 1. Thyroid = 1.1366E-04 Rem
Other organ doses are calculated in the Attachment 5 spreadsheet in Worksheet entitled "Accident Source Term" via the following steps.
- 1. Sum each organ submersion DCFs (Example: Gonad Submersion DCF sum in Cell V32 is the sum of Cells V7 -
V27). 2. Sum each organ inhalation DCFs (Example: Gonad Inhalation DCF sum in Cell AI32 is the sum of Cells AI7 - AI27).
CALCULATION NO. MNGP-018 PAGE 30 of 38 REV 0 3. The total combined organ DCF sum is the sum of the total Submersion DCF plus the total Inhalation DCF (Example: the Gonad total combined DCF sum in Cell AI33 is the sum of Gonad Submersion DCF in Cell V32 plus the Gonad Inhalation DCF in Cell AI32).
- 4. The RADTRAD calculated Thyroid dose (1.1366E-04) and the Effective dose (2.1128E-02) are place d is Cells AN34 and AP34, respectively.
- 5. The remaining organ doses are calculated by multiplying the Effective dose by the ratio of the summed organ dose DCF divided by the summed Effective organ dose DCF (Example: Gonad dose (Cell AI35) equals Cell AP34 (2.1128E-02) times [ (Cell AI33, 1.691E-04) / (Cell AP33, 7.550E-04)
] = 4.73E-03.
CALCULATION NO. MNGP-018 PAGE 31 of 38 REV 0 6.3 This section justifies a reduction factor of 4.5 that can be applied to Regulatory Guide 1.145 /Qs as calculated in Section 4.13 to obtain more realistic dispersion factor estimates.
Report PNL-10286 (Reference 2.24), Figure 9 (Attachment 10), shows the biases in Regulatory Guide 1.145's methodology for calculating /Qs compared to data from site tracer tests. Figure 9 shows overestimates of 1 to 2 orders of magnitude for low wind speeds. The lower speeds typically occur with stable to neutral conditions and result in larger X/Q values. At greater wind speeds, the biases decrease. Under prediction may occur at the highest speeds; however, the X/Q values tend to be numerically much smaller than low speed /Qs due to increased wind speed and turbulence (unstable meteorology). Consequentially, the low speeds are more limiting with respect to assessing doses.
The data in PNL-10286, Figure 9, was analyzed to demonstrate the amount of bias in Regulatory Guide 1.145's methods and provide a basis for applying this conservatism to engineering evaluations. The analysis proceeded as follows:
- 1. Three data bins, each of width 0.5 m/s, were used to examine the biases (X/Q overestimates) for low speeds from 0.5 to 2.0 m/s. See Attachment
- 11. 2. The data were visually extracted and tabulated in an Excel spreadsheet as Bins #1 (0.5 to 1 m/s), #2 (1 to 1.5 m/s) and #3 (1.5 to 2 m/s). See Attachment s 10 and 11. 3. The geometric mean was calculated for each bin.
Also, the arithmetic mean and median were determined for perspective.
- 4. The geometric mean is reasonable for this type analysis and follows from Ramsdell in Reference 2.
- 24. Additional justification is provided in the Notes to Attachment
- 11. 5. The geometric means of the biases for low speed Bins #1, #2 and #3 range from 14.7 to 100.
The geometric mean for the range 0.5 to 2 m/s is 30. See Attachment 11.
- 6. The above procedure was repeated for higher speed ranges comprising Bins #4 to #9. Biases may be observed in Attachment 11 to decrease with increasing speed.
Underestimates are more likely at the higher wind speeds.
- 7. The overall geometric mean of the bias across all speeds (0.5 to 12 m/s) is 4.76. See Attachment 11. The overall geometric mean bias (4.76) for the entire ensemble of data appears to be a reasonable, yet conservative factor to apply to /Qs. This value is very conservative relative to the 0.5 to 2 m/s speeds where a much greater overestimate bias occurs (i.e., geometric mean of 30). Thus, use of a is conservative and justifiable for dose values developed in Sections 6.1 and 6.2 and summarized in Section 7.
The bias factor is most applicable to the /Q at the EAB where, normally, conditions are assumed invariant for a two hour interval. It can be applied to longer time intervals as long as the downwind direction, wind speed, and stability are assumed to remain constant. The bias factor use would be valid for bounding type accident assessments where dispersion conditions are not postulated to change for the accident as assumed in this calculation
.
CALCULATION NO. MNGP-018 PAGE 32 of 38 REV 0 6.4 DSC Leakage Rate:
Based on a licensing basis that postulates no confinement failure and a satisfactory "leak tight" helium leak test on each canister (DSC 11 - 15), References 2.23 and 2.30, there is no basis and no method for theorizing a potential leak flow rate based on leak testing requirements.
To arrive at a hypothetical accident release for the DSCs 11-15 with noncompliant dye penetrant examinations
, the following methodology is utilized.
- 1. This hypothetical accident release is not triggered by a cask drop event, fire or any known material stress/corrosion failure process.
- 2. As such, a reasonable upper limit realistic leak diameter (hole size) is postulated to be no larger than the maximum allowable leak diameter associated with packaging and transport of radioactive materials.
- 3. Review of the Table 6.1 DSC activity and Attachment 5 identifies isotopes of interest and is used to determine the specific activity within the DSCs. 10 CFR 71 is then utilized to determine the maximum allowable "package" activity limit
. As shown in the calculation below, Kr-85 is the limiting isotope remaining within the subject DSCs. 4. ANSI N14.5
-1997 may then be used to calculate an allowable release rate (consistent with 10 CFR 71) and via Table B.2 to back-calculate an associated leak diameter (hole size).
- 5. The leak diameter calculated in step 4 above can then be used to calculate a hypothetical accident release based upon the critical mass flux release rate resulting from limiting pressure/temperature conditions within the DSCs from other postulated events. For this assessment, the limiting event is considered to be the Blocked Vent event with elevated internal pressure and temperature conditions within the subject DSCs.
Determination of Leakage Area (Hole Size) - ANSI N14.5-1997 Methodology:
The following procedure will be assumed to determine the postulated accident release hole size. This procedure is deemed reasonable in that the DSC is not subject to a cask drop or fire. The leak is thus based on the 10 CFR Part 71 maximum "package" activity limit (using Kr-85 as the representative isotope, R A values bounds other Table 6.1 isotopes) as a guideline. In accordance with the NUHOMS FSAR, Reference 2.23, Section 1
- 1. Determine C A , average activity in the DSC in Ci/cm 3 per the total activity released into the DSC divided by the DSC free volume.
Determine C A: From Attachment 5, Cell Q28, the activity released into the DSC is calculated to be 1.3834E+04 Ci.
DSC free volume = 5.9813E+06 cm3 (365,000 in
- 3) from DI Item 4.16.
CA = 1.3834E+04 / 5.9813E+06 = 2.3129E-03 Ci/cm 3
CALCULATION NO. MNGP-018 PAGE 33 of 38 REV 0 2. Determine R A , allowable release rate under accident conditions in Ci/sec using the Kr-85 A2 value per Appendix A of 10 CFR Part 71. This is assumed reasonable since the bulk of the activity released from the fuel is Kr-85 (1.3446E+04 Ci, Attachment 5, Cell Q19) in accordance with the ORIGEN
-ARP analysis.
From ANSI N14.5-1997 (Reference 2.25)
RA = 1.65x10-6
- A2 per second in accordance with Section 5.4.2 of ANSI N14.5-1997.
A2 for Kr-85 = 10*A 2 in accordance with Section 6.1 of ANSI N14.5-1997 From 10CFR71 A2 (Kr-85) = 270 Ci per Appendix A of 10 CFR Part 71 (Reference 2.26)
Thus Kr-85A2 = 2700 Ci/sec RA = 2700
- 1.65x10-6 = 4.455x10-3 Ci/sec 3. Determine L A allowable leakage rate of the medium (cm 3/s) under hypothetical accident conditions per Section 6.1 of ANSI N14.5-1997:
LA = RA /CA = 4.455E-03 Ci/sec / 2.3129E-03 Ci/cm 3 = 1.926 cm3/sec 4. Table B2 of ANSI N14.5-1997 lists representative DSC leakage rates vs pressure drop conditions at given leak hole diameters. Interpolating in Table B2 (Column 2 for low dP) of ANSI N14.5-1997 with the step 3 allowable leak rate of 1.926 cm3/sec yields a hole size of approximately 0.0106 cm in diameter, conservatively use 0.011 cm (Area = 9.5033E-05 cm2 = 1.0229E-07 ft 2). Determination of Critical Mass Flux (Leakage Rate) - Based on Blocked Vent Accident Conditions: Per Reference 2.27, Equation 2.60, the critical mass flux for an ideal gas through an orifice with an upstream pressure P 0 is defined as:
= ()/()
where:
k = specific heat ratio
CALCULATION NO. MNGP-018 PAGE 34 of 38 REV 0 g0 = conversion constant or 32.174 ft-lbm/lbf-sec2 (page 8 Reference 2.28) P0 = vessel pressure, psi 0 = gas density in lbm/ft 3 For this case, the gas is composed of DSC He fill gas, fuel gap He gas, and fuel gap fission gas.
Per DI Item 4.
22 (Reference 2.30, Table T.4-24), the gas mass for a single DSC @ accident conditions is listed as:
He in fuel: 83 gr-mole He in DSC (fill gas) = 192.9 gr mole Fission gases in fuel: 369.7 gr-mole Total: 645.6 gr-mole Thus the corresponding gas mass fractions are:
Fission gas mass fraction = 369.7/645.6 = 0.5726 He gas fraction = 1 - 0.5726 = 0.4274
Using ideal gas density calculations, the English units basis for an ideal gas is given on page 130 of Reference 2
.28 as:
P= 14.7 psia T= 492°R Specific volume = 359 ft 3/lbm- mole Per DI Item 4.20 (Reference 2.33, Item A.22), the DSC maximum accident pressure is 46 psig (60.7 psia). Per DI Item 4.
21 (Reference 2.33, Item A.23), the DSC maximum accident temperature is 405 °F or 865 °R.
Thus, for 1 ft 3 Basis of He, at an Atomic Weight of 4.0026 (Appendix B, Reference 2.28
): 1 ft3 * * . . * * . = 0.0262 lbm
62 lbm/ft3 Similarly for the fission gases, assuming the high density Xe as representative, Atomic Weight of 131.30 (Appendix B, Reference 2.28
): 1 ft3 * * . . * * . = 0.8590 lbm 8590 lbm/ft3 CALCULATION NO. MNGP-018 PAGE 35 of 38 REV 0 Using ideal gas law:
CP = CV + R Dividing above by C V, we get C P/CV = 1 + R/C V
For a monoatomic gas, C v = 1.5 R
Thus, k = C P/CV = 1 + R/C V = 1 + R/(1.5R) = 1 + 1/1.5 = 1.6667
Thus k(He) = k(Xe) = 1.6667
Using the DSC gas mass fractions calculated above:
ave. = 0.8590
- 0.5726 + 0.0262
- 0.4274 = 0.5031 lbm/ft3 Solving for G C : GC = ()/() = 0.5625 GC = 0.5625 * (1.6667
- 32.174
- 60.7
- 0.5031
- 144)0.5 GC = 273.2 lbm/ft2 -sec Calculating volumetric flow rate:
VC = 273.2 lbm/ft2 -sec
- 1.0229E-07 ft 2 * (1 / 0.5031 lbm/ft3)
- 60 sec/min = 3.333E-03 CFM Converting to cc/sec:
3.333E-03 ft3/min
- 1 min/60 sec
- 28316.85 cc/ft 3 = 1.573 cc/sec.
Note that this leakage rate, which will be used in the dose analysis, is less than the ANSI N14.5
-1997 accident allowable leakage rate of 1.926 cc/sec calculated above but is representative of the conservative accident conditions assumed.
This assumed leakage rate is very conservative as compared to DSC Leak Accident leak rates postulated in Reference 2.21, 1.3E-05 cc/sec and Reference 2.32, 1.58E-05 cc/sec.
In addition to the DSC leak rate calculated above, three additional leak rates are assumed based on hole sizes of 4.0E-03 cm 2, 2.5E-2 cm2, and 1.0E-01 cm
- 2. Dose calculation results are documented in Section 7.
CALCULATION NO. MNGP-018 PAGE 36 of 38 REV 0 The activity contained within the fuel rods of the subject DSC following 15.53 years of decay is tabulated in Table 6.1.
For a postulated accidental release from an affected DSC having noncompliant dye penetrant examinations (PTs)
, offsite radiological dose consequence analyses were performed. The results of these dose consequence analyses are tabulated below for each organ. Table 7.1 is based on RG 1.145 calculated . Table 7.1 -
Organ Dose Gonad Breast Lung Red Marrow Bone Surface Thyroid Remainder Effective Skin Dose (Rem) 4.73E-03 2.28E-06 2.99E-02 2.58E-02 3.23E-01 1.14E-04 1.19E-02 2.11E-02 1.48E-11 Dose Limit* (Rem) 50 50 50 50 50 50 50 5 50
Per NUREG-1567 page 9-14 (Reference 2.12), the Lens Dose equals the sum of the Skin and Effective dose not to exceed 15 Rem. From Table 7.1 above the calculated Lens Dose = 2.11E-02.
As noted in Section 6.4 addition al hole sizes were considered. The thyroid and effective dose for these cases is summarized below along with the base hole size for comparison, Table 7.1a.
Table 7.1a - - Additional Hole Sizes RG 1.145 /Q Values Leakage Thyroid Effective Diameter Area Area Rate Dose Dose (cm) (cm2) (ft2) (CFM) (Rem) (Rem TEDE) 1.10E-02 9.5033E-05 1.0229E-07 3.333E-03 1.1366E-04 2.1128E-02 4.0000E-03 4.3056E-06 1.403E-01 4.7750E-03 7.0204E-01 2.5000E-02 2.6910E-05 8.768E-01 2.9533E-02 4.2875E+00 1.0000E-01 1.0764E-04 3.507E+00 1.1392E-01 1.6534E+01
CALCULATION NO. MNGP-018 PAGE 37 of 38 REV 0 Calculation and Conclusion Conservatisms:
- 1. Table 7.2 is based on RG 1.145 reduced by the 4.5 factor calculated in Section 6.3. The Table 7.2 data is more realistic for the proposed comparison.
Table 7.2 - Organ Dose with Realistic Dispersion Factor Data Gonad Breast Lung Red Marrow Bone Surface Thyroid Remainder Effective Skin Dose (Rem) 1.05E-03 5.08E-07 6.64E-03 5.74E-03 7.17E-02 2.53E-05 2.65E-03 4.70E-03 3.28E-12 Dose Limit* (Rem) 50 50 50 50 50 50 50 5 50
Per NUREG-1567 page 9-14 (Reference 2.12), the Lens Dose equals the sum of the Skin and Effective dose not to exceed 15 Rem.
From Table 7.2 above the calculated Lens Dose = 4.70E-03.
As noted in Section 6.4 addition hole sizes were considered. The thyroid and effective dose for these cases is summarized below along with the base hole size for comparison, Table 7.2
- a. Table 7.2a - Organ Dose with Realistic Dispersion Factor Data
- Additional Hole Sizes RG 1.145 /Q Values / 4.5 Leakage Thyroid Effective Diameter Area Area Rate Dose Dose (cm) (cm2) (ft2) (CFM) (Rem) (Rem TEDE) 1.10E-02 9.5033E-05 1.0229E-07 3.333E-03 2.5258E-05 4.6951E-03 4.0000E-03 4.3056E-06 1.403E-01 1.0611E-03 1.5601E-01 2.5000E-02 2.6910E-05 8.768E-01 6.5629E-03 9.5278E-01 1.0000E-01 1.0764E-04 3.507E+00 2.5316E-02 3.6742E+00
- 2. This postulated accident assumes an Occupancy Factor of 1.0 (individual always present at EAB). Tables 7.1 and 7.2 dose values would reduce linearly with any change in assumed occupancy factor.
CALCULATION NO. MNGP-018 PAGE 38 of 38 REV 0 3. This postulated accident is a static event (i.e., not the result of a dynamic event such as a DSC drop or a fire) although 100% of the fuel rods are assumed damaged. Crediting less fuel damage would be a reasonable basis for analysis. Caution would need to be observed as any reduction in failed fuel would reduce the Tables 7.1 and 7.2 dose results but the reduction amount woo term would be altered by a less than 100% failed fuel assumption, but not linearly. Thus the G C and VC terms also would not reduce linearly with the failed fuel percentage.
- 4. The -1), was conservatively calculated based on a fall height of 5.8 inch (fuel channel height). Based on the methodology of Reference 2.21, the characteristic height (3.35 36 inch) of the fuel channel is apprless conservative natural deposition removal coefficient value of 34.65 Hr-1. 5. This postulated accident assumes a volumetric critical flux release for the entire 30 day duration of the postulated accident. This analysis approach is surely conservative since at some point in the 30 da y accident time frame the DSC pressure may be reduced below the critical pressure and maximum critical flux flow would no longer occur. The degree of conservatism of assuming constant leakage at the critical flow rate may be assessed by determining the time at which the DSC would reach 14.7 psia and review of the RADTRAD results to determine incremental dose accrual after the calculated time for the DSC to reach 14.7 psia (see Reference 2.27, Eqn 2.78). See AAC DVCS, Attachment 1 4, Comment No. 3.
- 6. In the Reference 2.3 default "INP" file the Cs-137 submersion DCFs are noted as including the impact of the Ba-137m daughter. Review of FGR 12 for Ba-137m submersion DCFs shows greater DCF values for Ba-137m than the combined value given in Reference 2.3 under the Cs-137 isotope. Conservatively, the Reference 2.3 DCFs for Cs-137 and the FGR
12 DCFs for Ba-137m are used.
158158 West Gay Street l Suite 400 West Chester, PA 19380 USA jensenhughes.com O: +1 610-431-82608260 RISK ASSESSMENT OF MNGP DSC 11-15 WELDS USING NUREG-1864 METHODOLOGY RESULTS TO DSC 11-15 Prepared For Xcel Energy 414 Nicollet Mall, 414-7 Minneapolis, MN 55401 Revision: 0 Project #: 1RCA16045.000 Project Name: Risk Assessment of MNGP DSC 11-15 Welds Using NUREG-1864 Methodology Report #: 016045-RPT-01 016045-RPT-01 Approval Summary Revision 0 Page ii Risk Assessment of MNGP DSC 11-15 Welds Using NUREG-1864 Methodology
Project No. 1RCA16045.000
Report #: 1RCA16045.000-001
Revision No.
0
Preparer: Date: 8-29-2017 Richard Anoba Preparer:
Date:8-29-2017 Matt Johnson Reviewer:
Date: 8-29-2017 Grant Teagarden Reviewer:
Date: 8-29-2017 Vincent Andersen Approver: Date: 8-29-2017 Richard Anoba Digitally signed by Matt Johnson Date: 2017.08.29 10:08:24-05'00' Richard Anoba Digitally signed by Richard Anoba
DN: C=US, E=ranoba@jensenhughes.com, O=Jensen Hughes, OU=Risk
Informed Engineering, CN=Richard Anoba
Location: Raleigh, NC
Reason: I am the author of this document
Contact Info: 650-678-6784
Date: 2017.08.29 14:21:12-04'00' Richard Anoba Digitally signed by Richard Anoba
DN: C=US, E=ranoba@jensenhughes.com, O=Jensen Hughes, OU=Risk Informed Engineering, CN=Richard Anoba
Location: Raleigh, NC
Reason: I am approving this document
Contact Info: 650-678-6784
Date: 2017.08.29 14:21:45-04'00' Grant Teagarden Digitally signed by Grant Teagarden Date: 2017.08.29 11:39:17 -07'00' 016045-RPT-01 Revision Record Summary Revision 0 Page iii REVISION RECORD SUMMARY Revision Revision Summary 0 Initial Issue
016045-RPT-01 Table of Contents Revision 0 Page iv TABLE OF CONTENTS REVISION RECORD SUMMARY ................................................................................................ iii
1.0INTRODUCTION
............................................................................................................... 11.1Purpose ................................................................................................................. 11.2Background ........................................................................................................... 11.3Scope .................................................................................................................... 12.0DEVELOPMENT OF METHODOLOGY TO APPLY/EXTRAPOLATE NUREG-1864 RESULTS TO DSC 11-15 PROCESS .............................................................................. 22.1Overview of Methodology ...................................................................................... 22.2Initiating Events ..................................................................................................... 22.2.1Flood .......................................................................................................... 32.2.2Tsunamis ................................................................................................... 42.2.3Volcanic Activity ......................................................................................... 42.2.4Intense Precipitation .................................................................................. 42.2.5Storage Tanks, Transformers, Barges, Trucks, Railcars, and Nearby Industrial Facilities ..................................................................................... 42.2.6Dropped Fuel Assembly ............................................................................ 42.2.7Dropped Transfer Cask ............................................................................. 42.2.8Seismic ...................................................................................................... 52.2.9High Winds ................................................................................................ 52.2.10Meteorites .................................................................................................. 62.2.11Lightning Strikes ........................................................................................ 62.2.12Aircraft ....................................................................................................... 62.2.13Blocked Vent.............................................................................................. 62.3Multipurpose Canister (MPC) Failure Model ......................................................... 62.4Fuel Assembly Failure Model .............................................................................. 112.5Secondary Containment Isolation Model ............................................................. 112.6Consequence ...................................................................................................... 142.7Risk Quantif ication and Results...........................................................................
153.0APPLICATION OF METHODOLOGY TO APPLY/EXTRAPOLATE NUREG-1864 RESULTS TO DSC 11-15 ............................................................................................... 243.1Overview of Application ....................................................................................... 243.2Initiating Events ................................................................................................... 243.2.1Flood ........................................................................................................ 243.2.2Tsunamis ................................................................................................. 24 016045-RPT-01 Table of Contents Revision 0 Page v 3.2.3Volcanic Activity ....................................................................................... 243.2.4Intense Precipitation ................................................................................ 243.2.5Storage Tanks, Transformers, Barges, Trucks, Railcars, and Nearby Industrial Facilities ................................................................................... 253.2.6Dropped Fuel Assembly .......................................................................... 263.2.7Dropped Transfer Cask ........................................................................... 263.2.8Seismic .................................................................................................... 263.2.9High Winds .............................................................................................. 273.2.10Meteorites ................................................................................................ 273.2.11Lightning Strikes ...................................................................................... 273.2.12Aircraft Accidents ..................................................................................... 273.3Dry Storage Canister (DSC) Failure Model ......................................................... 303.3.1Mechanical Failures ................................................................................. 303.3.2Thermal Failures ...................................................................................... 313.4Fuel Assembly Failure Model .............................................................................. 323.5Secondary Containment Isolation Model ............................................................. 323.6Consequence Model ........................................................................................... 333.6.1Fuel Type and Exposure
.......................................................................... 333.6.2Radionuclide Inventory ............................................................................ 333.6.3Source Term ............................................................................................ 343.6.4Initial Plume Dimensions ......................................................................... 353.6.5Plume Heat Content ................................................................................ 353.6.6Population ................................................................................................ 353.6.7Site Weather ............................................................................................ 383.7Results ................................................................................................................ 404.0SUMMARY ......................................................................................................................
454.1Summary of Methodology Development ............................................................. 454.2Summary of Methodology Application ................................................................. 45Conclusions ......................................................................................................... 454.3 4
55.0REFERENCES
................................................................................................................ 47 016045-RPT-01 Introduction Revision 0 Page 1
1.0 INTRODUCTION
1.1 Purpose This report documents the risk assessment of non-compliant weld inspections of five (5) spent fuel dry storage casks (DSCs 11 thru 15) at the Monticello Nuclear Generating Plant (MNGP) plant. This risk assessment employs the approaches used by the NRC in NUREG-1864 [1]. The purpose of this analysis is to compare the calculated risk of the alternative of leaving these casks as-is in their current stored location versus the alternative of transferring these casks back into the reactor building for inspection and then returning them to their storage location. 1.2 Background Xcel Energy is planning to submit an Exemption Request for five (5) NUHOMS Dry Shielded Canisters (DSCs) that were placed in service at the MNGP Independent Spent Fuel Storage Installation (ISFSI) with non-compliant dye penetrant examinations (PTs). Xcel Energy requested JENSEN HUGHES to develop a methodology to apply/extrapolate NUREG-1864 [1] results to MNGP DSC 11-15 and to apply the methodology to arrive at a comparative evaluation of the risks. The desired outcome is a report that provides an evaluation of the risks similar to Table 19 of NUREG-1864 [1] for the applicable initiating events and conditions facilitate decision-making regarding the following two proposed plan alternatives: Alternative 1: DSC 11-15 continued storage as-is Alternative 2: DSC 11-15 transfer, exam, return, and continued storage.
1.3 Scope The objective of this analysis is to develop and apply a methodology to compare the risk of moving the subject DSCs with non-compliant PT examinations (for the purpose of conducting further non-destructive examination) to the risk of leaving the subject DSCs in service for twenty years with non-compliant PT examinations. The scope of this work does not include any structural or radiological analyses, nor development of a PRA model. This report documents the analysis according to the following two major tasks: Task 1 - Develop a methodology to apply/extrapolate NUREG-1864 [1] results (Section 2 of Report) Task 2 - Apply methodology to apply/extrapolate NUREG-1864 [1] results (Section 3 of Report) Cask DSC 16 has a similar non-conforming weld inspection; DSC 16 has been previously exempted from certain 10 CFR 72 regulations [24] and is not part of this risk analysis.
016045-RPT-01 Development of Methodology to Apply/Extrapolate NUREG-1864 Results to DSC 11-15 Process Revision 0 Page 2 2.0 DEVELOPMENT OF METHODOLOGY TO APPLY/EXTRAPOLATE NUREG-1864 RESULTS TO DSC 11-15 PROCESS 2.1 Overview of Methodology The approach and methodologies in NUREG-1864 [1] are used as the basis for this risk assessment. NUREG-1864 [1] documents the NRC risk assessment of a spent fuel dry cask storage system at a U.S. boiling water reactor (BWR) site. The NUREG-1864 [1] study is for the Holtec International HI-STORM 100 cask system and covers the onsite handling, transfer, and storage phases of the cask life cycle. The analysis covers a broad spectrum of postulated initiating events and hazards (e.g., drop scenarios, external hazards) and calculates the risk associated with the postulated initiating events JENSEN HUGHES has developed a methodology to apply/extrapolate NUREG-1864 [1] methodologies and results by leveraging the configuration similarities between the MNGP model and the NUREG-1864 [1] model AND applying MNGP site specific (fuel loaded, flood hazard, seismic risk, etc.) and technology specific configuration (horizontal storage versus vertical) inputs where applicable. The overall goal is to compare the risks of the two alternatives using the plant-specific adaptation the NUREG-1864 methodology. The methodologies of this risk assessment are discussed in this section and Section 3 discusses the analysis. The NUREG-1864 [1] analysis is presented according to the following main analysis areas: Initiating event identification and frequencies Release from MPC Release from secondary containment Consequence assessment As such, the methodology used in this risk assessment is presented below according to the above analysis areas. In addition, the following analysis topics are discussed in this section: Risk quantification and results Risk acceptance criteria 2.2 Initiating Events Section 3 of NUREG-1864 [1] documents the initiating events analysis for the pilot PRA study. The analysis includes initiating event identification, screening, and quantification. Table 1 provides a summary of initiating events addressed in NUREG-1864 [1]. Table 1 also identifies MNGP data sources that contain plant-specific information of interest.
016045-RPT-01 Development of Methodology to Apply/Extrapolate NUREG-1864 Results to DSC 11-15 Process Revision 0 Page 3 Table 1: Summary of Initiating Events Addressed in NUREG-1864 Initiating Events Screened in NUREG-1864 Section of NUREG-1864 Addressed Monticello Data Sources Floods x 3.2.1 IPEEE, ISAR, and JENSEN HUGHES Report 1SML16012.000-1, Table 4-1. Tsunamis x 3.2.2 IPEEE, ISAR and any subsequent analysis Volcanic Activity x 3.2.4 IPEEE, ISAR and any subsequent analysis Intense Precipitation x 3.2.5 IPEEE, ISAR Storage Tanks, Transformers, Barges, Trucks, Railcars, and Nearby Industrial Facilities x 3.2.6 IPEEE, ISAR and any subsequent analysis Dropped Fuel Assembly 3.3.1 ISAR and Monticello-specific information Dropped Transfer Cask 3.3.2 ISAR and Monticello-specific information Seismic Events 3.3.3 ISAR ML14136A289, S&A Report 14C4229-RPT-001 Rev. 3 High Winds 3.3.4 ISAR NUREG/CR-4461, Rev 2, Table 6-1 NEI 17-02, Section 6 Meteorites 3.3.5 IPEEE, ISAR and any subsequent analysis Lightning Strikes 3.3.6 IPEEE, ISAR and any subsequent analysis Aircraft 3.3.7 IPEEE, ISAR and any subsequent analysis This risk assessment re-assesses the NUREG-1864 [1] initiating event screening process and frequency estimation to consider MNGP specific attributes.
2.2.1 Flood Section C.2.3.1 of the MNGP IPEEE [5], the MNGP ISAR [7] are used to re-assess whether to confirm that external floods can be screened out from further consideration due to non-significant or no impact on the MNGP ISFSI and associated DCS/HSM modules. If this hazard initiator cannot be screened out, then report 1SML16012.000-1 or other current analyses will be used as the primary source of the external flooding hazard and the initiator carried forward into the analysis to consider the impact of the hazard on the potential for radionuclide release from
the casks.
016045-RPT-01 Development of Methodology to Apply/Extrapolate NUREG-1864 Results to DSC 11-15 Process Revision 0 Page 4 2.2.2 Tsunamis Section 3.2.2 of NUREG-1864 [1] addresses the impact of tsunamis on the ISFSI of the reference site. Tsunamis were screened out on the basis that the reference site is far enough inland that it will not be affected by a tsunamis. Section C.2.4 of the MNGP IPEEE [5] and the ISAR [7] are used to confirm that tsunamis can be screened out from further consideration due to no impact on the MNGP ISFSI and associated DCS/HSM modules, as well as the TC/DSC while in transfer or RX Building. If the MNGP documents do not address tsunamis, then perform a qualitative assessment similar to NUREG-1864 [1]. 2.2.3 Volcanic Activity Section 3.2.3 of NUREG-1864 [1] addresses the impact of volcanic activity on the ISFSI of the reference site. Volcanic activity was screened out on the basis that the reference site is far from volcanic regions and well out of the influence of volcanic activity. Section C.2.4 of the MNGP IPEEE [5] and the ISAR [7] are used to confirm that volcanic activity can be screened out from further consideration due to non-significant or no impact on the MNGP ISFSI and associated DCS/HSM modules. 2.2.4 Intense Precipitation Section 3.2.4 of NUREG-1864 [1] addresses the impact of intense precipitation on the ISFSI of the reference site. intense precipitation was screened out on the basis that the ISFSI of reference site is designed so that graded land and drains conduct water away from the storage pads. Section C.2.3.2 of the MNGP IPEEE [5] and the ISAR [7] are used to confirm that intense precipitation can be screened out from further consideration due to non-significant or no impact on the MNGP ISFSI and associated DCS/HSM modules. 2.2.5 Storage Tanks, Transformers, Barges, Trucks, Railcars, and Nearby Industrial Facilities Section 3.2.5 of NUREG-1864 [1] addresses the impact of storage tanks, transformers, barges, trucks, railcars, and nearby industrial facilities on the ISFSI of the reference site. Storage tanks, transformers, barges, trucks, railcars, and nearby industrial facilities were screened out based on proximity of these hazard sources to the ISFSI of the reference site. Section C.2.4 of the MNGP IPEEE [5], the ISAR [7], and the 72.212-A [26] are used to re-assess whether to confirm that storage tanks, transformers, barges, trucks, railcars, and nearby industrial facilities can be screened out from further consideration due to non-significant or no impact on the MNGP ISFSI and associated DCS/HSM modules. If one or more of these hazard initiators cannot be screened out, then a reasonable estimate of the hazard frequencies will be determined from the MNGP IPEEE and/or from industry studies and the initiator(s) carried forward into the analysis to consider the impact of the hazard on the potential for radionuclide release from the casks. 2.2.6 Dropped Fuel Assembly Section 3.3.1 of NUREG-1864 [1] provides the basis for calculating frequency of a dropped fuel assembly during transfer operations to the ISFSI of the reference site. The MNGP alternatives considered in this report do not involve individual fuel assembly handling and dropped fuel assembly events can be screened out. 2.2.7 Dropped Transfer Cask Section 3.3.2 of NUREG-1864 [1] provides the basis for calculating frequency of a dropped transfer cask during transfer operations to the ISFSI of the reference site. The drop rate used in NUREG-1864 [1] is taken from NUREG-1774 [20] and known to be potentially conservative and 016045-RPT-01 Development of Methodology to Apply/Extrapolate NUREG-1864 Results to DSC 11-15 Process Revision 0 Page 5 based on events that do not necessarily apply to the reference plant or to the MNGP plant configuration and procedures. The ISAR [7] and other plant-specific documents are used to assess Monticello specific aspects and to adjust the cask drop failure rate (NUREG-1774 [20] data is used in this re-assessment), haul path, and to confirm that the NUREG-1864 [1] frequency estimate can be applied to MNGP. 2.2.8 Seismic Section 3.3.3 of NUREG-1864 [1] addresses the impact of seismic events on the ISFSI of the reference site. The seismic hazard curve used in NUREG-1864 [1] is provided in Figure 8 of the report and obtained from NUREG-1488 [21]. The seismic frequency used for risk quantification was based on the minimum ground acceleration of 1.35g Peak Ground Acceleration (PGA) necessary to cause storage cask tipping of a HI-STORM 100 cask on the ISFSI. Minimum required ground acceleration to result in cask tipping was calculated as 9 times the design basis earthquake ground acceleration of 0.15g PGA. The seismic initiating event is retained in the risk assessment; however, NUREG-1488 based seismic hazard curves for U.S. nuclear power plant sites have been deemed obsolete for risk-informed interactions with the NRC following Generic Issue 199 [22]. As such, this risk assessment should use the latest available MNGP seismic hazard curve produced using methodologies accepted by the NRC. The seismic hazard curves for MNGP are provided in Figure 2.3.7-1 of the S&A Seismic hazard screening report [3]. The PGA hazard curve could be used in this assessment. Due to the differences in design, the NUHOMs canisters stored at MNGP in the horizontal position are not subject to tip over. Consequently, a plant-specific seismic-induced damage mechanism (if plausible) will have to be applied to determine the
seismic hazard frequency. 2.2.9 High Winds Section 3.3.4 of NUREG-1864 [1] addresses the impact of high wind events on the ISFSI of the reference site. The tornado hazard curve used in NUREG-1864 [1] is based on data provided in Table 8 and Figure 9 of the report. The tornado hazard was screened out based on extremely low frequencies associated with wind speed required to cause a storage cask to slide on the storage pad (400 mph), to cause storage cask to tip over on storage pad (600 mph), and to propel a heavy object onto a storage cask to cause damage (900 mph). The high wind assessed in NUREG-1864 [1] is a tornado wind. This wind hazard remains the applicable significant wind for the MNGP site (e.g., hurricanes do not apply to the MNGP site).
The tornado hazard curves for MNGP can be generated using plant-specific data from Table 6-1 of NUREG/CR-4461 [2]. The Fujita Scale data (annual exceedance frequency versus wind speed) is entered into MS Excel and used as the input to develop the tornado hazard curve. The trending function in MS Excel is used to generate the tornado hazard curve for MNGP. Section 6.29 of the ISAR [7] indicates that the design basis tornado is 360 mph. A preliminary review of the ISAR [7] indicates that cask tipping can be screened out for the MNGP configuration. If this is confirmed, then review the plant-specific documentation to identify other tornado induced failure mechanisms that can be used to estimate the tornado initiating event frequency. Review the ISAR [7] and other plant-specific documents to determine the wind speed required to cause a storage cask to slide on storage pad, to cause storage cask damage on storage pad, and to propel a heavy object onto a storage cask to cause damage. Use the tornado hazard curve for MNGP to determine the tornado frequency that would be used for quantification. Compare the NUREG-1864 [1] value against the MNGP value to determine if the NUREG-1864 [1] value is bounding. If NUREG-1864 [1] is not bounding, then use plant-specific information to estimate the impact. Due to the differences in design, the NUHOMs canisters stored at MNGP in the horizontal position are not subject to tip over. Consequently, a plant-specific tornado-induced 016045-RPT-01 Development of Methodology to Apply/Extrapolate NUREG-1864 Results to DSC 11-15 Process Revision 0 Page 6 damage mechanism (if plausible) will have to be applied to determine the tornado hazard frequency. 2.2.10 Meteorites Section 3.3.4 of NUREG-1864 [1] addresses the impact of meteorites on the ISFSI of the reference site. This initiating event is retained in this risk assessment and the NUREG-1864 [1]
frequency estimate for meteorite strike per area is used for MNGP. 2.2.11 Lightning Strikes Section 3.3.6 of NUREG-1864 [1] addresses the impact of lightning strikes on the ISFSI of the reference site. Lightning strike induced radionuclide release accidents are determined in NUREG-1864 [1] to be non-credible. NUREG-1864 [1] information, the MNGP IPEEE [5], ISAR [7] and other plant-specific documents are used in this risk assessment to document that lightning strike induced accidents are non-credible radionuclide release accidents for the MNGP NUHOMS dry cask storage system.
2.2.12 Aircraft Section 3.3.7 of NUREG-1864 [1] addresses the impact of aircraft impact on the ISFSI of the reference site. This hazard is maintained in this risk assessment. The MNGP IPEEE [5], ISAR
[7] and other plant-specific documents are used to re-assess the aircraft impact frequency estimate for MNGP. 2.2.13 Blocked Vent Blocked vents are evaluated for the MPC failure model due to thermal events. The DSC failure model thermal event discussion is contained in Section 3.3.2. 2.3 Multipurpose Canister (MPC) Failure Model Section 4.3.2 of NUREG-1864 [1] addresses the probabilities of MPC failures for the reference site. Table 12 of NUREG-1864 [1] provides a summary of MPC probabilities for the reference site. The MPC failure probabilities are a function of mechanical impact load due to various event scenarios. Many of the event scenarios involve load drops at various heights. The drop heights on Table 12 of NUREG-1864 [1] were derived from Table 1 of NUREG-1864 [1], which defines the stages of dry cask operations for the reference site. Table 1 of NUREG-1864 [1], seems to indicate that the MPC failure probability is a strong function of drop height. Table 2 of the report provides a comparison of the reference site dry cask operations on Table 1 of NUREG-1864 [1] against the MNGP dry cask operations. Information in Table 2 [34] of this report and other MNGP documentation are used to assess plant-specific DSC failure probabilities given an initiating event challenge. The DSC provides the equivalent function of the MPC evaluated in NUREG-1864 [1]. Certain accident scenarios (e.g., meteorite strike) in NUREG-1864 [1] use a conditional failure probability of 1.0 for the MPC; this same approach is used in this risk assessment for these scenarios.
016045-RPT-01 Development of Methodology to Apply/Extrapolate NUREG-1864 Results to DSC 11-15 Process Revision 0 Page 7 Table 2: Stages of Dry Cask Storage Operation - NUREG-1864 Compared to MNGP Alternatives NUREG 1864: Table 1. Stages of the Dry Cask Storage Operation Monticello Operations to Consider for DSC 11-15 Stages Height (A) MNGP Step (Reference Lesson Plan M-9014L-058 for Steps) MNGP 9500-Series Procedure & Part(s) Listed with Description Height m ft 1 Loading fuel assemblies into the MPC (B) 4.8 16 6 N/A, DSC 11-15 already loaded, sealed and placed in HSM. 9505 Rev 12, Part I 16.75' Ref. 36 2 Placing the MPC lid onto the MPC and engaging the lift yoke on the transfer cask (C) 0 0 7 N/A, DSC 11-15 already loaded, sealed and placed in HSM. 9506 Rev 17, Parts B & C 0 3 Lifting the transfer cask out of the cask pit 13 42.5 8 N/A, DSC 11-15 already loaded, sealed and placed in HSM. 9506 Rev 17, Part D39.5' Ref. 37 4 Moving the transfer cask over a railing of the spent fuel pool 0.9 3 8 N/A, DSC 11-15 already loaded, sealed and placed in HSM. 9506 Rev 17, Part D 5 Moving the transfer cask to the preparation area (1 st segment) 0.3 1 8 N/A, DSC 11-15 already loaded, sealed and placed in HSM. 9506 Rev 17, Part D 6.5"-8.5" Ref. 38 6 Moving the transfer cask to the preparation area (2 nd segment) 0.3 1 8 N/A, DSC 11-15 already loaded, sealed and placed in HSM. 9506 Rev 17, Part D 6.5"-8.5" Ref. 38 7 Moving the transfer cask to the preparation area (3 rd segment) 0.3 1 8 N/A, DSC 11-15 already loaded, sealed and placed in HSM. 9506 Rev 17, Part D 6.5"-8.5" Ref. 38 8 Lowering the transfer cask onto the preparation area (D) 0.3 1 8 N/A, DSC 11-15 already loaded, sealed and placed in HSM. 9506 Rev 17, Part D 6.5"-8.5" Ref. 38 9 Preparing (draining, drying, inerting, and sealing) the MPC for storage 0 0 9-19 N/A, DSC 11-15 already loaded, sealed and placed in HSM. 9506 Rev 17, Parts F through Q 0 The following sequences represent PAUT of a DSC, followed by moving the DSC and placing in the HSM storage module Perform PAUT (Phased Array Ultrasonic Test) of DSC while in the TC on the refueling floor.
0 016045-RPT-01 Development of Methodology to Apply/Extrapolate NUREG-1864 Results to DSC 11-15 Process Revision 0 Page 8 Table 2: Stages of Dry Cask Storage Operation - NUREG-1864 Compared to MNGP Alternatives NUREG 1864: Table 1. Stages of the Dry Cask Storage Operation Monticello Operations to Consider for DSC 11-15 Stages Height (A) MNGP Step (Reference Lesson Plan M-9014L-058 for Steps) MNGP 9500-Series Procedure & Part(s) Listed with Description Height m ft 20 Install the TC lid 9507 Rev 19, Part B 10 Installing the short stays and attaching the lift yoke (D) 0 0 21 Similar: Attaching the lift yoke (D) 9507 Rev 19, Part C 0 11 Lifting the transfer cask 0.6 2 21 Same 9507 Rev 19, Part C 6.5" - 8.5" Ref. 38 12 Moving the transfer cask to exchange bottom lids of the transfer cask (1 st segment) 0.6 2 N/A for NUHOMS system 13 Moving the transfer cask to exchange bottom lids of the transfer cask (2 nd segment) 0.6 2 N/A for NUHOMS system 14 Replacing the pool lid with the transfer lid 0.1 0.25 N/A for NUHOMS system 0 15 Moving the transfer cask near the equipment hatch 0.6 2 21 Same 9507 Rev 19, Part C 6.5" - 8.5" Ref. 38 16 Holding the transfer cask 0.6 2 21 Same 9507 Rev 19, Part C 6.5" - 8.5" Ref. 38 17 Moving the transfer cask to the equipment hatch 0.6 2 21 Same 9507 Rev 19, Part C 6.5" - 8.5" Ref. 38 18 Lowering the transfer cask to the over-pack through the equipment hatch 24.4 80 22-23 Similar: Lowering the transfer cask to the transfer trailer (TT) through the equipment hatch. Based on 1027.67 + 0.71 carry height minus 935 (assuming TT is missing). Height to the TT trunnion is about 8 fewer feet - 85.4' 9507 Rev 19, Part C93.4' Ref. 39 19 Preparing (remove short stays, disengage lift yoke, attach long stays) to lower the MPC 0 0 N/A for NUHOMS system 20 Lifting the MPC and opening doors of transfer lid 5.8 19 N/A for NUHOMS system 21 Lowering the MPC through the transfer cask into the storage cask 5.8 19 N/A for NUHOMS system
016045-RPT-01 Development of Methodology to Apply/Extrapolate NUREG-1864 Results to DSC 11-15 Process Revision 0 Page 9 Table 2: Stages of Dry Cask Storage Operation - NUREG-1864 Compared to MNGP Alternatives NUREG 1864: Table 1. Stages of the Dry Cask Storage Operation Monticello Operations to Consider for DSC 11-15 Stages Height (A) MNGP Step (Reference Lesson Plan M-9014L-058 for Steps) MNGP 9500-Series Procedure & Part(s) Listed with Description Height m ft 22-23 Down-end onto the TT and disengage the lift yoke [Once the bottom trunnion is seated a long TC would rotate and drop 148.5 inches until the top trunnion is seated. 138.5" for the short TC.]
9507 Rev 19, Part C 148.5" Ref. 40 22 Moving the storage cask into the airlock on Helman rollers 0 0 25 Similar: Moving the storage cask into the airlock on TT (Based on max 43" trailer deck height per NUH-07-0218) 9510 Rev 13, Part A 64.5" Ref. 41 23 Moving the storage cask out of the airlock on Helman rollers 0 0 25 Similar: Moving the storage cask out of the airlock on TT 9507 Rev 19, Part D 64.5" Ref. 41 24 Moving the storage cask away from the secondary containment on Heiman rollers 0 0 25 Similar: Moving the storage cask away from the secondary containment on TT 9507 Rev 19, Part D 64.5" Ref. 41 25 Preparing (installing lid, vent shield cross-plates, vent screens) the storage cask for storage 0 0 N/A for NUHOMS system 26 Lifting the storage cask above the Heiman rollers with the over-pack transporter 0.1 0.25 N/A for NUHOMS system 27 Moving the storage cask above a cushion on the preparation area <0.1 <0.25 N/A for NUHOMS system 28 Holding the storage cask above the cushion while attaching a Kevlar belt <0.1 <0.25 N/A for NUHOMS system 29 Moving the storage cask above the concrete surface of the preparation area 0.3 1 25 Same (Based on max 43" trailer deck height per NUH-07-0218) 9508 Rev 17, Part D 64.5" Ref. 41 30 Moving the storage cask above the asphalt road 0.3 1 25 Same 9507 Rev 19, Part D 64.5" Ref. 41 31 Moving the storage cask above the gravel surface around the storage pads 0.3 1 25 Same 9507 Rev 19, Part D 64.5" Ref. 41 016045-RPT-01 Development of Methodology to Apply/Extrapolate NUREG-1864 Results to DSC 11-15 Process Revision 0 Page 10 Table 2: Stages of Dry Cask Storage Operation - NUREG-1864 Compared to MNGP Alternatives NUREG 1864: Table 1. Stages of the Dry Cask Storage Operation Monticello Operations to Consider for DSC 11-15 Stages Height (A) MNGP Step (Reference Lesson Plan M-9014L-058 for Steps) MNGP 9500-Series Procedure & Part(s) Listed with Description Height m ft 32 Moving the storage cask above the concrete storage pad 0.3 1 25 Same 9508 Rev 17, Part D 64.5" Ref. 41 33 Lowering the storage cask onto the storage pad 0.3 1 N/A for NUHOMS system 70.4" Ref. 42 26-34 Remove TC lid, align to HSM, grapple DSC and insert
into HSM and install the door 9508 Rev 17, Parts D through K 70.4" Ref. 42 The sequence above would be performed in reverse to remove DSC 11-15 from the HSM and return to the refueling floor for PAUT. 34 Storing the storage cask on the storage pad for 20 years 0 0 Same 0 (A) Height is the distance the cask would fall if the support system failed.
(B) Prior to Stage 1, the MPC is inserted into the transfer cask, and after other preparations, lowered into the cask pit. The storage over-pack is placed on Heiman rollers and moved under the equipment hatch. (C) The lift yoke attaches to the trunnions of the transfer cask. (D) Stays attach to the lift yoke on one end and cleats of the MPC on the other end. (A) Height is the distance the cask would fall if the support system failed (B) Prior to Stage 1, the DSC is inserted into the transfer cask, and after other preparations, lowered into the cask pit. The transfer trailer is moved under the equipment hatch. (C) The lift yoke attaches to the trunnions of the transfer cask. (D) Lift yoke is disconnected during preparation and re-attached prior to next movement.
016045-RPT-01 Development of Methodology to Apply/Extrapolate NUREG-1864 Results to DSC 11-15 Revision 0 Page 11 2.4 Fuel Assembly Failure Model Section 4.4 of NUREG-1864 [1] addresses the probabilities of fuel and cladding failures due to dynamic loadings (i.e., drop scenarios) for the reference site. As discussed in sections 2.2.6 and 3.2.6 of this report, fuel assembly drops have been screened out based on the defined scope of this PRA. The fuel can fail given a drop of the DSC. In NUREG-1864 [1] fuel failure is included in the overall failure of the MPC given a drop, and this is applied to this evaluation as well. Table 3 shows the probability of release given failure of the fuel and MPC based on the NUREG-1864
[1] template, and for MNGP, Tables 14 and 15 contain the probability of release from the fuel and DSC. 2.5 Secondary Containment Isolation Model Section 5.0 of NUREG-1864 [1] describes the secondary isolation model for the reference site that is applied for accident scenarios initiating within the reactor building. A logic model was developed to quantify the failure probability of the Secondary Isolation System. Figure 17 of NUREG-1864 [1] provides a flow diagram for the Secondary Isolation System for the reference plant. This aspect of the analysis applies to MNGP, as well. This aspect of the analysis is an assessment of the secondary containment isolation system and the configuration of the secondary containment boundary at the time of postulated accidents. MNGP documents and drawings are used to determine the plant-specific differences for the Secondary Containment Isolation System. Section 5.3 of the Monticello USAR [12] provides a description of the Secondary Containment Isolation System. The flow diagrams, provided in References [13]
through [17], provide the details of the Monticello Secondary Isolation System. Simplified flow diagrams for Monticello Secondary Isolation System is provided in Figures 1 through 3.
Compare the NUREG-1864 [1] system against the MNGP system to determine if the NUREG-1864 [1] system failure probabilities are bounding based on system considerations.
016045-RPT-01 Development of Methodology to Apply/Extrapolate NUREG-1864 Results to DSC 11-15 Revision 0 Page 15 reflected in Table 18 of that document. However, the other cases may be applicable to represent the lower expected release associated with releases from a DSC on the pad. Review, select, or if required adjust, the consequence result from Table E.3 of NUREG-1864 [1] that best reflects or bounds the applicable MNGP configurations (e.g., based on stage). 2.7 Risk Quantification and Results This risk assessment does not require building or quantification of PRA event tree and fault tree models. The initiating event frequencies and associated conditional probabilities of releases and consequences are easily multiplied in a spreadsheet to determine the risk results. Structural failure probabilities given a drop from NUREG-1864 [1] are assessed as reasonable for MNGP with scaling factor used for the presence of weld flaws. Offsite consequence results from NUREG-1864 [1] are evaluated as reasonable to represent MNGP specifics; building and quantification of detailed offsite consequence models is not performed as part of this risk assessment. Section 6.0 of NUREG-1864 [1] describes the results of the risk calculations for the reference site. Table 19 of NUREG-1864 [1] provides the summary of risk results for the reference site. The tabular results presentation approach of NUREG-1864 [1] Table 19 is used in this risk assessment for MNGP. Table 3 below provides the template used in this assessment for tabulating the MNGP risk results similar to the NUREG-1864 [1] results. The shaded rows are not applicable to MNGP, based on the comparison on Table 2. The MNGP specific process steps in Table 2 are related to the NUREG-1864 [1] stages. The results information is presented as follows in each of the columns: Stage: This column lists individual stages of cask onsite loading, transportation and storage. This allows traceability to the NUREG-1864 analysis approach as well allows identification of different challenges (i.e., initiating events) at different stages. Initiating Event: This column lists the challenges (i.e., initiating event) by stage that are considered further in this risk assessment. Initiating Event Frequency: This column lists the frequency of occurrence per calendar year of the initiating events. Probability of Release from Fuel Rod and MPC: This column provides the probability of release from the fuel rod and MPC given the initiating event. The values in this column are conditional probabilities given the associated initiating event frequency. Consequences: This column provides the probability of public consequences in terms of latent individual cancer fatalities within 10 miles. The values in this column are conditional probabilities given radionuclide release from the MPC in NUREG-1864 (and for the MNGP DSC for this evaluation) Risk: This column provides the occurrence frequency, in terms of latent individual cancer fatalities within 10 miles, for each of the analyzed initiating event induced radionuclide release scenarios. The values in this column are calculated by multiplying the initiating event frequency, release probability and consequence probability. The results are presented in units of per calendar year.
016045-RPT-01 Development of Methodology to Apply/Extrapolate NUREG-1864 Results to DSC 11-15 Revision 0 Page 16 Table 3: Template for MNGP Risk Results TabulationStages Initiating Event Initiating Event Frequency (Sections 2.2 and 3.2) Probability of Release from fuel rod and MPC (Sections 2.3 and 3.3) Probability of Release from Containment (Sections 2.5 and 3.5) Consequences (Sections 2.6 and 3.6)
RiskNUREG-1864 MNGP Description 1 6 Loading fuel assemblies into the MPC (B) Fuel assembly dropped 2 7 Placing the MPC lid onto the MPC and engaging the lift yoke on the transfer cask (C) 3 8 Lifting the transfer cask out of the cask pit Transfer cask dropped 4 8 Moving the transfer cask over a railing of the spent fuel pool Transfer cask dropped 5 8 Moving the transfer cask to the preparation area (1 st segment) Transfer cask dropped 6 8 Moving the transfer cask to the preparation area (2 nd segment) Transfer cask dropped 7 8 Moving the transfer cask to the preparation area (3 rd segment) Transfer cask dropped 8 8 Lowering the transfer cask onto the preparation area (D) Transfer cask dropped 9 9-19 Preparing (draining, drying, inerting, and sealing) the MPC for
016045-RPT-01 Development of Methodology to Apply/Extrapolate NUREG-1864 Results to DSC 11-15 Revision 0 Page 17 Table 3: Template for MNGP Risk Results TabulationStages Initiating Event Initiating Event Frequency (Sections 2.2 and 3.2) Probability of Release from fuel rod and MPC (Sections 2.3 and 3.3) Probability of Release from Containment (Sections 2.5 and 3.5) Consequences (Sections 2.6 and 3.6)
RiskNUREG-1864 MNGP Descriptionstorage Perform PAUT of DSC while in the TC on the refueling floor. 20 Install the TC Lid 10 21 Installing the short stays and attaching the lift yoke (D) 11 21 Lifting the transfer cask Transfer case dropped 12 Moving the transfer cask to exchange bottom lids of the transfer cask (1 st segment) Transfer cask dropped 13 Moving the transfer cask to exchange bottom lids of the transfer cask (2 nd segment) Transfer cask dropped 14 Replacing the pool lid with the transfer lid Transfer cask dropped 15 21 Moving the transfer cask near the equipment hatch Transfer cask dropped 16 21 Holding the transfer cask Transfer cask dropped 016045-RPT-01 Development of Methodology to Apply/Extrapolate NUREG-1864 Results to DSC 11-15 Revision 0 Page 18 Table 3: Template for MNGP Risk Results TabulationStages Initiating Event Initiating Event Frequency (Sections 2.2 and 3.2) Probability of Release from fuel rod and MPC (Sections 2.3 and 3.3) Probability of Release from Containment (Sections 2.5 and 3.5) Consequences (Sections 2.6 and 3.6)
RiskNUREG-1864 MNGP Description17 21 Moving the transfer cask to the equipment hatch Transfer cask dropped 18 22-23 Lowering the transfer cask to the over-pack through the equipment hatch Transfer cask dropped 19 Preparing (remove short stays, disengage lift yoke, attach long stays) to lower the MPC MPC drop 20 Lifting the MPC and opening doors of transfer lid MPC drop 21 Lowering the MPC through the transfer cask into the storage cask MPC drop 22-23 Down-end onto the TT and disengage the lift yoke [Once the bottom trunnion is seated a long TC would rotate and drop 148.5 inches until the top trunnion is seated. 138.5" for the short TC.] Transfer cask dropped 22 25 Moving the storage cask into the airlock on Helman rollers (TT for MNGP) 23 25 Moving the storage cask out of the 016045-RPT-01 Development of Methodology to Apply/Extrapolate NUREG-1864 Results to DSC 11-15 Revision 0 Page 19 Table 3: Template for MNGP Risk Results TabulationStages Initiating Event Initiating Event Frequency (Sections 2.2 and 3.2) Probability of Release from fuel rod and MPC (Sections 2.3 and 3.3) Probability of Release from Containment (Sections 2.5 and 3.5) Consequences (Sections 2.6 and 3.6)
RiskNUREG-1864 MNGP Descriptionairlock on Helman rollers (TT for MNGP) 24 25 Moving the storage cask away from the secondary containment on Heiman rollers (TT for MNGP) 25 Preparing (installing lid, vent shield cross-plates, vent screens) the storage cask for storage 26 Lifting the storage cask above the Heiman rollers with the over-pack transporter Storage cask dropped 27 Moving the storage cask above a cushion on the preparation area Storage cask dropped 28 Holding the storage cask above the cushion while attaching a Kevlar belt Storage cask dropped 29 25 Moving the storage cask above the concrete surface of the preparation area Storage cask dropped 30 25 Moving the storage cask above the asphalt road Storage cask dropped 016045-RPT-01 Development of Methodology to Apply/Extrapolate NUREG-1864 Results to DSC 11-15 Revision 0 Page 20 Table 3: Template for MNGP Risk Results TabulationStages Initiating Event Initiating Event Frequency (Sections 2.2 and 3.2) Probability of Release from fuel rod and MPC (Sections 2.3 and 3.3) Probability of Release from Containment (Sections 2.5 and 3.5) Consequences (Sections 2.6 and 3.6)
RiskNUREG-1864 MNGP Description31 25 Moving the storage cask above the gravel surface around the storage pads Storage cask dropped 32 25 Moving the storage cask above the concrete storage pad Storage cask dropped 33 Lowering the storage cask onto the storage pad Storage cask dropped 26-34 Remove TC lid, align to HSM, grapple DSC and insert into HSM and install the door 34A Storing the storage cask on the storage pad for 20 years Tipped be Seismic Event 34B Storing the storage cask on the storage pad for 20 years Struck by aircraft 34C Storing the storage cask on the storage pad for 20 years Struck by meteorite 34D Storing the storage cask on the storage pad for 20 years Heated by aircraft fuel 016045-RPT-01 Development of Methodology to Apply/Extrapolate NUREG-1864 Results to DSC 11-15 Revision 0 Page 21 The risk analysis is performed to model the following two proposed alternatives: Alternative 1: DSCs 11-15 Remain As-Is in the HSM Alternative 2: DSCs 11-15 Transferred to RB for Inspection and then back to the HSM Table 4 provides a summary of the risk analysis quantification approach used to apply the NUREG-1864 [1] results to model the MNGP DSCs 11-15 non-compliant weld inspection
issues. Table 4: MNGP Dry Cask Risk Assessment Approach Topic MNGP Alternative 1 Leave DSCs 11-15 As-Is in the HSM MNGP Alternative 2 Transfer DSCs 11-15 to RB for Inspection and then Transfer Back to the HSM PRA Case Adjustments to base case risk scenarios are made to reflect that 5 casks on the pad have weld inspection issues Adjustments to base case risk scenarios are made to reflect that the 5 casks on the pad with weld inspection issues are transported back into RB for inspection, inspected and then transported back out to pad Delta and Absolute Risk From the above two cases, the absolute risk of the two alternatives as well as the delta risk between the alternatives is determined. The case for each alternative incorporates the non-compliant weld inspections for DSCs 11-15, as well as the additional postulated accident scenarios (i.e., additional drop scenarios) for Alternative 2 given the transfer of the casks back into the reactor building for inspection). The modeling adjustments for the non-compliant weld inspections are treated by assuming the presence of weld flaws degrades the capacity of the lid welds to resist failure. Postulated thermal cycling induced through-wall cracks in the cask welds and resulting release accidents are non-credible scenarios over the life of the cask on the ISFSI for both alternatives. These cases for the two alternatives allow calculation of the absolute risk for each alternative, as well as the delta risk for each alternative. The primary risk metric used in this risk assessment is latent cancer fatality to the public (/yr). NUREG-1864 [1] analysis determined that acute fatalities to the public are not applicable to dry cask storage accident scenarios; that determination is applicable to MNGP, as well.
Risk Acceptance Criteria Reference [9] provides proposed guidance for "Risk-Informed Decision-Making for Nuclear Material and Waste Applications." Reference [9] indicates that for exemptions and changes to the licensing basis of a facility that would tend to increase risk, very general guidance can be adapted from the RG 1.174 [10]. Specific requirements may be relaxed if the initial risk is already low and the incremental increases from a change are also small. Table 5 provides the Quantitative Health Guidelines (QHGs) proposed for determining negligible risk.
016045-RPT-01 Development of Methodology to Apply/Extrapolate NUREG-1864 Results to DSC 11-15 Revision 0 Page 22 Table 5: Proposed Criteria for Acceptable Risk ChangeQuantitative Health Guidelines Risk Metric Criteria for Risk Change QHG-1 Public individual risk of acute fatality Negligible if 5x10-7 /yr QHG-2 Public individual risk of latent cancer fatality Negligible if 2x10-6 /yr QHG-3 Public individual risk of serious injury Negligible if 1x10-6 /yr QHG-4 Worker individual risk of acute fatality Negligible if 1x10-6 /yr QHG-5 Worker individual risk of latent cancer fatality Negligible if 1x10-5 /yr QHG-6 Worker individual risk of serious injury Negligible if 5x10-6 /yr Table 4.2 of Reference [9] suggests that a 10% change in QHG would be acceptable. This is consistent with the criteria provided in Figures 4 and 5 for RG 1.174 [10]. Figure 4 of this report provides an adaptation of the RG. 1.174 [10] for QHG-2. Table 18 of NUREG-1864 [1] indicates that the risk metrics associated with QHG-1 and QHG-2 were quantified for the reference site.
Table 18 also indicates that the contribution for QHG-1 was negligible for the reference site. It is reasonable to assume that the MNGP results for QHG-1 would be similar. Consequently, the primary focus of this risk analysis is with respect to QHG-2. As such, the risk criteria shown in Figure 4 are used in this analysis to assess the acceptability of the proposed alternatives.
Acceptability is shown as a measure of the calculated delta risk with respect to the absolute risk.
016045-RPT-01 Development of Methodology to Apply/Extrapolate NUREG-1864 Results to DSC 11-15 Revision 0 Page 23 Public individual risk of latent cancer fatalityPublic individual risk of latent cancer fatality Region III Region II 10-6 10-7 10-7 10-8 Region I RegionI*NoChangesAllowed RegionII *SmallChanges *TrackCumulativeImpacts RegionIII *VerySmallChanges *MoreFlexibilitywithRespecttoBaselineRiskofLatentCancerFatality *TrackCumulativeChanges Figure 4 - Proposed Risk Criteria for NUREG-1864 Comparison 016045-RPT-01 Application of Methodology to Apply/Extrapolate NUREG-1864 Results to DSC 11-15 Revision 0 Page 24 3.0 APPLICATION OF METHODOLOGY TO APPLY/EXTRAPOLATE NUREG-1864 RESULTS TO DSC 11-15 In this section of the report, JENSEN HUGHES documents the application of the methodology developed in Section 2 to arrive at a comparative evaluation of the risks for (1) transfer, examination, and return of the canisters for continued storage versus (2) continued storage of the non-compliant canisters in horizontal storage modules (HSMs). This task includes a review of all relevant and available documents and the collection of appropriate data to support the
evaluation. 3.1 Overview of Application In Section 2.0 of this report, JENSEN HUGHES documents the application of the methodology developed in Section 3. 3.2 Initiating Events Section 2.2 of this report provides the methodology to address in the MNGP adaptation of NUREG-1864 [1].
3.2.1 Flood Section 2.2.1 of NUREG-1864 [1] addresses the external flood impact on the ISFSI of the reference site. External floods are screened out from further on the basis that the flood waters for the combined maximum storm, sustained winds, and dam failures would be insufficient to reach the storage cask on the storage pad. As discussed in the MNGP IPEEE [4], the probable maximum flood for MNGP corresponds to a peak elevation of 939.2, which is 9 feet above plant grade. A recent flood re-evaluation report
[25] concluded that this flood elevation bounds the actual hazard at MNGP. The MNGP ISFSI is located at 943 feet above MSL [6], thus the probable maximum flood will not reach the bottom of the casks, thus, floods events are screened out of this evaluation.
3.2.2 Tsunamis Section 3.2.2 of NUREG-1864 [1] addresses the impact of tsunamis on the ISFSI of the reference site. Tsunamis are screened out on the basis that the reference site is far enough inland that it will not be affected by a tsunami. Tsunamis are not explicitly addressed in the MNGP IPEEE [4]. Consistent with NUREG-1864 [1], tsunamis can be screened out because the site is far enough inland that it will not be affected by tsunamis. 3.2.3 Volcanic Activity Section 3.2.3 of NUREG-1864 [1] addresses the impact of volcanic activity on the ISFSI of the reference site. Volcanic activity was screened out on the basis that the reference site is far from volcanic regions and well out of the influence of volcanic activity. The MNGP IPEEE [4] screened out volcanic activity generically with the statement that such events do not apply to Monticello. There are no volcanoes nearby. Volcanic activity hazards are screened out from further consideration in this evaluation. 3.2.4 Intense Precipitation Section 3.2.4 of NUREG-1864 [1] addresses the impact of intense precipitation on the ISFSI of the reference site. Intense Precipitation was screened out on the basis that the ISFSI of reference site is designed so that graded land and drains conduct water away from the storage 016045-RPT-01 Application of Methodology to Apply/Extrapolate NUREG-1864 Results to DSC 11-15 Revision 0 Page 25 pads. Intense precipitation is included in flood analysis discussed in Section 3.2.1 of this report, which concluded that a flood event could not impact the MNGP ISFSI. Consequently, intense precipitation is screened out from further consideration in this evaluation. 3.2.5 Storage Tanks, Transformers, Barges, Trucks, Railcars, and Nearby Industrial Facilities Section 3.2.5 of NUREG-1864 [1] addresses the impact of storage tanks, transformers, barges, trucks, railcars, and nearby industrial facilities on the ISFSI of the reference site. Storage tanks, transformers, barges, trucks, railcars, and nearby industrial facilities were screened out based on proximity of these hazard sources to the ISFSI of the reference site. Storage tanks, transformers, barges, trucks, railcars, and nearby industrial facilities potentially pose fire and/or explosive hazards to the MNGP ISFSI and to the storage cask when being transported from the reactor building to the ISFSI. The ISFSI Fire Hazards Analysis (72.212-A)
[26] evaluated the heat flux from potential fire sources and compared the heat flux to the design capacity of the storage cask and the HSM. A similar evaluation was performed to evaluate the potential for explosive shockwaves to damage the storage cask or the HSM. Appendix A.1 [26] contains the list of potential fire/explosion sources that were evaluated. The following summarizes the results of the fire hazards analysis: Fires: No fire source can produce sufficient heat flux to damage the storage cask, whether on the haul path or at the MNGP ISFSI, with the exception of diesel and gasoline delivery trucks which present a hazard to the storage cask on the haul route (if allowed on the haul route). Due to the potential of a damaging fire from diesel or gasoline delivery trucks, administrative controls are in place to keep delivery trucks sufficient distance from the haul path. During construction operations (i.e., when additional HSMs are added to the MNGP ISFSI over time), multiple vehicles burning simultaneously could damage the HSM. Administrative controls limit the number of construction vehicles allowed near the HSM during construction operations, and provide for fire watches when a single vehicle is allowed near the HSM. A fire involving the fuel load at the transfer trailer was evaluated, and the conclusion was that the fire would not result in fuel cladding temperature near the short or long-term limits. Explosions: No explosive source can produce a sufficient blast shockwave that would damage the storage cask or HSM with the exception of the diesel and gasoline delivery trucks, as discussed for fires. The same administrative controls are used to ensure such vehicles are sufficient distance from the storage cask when on the haul route and sufficient distance from the MNGP ISFSI. The administrative controls preclude a sufficient fire or blast from damaging the storage cask, whether on the haul route or at the MNGP ISFSI, and preclude damage to the HSM. Although administrative controls are considered effective enough to screen the hazard deterministically, plant staff could fail to implement the controls and a delivery truck could approach the storage cask on the haul path or approach the MNGP ISFSI. Should this occur, an accident or event that triggers an explosion would need to occur, and the truck would need to be close to the storage cask or ISFSI for any damage to result. Such an event could be an accident involving the transfer trailer and delivery vehicle, which, if the vehicles approached each other on the haul path, would presumably be avoided by each driver to the maximum extent possible. At the
ISFSI, the HSM would provide shielding for the DSC. The likelihood of the additional event or accident, combined with the likelihood of failure to follow the administrative controls, is considered low enough that failure to follow the administrative controls with subsequent accident causing a fire/explosion is assumed low and can be screened out from this evaluation.
016045-RPT-01 Application of Methodology to Apply/Extrapolate NUREG-1864 Results to DSC 11-15 Revision 0 Page 26 3.2.6 Dropped Fuel Assembly The scope of this assessment does not include a dropped fuel assembly. The fuel assemblies do not have to be removed from the DSC in either Alternative 1 or Alternative 2. 3.2.7 Dropped Transfer Cask This hazard does not apply to Alternative 1 but does apply to Alternative 2. This hazard involves consideration of two attributes: 1) number of lifts; and 2) postulated inadvertent drop rate. Alternative 2 includes two separate lifts per cask. One lift transports the Storage Cask up the equipment hatch and across the refueling floor to an area where inspections/welding is normally performed. The other lift is the reverse of this first lift to lower the cask back down the equipment shaft so it can be transported back to the ISFSI. No lifts occur on the ISFSI or transport from the
RB to the ISFSI. The drop rate frequency in NUREG-1864 [1] is developed using data in NUREG-1774, Survey of Crane Operating Experience at U.S. Nuclear Power Plants from 1968 through 2002 [20]. The data used in the NUREG-1864 analysis is the NUREG-1774 drop rate frequency for "very heavy" loads. NUREG-1774 defines "very heavy" loads as those greater than 30 tons. NUREG-1774 review of very heavy load lifts at nuclear plants showed no records of crane equipment related failures; the three very heavy load drops identified in NUREG-1774 were assessed as due to human error and not due to crane equipment failure. The NUREG-1774 "very heavy" loads inadvertent drop frequency is calculated as 3 very heavy load drops in 54,000 lifts, which is a drop rate of 5.56E-05/lift. NUREG-1864 recognizes that this frequency may be conservative because, among other reasons, the NUREG-1774 very heavy load drop incidents involved mobile cranes whereas the reference plant uses a fixed single failure proof crane. However, discussion in NUREG-1774 indicates that although a fixed single failure proof crane contributes to a lower drop frequency it does not preclude the potential of a load drop. The three very heavy load drop events identified in NUREG-1774 were all due to operator error that caused the nylon (in one case) and Kevlar (in two cases) slings to fail. Although MNGP does not use Kevlar or other fabric slings for DSC movements, rigging errors (although very remote given the controls and apparatus) can still be postulated. As such, the NUREG-1774 "very heavy" load drop frequency of 5.56E-5/lift, although likely conservative, is considered reasonable for the purposes of this risk assessment.
3.2.8 Seismic As discussed in Section 2.2.8 of this report, due to the differences in design, the NUHOMs canisters stored at MNGP in the horizontal position and are not subject to tip over. For transfer operations, consistent with NUREG-1864 [1], time based initiating events are considered unlikely during the short amount of time of transfer and tipping of the transfer trailer or seismic induced drop of the DSC inside the reactor building is screened out based on low likelihood. During storage, a seismic event could induce an acceleration load on the DSCs. The ISAR [7] evaluated large magnitude accelerations on the DSC for drop scenarios, and concluded the DSC has margin to withstand high acceleration events of 25g for a corner drop and 75g for a vertical or side drop. The frequency of seismic events that would induced an equivalent acceleration on the DSC would be very low, based on the hazard curve in Reference 3, and thus seismically induced failure for reasons other than tipping during storage is not considered a plausible failure mode for the DSC.
016045-RPT-01 Application of Methodology to Apply/Extrapolate NUREG-1864 Results to DSC 11-15 Revision 0 Page 27 3.2.9 High Winds As discussed in section 2.2.8 of this report, due to the differences in design, the NUHOMs canisters stored at MNGP are in the horizontal position are not subject to tip over. A tornado wind induced missile could impact the storage cask or the HSM. For transfer operations, consistent with NUREG-1864 [1], time based initiating events are screened out from further consideration due to very low likelihood of occurrence during the comparatively short amount of DSC cask transfer time of transfer (e.g., the conditional probability of a sufficiently extreme tornado occurring onsite during the short time period of a DSC transfer outdoors is on the order of E-10). As such, high-wind induced tipping of the transfer trailer or impact due to tornado missile during transfer operations is screened out from further consideration as a non-significant risk contributor. During storage, a wind-induced missile could impact the HSM. The minimum wall thickness for the HSM exterior walls is at least 30 inches. The end module shield walls are 24 inches. The evaluation in the ISAR concludes that missiles (including a "massive" vehicle missile) cannot affect the DSC. Based on the evaluation in the ISAR [7], high winds and high wind induced missile impacts are screened out from further consideration in this assessment. 3.2.10 Meteorites Section 3.3.5 of NUREG-1864 [1] addresses the impact of meteorites on the ISFSI of the reference site. This initiating event is retained in this risk assessment and the NUREG-1864 [1] frequency estimate for meteorite strike per area is used for MNGP. The frequency of a meteorite strike per area from NUREG-1864 [1] was applied to the MNGP specific area of the five DSCs considered in this evaluation. Meteorites that strike the transfer cask while on the haul path were not considered, based on the small fraction of time spent on the haul path relative to the amount of time spent in storage at the MNGP ISFSI. The five DSCs are enclosed in five HSMs. Consistent with NUREG-1864 [1], the roof dimensions are used to define the strike area. The roof dimension of an HSM is 9'-8" by 20'-8" (.0029464 km by .0062992 km) which yields a surface area for a single HSM of 1.85599E-05 sq km. The frequency is 4E-09/ km2, which yields of a strike frequency on a single HSM of 7.42E-14/yr.
Multiplying by 5 yields a strike frequency on the area of the 5 HSMs of 3.71E-13/yr. 3.2.11 Lightning Strikes Lightning strike induced radionuclide release accidents are determined in NUREG-1864 [1] to be non-credible. The ISAR [7] evaluated potential lightning strikes and states that the HSM will not be damaged by current flow through the concrete and will not affect the normal operation of the HSM. Lightning strikes are screened out from further consideration this assessment. 3.2.12 Aircraft Accidents Section 3.3.7 of NUREG-1864 [1] addresses the impact of aircraft impact on the ISFSI of the reference site. This hazard is maintained for MNGP in this risk assessment. Consistent with NUREG-1864 [1], airport sites within approximately 29 miles of MNGP were evaluated for the potential for an accident impacting the five HSMs. The five airport sites and the number of flights are shown in the Table 6. The source of the data are the FAA master records for each airfield, which can be accessed using the following website: http://www.gcr1.com/5010web/ The distance from MNGP (45.34° N/98.34°W) to each airport in Table 6 was approximated using google maps distance measurement feature at the following website:
016045-RPT-01 Application of Methodology to Apply/Extrapolate NUREG-1864 Results to DSC 11-15 Revision 0 Page 28 Table 6: Airfields near MNGPAirfield Appx. Distance from MNGP (mi) Annual Number of FlightsMaple Lake Municipal airport (MGG) 9 20,800 Buffalo Municipal Airport (CFE) 13 22,350 St. Cloud Regional Airport (STC) 18 28,316 Princeton Municipal Airport (PNM) 20 13,300 Crystal Airport (MIC) 30 42,351 NUREG-1864 [1] utilized NUREG-0800, Rev 2, to provide the methodology for computing the aircraft impact frequency. The current revision of NUREG-0800 is Revision 4 [28]. Revision 4 documents the same approach as Revision 2. Using equation 4 from Reference 1, the values of the following are needed to complete the estimate: Effective target area for a plane to strike the target on takeoff or landing (based on the effective area of 5 HSMs and a shadow and skid length). Probability per square kilometer (this evaluation uses miles) of crash of aircraft at each airfield. This value is provided in NUREG-0800 [28] for distances of up to 10 miles. For airfields greater than 10 miles away, the probability was extrapolated from the values in
Reference 28. Number of movements at each airfield. The number of movements is taken from the FAA master record for each of the airfields in Table 6. The product of the probability per mile of crash of an aircraft from a given field is multiplied by the number of movements from that field. The resulting probabilities for all movements of aircraft from each airfield are summed, and the sum is multiplied by the effective target area for the
HSMs. Reference 28 provides the ability to estimate the frequency based on the type of aircraft operation. The FAA master records for each facility provide the aircraft type. The aircraft types for each airfield are shown in Table 7:
016045-RPT-01 Application of Methodology to Apply/Extrapolate NUREG-1864 Results to DSC 11-15 Revision 0 Page 29 Table 7: Aircraft Types per Airfields near MNGPAirfield Air CarrierAir TaxiGA LocalGA In Transit MilitaryMaple Lake Municipal airport (MGG) 0 1000 15000 4000 800 Buffalo Municipal Airport (CFE) 0 130 11000 11000 220 St. Cloud Regional Airport (STC) 241 508 11558 12911 3098 Princeton Municipal Airport (PNM) 0 0 6500 6500 300 Crystal Airport (MIC) 0 549 19250 22458 94 The probability of crash per aircraft movement in Reference 28 is provided for Air Carrier, General Aviation, U.S. Navy/U.S. Marine, and U.S.A.F. For the purposes of this evaluation, the number of operations in Table 6 are assigned to these four categories. The Air Carrier crash probability in Reference 28 is considered applicable to Air Carrier and Air Taxi flights. The General Aviation crash probability is considered applicable to GA local and GA Intransit flights.
There are two military categories (U.S. Navy/U.S. Marine, and U.S.A.F). The U.S.A.F crash probability is higher than the U.S. Navy/U.S. Marine crash probability in Reference 28, so all Military flights were considered applicable to the U.S.A.F crash probability. The effective target area of the plane is not known and would change for each type of plane that could impact the five HSMs. Considering the uncertainty regarding the dimensions of the aircraft operating from each airfield, the same aircraft dimensions used in NUREG-1864 are assumed here. Similarly, the shadow and skid length is assumed the same as used in NUREG-1864. These dimensions are: Diameter of Engine: 5.2 ft. Centerline Spacing Between Engine: 13.8 ft. Width of 5 HSMs: (5 times 9'-8"): 48.3 ft. Shadow and Skid Length (from NUREG-1864): 200 ft. The sum of the plane and HSM dimensions listed above is 67.3 ft. (.01275 miles). The shadow and skid length is 200 ft. (.03788 miles). The area in square miles is 4.83E-04. Using the data in Table 7 and the probabilities in Reference 28, the aircraft type frequencies from each site are summed and multiplied by the target area. The resulting frequency is 7.43E-
- 08. Overflight crash hazards can be calculated using the same method in NUREG-1864 [1]. Consistent with NUREG-1864 [1], overflights are assumed to be a large aircraft. The frequency of an overflight crash is 4E-07 crashes per square mile per year. The target area is the same as described for accidents originating at nearby airfields. Multiplying the two gives an overflight crash frequency of 1.93E-10. The assumption made in NUREG-1864 [1] is that only large aircraft travelling at high velocity can fail the MPC, which requires an overflying airplane larger than a Gulfstream IV jet. The vast majority of air traffic landing or departing airports near MNGP is general aviation traffic, which 016045-RPT-01 Application of Methodology to Apply/Extrapolate NUREG-1864 Results to DSC 11-15 Revision 0 Page 30 are typically smaller airplanes travelling at slow velocities relative to large aircraft on overflight routes. St. Cloud Regional Airport (STC) does have some air carrier flights; however, even if the airplanes are larger than a Gulfstream IV, the airplanes would be travelling at slow velocity because they would be approaching to land or departing on takeoff. Based on this, the probability of DSC failure given aircraft strike is set equal to the probability of an overflight aircraft impact, which does not depend on the airports nearby, which is consistent with the evaluation of aircraft impacts in NUREG-1864 [1]. Therefore, the total initiating event frequency for aircraft impact on the subject DSCs is 7.45E-08/yr (i.e., 7.43E-08 + 1.93E-10). Dividing the overflight impact frequency by the total aircraft impact frequency gives the conditional probability the aircraft accident is a large aircraft on overflight, which is 2.59E-03. 3.3 Dry Storage Canister (DSC) Failure Model For each of the two modeled alternatives, the DSC failure model for this evaluation considers the following failure mechanisms: Mechanical failure of the DSC given a drop, including Failure of the DSC shell (NUREG-1864 [1] MPC failure probabilities assumed applicable to the DSC) Failure of the DSC lid welds (for a DSC with and without postulated weld flaws) Mechanical failure of the DSC, given a meteorite strike, or overflight aircraft accident Thermal failure scenarios, caused by Blocked air inlet and outlet vents Aircraft fuel fires 3.3.1 Mechanical Failures The DSC can fail given a drop or upon a meteorite or large aircraft strike. Alternative 1 includes the risk of meteorite and large aircraft strike, while Alternative 2 includes the additional drop failure mode. For meteorite strikes and large aircraft strikes, the failure probability is assumed to be 1.0. This is consistent with the assumption in NUREG-1864 [1] for these events. For drops, Alternative 2 postulates four different drops with two unique drop heights. NUREG-1864 [1]
evaluated the MPC and concluded that the MPC lid welds were rugged based on the weld type and the redundancy in the welds, and only evaluated the MPC shell for mechanical failure given a drop. For the DSC, considering the potential for weld flaws, both the DSC shell and lid welds need evaluation to estimate the probability of failure given a drop. For mechanical failures of the DSC, consistent with the evaluation of MPC lid welds as robust in NUREG-1864 [1], it is assumed that the DSC lid welds would be robust if they were at nominal conditions with all requisite inspections completed satisfactorily, with higher capacity to resist failure given a drop than the DSC shell welds. To account for the presence of weld flaws, the DSC lid welds with potential weld flaws are assumed to have the same capacity as the shell (rather than be robust if the weld were at nominal conditions). The DSC shell is assumed to be similar to the MPC shell such that the probabilities of failure from NUREG-1864 [1] for the MPC shell given a drop can be directly applied to the DSC shell and, with the assumption that the presence of weld flaws decreases the lid weld capacity to resist failure to be equal to the capacity of the shell, the probabilities can also be applied to the lid welds. This is reasonable, considering the following: The MPC and DSC are comprised of the same type of stainless steel (SA240 304). The MPC and DSC are of similar dimensions and the steel is of similar thickness.
016045-RPT-01 Application of Methodology to Apply/Extrapolate NUREG-1864 Results to DSC 11-15 Revision 0 Page 31 The MPC and DSC are designed to the requirements of the same ASME class. The drop heights for the transfer operations in NUREG-1864 [1] and for the applicable MNGP Alternative 2 heights are similar. The drop surfaces are similar (refueling floor, concrete or transfer trailer/storage overpack, with concrete drop being limiting and used in this evaluation). The MPC lid welds were evaluated to be robust and not analyzed for failure in NUREG-1864 [1], meaning they have higher nominal capacity to resist failure relative to the capacity of the shell. The combined failure probability for the DSC shell and lid welds is the sum of the probability of failure of the shell and failure of the lid welds. This sum equates to twice the failure probability of the MPC evaluated in NUREG-1864 [1] for a given drop height equivalent to the applicable drop height in MNGP Alternative 2. The stages applicable to the MNGP Alternative have probabilities for release from the fuel rod and MPC of 1E-06 and 1.96E-02, thus, for the MNGP DCSs with potential weld flaws, the probabilities are 2E-06 and 3.92E-02, respectively. 3.3.2 Thermal Failures Thermal failures can result from heating of the DSC via blocked air inlet and outlet vents or a fire affecting the HSM. Each of these thermal scenarios is evaluated for potential impact to the DSC, given the potential condition of weld flaws for DSC lid welds. For a DSC at design capacity, the ISAR [7] evaluated the blocked inlet and outlet air vent scenario using a set of conservative assumptions. The evaluation concluded that neither the fuel cladding nor the DSC would exceed temperature limits with vents blocked for up to five days. The evaluation concluded that the decay heat from the fuel would be transferred to the HSM and because the HSM has a very slow heat-up time, the heat transfer can be considered steady state. Although the fuel and DSC temperature limits would not be exceeded, the HSM temperature limit may be exceeded, and a daily surveillance is conducted to ensure that inlet and outlet vents are clear of debris or thermal performance is monitored which would limit the heat-up time. NUREG-1864 [1] evaluated a blocked vent scenario for the 20-year duration of the storage phase. The conclusion was that although some fuel damage would occur, the MPC would not
fail. NUREG-1864 [1] evaluated a jet fuel fire from an aircraft crash and concluded that the MPC would not fail given a three-hour duration fire that occurs after 20 years of blocked vents. The statement is made that a 30-minute fire scenario is more realistic, which means the three-hour jet fuel fire scenario is very conservative. The NUREG-1864 [1] model did not explicitly evaluate MPC lid welds. The model evaluated the MPC shell including the axial and circumferential shell welds for limit load and creep rupture.
The load limit model never contributed to failure for any of the heat-up scenarios. Creep rupture was evaluated with the presence of weld flaws. The creep rupture model calculated the time to rupture given a time at temperature and the stress at that temperature. The presence of weld flaws was modeled as a stress magnification factor which increases the stress for a given temperature and reduced the time to creep rupture failure. Even with flaws assumed in the axial and circumferential welds on the MPC shell, 20 years of blocked vents or a 3-hour aircraft fuel fire did not result in the failure MPC. This evaluation shows that the MPC, including shell welds, which are considered by NUREG-1864 to be less robust than lid welds, is very robust against failure due to thermal scenarios. For the MNGP DSC, the capacity of the DSC to resist failure given a blocked vent and/or fire scenario, and the presence of potential weld flaws in the lid welds, is considered to be similar to the capacity of 016045-RPT-01 Application of Methodology to Apply/Extrapolate NUREG-1864 Results to DSC 11-15 Revision 0 Page 32 the MPC shell with presence of shell weld flaws evaluated for the MPC shell weld creep rupture model. Thus, the MNGP DSC is assumed to be relatively robust to thermal failures. Considering the daily surveillance requirement to ensure the air inlet and outlet vents are clear of debris or thermal performance monitoring, DSC failure from a blocked vent is deemed incredible and is screened out from this evaluation. For potential aircraft accidents involving a jet fuel fire a realistic fire scenario will be less than 3 hours3.472222e-5 days <br />8.333333e-4 hours <br />4.960317e-6 weeks <br />1.1415e-6 months <br /> and suppression of the fire will preclude sufficient heat-up that leads to creep rupture, and this event is deemed incredible and screened out from this evaluation. No thermal scenarios are included in the risk of Alternative 1 and Alternative 2. 3.4 Fuel Assembly Failure Model Fuel assemblies can fail given a drop of the DSC. In NUREG-1864 [1], fuel failure is included in the overall failure of the MPC given a drop, and this is applied to this evaluation as well. The probability of fuel failure given a drop scenario from NUREG-1864 [1] is assumed applicable because the drop scenarios are similar. For shorter drops, fuel failure is not expected and is bound by the low probability of failure of the DSC. For long drops, the probability of fuel failure is likely, but the DSC has to fail to generate a release. 3.5 Secondary Containment Isolation Model The MNGP Secondary Containment Isolation system functions to isolate the containment from the environment and activate the SGTS (Standby Gas Treatment System), which is the same function modeled in NUREG-1864 [1] for the reference site. For the purposes of this evaluation, the Secondary Containment Isolation model in NUREG-1864 [1] is considered applicable to the MNGP Secondary Containment Isolation system based on the following similarities to the model in NUREG-1864 [1]: MNGP and the NUREG-1864 reference plant are similar reactor containment designs (i.e., GE BWR Mark I). The MNGP reactor building is maintained at a slightly negative pressure relative to the environment, which minimizes the amount of exfiltration [12]. The MNGP secondary containment isolation system is highly redundant, and consists of multiple detector trains, redundant isolation dampers at the secondary containment boundary, and trip signals to isolate fans on initiation of SGTS [12]. Redundant isolation dampers close on detection of radiation in the reactor building exhaust plenum or in the area of the spent fuel pool and provide for SGTS initiation [12]. The NUREG-1864 [1] model is for a two-unit site. The model in Figure 18 [1] shows that the reference site has more isolations to complete to isolate the secondary containment (due to the shared nature of the reactor building shown in Figure 17 [1]), which would tend to increase the probability of failing to isolate relative to MNGP. The NUREG-1864 [1] model includes credit for both SGTSs, whereas MNGP only has one SGTS with two trains, which would tend to decrease the probability of failure relative to MNGP. Without a detailed fault tree of the secondary containment isolation model in NUREG-1864 [1] and without a detailed fault tree model for the MNGP secondary containment isolation and SGTS, an exact comparison is not feasible.
However, given the similarities discussed above, the NUREG-1864 [1] probability for failure to isolate the secondary containment is assumed applicable to the MNGP configuration in this evaluation. There are different probabilities of failing to isolate the secondary containment in the results of this assessment. The two probabilities of failing to isolate the secondary containment are 1.0 for noble gas releases, and 1.57E-04 for all other releases. Noble gases are not captured by the exhaust filters and thus the probability of failing to isolate the secondary containment is 1.0 for noble gas releases.
016045-RPT-01 Application of Methodology to Apply/Extrapolate NUREG-1864 Results to DSC 11-15 Revision 0 Page 33 3.6 Consequence Model This section applies the methodology outlined in Section 2.6. Section 6.0 of NUREG-1864 [1] describes the consequence analysis for the reference site. Important consequence related parameters and inputs noted in NUREG-1864 [1] include the following: Fuel type and exposure Radionuclide inventory Source term (e.g., release fraction) Initial plume dimensions Plume heat content Population density/distribution Site weather Each of these inputs is reviewed for applicability and comparison to the MNGP configurations. 3.6.1 Fuel Type and Exposure The NUREG-1864 [1] consequence results are based on a single core containing 68 BWR fuel assemblies that were high burnup (i.e., 50 GWd/MTU) with 10 years of cooling. The MNGP cask contains 61 BWR fuel assemblies with a core average exposure of 41 GWd/MTU (i.e., not high burnup fuel, where high burnup fuel is defined in NUREG-1864 [1] as greater than 45 GWd/MTU), with 15.5 years of cooling [35]. Thus, the MNGP fuel type and exposure parameters are all bounded by the NUREG-1864 [1] parameters. 3.6.2 Radionuclide Inventory The NUREG-1864 [1] nuclide inventory for a single cask is provided in Table E-1 of that document. It is reproduced in Table 8 below, with the MNGP values taken from Table 6-1 of Reference [8] included for comparison. On average, the cask radionuclide inventory activity for the NUREG-1864 [1] cask is 7.0 times that of MNGP cask. Therefore, the NUREG-1864 [1] inventory is found to bound that of the subject MNGP casks with significant margin. Table 8: Cask Radionuclide Inventory ComparisonRADIONUCLIDE NUREG-1864 (CI)
MNGP DSC(CI) RATIO(1864 / MNGP) CONCLUSION Am-241 3.2504E+04 1.6758E+04 1.94 Bounded Am-242m 5.32E+02 6.7903E+01 7.83 Bounded Am-242 N/A 6.7598E+01 -- Indeterminate Am-243 8.30E+02 1.1141E+02 7.45 Bounded Ba-137m N/A 7.1753E+05 -- Indeterminate Ce-144 1.374E+03 N/A -- Indeterminate Cm-242 N/A 5.5901E+01 -- Indeterminate Cm-243 8.16E+02 6.5201E+01 12.52 Bounded Cm-244 1.53000E+05 5.3060E+03 28.84 Bounded Co-60 3.133E+03 6.7047E+02 4.67 Bounded Cs-134 1.38720E+05 5.8502E+03 23.71 Bounded Cs-137 1.496000E+06 7.5991E+05 1.97 Bounded Eu-154 1.12200E+5 1.3159E+04 8.53 Bounded Kr-85 7.4800E+04 3.9393E+04 1.90 Bounded Np-239 N/A 1.1141E+02 -- Indeterminate Pm-147 9.1120E+04 2.9001E+04 3.14 Bounded Pu-238 1.07440E+05 1.8304E+04 5.87 Bounded 016045-RPT-01 Application of Methodology to Apply/Extrapolate NUREG-1864 Results to DSC 11-15 Revision 0 Page 34 Table 8: Cask Radionuclide Inventory ComparisonRADIONUCLIDE NUREG-1864 (CI)
MNGP DSC(CI) RATIO(1864 / MNGP) CONCLUSION Pu-239 5.060E+03 2.5568E+03 1.98 Bounded Pu-240 9.384E+03 4.1678E+03 2.25 Bounded Pu-241 1.414400E+06 4.2150E+05 3.36 Bounded Ru-106 7.888E+03 N/A -- Indeterminate Sm-151 N/A 2.7539E+03 -- Indeterminate Sr-90 9.18000E+05 5.8899E+05 1.56 Bounded Y-90 918000E+05 5.8909E+05 1.56 Bounded (1) The MNGP DSC radionuclide inventory reflects the activity at end of core and does not include the 15.5 years of decay time. Inclusion of the decay time would be expected to increase the calculated 1864 / MNGP ratio for most radionuclides, adding additional margin.
3.6.3 Source Term NUREG-1864 [1] presents MACCS2 conditional consequence results in Table E.3 for six (6) cases. The six cases reflect the consequences associated with a 100 ft. cask drop, but with differing attributes of fuel damage, release height, and release filtering. Section 6.2.2 of NUREG-1864 identifies one other consequence result for a release of only noble gases. Table 9 below summarizes these seven consequence results (latent individual cancer fatalities within 10 miles). The final NUREG-1864 [1] risk assessment as reflected in Table 18 of that document primarily relies upon the bounding consequence result of 3.6E-04/yr (see Case 1 in Table 9 below). However, the other cases may be applicable to certain configurations (e.g., to represent the lower expected release associated with releases from a DSC on the pad). For example, Case 6 would be judged to best represent a DSC located on the pad which was not subject to a cask impact failure (e.g., drop), where the release pathway would be through a weld flaw.
Absent a cask impact, the fuel pellet rim fracture factor of 1.24E-04 (the value representative of the fuel pellet as a whole) is judged most representative. It is noted that, per NUREG-1864, the formation of a rim on the fuel pellet is primarily a phenomenon associated with only high burnup fuel and would be conservative for MNGP fuel. Without an impact, a release of noble gases, fuel particles, and CRUD (i.e., Chalk River unidentified deposits) would be conservative given that the release would generally be limited to only CRUD. The torturous release pathway of the weld flaw would also be expected to reduce the release in a manner akin to a filtered release. In summary, review of the NUREG-1864 source term attributes indicate that they would adequately represent or bound those of the MNGP configuration, depending upon the case selected. Table 9: NUREG-1864 Consequence Result CasesCASE FUEL PELLET RIM FRACTURE FACTOR RELEASE FRACTION CONTRIBUTORS RELEASE HEIGHT (M) RELEASE FILTERING INDIVIDUAL CANCER FATALITY RISK
(/YR) 1 1.0 Noble Gas, Fuel Particles, CRUD 50 Not Filtered 3.6E-04 2 1.0 Noble Gas, Fuel Particles, CRUD 120 Not Filtered 2.1E-04 3 1.0 Noble Gas, Fuel Particles, CRUD 50 Filtered 5.2E-05 4 1.24E-4 Noble Gas, Fuel Particles, CRUD 50 Not Filtered 4.3E-06 5 1.24E-4 Noble Gas, Fuel Particles, CRUD 120 Not Filtered 2.6E-06 6 1.24E-4 Noble Gas, Fuel Particles, CRUD 50 Filtered 4.3E-07 016045-RPT-01 Application of Methodology to Apply/Extrapolate NUREG-1864 Results to DSC 11-15 Revision 0 Page 35 Table 9: NUREG-1864 Consequence Result CasesCASE FUEL PELLET RIM FRACTURE FACTOR RELEASE FRACTION CONTRIBUTORS RELEASE HEIGHT (M) RELEASE FILTERING INDIVIDUAL CANCER FATALITY RISK
(/YR) 7 1.0 Noble Gas only 50 Not Filtered 1.0E-10 3.6.4 Initial Plume Dimensions The initial plume dimensions used in the MACCS2 code are dependent upon the building wake effects, as calculated based on the building width and height (typically 40m to 50m). For the NUREG-1864 [1] consequence analysis, a structure size of 4m wide by 4m high was used, approximating the dimensions of a cask. It should be noted that this value was used even for drops postulated to occur inside the site reactor building where plume effects would actually be based on the reactor building structure. The MNGP NUHOMs storage system is comprised of individual HSMs (10' wide, 20' long, 15' tall) situated back-to-back and side-to-side to comprise a 2x15 array that is 150' wide across HSM fronts, 40' wide (two HSMs back-to-back) and 15' tall
[27, 32, 33]. For the purposes of the consequence assessment, the NUHOMs storage system configuration is judged to be adequately represented by the initial plume dimension parameters used in the NUREG-1864 [1] consequence calculation. Likewise, a release from the DSC during transfer (along with haul path) and from the reactor building would also be represented by the NUREG model. 3.6.5 Plume Heat Content The NUREG-1864 [1] notes that the plume heat content for a cask release is estimated to be that of the spent fuel. For ten-year old spent fuel, NUREG-1864 [1] estimates the maximum decay heat load to be 264 watts per assembly. For the Monticello DSCs, the fuel is over 15 years old and the maximum decay heat load (i.e., approximately 220 watts per assembly per Reference [35]) is bounded by the NUREG-1864 [1] estimate. NUREG-1864 [1] notes that the plume resulting from the release will not be thermally hot enough to produce significant plume rise. 3.6.6 Population The NUREG-1864 [1] consequence results of interest in the risk assessment are presented in the form of Individual Risk of Cancer Fatality. These results are developed by taking the population-weighted health effect risk and dividing by the total population in the region to develop the metric for individual risk. Since the consequence metric is based on individual risk, the metric should be relatively insensitive to absolute population differences between the NUREG-1864 [1] site and the MNGP site. The NUREG-1864 [1] population is based on year 2000 population data using the SECPOP code, and is assumed to be the Hatch site. For this current assessment, the same SECPOP code (ver. 4.2) [31] was used to develop the 10-mile radius population distribution for the Hatch site for year 2000 and the MNGP site for year 2010 for comparison purposes. Tables 10 and 11 present the results for Hatch and Monticello, respectively. The 10-mile population for Hatch is 8,539, while that of Monticello is 58,869. The larger population of Monticello however, does not automatically translate into a linear increase in individual risk consequence results. The following insights are noted from comparing the population distribution for the two sites: Both Hatch and Monticello have some population variation as a function of direction from the site. Prevailing wind directions could thus have impacts or radiological dispersion.
This is evaluated in the comparison of weather.
016045-RPT-01 Application of Methodology to Apply/Extrapolate NUREG-1864 Results to DSC 11-15 Revision 0 Page 36 The Hatch and Monticello population distribution has some variation as a function of distance from the site. For releases at ground level, a closer population would tend to increase radiological impacts. A release from the pad would be expected to have deposition closer to the site rather than further from the site. For elevated releases, a closer population would tend to decrease radiological impacts given that the radiological plume would tend to travel over areas near the site with deposition occurring more once the plume expands in the vertical direction down to the grade elevation. Review of Tables 10 & 11 show that the percent of total population within 3 miles of the site is higher for Hatch than for Monticello (i.e., 10% as compare to 8%, respectively). Therefore, with respect to releases from the pad, the Hatch attributes for population distance distribution are judged to bound those for MNGP. With respect to releases from the reactor building, the population distribution for Hatch is judged to adequately
represent those for MNGP, considering the many potential areas of variability associated with atmospheric dispersion and deposition.
016045-RPT-01 Application of Methodology to Apply/Extrapolate NUREG-1864 Results to DSC 11-15 Revision 0 Page 37 Table 10: Hatch Population Distribution (Year 2000) Sector 0-0.3 miles 0.3-1 miles 1-2 miles 2-3 miles 3-4 miles 4-5 miles 5-7 miles 7-10 miles 0-10 miles % of Total N 0 0 20 4 19 103 65 251 462 5% NNE 0 0 0 0 0 18 48 282 348 4%
NE 0 0 0 0 21 30 57 276 384 4%
ENE 0 0 0 0 0 2 19 101 122 1%
E 0 0 0 0 0 23 0 21 44 1%
ESE 0 0 0 27 0 14 72 199 312 4%
SE 0 0 0 64 13 63 136 104 380 4%
SSE 0 0 0 30 71 110 217 321 749 9%
S 0 0 43 127 53 40 374 1855 2492 29%
SSW 0 0 54 78 72 72 87 201 564 7%
SW 0 0 89 0 38 35 89 126 377 4% WSW 0 0 0 144 0 55 77 317 593 7%
W 0 0 91 0 0 0 20 88 199 2%
WNW 0 0 37 5 223 0 42 155 462 5%
NW 0 0 0 0 4 0 155 177 336 4%
NNW 0 0 1 20 54 47 168 425 715 8% Total 0 0 335 499 568 612 1626 4899 8539 100% % of Total 0 0 4% 6% 7% 7% 19% 57% 100%
% Cum 0 0% 4% 10% 16% 24% 43% 100%
016045-RPT-01 Application of Methodology to Apply/Extrapolate NUREG-1864 Results to DSC 11-15 Revision 0 Page 38 Table 11: Monticello Population Distribution (Year 2010) Sector 0-0.3 miles 0.3-1 miles 1-2 miles 2-3 miles 3-4 miles 4-5 miles 5-7 miles 7-10 miles 0-10 miles % of Total N 0 0 4 0 668 719 334 445 2170 4% NNE 0 0 0 2 155 190 397 1241 1985 3%
NE 0 0 0 11 262 621 675 1229 2798 5%
ENE 0 0 57 32 1356 2506 1665 2010 7626 13%
E 0 0 205 110 276 1876 3293 2191 7951 14%
ESE 0 0 63 586 214 827 1314 1662 4666 8%
SE 0 0 886 1278 2271 2369 3428 480 10712 19%
SSE 0 0 349 46 1277 639 152 116 2579 5%
S 0 22 3 58 130 263 226 4508 5210 9%
SSW 0 0 143 220 0 124 174 703 1364 2%
SW 0 0 0 0 322 92 323 643 1380 2% WSW 0 0 0 94 84 29 108 656 971 2%
W 0 0 183 22 164 11 358 624 1362 2%
WNW 0 0 98 1 32 232 274 524 1161 2%
NW 0 0 0 0 29 39 36 470 574 1%
NNW 0 0 0 0 44 1044 2311 961 4360 8% Total 0 22 1991 2460 7284 11581 15068 18463 56869 100% % of Total 0 0.04% 4% 4% 13% 20% 26% 32% 100% % Cum 0 0% 4% 8% 21% 41% 68% 100% 3.6.7 Site Weather The NUREG-1864 [1] consequence results are based MACCS2 code sampling of an annual weather file, including data for wind speed, wind direction, atmospheric stability, and rainfall. The consequence results are based on the mean values generated from the weather sampling. It is not possible to compare all weather variables fully without a MACCS2 code calculation.
However, some comparison can be made using the Exposure Index approached documented in NUREG-1437 [29]. NUREG-1437 (p. 5 - 25) notes the following:
. NUREG-1437 [29] proceeds to develop the Exposure Index (EI) approach to estimate consequence results based on population and wind direction frequency (i.e., fraction of time per year that wind blows in each compass sector direction). An EI value is developed by multiplying the wind direction frequency by the downwind population out to a certain distance. Wind direction frequency for each site is available in Table A.4-1 of NUREG/CR-2239 [30].
016045-RPT-01 Application of Methodology to Apply/Extrapolate NUREG-1864 Results to DSC 11-15 Revision 0 Page 39 Table 12 and 13 present the EI results for the Hatch and Monticello sites, respectively. The following insights are noted from comparing the population distribution for the two sites: For Hatch, the wind frequency in the most populated direction (south) is below average. Overall, the EI value is below average by approximately 14% indicating that the wind tends to blow towards lower populated sectors. For Monticello, the wind frequency in the most populated direction (southeast) is above average. Overall, the EI value is above average by approximately 7% indicating the wind tends to blow towards higher populated sectors. The differences between the two sites as explored with the exposure index are not considered substantial. Experience with the MACCS2 code has shown that using different annual weather data files (e.g., 2008 vs. 2009) for a given site will often result in differences of +/- 5% on calculated mean dose impacts. Such differences derive from the
inherent variability of weather parameters from year to year. The variation exhibited between the two sites is not significantly greater. Therefore, with respect to site weather, the NUREG-1864 consequence results are judged to adequately represent those of MNGP. Table 12: Hatch 10
-Mile Exposure IndexSector NUREG/CR-2239 Wind Rose 0-10 Mile Pop Exp Index % of Total N 0.055 462 25 6% NNE 0.069 348 24 5%
NE 0.082 384 31 7%
ENE 0.073 122 9 2%
E 0.075 44 3 1%
ESE 0.077 312 24 5%
SE 0.072 380 27 6%
SSE 0.049 749 37 8%
S 0.04 2492 100 22%
SSW 0.038 564 21 5%
SW 0.051 377 19 4% WSW 0.067 593 40 9%
W 0.081 199 16 4%
WNW 0.068 462 31 7%
NW 0.057 336 19 4%
NNW 0.044 715 31 7% Total 0.998 8539 459 100% EI if ave 534 % off ave -13.9%
016045-RPT-01 Application of Methodology to Apply/Extrapolate NUREG-1864 Results to DSC 11-15 Revision 0 Page 40 Table 13: Monticello 10-Mile Exposure IndexSector NUREG/CR-2239 Wind Rose 0-10 Mile Pop Exp Index % of Total N 0.089 2170 193 5% NNE 0.091 1985 181 5% NE 0.063 2798 176 5% ENE 0.055 7626 419 11% E 0.03 7951 239 6% ESE 0.089 4666 415 11% SE 0.104 10712 1114 29% SSE 0.119 2579 307 8% S 0.036 5210 188 5% SSW 0.041 1364 56 1% SW 0.029 1380 40 1% WSW 0.051 971 50 1% W 0.031 1362 42 1% WNW 0.055 1161 64 2% NW 0.052 574 30 1% NNW 0.065 4360 283 7% Total 1 56869 3797 100% EI if ave 3554 % off ave 6.8% The consequence probabilities in NUREG-1864 [1] apply to failure of a single MPC. The consequence probabilities are judged to be reasonable to represent failure of the MNGP DSCs in this evaluation. Alternative 1 and 2 include hazards that can affect up to five (5) DSCs (aircraft strikes, meteorite strike) during years of storage. The consequences applied represent failure of a single DSC but the frequency of aircraft strike and meteorite strike assumes that any one of the five DSCs can be impacted by the hazard.
3.7 Results The risk of Alternative 1 and Alternative 2 has been evaluated to determine the absolute value of latent cancer fatality risk for DSCs 11-15 and the relative risk of the alternatives considering the potential presence of flaws in the DSC lid welds. The major assumptions used in this quantitative evaluation are: 1. Consistent with NUREG-1864 [1], time based initiating events (seismic events, high winds, floods) are assumed not to occur during transfer of the DSC to and from the Fuel Building and during inspection of the welds, based on the short amount of time that occurs during transport and inspection. 2. Tipping of the HSM due to a seismic or high wind event is assumed incredible, based on the horizontal configuration of the HSMs. 3. Sliding of the HSM is assumed to have no impact on the DSC. The likelihood of sliding is low based on the size and weight of the HSMs and the low likelihood that a wind or seismic 016045-RPT-01 Application of Methodology to Apply/Extrapolate NUREG-1864 Results to DSC 11-15 Revision 0 Page 41 event occurs. If sliding occurred, no damage would occur to the DSC unless the HSM was slid into another object or off of the ISFSI pad. 4. The failure probabilities for the DSC shell given a drop are based on similarity to the MPC shell evaluated in NUREG-1864 [1]. DSC lid welds with flaws, are assumed to be the same capacity of the shell, which gives an overall DSC failure probability of twice the MPC shell failure probability from NUREG-1864 [1] for applicable drops. This is based on the evaluation in NUREG-1864 [1] that lid welds are robust compared to the MPC shell, based on weld type and weld redundancy, so the presence of weld flaws degrades the capacity from robust to be equal to the shell capacity. 5. Thermal scenarios were considered incredible based on the evaluation of the MPC in NUREG-1864 [1], which included weld flaws for shell welds, which indicate robust design capacity against thermal events, and assumed equivalence of the DSC shell and lid welds to the MPC, and the short time duration of blocked vent and aircraft fire events. 6. Consistent with NUREG-1864 [1], the conditional probability of failure of the fuel cladding and the DSC is assumed to be 1.0 for large aircraft overflight strikes, and meteor strikes.
The presence of potential weld flaws does not impact the resulting risk calculations. If detailed evaluations showed the potential for DSC survival given a large aircraft or meteorite strike, the potential for weld flaws may impact the resulting probability of release, but given
the uncertainty and the potential magnitude of these two events, it is assumed that the DSC will fail regardless of the presence of potential weld flaws. A summary of each of the contributors to each Alternative is shown in Tables 14 and 15. A summary of the results in the risk tables is as follows: Storage stage risk is shown for one year of storage of all five DSCs in the individual rows in Tables 14 and 15 for Alternative 1 and 2. When included in the totals in Table 14 and 15, storage risk is multiplied by twenty years. Transport stage risk is shown for a single DSC in Table 15 for Alternative 2. When included in the total in Table 15, transport risk is multiplied by 5 DSCs and added to the risk of storage of all five DSCs for twenty years. For the storage risk of both alternatives, the frequency of aircraft strike and meteorite strike is adjusted to use the target area of all DSCs, so risk reflects a release from any of the five (5) DSCs, but not simultaneous release of multiple DSCs. For the unique transport stages of Alternative 2, although there are only two lifts, the probability of failure of the fuel rod and DSC with subsequent release depends on the distance the DSC would fall, thus, the two lifts are separated into four rows in the results in Table 15 in order to capture the different release probabilities. The results are slightly conservative, because a split fraction of the lift distance could be applied to the two rows as opposed to 100% of the lift frequency being applied to each, but the impact of the fraction on the results would be small because the higher probability of release given a drop down the equipment hatch is much more significant to the results compared to the probability of release given the very small drop distance when the DSC is moved over the refueling floor. Tables 14 and 15 include stages for each alternative. These stages are named for convenience and are in chronological order for Alternative 2. Also for Alternative 2, where applicable, the MNGP specific process steps from Table 2 are included.
016045-RPT-01 Application of Methodology to Apply/Extrapolate NUREG-1864 Results to DSC 11-15 Revision 0 Page 42 Table 14: Alternative 1 RiskStages Initiating Event Initiating Event Frequency per Year (Sections 2.2 and 3.2) Probability of Release from fuel rod and MPC (Sections 2.3 and 3.3) Probability of Release from Containment (Sections 2.5 and 3.5) Consequences (Sections 2.6 and 3.6) RiskPer Yea r1-1 Storing the DSCs (for 1 year) Struck by Aircraft (overflight) 7.45E-08 2.59E-03 (conditional probability of aircraft being large plane on overflight) 1.00 (no secondary containment at the HSM) 3.60E-04 6.95E-14 1-2 Storing the DSCs (for 1 year) Struck by Meteorite 3.71E-13 1.00 1.00 (no secondary containment at the HSM) 3.60E-04 1.34E-16 Total Risk of Alternative 1 for all 5 DSCsSum of stages 1-1 and 1-2 multiplied by 20 years 1.39E-12
016045-RPT-01 Application of Methodology to Apply/Extrapolate NUREG-1864 Results to DSC 11-15 Revision 0 Page 43 Table 15: Alternative 2 Risk Stages (and MNGP Table 2 Steps) Initiating Event Initiating Event Frequency Per Year (Sections 2.2 and 3.2) Probability of Release from fuel rod and MPC (Sections 2.3 and 3.3) Probability of Release from Containment (Sections 2.5 and 3.5) Consequences (Sections 2.6 and 3.6) RiskPer Yea r2-1 (25-34) Removing the DSC from the HSM and transporting to the Equipment Hatch None 2-2 (22-23) Lifting the DSC from the TT and raising through the Equipment Hatch (weld flaws assumed) Dropped DSC 5.65E-05 3.92E-02 1.57E-04 1.0 (Noble Gases) 3.60E-04 1.0E-10 1.23E-13 2.18E-16 2-3 (21) Lifting/moving the DSC over the Refuel Floor to the Inspection Area (weld flaws assumed) Dropped DSC 5.65E-05 2E-06 1.57E-04 3.60E-04 6.29E-18 2-4 Inspecting and Repairing Welds, if necessary None 2-5 (20-21) Lifting/moving the DSC over the Refuel Floor from the Inspection Area to the Equipment Hatch (weld flaws repaired) Dropped DSC 5.65E-05 1E-06 1.57E-04 3.60E-04 3.15E-18 2-6 (22-23) Lifting the DSC and lowering down the Equipment Hatch to the TT (weld flaws repaired) Dropped DSC 5.65E-05 1.96E-02 1.57E-04 1.0 (Noble Gases) 3.60E-04 1.0E-10 6.17E-14 1.09E-16 2-7 (25-34) Transporting the DSC to the HSM and re-inserting into the HSM None 2-8 Storing the DSCs (for 1 year) Struck by Aircraft (overflight) 7.45E-08 2.59E-03 (conditional probability of aircraft being large plane on overflight) 1.00 (no secondary containment at the HSM) 3.60E-04 6.95E-14 2-9 Storing the DSCs (for 1 year) Struck by Meteorite 3.71E-13 1.00 1.00 (no secondary containment at the HSM) 3.60E-04 1.34E-16 Total Risk of Alternative 2 for all 5 DSCsSum of stages 2-1 to 2-7 multiplied by 5 DSCs plus sum of stages 2-8 and 2-9 multiplied by 20 years 2.32E-12The total risk of each alternative is very low and is several orders of magnitude lower than the acceptance criteria in Table 5. These results are on the same order of magnitude as the results in NUREG-1864 [1], which is reasonable considering the similarity in cask designs and the 016045-RPT-01 Application of Methodology to Apply/Extrapolate NUREG-1864 Results to DSC 11-15 Revision 0 Page 44 overall low risk of release given an initiating event. The difference in risk between the two alternatives is 9.26E-13. Alternative 2 is higher in risk by a factor of 1.66. The summary of results in shown in Table 16: Table 16: Summary of ResultsAlternative 2 Alternative 1 RiskAcceptance criteriaResult 2.32E-12 1.39E-12 9.26E-13 <1E-08 "Very Small" change in risk 016045-RPT-01 Summary Revision 0 Page 45 4.0 SUMMARY 4.1 Summary of Methodology Development The methodology in NUREG-1864 [1] has been adapted to develop a simplified Probabilistic Risk Assessment for the MNGP dry cask storage system. The methodology was to: Determine the stages of operation for two proposed alternatives to evaluate the latent cancer fatality risk of five MNGP DSCs (11-15) with non-compliant weld dye penetrant examinations, Screen the initiating events that could affect the integrity of the DSCs, Estimate initiating event frequencies based on MNGP specifics, Assess the failure probability of the DSC given an initiating event, and Quantify the latent cancer fatality risk given a failed DSC for each alternative. 4.2 Summary of Methodology Application The evaluation in NUREG-1864 [1] has been applied/extrapolated to determine the risk of Alternatives 1 and 2 for MNGP DSCs 11-15. The stage evaluated for Alternative 1 is storage for 20 years; the potential initiating events applicable to these stages are aircraft impacts and meteorite strikes on any 1 of the 5 DSCs. All other initiating events are screened out as not applicable or non-risk significant. The stages applicable to Alternative 2 are to: Remove the DSC from the HSM and transport to the Fuel Building, Lift the DSC up the Equipment Hatch to the Refuel Floor, Perform weld inspections, Lift the DSC and lower it down the Equipment Hatch to the transfer trailer, Transport it back the HSM and re-insert the DSC into the HSM, and Store the DSC for 20 years. Like Alternative 1, Alternative 2 also includes aircraft impact and meteorite strike hazards during the storage stage with the DSCs in the HSMs. The initiating event unique to Alternative 2 is a potential drop while lifting the DSC during the movements in the reactor building. These stages apply to all 5 DSCs that need inspection, thus, the risk of Alternative 2 assumes all 5 are moved for inspection. Latent cancer fatality risk has been quantified for both alternatives and shown to be well below potential risk acceptance guidelines for latent cancer risk, and shown to be not significantly different between alternatives. Risk is presented for all 5 DSCs for a period of 20 years.
4.3 Conclusions In conclusion, the risk of Alternative 1 and Alternative 2 are both very small relative to the criteria in Table 5. The difference in risk between the alternatives is not significant. With regards to the welds with non-compliant PT examinations, the risk of Alternative 2 includes higher failure probabilities given a cask drop for the drops that occur prior to inspection. Alternative 1 risk, as estimated in this evaluation, is not affected by the potential presence of weld flaws because the included initiating events that can fail the DSC are assumed to fail the DSC with probability of 1.0, consistent with NUREG-1864 [1], for aircraft strikes and meteorite strikes, based on the uncertainty of and potential magnitude of such events.
016045-RPT-01 Summary Revision 0 Page 46 The magnitude of risk of either alternative is similar to the magnitude of risk of the reference site in NUREG-1864 [1], with differences attributable to the number of stages applied to the risk model for MNGP and the different frequency of the initiating events at each site. Overall, the differences are small, in the context of the total quantified risk and the risk acceptance criteria.
016045-RPT-01 References Revision 0 Page 47
5.0 REFERENCES
- 1. U.S. Nuclear Regulatory Commission. , NUREG-1864, March 2007. 2. U.S Nuclear Regulatory Commission. . NUREG/CR-4461, Revision 2, February 2007. 3. Stevenson and Associates. NEI 05-04, Rev. 3, May 2014. 4. Northern States Power Company. NSPLMI 95-001, July 2010. 5. Hughes Associates. 1SML16012.000-1, June 2013. 6. Monticello Nuclear Generating Plant, 72.212, Revision 7, June 2013. 7. Transnuclear Incorporated, Updated Final Safety Analysis Report for the Standardized NUHOMS Horizontal Modular Storage System , NUH-003, Revision 12, January 2012. 8. Applied Analysis Corporation, Accident Dose Assessment for MNGP DSCs 11-15, MNGP-018, Revision 0, June 2017. (Approval pending) 9. Office of Nuclear Regulatory Research, Office of Nuclear Material Safety and Safeguards, Office of Federal and State Materials and Environmental Management Programs, and U.S.
Nuclear Regulatory Commission, Revision 1, February 2008. 10. U.S. Nuclear Regulatory Commission, , RG 1.174, Revision 2, May 2011. 11. Monticello Nuclear Generating Plant, Application for Renewed Operating License, Appendix E -Environmental Report, 2005. 12. Monticello Nuclear Generating Plant, , USAR-05.03, Revision 34P. 13. MNGP Drawing NH-36267, Sheet 1 of 4, 14. MNGP Drawing NH-36267-3, Sheet 4 of 4, 15. MNGP Drawing NH-36807, 16. MNGP Drawing NH-36808, 17. MNGP Drawing NH-36881, 18. Not Used. 19. Not Used.
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016045-RPT-01 References Revision 0 Page 48 21. Nuclear Regulatory Commission. , NUREG-1488, April 1994. 22. Nuclear Regulatory Commission. , Information Notice 2010-18, September 2, 2010. 23. Not Used. 24. U.S. Nuclear Regulatory Commission letter to Peter Gardner (Site Vice President, Monticello Nuclear Generating Plant), Exemption from Certain Provisions of Standardized NUHOMS Dry Shielded Canister 16 at Monticello Nuclear Generating Plant Independent Spent Fuel Storage Installation (CAC No. L25058), dated June 15, 2016 (Adams Accession No. ML16167A036). 25. Flood Hazard Reevaluation Report for the Monticello Nuclear Generating Plant, dated 4/25/2016 (ADAMS Accession No. ML16145A180). 26. Independent Spent Fuel Storage Installation Fire Hazards Analysis, 72.212-A, Revision 5. 27. NUH-03-7104 Sheet 2 of 3, Roof, Revision 3. 28. Nuclear Regulatory Commission. , NUREG-0800, U.S. Nuclear Regulatory Commission, Revision 4 (ADAMS accession no. ML100331298). 29. Nuclear Regulatory Commission. NUREG-1437, U.S. Nuclear Regulatory Commission, Revision 1. 30. Nuclear Regulatory Commission. , NUREG-2239 (ADAMS Accession ML072320420). 31. Nuclear Regulatory Commission. , NUREG/CR-6525, Rev. 1. Code version 4.2. 32. NUH-03-7101 Sheet 1 of 6, General Arrangement, Revision 2 33. NUH-03-7104 Sheet 2 of 5, Base, Revision 3 34. SNFP-MT-2943-TN-0206, QF0939 - PIT #2 - Monticello Risk Informed Weld Issue - Table 1 35. SNFP-MT-2943-TN-0206, QF0939 - PIT #3 - Monticello Risk Informed Weld Issue - Cask Loading Reports 36. EC 783 Design Description, Drawing NUH61BTH-3000 Rev. 9, Drawing NUH-06-8003 Rev. 6
- 37. Drawing NF-36309 Rev. B, Procedure 9506 Rev. 17, Drawing NF-36393 Rev. B 38. Procedure 9506 Rev. 17
- 39. Drawing NF-36309 Rev. B, Procedure 9506 Rev. 17, Drawing NF-36306 Rev. 76
- 40. Drawing NUH-07-2101 Rev. 4 or NUH-06-8003 Rev. 6
- 41. Drawing NUH-07-2108 and DIT No. ISFSI-070
- 42. Drawing NUH-03-7103 Rev. 5, Drawing NUH61BTH-3000 Rev. 9, Drawing NF-201626 Rev. 0
TN Americas LLC State of Maryland County of Howard ) ) SS. ) AFFIDAVIT PURSUANT TO 10 CFR 2.390 Enclosure 1 to E-49704 I, Jayant Bondre, depose and say that I am Chief Technology Officer of TN Americas LLC, duly authorized to execute this affidavit
, and have reviewed or caused to have reviewed the information which is identified as proprietary and referenced in the paragraph immediately below. I am submitting this affidavit in conformance with the provisions of 10 CPR 2.390 of the Commission's regulations for withholding this information.
The information for which proprietary treatment is sought is listed below:
- Calculation ll 042-0400, Rev. 0, "Site-Specific Thermal Evaluation of 61BTH Type 1 DSCs Stored in HSM-H at Monticello Nuclear Generating Plant" This information has been appropriately designated as proprietary
. I have personal knowledge of the criteria and procedures utilized by TN Americas LLC in designating information as a trade secret, privileged
, or as confidential commercial or financial information
. Pursuant to the provisions of paragraph (b) ( 4) of Section 2.3 90 of the Commission's regulations, the following is furnished for consideration by the Commission in determining whether the information sought to be withheld from public disclosure
, included in the above referenced documents
, should be withheld.
- 1) The information sought to be withheld from public disclosure involves the determination ofbounding DSC shell temperatures and internal pressures during storage operations
, an analysis which is owned and has been held in confidence by TN Americas LLC. 2) The information is of a type customarily held in confidence by TN Americas LLC and not customarily disclosed to the public. TN Americas LLC has a rational basis for determining the types of information customarily held in confidence by it. 3) Public disclosure of the information is likely to cause substantial harm to the competitive position of TN Americas LLC because the information consists of thermal analyses associated with the NUHOMS 61 BTH Type 1 DSCs, the application of which provide a competitive economic advantage.
The availability of such information to competitors would enable them to modify their product to better compete with TN Americas LLC, take marketing or other actions to improve their product's position or impair the position of TN Americas LLC's product, and avoid developing similar data and analyses in support of their processes
, methods or apparatus.
Further the deponent sayeth not. ChiefTechnology Officer, TN Americas LLC Subscribed and sworn before me this 14th day of September, 2017. Notary Public My Commission Expires _jD_j j.ft_j J3__ RONDA JONES NOTARY PUBLIC STATE OF MARYLAND My Ccmmission l!xpit* OctGGir 18, 2019 Page 1 of 1