ML20235J095
| ML20235J095 | |
| Person / Time | |
|---|---|
| Site: | 07105814 |
| Issue date: | 09/18/1987 |
| From: | Engel W ENERGY, DEPT. OF |
| To: | Macdonald C NRC OFFICE OF NUCLEAR MATERIAL SAFETY & SAFEGUARDS (NMSS) |
| References | |
| 28588, NUDOCS 8710010303 | |
| Download: ML20235J095 (100) | |
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September 18, 1987 Qb%
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RECEIVED \\g Chief, Transportation Certification Branch d'
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Z SEP241987 IT Nuclear Regulatory Commission Washington, D.C.
20555 i\\ y m ue m nco m c m /
comica NRBK-4' IRRADIATED SPECIMEN SHIPPING CONTAINER - SUPP-NT M on
/4 FORWARDING-OFANDREQUESTFORRENEWALOFTHENRCCERT INFORMATION FOR THE SAFETY ANALYSIS REPORT FOR PACKAGI 't)N-COMPLIANCE Reft (a)
NRC letter FCTC:CEW 71-5814 dated April 28, 1983 (b)
NR letter Z#85-1605 dated April 1, 1985 (c)
NRC letter FCTC:WHL 71-5814 dated April 29, 1985
Background:
In reference (a), the NRC renewed the Certificate of compliance for the NRBK-41 cask for a period of two years in lieu of the usual five years.
The NRC stated that to renew the certification beyond the two years, it would be necessary to show that the lead shielding would not melt in the hypothetical fire accident of 10 CFR Part 71, or that the loss of all shielding material which did melt during the fire accident would not cause radiation levels to exceed the limits specified in 10 CFR Part 71 for accident conditions.
In reference (b), Naval Reactors requested that the NRC extend its certification of the NRBK-41 cask for an additional three years to April 30, 1988 to provide time for completion of cask modifications necessary to meet the NRC requirement.
Reference (c) granted this extension.
Discussion:
This letter submits, for NRC information and review, supplemental information to the NRBK-41 Cask Safety Analysis Report for Packaging.
This supplemental information concerns the addition of a thermal shield to the NRBK-41 cask to minimize lead melting in the hypothetical fire and the redesign of the inner container to provide assurance that a leak-tight joint can be maintained in a 30 foot drop onto an unyielding surface and the fire accident.
The following analyses are presented in Attachments A through E:
a.
Attachment A:
A thermal analysis is presented for the modified cask to determine the amount of lead melting that will occur in the hypothetical fire.
Enclosure (1) is the drawing for ew thermal shield.
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b.
Attachment B:
A shielding analysis is presented which shows that, if the' decay heat limits and the normal condition on-contact dose rate limits are met, the dose rate after the hypothetical fire accident would be less than the 10 CFR Part 71 limit.
c.
Attachment C:
A structural analysis is presented which
'shows that the new MIN-41 inner container will maintain a leak-l 1
tight. joint in the hypothetical accident and fire conditions.
Enclosures (2) through (6) are the drawings for the new MIN-41 inner container.
d.
Attachment D:
A criticality analysis is presented which shows that the current restriction for shipping unmoderated is unnecessary.
's will enable organic packing material to be used during shipmencs.
Also, the NRBK-41 is shown to meet the Fissile Class I requirements of 10 CFR Part 71.
e.
Attachment E:
A copy of the proposed revision to the DOE Certificate of Compliance is attached as Attachment E.
The analyses included with this letter are in the form of supplements to the existing NRBK-41 SARP.
NR has initiated.
efforts to compl?tely rewrite the SARP to fully upgrade it to be consistent with current SARP format and content.
Since NR
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expects this upgraded SARP to be available for NRC review in less j
than three years, NR requests that the expiration date of the NRC certificate for this cask be April, 1991 instead of the usual five years from the date of issuance.
NRC is requested to issue a revised Certificate of Compliance by December 31, 1987 to support use of the modified NRBK-41 casks.
I J. R hs W. P. Engel Naval Reactors Attachments:
A:
Thermal Analysis of a Modified NRBK-41 Cask for Normal Conditions of Transport and the Hypothetical Fire Accident B:
Shielding Analyses of a Modified NRBK-41 Cask for Normal Conditions of Transport and Following the Fire Accident C:
Structural Evaluation of the Redesigned Inner Container for Use with the NRBK-41 Cask D:
Criticality Analyses of an NRBK-41 Cask for Normal Conditions of Transport and Hypothetical Accident Conditions E:
Proposed Revision 3 of the DOE Certificate of Compliance for the NRBK-41 Shipping Cask l
___-________________A
3 NR:RM:LFPoletti-S#87-2738
Enclosures:
(1) Drawing 1755E01, Cask NRBK-41 Thermal Shield Assembly (2) Drawing 2D77456, Cask Sealed Container Assembly (3) Drawing 5076B68, Cask Sealed Container Screw Detail (4) Drawing 7593C55, Cask Sealed Container Cap Detail (5) Drawing 7593C78, Cask Sealed Container Plug Weldment Detail (6) Drawing 2D77445, Cask Sealed Container Housing i
Detail CC: (without attachments and enclosures)
Julio Torres, DP-4, DOE Manager, PNRO General Manager, Bettis Manager, shipping Container Analysis, Bettis Manager, Refueling Engineering and Operations, Bettis Manager, NRF Project, Bettis Manager, ECF Programs, Bettis Manager, IBO l
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Attachment A to Sf87-2738..
Page 1 i
THERMAL' ANALYSES OF A MODIFIED NRBK-41 CASK FOR NORMAL CONDITIONS OF TRANSPORT AND THE-HYP0THETICAL 1475'F FIRE ACCIDENT WP:RE0871096A
Attachment A to Sf87-2738 Page 2
References:
(1)
Y. S. Touloukian, "Thermophysical Properties of Matter,"
New York:
Plenum, 1970 (2)
Brown and Marco, " Introduction to Heat Transfer,"
.New York: McGraw-Hill, 1951 Introduction and NRBK-41 Cask Description A thermal analysis for a modified NRBK-41 shipping cask was performed using the TRUMP computer program. TRUMP is used to determine the cask steady state temperature distribution during normal conditions of transport cs well as the cask temperature distribution and the amount of the lead melted during a hypothetical fire transient.
A representation of the node network used for the TRUMP model of the NRBK-41 shipping cask is shown in Figure 1.
The basic cask is 39.40 inches high by 26.50 inches in diameter.
Irradiated fuel and non-fuel specimens are placed into an inner container which in turn is loaded into the 5-inch diameter by 16-inch high central cavity for shipment. A steel-encased lead plug is located above the specimen chamber. Approximately 10 inches of lead surrounds the specimen chamber on the sides and bottom of the cask for radiation shieldino The supporting structure (the cross-hatched area in Figure 1} purposes.
is fabricated from stainless steel.
The cask is being modified to include a 1/4-inch thermal barrier nounted concentrically around the cask outer diameter. The thermal barrier is represented by nodes 510-524 of Figure 1.
An air gap of 0.06 inch separates the thermal barrier inside surface from the cask.
The spacing between the thermal barrier and the cask is established and maintained by a 60 mil diameter steel wire wrapped spirally around the outside of the cask. The wire strands have a vertical pitch of 6 inches with one strand at the top and bottom of the cask as shown in Figure 2.
In TRUMP, the wire is modeled as a rectangular cross section which connects specific nodes on the thermal shield to adjacent nodes on the cask. The wire spacers are assumed to be in perfect contact with the thermal shield and the cask surface. A secondary base plate with a 1/8-inch recess is attached to the bottom of the cask.
The secondary base plate is represented by nodes 500-509 of Figure 1.
Localized deformation due to the hypothetical drop accident and puncture accident were not considered in the thermal analysis since the purpose of this assessment is to demonstrate the effectiveness of the heat shield. Thus, an axisymmetric model was used, consistent with the NRC's position as discussed in a meeting between the NRC and Naval Reactors on February 11, 1985.
TRUMP Model Table 1 presents a listing of a typical TRUMP deck used in the analysis.
In the model a 5 mil air gap separates the central fuel cavity from the steel inner liner and the top plug steel-to-steel interfaces. An air gap of 0.37 inch separates the top steel cover from the lead plug. A gap of 0.0625 inch between the outer diameter of the lead and the inside surface of the outer
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1 i
1 Attachment A to l
S#87-2738 Page 3 l
shell of the cask is present based on as-manufactured diraensions. This gap l
was not included in the TRUMP fire transient model in order to maximize the
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amount of heat transferred into the interior of the cask.
Lastly, to simulate the expansion regions in the top and bottom of the cask, nodes 308-310, 312-314, 404, 405, 409, and 410 were assumed to be voided.
l The TRUMP computer model also includes the PVC wrapping used for each cask I
shipment as well as the surfaces of the closed truck. For conservatism, a gap j
of 500 mils was assumed to separate the 12 mil PVC wrapping from the outer surface of the cask. TRUMP nodes were modeled some distance away from the cask and PVC surface to simulate a 10 ft. wide by 7 ft. high truck bed size.
The enlissivity of the PVC wrapping and truck surfaces used in the analysis were 0.87 and 0.5, respectively [ Reference (1)].
Heat transfer by radiation and conduction was considered across the 500 mil gap between the PVC and cask surface. The PVC wrapping was treated in the analysis as an opaque surface and therefore the only effect of the wrapping is to add an extra temperature drop because of the additional gap.
Natural convection is expected to occur in the enclosed space between the truck inner surface and the surface of the PVC. To estimate.the heat transfer coefficient in this space, a temperature difference of 20*F between the truck surface and the surface of the PVC was assumed.Thegggttransfercoefficientwascalculatedusingtherelationship h = 0.19 (AT) <
(Reference (2)] which yielded a' coefficient of about 0.5 BTU /hr-ft2 *F.
Use of this coefficient in the TRUMP calculations resulted in a calculated aT of 18'F which supports the 20*F assumption. Separate problems performed to determine the sensitivity of this parameter show that if the con-vective heat transfer in the enclosed space was neglected, the surface temper-ature of the cask and PVC wrapping would be increased by 7-10*F over the base case results for a decay heat rate of 1000 BTV/hr. Sensitivity studies were also performed to determine the effect of the gap size between the PVC and the cask surfaces for normal conditions of transport.
It was found that if the gap were decreased in size to 100 mils, the surface temperature of the cask would decrease by 13-16*F for a decay heat rate of 1000 BTU /hr.
Normal Conditions of Transport Analysis Paragraph 71.43(g) of 10CFR71 requires that a package "be designed, con-structed, and prepared for transport so that in still air at 38'C (100"F) and in the shade, no accessible surface of a package would have a temperature exceeding 50*C (122*F) in a non-exclusive use shipment or 82*C (180*F) in an exclusive use shipment." Using the TRUMP model previously described and the prescribed ambient conditions, cargo decay heat loads that would ensure that the specified temperature limits would not be exceeded were determined. For an exclusive use shipment, the decay heat load must be limited to 900 BTU /hr.
For a non-exclusive use shipment, a limit of 250 BTU /hr applies.
4 Figure 3 provides the temperature gradient through the NRSK-41 cask at the location of maximum axial temperature. This temperature distribution is rep-resentative of normal conditions of transport in that an ambient temperature of 100*F, the solar load specified by 10 CFR 71, and an internal decay heat rate of 1000 BTV/hr (conservatively high) were considered in determining the
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I Attachment A to Sf87-2738 Page 4 profile. A temperature.for.the cask cavity wall of 250*F is shown on Figure 3.. Because of the tight fit of the inner container within the cask cavity.
a temperature of 300*F can be assumed for the inner container and used in evaluating performance in the hypothetical drop accidents in Attachment C.
Fire Transient Analysis Only the case of an internal heat load of 900 BTU /hr, corresponding to the limit for an exclusive use shipment,.was considered in the calculations. This produced peak initial temperatures and subsequently resulted in maximizing the amount of lead melting'in the fire accident.
The fire transient is modeled in three separate parts. The first is the steady state portion followed by the cask being subjected to a 1475*F fire for i
a period of 30 minutes, and then followed by a cooldown period. The steady state calculation as discussed above provides the temperature distribution of the cask prior to the start of the fire. The fire is assumed to incorporate the effects of both radiative and convective heat transfer as specified by 10CFR71. The fire is assumed to completely engulf the cask so that the sides, top, and bottom of the cask are simultaneously subjected to the fire.
10CFR71 specifies emissivity values to be used for the fire and the package surface.
The emissivity values used in the analysis are 0,9 and 0.8 for the fire. and the cask surface, respectively. The convective heating was determined on the taasis of still ambient air at 1475"F.
Following the fire, the cask is permitted to cool down in 100*F still air. The initial conditions prior to the initiation of the fire are similar to those considered for normal conditions of transport except that no solar load is applied to the cask surfaces. This is in accordance with 10CFR71.
In the calculations, it was assumed that the PVC wrapping would immediately be removed (i.e., consumed) upon exposure to thu fire and, therefore, other than for the establishment of the initial cask temperature, the PVC is not considered in the fire transient analysis. For the fire transient analysis, a 500 mil PVC gap was used to establish the initial cask temperature gradient prior to initiation of the fire. A gap size of 500 mils will produce higher temperatures and will ultimately lead to more predicted lead melting.
The volume of melted lead is tabulated below for each TRUMP node during the fire and subsequent cooldown.
l Attachment A to S#87-2738 Page 5 Percent of Lead Melted Modified NRBK-41 Cask Wrapped With PVC and Shipped in a Closed Truck -
0 Decay = 900 BTU /HR 500 Mil Gap TRUMP
% Volume Node Melted 307 100 306 100 16 100 28 100 75 81 17 24
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311 14 The volume of melted lead that is shown in the preceding table is assumed to be lost from the cask. Therefore, the shielding calculations presented in Attachment B for Hypothetical Accident Conditions do not consider the presence of this lead.
The peak temperature of the inner container during the fire is also obtained from the TRUMP results.
The maximum temperature at any time during the fire and subsequent cooldown for nodes 110 through 114 (Figure 1) is 580*F.
This temperature is used in Section C of Attachment C in determining the pressure increase inside the sealed inner container during the fire accident.
Summary and Conclusions A thermal analysis of a modified NRBK-41 shipping cask design was performed using the TRUMP computer program for normal conditions of transport and for the hypothetical fire conditions specified in 10 CFR 71. The modifications to the NRBK-41 cask include the addition of a heat shield around the cylindrical part of the cask plus an additional plate attached to the base of the cask.
l The purpose of these cask modifications is to reduce the amount of lead melt-ing predicted to occur during the fire transient for a cask configuration in i
which no heat shield or base plate was considered. The results of the study l
indicated that the NRBK-41 modifications are successful in achieving this j
goal.
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. to S#87-2738 Page 1
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1; l
1 SHIELDING ANALYSES OF A MODIFIED NRBK-41 CASK FOR NORMAL CONDITIONS OF TRANSPORT AND FOLLOWING THE 1475'F FIRE ACCIDENT l
__u___-.____m_
Attachment B to Sf87-2738 Page 2 NRBK-41 Shielding Analysis 4
Introduction This attachment presents the shielding analysis performed on the NRBK-41 Shipping Container. This analysis shows that if the surface temperature limits for exclusive or non-exclusive use shipments and on-contact radiation limits of 10CFR71 are met for the normal transport condition, radiation levels will not exceed the radiation limits of 10CFR71 two meters from the cask for j
the normal condition and one meter from the cask following the hypothetical i
fire accident.
Discussion Cask Description and Ccmputer Model The NRBK-41 cask is a cylindrical unit with an inner cavity 16 inches high and j
radius of 2.5 inches. The sides of the cask contain 10.0 inches of lead I
shielding, with an inner steel canning of thickness 0.2815 inches, and an
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outer steel canning of thickness 0.5 inches. A cylindrical thermal shield made of steel with a thickness of 0.25 inches provides additional shielding.
Below the cavity the shield stackup consists of 1.0 inch of steel, 9.12 inches of lead and 2.0 inches of steel provided by the base plate and thermal shield. Above the cavity is the lead-filled plug which contains 9.63 inches of lead with 0.5 inches of steel canning on the bottom, and a cover plate with j
1.25 inches of steel. These and other critical dimensions were used to
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develop a model in the SPAN computer code (Reference (1)).
Figure 1 shows the j
SPAN plot of the NRBK-41 cask.
Thermal shields will be installed around the circumference and on the bottom of the cask to limit lead loss during the hypothetical 1475* fire accident.
The volume of lead lost during the fire accident, as determined by the thermal analysis, was incorporated into the shield model at the appropriate locations.
Figure 2 shows the NRBK-41 cask with the lead loss. The majority j
of the lead loss is from the upper corner of the cask. A small amount of lead is also lost along the side of the cask.
l Gamma radiation levels were calculated using buildup factors for a material J
with an atomic number equal to 82 for the normal cask configuration since this is equivalent for the iron / lead / iron material array through the cask.
Buildup factors for a material with an atomic number equal to 74 were used for the accident radiation levels which were calculated through the upper corner.
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Shipping Requirements Shipments made in the NRBK-41 cask must meet the following requirements:
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Attachment B to S#87-2738 Page 3 a.
Radiation Levels 1.
During normal shipments, the radiation level on contact with the cask can not exceed 200 mrem / hour; at 2 meters from the cask surface, the levels can not exceed 10 mrem / hour.
2.
Following an accident, the radiation levels 1 meter from the cask surface can not exceed 1000 mrem / hour.
b.
Surface Temperature Limits:
1.
For an exclusive use shipment, the surface temperature can not exceed 180'F. For 'the NRBVs-41 cask, this means that the cargo cannot release more than 900 BTU / hour.
2.
For a non-exclusive use shipment the surface temperature can not exceed 122*F. For the NRBK-41 cask, this means that the cargo cannot release more than 250 BTU / hour.
Source Model l
The NRBK-41 cask is used for shipping various irradiated test samples, fuel and non-fuel.
Investigation into the types of materials that may be shipped i
in the casks resulted in the identification of over 30 elements and numerous alloys. Also, since shipments involve samples which are tested prior to use in the NR program, it is possible that materials not currently planned for use may be tested, and shipped, within the next few years. Therefore, it is not feasible to analyze the materials individually.
Samples shipped in the NRBK-41 casks include activated materials and irradiated fuel. Shipments made in the casks must meet radiation limits of 10CFR71 and thermai limits for exclusive or non-exclusive use shipments.
To determine the criterion necessary to meet both radiation and thermal limits without analyzing specific materials, various gamma energies and neutron radiations were evaluated separately.
1.
Gamma Radiation Levels Radiation levels were calculated on contact with and 2 meters away from the cask for the normal shield configuration and 1 meter away for the hypothetical fire accident using various discrete gamma energies.
The gamma energies analyzed are those energy groups which are representative of activation and fission product radiations and are listed in Table 1.
The thermal limit on the surface temperature of the cask imposes a 900 BTU /hr and 250 BTU /hr heat generation limit on the cargo for the exclusive and non-exclusive use shipments, respectively.
The heat generation of the cargo was calculated by converting the gamma energy (assuming 1 gamma emission per disintegration) to BTU's per hour. The source strength was limited to produce 900 BTU /hr for each gamma energy. For those cases 1
1
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Attachment B to Sf87-2738 Page 4 where this source strength resulted in surface radiation levels which exceeded 200 mrem /hr, the source strength was reduced to produce 200 mrem /hr.
This occurred for gamma energies of 1.75 Mev or greater. For gamma energies of 1.55 Mev or less, the heat generation limit of 900 BTU /hr established the source strength. The same source strength was then used to calculate radiation levels at 2 meters (normal shipment) and one meter (accident configuration).
o 3
In the cases where the on-contact radiation levels were limiting, no 1.3 design uncertainty factor was applied to the on-contact calculation. This maximizes the source with respect to tiie on-contact radiation level to produce a maximum radiation level for the accident condition and at 2 meters.
The 1.3 factor was applied to the radiation levels calculated at 2 meters.
Radiation levels for the normal transport condition were calculated with the source concentrated in the center of the cask cavity and the source volume modeled as a material with one-half the density of stainless steel. The source volume is a. cylinder 6 inches high and radius of 1.5 inches.
Investigation into typical cargos showed that, in general, the containers are not filled to capacity and consist of a few test samples contained in holders or other packing material.
The test samples are typically about 6 inches by 1 to 2 inches, with a thickness of about one quarter inch. This source model was developed to produce realistic radiation levels for the normal transport condition.
Then, after determining the source strength required to produce 200 mrem /hr on contact with the side of the cask, this same source strength was used to calculate s3, all other radiation levels.
Radiation levels calculated at the top and bottom of the cask are about 15% and 40%, respectively, of those at the cask side.
Radiation levels along the side of the cask during the accident are not significantly different from the normal condition due to the small amount of lead lost. As seen in Figure 2, most of the lead is lost from the upper corner of the cask; however, a significant amount of shielding still remains.
Radiation levels through the upper corner were maximized by concentrating the source in a one-inch cube in the upper corner of the cask cavity and assuming no self shielding. As shown in Table 1, all levels remain below the 1000 mrem /hr limit of 10CFR71 at 1 meter from the cask.
Two other source configurations were evaluated.
For one model, the source was distributed evenly throughout the cask volume with varying degrees of self shielding. The other was similar to the model used for the accident condition in that the source was concentrated in a one-inch cube directly opposite the detector point.
No self shielding was assumed in this case. For both models a 200 mrem /hr on-contact dose rate resulted in radiation levels that were less than 10 mrem /hr 2 meters from the cask.
The radiation limits for the accident condition, using the same corner source model described above, also met the 1000 mrem /hr limits.
u_____.___________..___-____.
Attachment B to S#87-2738 Page 5 2.
Neutron Radiation Levels A comparative analysis was performed to determine acceptable neutron.
radiation levels. The cask model discussed earlier was converted for use with the SPIC option of SPAN.
Neutron radiation levels were then calculated at the various locations around the cask for the normal and accident cask configurations to determine levels relative to the on-contact level.
Irradiated fuel samples are expected to be the only cargo shipped in the NRBK-41 cask to produce neutron radiatien.
The source configurations for the normal and accident conditions are the same as those described above for the gamma radiation calculations.
The realistic model was used to calculate normal transport levels and normalized to 200 mrem /hr on contact.
The same source strength was then concentrated in the upper corner to maximize the accident radiation levels. As was the case for the gamma radiations, the neutron levels 2 meters from the cask for the normal condition and at 1 meter from the cask for the accident condition are below the radiation limits of 10CRF71.
Results Table 1 shows radiation levels calculated for the various gamma energies on contact with the cask side, 2 meters from the cask side and, for the hypothetical accident, 1 meter from the cask corner. For gamma energies of I
1.55 Mev or less, the source strength is limited by the heat generated by the cargo.
The radiation limits presented for gamma energies of 1.55 Mev or less are for a shipment that generates 900 BTU /hr, which would therefore be the maximum levels expected from this type of cargo.. A 1.3 design uncertainty factor was applied to the normal transport radiation levels for these radiations.
Radiation levels for the non-exclusive use shipment would be about 1/3 of these levels.
Shipments of materials with gamma energies of 1.75 Mev or greater are limited by radiation levels on contact with the cask side. The source strength that produces a radiation level of 200 mrem /hr on contact with the cask side was then used to conservatively calculate the levels for the accident condition.
Radiation levels for the accident condition were maximized by concentrating I
the source directly opposite the detector point in the area of the most reduced shielding (i.e., upper corner). Due to the conservative method used to calculate the accident radiation levels, no 1.3 design uncertainty factor was applied. The 1.3 factor was applied to the levels calculated at 2 meters from the cask for the normal transport.
Neutron radiation levels are also provided in Table 1.
Neutron levels were calculated in a manner similar to that discussed above for gamma radiations of 1.75 Mev or greater.
I l
Attachment B to Sf87-2738 Page 6 This analysis shows that if the thermal limits (in terms of BTU /hr generated by the cargo) and the surface radiation limit are met, then the radiation limits 2 meters from the cask for the normal transport and 1 meter from the cask for the hypothetical accident condition will also be met.
Since neutror.
radiations and each single gamma energy meets this criterion, a mixture of gamma and neutron radiations will also meet the criterion, remembering that noth thermal limits and surface radiation limits must be met.
Conservatism The shield model used to calculate normal transport radiation levels was developed to provide realistic results. The source model approximates a bundle of test specimens, and a buildup factor with an equivalent atomic number of 82 was used. Also, for those cases where the limiting requirement was the on-contact radiation level (i.e., for gamma energies of 1.75 Mev and greater), no 1.3 design uncertainty factor was applied to the radiation level. By not maximizing radiation levels for the normal condition, a larger source strength was established for the accident condition.
Using this source strength, the radiation level for the hypothetical fire accident was maximized by concentrating the entire source in a 1-inch cube in the upper corner directly opposite the detector point. This source model produces levels that are about 5 times higher than those calculated for the source model used in the normal transport configuration. A buildup factor with an equivalent atomic number of 74 was used in these calculations since the shield configuration consists of less lead and more steel than the shield array through the cask side.
No 1.3 design uncertainty factor was applied to the calculated radiation levels for the hypothetical accident. All radiation levels were below the radiation limits of 10CFR71.
Heat input from beta radiation was not included in this evaluation.
Since beta radiation does not contribute to the radiation levels, this is conservative for calculating expected radiation levels where the heat input limits the cargo, and does not affect the results where radiation levels are limiting.
For neutron radiation calculations, the SPIC option of the SPAN computer code was used.
The same source configurations discussed for gamma calculations were used, which maximizes the source strength for calculating the accident radiation levels. The neutron radiation levels also meet all radiation levels specified in 10CFR71.
References 1.
" SPAN 4: A Point Kernel Computer Program for Shielding", WAPD-TM-809(6),
j
- 0. J. Wallace, Volumes 1 and 2, October 1972 i
)
)
Attachment B to S#87-2738 Page 7 1
se m nattaint ecv l
]
j mio uno g'j
{
, s. w
< cnn,ns >
c,ce, neu,,c.u Q
sn manu w. e, d%k k \\%
9 3"
25" thermal shield 16.0" l
10.0" lead
!*- 5.
i
/
9.12" Inh \\k\\1%'\\\\
EN\\\\M 2 1 erm i
d i
1 Tigure 1: SPAN Model of NRBK-41 Cask
Attachment B to S#87-2738 l
Page 8
$PaH5 twaf EMift KEY
]
l l
voio me Q
y
.... 28.
< cannas, canoou arrou e rr l
F i
k e
m
)
2.5" x 1.75" l
s l
Y 6.0" x 0.21" l
l l
1 f4.0"x0.1" xxxxxxxxxxxxxxxxxsxxxxxsxsxxxsxssxxxxxsxxxxxww l
Figure 2: SPAN Model of NRBK-41 Cask Following 1475* Fire Accident L_______________
_ _ _ _ _ _ _ _ _ _ _ - to.
Sf87-2738 J
Page 9 1
Table 1
]
Radiation t.evels (mrem /hr)
Normal Transport (cask side)
Accident Condition-Energy _
On Contact 2 Meters 1 Meter From Top Corner) 0.301-
.11 x 10 '
.46 x 10 "'
.9 x 10 **~
0.632
.45 x 10 '
.13 x 10 '
.21 x 10
- j 0.801
.21 x 10 *
.57 x 10 "
.35 1
1.101 1.6'
.04 23 1.252 13.8
.36 95 1.551 118 3.0 392 i
1.752' 200 6.5 671 1.992 200.
6.5 537 2.202 200
- 6. 5 490 2.382 200 6.5 455 2.752 200 6.5 428 2.802 200 6.5 429 3.52 200 6.5 436 4.502 200 6.5 471 5.502 200 6.5 496 6.752 200 6.5 560 neutron 2 200 6.0 15.2 10CFR71 limit 200 10 1000 1.
Limited by thermal limits.
Levels calculated for an exclusive use shipments. Radiation levels for normal transport include a 1.3 design
' uncertainty factor.
2.
Limited by radiation levels. The on contact radiation level does not include a 1.3 design uncertainty factor to maximize source for other calculations. The radiation levels at 2 meters from the cask for the normal condition were calculated using the 1.3 factor.
3.
Radiation levels for the accident condition do not include a 1.3 design uncertainty factor.
l Attachment C to Sf87-2738 Page 1 STRUCTURAL EVALUATION OF THE REDESIGNED INNER CONTAINER FOR USE WITH THE NRBK-41 CASK i
i
Attachment C to Sf87-2738 Page 2 INTRODUCTION This attachment provides analyses which demonstrate the adequacy of the modified inner container design for use with the NRBK-41 cask. The modified inner container.is shown on W 2077456.
The modified inner container incorporates a metallic c-ring into the redesigned closure joint to improve the containment capability of the inner container during any of the hypothetical accidents specified by 10CFR71. The closure joint region of the inner container is also shown in Figure 6 of this attachment. As shown on }{ 2D77456, teeth on the plug engage teeth on the housing by means of a breech lock arrangement which is secured when the eight 1/2-20 UNF2A screws are tightened. Tightening the screws also forces the cap down, compressing the c-ring.
Initial leaktightness will be confirmed at i
assembly by verification leak testing as required by the Certificate of Compliance (C0C) for the NRBK-41 cask (refer to Attachment E of this letter.) The inner container will also be tested for leaktightness within one year prior to use to an increased sensitivity. This is also required by the C0C.
Since the Inconel X-750 c-ring can withstand temperatures up to 1800'F, even direct exposure to a 1475'F fire as defined by 10CFR71 would not be detri-mental to the integrity of the c-ring.
In addition, high radiation environ-ments have no effect on the c-ring.
Therefore, the only concerns are:
(1) the effect of a 30-foot drop of the outer NRBK-41 cask (Figure 1) on the inner container closure joint region shown in Figure 6, and (2) the ability of the inner container to withstand the pressure increase associated with exposure of the NRBK-41 cask to the 1475'F fire. Assessments of the effect of a 30-foot drop for all significant drop orientations and of the effect of the pressure increase due to exposure to the fire follow. The analyses provided in this attachment demonstrate that the inner container would continue to provide containment if the NRBK-41 cask were subjected to a 30-foot drop or a 1475'F fire.
Because of the support offered by the NRBK-41 cask cavity walls and the relatively close fit of the inner container within the cask cavity, the inner container could be subjected to lateral forces, axial forces, or both in a 30-foot drop. A corner impact of the inner container itself on a surface not parallel to a face of the inner container is not possible. Axial loadings of the inner container will be examined first.
httachment C to S?87-2738 Page 3 Ig COVER PROTECTION PLATE 3
/
?
I l
1 Q
t 27.16In, r%-
1 a$
)
w J
'ic lf ll 39.45 lN.
ll I
(
m
+.
A x
1 yg,[ j 6
k(
~
A.i -
- <. ys m
l Figure 1 - Modified NRBK-41 Cask i
Attachment C to S#87-2738 Page 4 A.
Axial loading of the Inner Container Thirty-foot free drops against the top, top and bottom corners, and bottom surfaces of the NRBK-41 cask will be examined. The peak cask decelera-tions developed during these impacts will be assumed to be applied to the inner container with no attenuation. Therefore, in all evaluations, all components will be considered to remain rigid and a static evaluation of the closure joint of the inner container will be performed.
The acceptance criteria that is used is that stresses in the joint members will not exceed the static yield strengths of the materials.
1.
Energy Absorption Thirty-foot free drops against the top, top and bottom corners, and bottom surfaces of the NRBK-41 cask will be evaluated since the worst axial loading is not obvious by inspection. Deceleration loadings developed during the drops and imparted to the nested inner container are determined using the incremental energy absorption technique presented below.
This technique has been used previously in SARPs for the S3G Core Basket Disposal Container (Reference (11)), the A1W-3 f
Holddown/ Support Barrel and Shroud Disposal Container (Reference (3)),
and the AlW-3 Core Barrel and Thermal Shield Disposal Container (Reference (2)).
In the latter SARP, this method is compared with IMPAC2, a lumped-mass, nonlinear spring computer program described in Reference (10), to verify the conservatism of its results relative to those of IMPAC2. The incremental energy absorption technique can be applied to each type of drop for the NRBK-41 cask by varying the method of predicting redistribution of the crushed material.
(a)
The following steps are performed for each energy absorption increment:
1.
Choose an initial crush increment to be used in solving for j
the final crush depth using the iterative incremental energy 1
absorption process. The initial increment should be chosen so that at least 20 energy absorption iterations will be required to completely absorb the package energy.
j i
2.
Determine the initial energy (E ) of the package j
g=fmv$
2 E
= Wh where m = mass of package
Attachment C to Sf87-2738 Page 5 vj = velocity at initial impact A9l 9I
=
W = weight of package h = drop height 2
g = acceleration due to gravity (386.4 in/sec )
Steps 1 and 2 determine initial conditions for the package at impact. The following steps are used during each energy absorption increment to calculate parameters for that increment.
3.
Determine the strain rate (c) by dividing the remaining veloc-ity (initially taken from Step 2 and subsequently from Step 9) by the remaining length of the section subjected to crush.
4 (or v )
v R
C "
L-d where vg or vR = velocity at beginning of increment (initi-ally v.
from step 2, subsequently vR from step 9?
L = original length of the section subjected to crush d = cumulative crush depth.
4.
Determine total strain (c) at this increment, and compare this total strain with the ultimate strain (c )*
u e=f,andc<cu If c is less than cu, continue the incremental process.
If e is greater than cu, the section fails in compression.
5.
Determine the cross-sectional area (A ) representing the c
redistributed material arrangement as a function of crush depth. Constant material volume is assumed, and an expres-sion is derived for A based on the geometry of the section subjected to crush. kppropriateequationsforA are c
developed for applicable container regions as crush calculations are performed.
6.
Determine the flow stress based on the strain (see Step 4) and strain rate (see Step 3) during the crush increment.
The following series of equations is used.
)
Attachment C to Sf87-2738 Page 6 4
For low strength (o s 80,000 psi) steel (using equations for best fit u
curves where the curves are taken from Reference (1))
The dynamic ultimate strength (o,dyn) is determined using the u
following equation which will be identified as Equation 1:
1.0608 + 0.00024596 in i
~
o,dyn = o stat x
_- 0.0014241(1n i)2 + 0.0016566(in i)3_
u u
The dynamic yield strength (o,dyn) is determined using the y
following equation identified as Equation 2:
1.3932 + 0.095498 in i
~
i Y
Y
+ 0.0088759(in i)2 + 0.00022646(in i)3 For intermediate to high strength (o 2 80,000 psi) steel (using u
equations for best fit curves where the curves are taken from Reference (1))
The dynamic ultimate strength ( u,dyn) is determined using the following equation identified as Equation 3:
~1.08099 + 0.01369745 in i u,dyn = o stat x
_- 0.00005576917(in i)2 - 0.000113955(in i)3 u
i The dynamic yield strength (o#a,dyn) is determined using the following equation identified s Equation 4:
"1.04475 - 0.003325876 in i c dyn = o. stat x
_, 0.0038075(in i)2 + 0.0007212455(ln i)3_
y y
+
l l
l l
L 1
________________-__---_J
Attachment C to S#87-2738 Page 7 Then, for cases where o,dyn > o,dyn, the dynamic flow stress (oFLOW) u y
is determined by:
c-c
" FLOW " y,dyn +
Y(o,dyn-c,dyn) u y
,u, y
- o,dyn +
(d "y,dyn
[o,dyn-o,dyn) u y
)
y g-E o,dyn y
.k Su E
/
~
and, for cases where "u.dyn < o dyn, the dynamic flow stress, FLOW =
y "u.dyn.
The equation for oFLOW for cases where o,dyn > o,dyn, evolves from u
y assuming the. stress-strain curve of Figure 2.
.,,,,,n l
a gow r
y,,,on l
r l
l_
E I
i 1
i l
l
.,( f")
.(g)
STRAIN Figure 2 - Bilinear Stress-Strain Curve 7.
Determine the package deceleration for the crush increment:
F xA c
FLOW c
Ng=p=
y where Fc = crush force.
Attachment C to Sf87-2738-3 Page 8 j
1 8.
Determine the energy absorbed during the crush increment and the total energy remaining after the crush increment:
E = W x ad x N c
g l
ER=Ej - IE c
where Ec = energy absorbed during the crush increment ad = additional crush depth per increment ER = total remaining energy IEc = cumulative energy absorbed during preceding crush increments.
9.
Determine instantaneous package velocity based on remaining package energy:
i R"/2xgxE R l
v W
10.
Determine the crush increment for the next iteration.
If the energy absorbed during the preceding crush increment is greater than the remaining energy, reestablish the crush increment at 1/10 the previous value, and then evaluate the next increment beginning at Step 3.
If the energy absorbed is less than the remaining energy, use the same crush increment and proceed to Step 3 to continue the energy absorption iteration.
The iterative process is halted when the new crush increment (see Step 10) is less than the specified accuracy. This assumes that the process was not halted in Step 4 due to totalstrain(c)exceedingultimatestrain(c)*
u 2.
Top Drop Analysis In this section, the maximum deceleration developed when the NRBK-41 cask is inverted and impacts an unyielding surface in a 30-foot drop is determined.
In Section A.5 of this attachment, the maximum deceleration developed in a flat top or top corner drop of the cask is applied to the inner container and stresses are calculated and compared to the appropriate material's yield strength.
7 l
'l Attachment C to S#87-2738
{
Page 9 For conservatism, in order that the deceleration and thus the load on the inner container will be maximized, all of the kinetic energy of the package will be considered to be absorbed by crushing the surface of the cask which impacts the unyielding surface. This surface, the coverprotectionplatewhichisitem9ofBagte11eMemorialInstitute drawing 41-0002, has a face area of 263.2 in.
It is identified on Figure 1.
In addition, the cask deceleration will be applied to the inner container without attenuation.
This scenario is conservative since absorption of portions of the total energy by deformation of the cask shell due to axial compressive loading and circumferential strain resulting from lead rearrangement is not considered. Also, the cask cover assembly is treated as a rigid member in a static situation which neglects the attenuating benefits of the lead-filled cover assembly design.
Material redistribution is represented, based on constant volume, as shown in Figure 3.
n L
DT 1r
, E. W TQ yo n
n Figure 3 - Top Drop Material Redistribution A0 -- original, contact area (area of cover protection plate)
= 263.2 in DT -- cumulative crush depth MRD -- depth over which material is redistributed
= 2.5 in. (combined thickness of cover protection plate, cover plate, and cask upper weldment topplate)
Attachment C to S#87-2738 Page 10 A
-- face area after crushing c
Since volume is constant, (A ) (MRD-DT) = (AO) (MRD) and c
~i MRD OT
=A0(MRD-DT)=A0((MRD-DT)+I}
Ac The remainder of the variables for this orientation are as follows:
E
-- Young's modulus for the 304L stainless steel cover protection plate material (from the NRBK-41 Cask SARP, Reference 12))
= 28.0 x 10,(psi; o, stat -- compressive yield strength for the 304L stainless y
steel material (from the SARP)
= 35,000 psi;
- u. stat -- ultimate strength for the 304L stainless steel material (fromtheSARP)
= 75,000 psi; u -- ultimate compressive strain c
= 0.63 in/in (Note: Use of this value is based on the justification provided in Reference (2), where it is shown that tests on an assortment of materials of various strengths lead to the conclusion that an ultimate compressive strain of at least 0.63 in/in is appropriate for most steels. Based on strength and ductility considerations, this strain value can also be applied to the 304L stainless steel material);
W -- cask loaded weight
= 9010 pounds; h -- cask drop height
= 360 in; DINC -- initial crush increment used
= 0.01 in; L -- total length of the package (Figure 3)
= 40.4 in;
Attachment C to j
Sf87-2738 i
Page 11 I
I For this postulated situation, the cover protection plate would I
be crushed approximately 0.21 in. The associated peak g-load would be 1768.5 g's based on the conservative approach used.
3.
Bottom Drop Analysis In this section, the maximum deceleration developed when the NRBK-41 I
cask is upright and impacts an unyielding surface in a 30-foot drop is
' determined.
In Section A.5 of this attachment, the maximum deceleration developed in a flat bottom or bottom corner drop of the cask is applied to the inner container and stresses are calculated and compared to the appropriate material's yield strength.
{
In a bottom drop orientation, the kinetic energy of the NR8K-41 cask package will be absorbed primarily by crushing the skid assembly.
For this evaluation, to maximize the g-loading, all the energy will be assumed to be dissipated by skid deformation.
Material redistribution is again represented basically as shown in Figure 3 although the skid cross-section differs significantly from thatofthecgverprotectionplate. The cross-sectional area of the skid is 83 in.
The skid material is ASTM A-36 carbon steel so material properties differ from those used in the top drop evaluation. The following apply (except as indicated, variables are as defined in Section A.2 of this attachment):
2 83 in ;
A0
=
L total length (height) of the skid webs 4.75 in;
=
MRD = L 6
28 x 10 ps$.
E
=
o stat = 36,000 psi; y
"u. stat = 80,000 psi (maximum value from the applicable ASTM specification to maximize the deceleration i
developed);
0.63 in/in; c
=
u 9010 pounds; W
=
f l
l
Attachment C to Sf87-2738 Page 12 h
360 inches;
=
0.02 in.
DINC
=
The results indicate that the skid webs would crush approximately 0.5 inches. The peak deceleration would be 734.5 g's.
4.
Corner Drop Analysis i
For this ant,1ysis, the package is oriented with its centerline inclined with respect to vertical, and.the impact, resulting from a 30-foot free fall, is against a corner of the NRBK-41 cask.
If the package lards with its center of gravity positioned directly above the corner, the accident is termed a stable corner impact, and the total free fall energy of the package is absorbed in the plastic deformation of the corr.er.
If the package center of gravity is misaligned with respect to a vertical line extending up from the corner at impact, the accident is referred to as an oblique corner impact.
For oblique impacts, the misalignment of the corner crushing force and the inertia force acting through the center of gravity (CG) creates an unbalanced moment that tends to rotate the package onto its side or end surface, depending upon the position of the center of gravity relative to the corner impact. As a result, a portion of the total package energy goes into the plastic deformation of the corner and the remainder is converted into rotational energy that is subsequently expended in crushing the end or side surfaces of the cask.
At impact, the NRBK-41 cask is treated as if it were pinned at the corner of impact to the unyielding contact surface.
It is assumed that the container is free to rotate or be displaced vertically, but that it cannot be translated in the horizontal direction.
The energy available for crush is equal to the kinetic energy of the i
package minus a portion of the energy which is converted into rotational energy. The energy available for crush is found using the following expression (Appendix 2.10.4, Section 3 of Reference (13)):
WP cos' y + I p
A yp2,y o
Variables are defined as follows:
Up -- energy available for crush
Attachment C to Sf87-2738 Page 13 UA -- kinetic energy resulting from a 30-foot drop W -- package weight P -- distance from point of impact to package CG (See Figure 4) y -- oblique drop angle (see Figure 4)
Io -- weight moment of inertia about the CG for the package STABLE CORNER OROP OBLIQUE DROP
/
N
/
s Y
,' Al s
/
/
's
/
f I
/
HUU/HUUU/
Figure 4 - Oblique Corner Impact I
Attachment C to S#87-2738 Page 14 For conservatism, energy absorbed in bending a cask trunnion, deformation of the skid, or bending of the protruding portion of the base plates will be neglected when evaluating the appropriate drop orientations.
Therefore, the remaining energy, V will be dissipated entirelybytheplasticdeformationoftheappropEiatecaskcorner.
The volume of material that is redistributed during crush while absorbing this remaining energy can be represented by an ungula as shown in Figure 5 (for cases where b is less than R).
B i
i N
s s'
~~~~~'~'
y CRUSH (SLANT FACE) AREA
\\
c/2
/
\\
'b c/2 1%.~
Figure 5 - Volume of Material Redistributed During Crush l
The volume of material redistributed at each crush increment can be l
determined and a crush area can be determined at each crush increment.
The volume of redistributed material as a function of crush depth (represented by dimension d in Figure 5) can be found from the following equation (from Reference (4)):
________.__m__..___
Attachment C to S#87-2738 Page 15 l
Vol -- volume of redistributed material
= g $,
[(j)(j)3 -(e)(B)}
where H = co I
e d
e = R cos a I
- 1/2 2
c/2 = / R2 _,2 2
=
I (cosa)2 esa B -- area of the base (a segment of a circle) 2
=h[2+-sin 2+];
and
-1
({) = cos-1 (1 R c s a)*
+ = cos It can be seen that equations for H e, c/2, B, and + are all in terms of known values R, d, a, and 8.
Therefore, these equations can be solved and the results substituted into the equation for'"Vol" to determine the volume of the displaced ungula, the crush volume, as a functionofcrushdepth(d).
Similarly, the crush area, which is the slant face area of the ungula (Figure 5) can be found from:
A
-- crush area c
2
-I({}-ec R cos Cos a for cases where (co a)isgreaterthanRindicatingthattheungula l
extends past the center of the base:
e is redefined as co a i
Attachment C to S#87-2738 Page 16 2
2 2
Vol=tana[(")(R)(e)+(e)(l)-(e)(R)(sin ~I(-f))+
(f)(f)3);
= (n) (R ) - (R )(cos-1 )+(e)(f) 2 2
A Cos 6 Forcaseswhere(ggf,)isgreaterthan2Rindicatingthatthecontact surface includes the entire container base:
2 Vol = 3R d -xR tan 8; 3
2 mR C
Cos 8
- Using the preceding equations (as applicable) for crush volume and crush area, the same basic approach and equations presented in Section A.1 of this attachment can be used to evaluate incremental energy absorption and resulting deformation of the container in a corner drop. However, due to the irregular cross-section of the redistributed material, a slight variation in some of the equations is required:
(1)
Material strain:
(c) = crush volume at crush depth (d),
crush volume at total depth where total depth is defined as the minimum cask body thickness normal to the impact surface.
(2)
Time per crush increment:
(at) =
crush increment (ad) velocity at beginning of crush increment (v ) '
j (i)=h ;
(3)
Strain rate:
(4)
Energy absorbed during a crush increment:
Ec " ( FLOW) (a Vol) where a Vol = volume of material displaced during a crush increment.
l j
i 1
Attachment C to Sf87-2738 Page 17 a) Top Corner Drops In this section, the maximum decelerations developed when the NR8K-41 cask is inverted and impacts an unyielding surface in a 30-foot drop in orientations such that a top corner of the cask is contacted are determined.
In Section A.5 of this attachment, the maximum deceleration developed in a top corner or flat top drop of the cask is applied to the inner container and stresses are calculated and compared to the appropriate material's yield strength.
For top corner drops of the NR8K-41 cask, the following apply:
CG distance from the package CG location along the container central axis to the package surface that will be contacting the unyielding surface (in this case the top) 24.29 in;
=
R modified cask outer radius 13.59 in;
=
h
-- drop height 360 in;
=
W weight of the package 9010 pounds;
=
I packagewegghtmogentofinertia o
1.566 x 10 lb-in ;
=
a, stat
-- compressive yield strength of 304L stainless steel y
material 35,000 psi;
=
- u. stat ultimate strength of 304L stainless steel material 75,000 psi;
=
y ultimate compressive strain (See Section A.2 c
of this attachment for justification of the use of this value) 0.63 in/in;
=
E -- Young'smgdulusfor304Lstainlesssteelmaterial 28.0 x 10 psi;
=
DINC --
Initial increment of crush depth 0.03 in.
=
l l
Attachment C to Si87-2738 L
Page 18 L
Two additional items of cask geometry need to be defined to. establish the minimum cask body thickness normal to the impact surface. The total cask wall thickness (WTH) is 1.0 inch (the cask is 0.75 in. in the upper region of the cask and the heat shield is 0.25 in.).
The thickness of the top portion (LTH) varies depending on the circumferential location of the top surface chosen. Adjacent to a trunnion, the thickness is at least 2.50 inches; in the remaining regions, the thickness is only 2.0 inches.
This includes the combined thickness of the cover top plate flange (0.75 in.) and the cover protection plate flange (0.50in.). The smaller value (2.00 in.) produces a higher g-load since, for the same package weight and equivalent crush increments, strain, strain rate, and therefore flow I
stress are larger. As a result, the smaller thickness will be used.
For shallow drop angles (small values of w - see Figure 4),
corner crushing would occur as predicted by the preceding set of equations. However, for larger angles (w > 30' for the angles considered) ultimate strain is exceeded before all the energy is absorbed.
This indicates that the remaining energy would have to be absorbed by deformation of some other region of the cask such as the lead filler, and the g-load developed would not be maximized..To avert this possibility, thicker walls were represented in the analysis-for w > 30*.
As can be seen from the results, this proves to be inconsequential since the maximum g-loads are developed for the shallowest drop angles.
Nevertheless, for w > 30*, WTH = 1.50 in.
The results of the top corner drop analysis are shown in the following table:
Wall Cumulative I
Orop angle (w),
thickness, crush Maximum degrees in.
depth, in.
g-load 5
1.0 0.70 1071.6 10 1.0 1.04 764.2 20 1.0 1.52 584.5 l
30 1.0 1.81 515.4 1
40 1.5 1.99 454.4 50 1.5 1.99 423.2 i
60 1.5 1.84 400.5 70 1.5 1.56 386.9 80 1.5 1.13 409.9 85 1.5 0.83 472.2
)
Attachment C to Sf87-2738 l
Page 19 l
l The axial components of the g-loads in the preceding table would be somewhat less than the values shown in the table.
Disregarding this fact and considering the values in the l
table as the axial g-loads, the maximum g-load in a top corner drop accident, for a drop angle of 5*, is 1071.6 g's.
b)
Bottom Corner Drops In this section, the maximum decelerations developed when the NRBK-41 cask is upright and impacts an unyielding surface in a 30-foot drop in orientations such that a bottom corner of the cask is contacted are determined.
In Section A.5 of this attachment, the maximum deceleration developed in a bottom corner or flat bottom drop of the cask is applied to the inner container and stresses are calculated and compared to the appropriate material's yield strength.
Variables for bottom corner drops are the same as for top corner drops (Section A.4.a. of this attachment) with the following exceptions:
CG 22.39 in.
=
WTH = 0.75 in. (cask is 0.5 in, and heat shield is 0.25in.)
LTH = 2.00 in. (regardless of circumferential location)
Similar to the top drop evaluation, for shallow drop angles, bottom corner crushing occurs as predicted by the method presented in Section A.4.
However, for larger angles (in this case, w > 20' for the angles considered) ultimate strain is exceeded before all the energy is absorbed. To maximize the g-load for w > 20* thicker walls (WTH=1.50 in.) were represented. Again, the greatest g-loads are developed for the shallow angles so this alteration of the wall thickness has no effect on the principal results.
I g,.
Attachment C to S#87-2738 Page 20
~.
The results of'the bottom corner drop analyses are shown in the following table:
n; Wall:
Cumulative Orop angle (w),
thickness, crush Maximum degrees in, depth in.
q-load 5
0.75 0.69 1065.7
-10 0.75 1.03 763.4 20 0.75 1.50 589.1 30 1.50 1.84 498.9 40 1.50 2.00 457.1-50 1.50 2.00
'428.6 60 1.50 1.86 411.5 70 1.50 1.58 402.3 80-1.50 1.16 431.6' 85 1.50 0.86 503.3 Again, these values can be considered in lieu of their axial components for conservatism. Therefore, the maximum g-load l
in a bottom corner drop accident, for a drop angle of 5*, is 1065.7 g's.
l l
4
_---___-__---._,x
Attachment C to Sf87-2738 Page 21 5.
Loads Imparted to Critical Joint Members and Stresses Developed Examination of Figure 6 reveals that the critical areas in the closure region in an axial impact of the inner container are the bearing area in the container housing where the cap seats after the c-ring is compressed, the eight screws, and the internal threads in the plug.
The following calculations are provided to determine the maximum stresses in these joint members based on the decelerations developed in Sections A.2 through A.4.
Yield strength is chosen as the stress limit to ensure that there would be no permanent joint deformation in a drop accident.
In the analyses, material properties for the inner container are degraded to reflect a temperature of 300*F (this temperature is developed in Attachment A).
SCREW Zh V
- 7 j
PLUG r
r-I I
/
HOUSING g
h I
/- CAP
\\
>N
^^R Y
C-RING Figure 6 - Inner Container Closure Joint
i l
1 Attachment C to l
Sf87-2738 1
Page 22 i
(a)
Evaluation of Loading on Inner Container Housing Due to Bottom and Bottom Corner Drops j
As can be seen in Figure 6, during assembly the eight screws force the cap down to compress the c-ring. After the c-ring is adequately compressed, the cap bottoms on a ledge within the i
housing.
In a bottom or bottom corner drop accident, the total I
peak bearing load on the cap and on the housing ledge is con-servatively assumed to equal the preload plus the combined inertial loads of the cap and plug. Since the housing is fabricated from 304 stainless steel material and the cap is fabricated from higher-strength 17-4 PH material, the load on the i
housing is more critical.
The maximum stress developed in the housing in the closure region is found as follows:
Do -- cap outer diameter (from W 7593C55) 4.295 in;
=
Dj - 4.08 in; ledge inner diameter (from W 2D77445)
=
AL -- bearing area w/4 (D* - D ')
=
j w/4(4.295*-4.08*)
=
1.414 in*;
=
W1 -- calculated weight of inner container components above ledge (cap and plug) 2.0 + 4.5 = 6.5 pounds;
=
P -- maximum preload 2750 pounds [ load of 1260 pounds to compress the c-ring
=
based on manufacturer's testing plus additional load to resist a potential pressure increase in a fire accident-see Section C of this attachment]
Ng -- maximum g-load developed in a bottom or bottom corner drop accident 1065.7 g's (from Sections A.3 and A.4.b of this attachment)
=
L -- peak bearing load P+(Wi x Ng)
=
2750 + (6.5 x 1065.7)
=
9677 pounds
=
Attachment C to S#87-2738 Page 23 y
'/)
b.-- peak bearing stress Lk 9677
'r 1.414 6844 psi
=
The allowable stress is defined based on the criteria of not permitting the material to yield:
compressive static yield strength of 304 stainless steel oy - material at 300*F ti.
si
~ 27,420 psi (based on the compressive yield strength at 70*F
=
'shown in the NRBK-41 cask SARP and degraded for the elevatedtemperatureof300*F)
- /
Since the peak bearing stress of 6644 psi is significantly less
~ *'
than the static yield strength of the material of 27,420 psi, there would be no yielding of the housing material in the closure region in a flat bottom or bottom corner drop accident and the joint would be maintained after this type of accident.
(b) Evaluation of Loading on Inner Container Housing Due to Top and Top Corner Drops The only difference in calculating the peak stress for these drop orientations from the calculation in Section A.5.(a) is the weight of the component that compresses the housing ledge and the applicable'g-load.
W2'-- calculated weight of the housing (W 2077456, item 1) which acts to compress the ledge 19 pounds;
=
d Ng -- maximum g-load developed in a top or top corner drop j
accident-1768.5 g's (from Sections A.2 and A.4.a of this attachment)
=
L' = P+(W2 xN) g 2750 + (19 x 1768.5)
=
36,351.5 pounds a:
=
36,351.5 "b "
1.414 25,708 psi
=
q.
q
Attachment C to Sl87-2738 Page 24 Again, this stress is less than the static yield strength of the material of 27,420 psi so there would be no yielding of the housing material in the closure region in a flat top or top corner drop accident and the joint would be maintained after the accident.
(c)
Evaluation of Loading of Screws Due to Bottom and Bottom Corner Drops The combined stress at the root of the screw threads due to the torsional shear stress in the threads and the compressive stress in the screws as a result of column loadings of portions of the screws protruding below the plug (see Figure 6) is addressed.
The column loading is a result of both preload and the inertial loadings of the plug and screws, and this total load is found as follows:
n -- number of threads per inch for screw 20
=
T -- maximum torque required to compress c-ring h[1+sina(cose-fsina)]
(Reference (5),
=
Equation 6-7) where:
F --
maximum joint preload per screw 20
= 344 pounds
=
8 p --
pitch 1/n = 1/20
=
0.05 in.
=
f --
maximum coefficient of friction with chrome-plated screws 0.50 (based on a final value of 0.50 for Armco type
=
17-4 PH with as-plated chromium from Reference (8),
Table 4, page 10-79 and experience with actual joint testing on the design shown on W 2262F51) lead angle of threads at basic pitch diameter for e --
1/2-20 UNF-2A threads l
1.95*
(Reference (7), Table 2.9)
=
l 4 --
one-half the included thread angle 30'
=
Attachment C to 5#87-2738 Page 25 e --
angle between the normal to thread surface and a line parallel to axis
-((1+ tan $+tana)t/2)=cos-((1+ tan 30+ tan 1.9 cos
=
2 2
2 e 2
o 30.04*
=
- Then, T = (344) (.05) jy.
(0.50)(sin 1.95)))
0.50 23 sin 1.95 (cos 30.04 -
50.13 in-lbs (a value of 55 in-lbs is used in subsequent
=
calculations);
Dmin -- minor diameter of a screw 0.4374 in (Reference (7), Table 2.21);
=
c -- distance from central axis of screw to extreme fiber (to minimize this value and maximize the shear stress, the minor diameter of the 1/2-20 UNF-2A screw is used)
Dmin/2 = 0.4374/2 = 0.2187 in;
=
J -- polar moment of inertia wDmin /32 (Reference (6), p. 453) 4
=
(w) ( 374)' = 0.0036 in';
=
t -- torsional shear stress at root of cap screw threads f
(Reference (5), p. 151)
=
(55) (0.2187)
= 3341 psi (0.0036)
W3 -- calculated weight of plug plus screws 4.5 + 1.0 = 5.5 pounds;
=
total axial load on each scrrw in a bottom or bottom corner P
T - drop accident (where the deceleration, Hg = t) 1065.7, is taken from section A.S.(a) of this attachmen p,IN ) (N ) = 344 + (5.5)(1065.7) 3 g
8 8
Attachment C to S#87-2738 Page 26
= -1076.7 pounds; Since the screw will protrude no more than 0.066 inch below the bottom of the plug the screw can be treated as a short column loaded in compression and the compressive stress for the screw can be determined.
A
--areaofcolumn(screw) r
=(w/4)(.4374)'
= (w/4) Dmin
= 0.150 in';
"c
-- stress due to column loading of screw T
"A r
1076.7
.150
= 7178 psi max -- maximum shear stress T
=((
)'+t*]*/*=[(
7f78 )2 3341*]*/*
+
= 4903 psi t(max) -- maximum combined stress at root of screw threads C
- 2'
- ' max 7f78+4903
=
= 8492 psi The allowable stress is equal to the static yield strength of the material:
Attachment C to Sf87-2738 Page 27 y -- screw material yield strength at 300*F o
= 101,730 psi (based on the yield strength at 70*F speci-fied by ASTM A-564, Comp. 630 Heat Treated to 1100'F, and degraded for the elevated temperature);
Since the maximum stress of 8492 psi is significantly less than the material's static yield strength of 101,730 psi, the screws would not yield in a bottom or bottom corner drop accident.
The maximum shear stress also needs to be checked to ensure that it is less than the shear yield strength of the material. As calculated
- T equals 4903 psi which is less than the max material's shear yield strength of 58,698 psi (0.577 times o ).
y (d) Evaluation of Loading of Screws Due to Top and Top Corner Drops For this evaluation, the approach presented in Section A.S.(c) applies with the only variables being the weight of the components contributing to the inertial load that acts on the screws and the deceleration developed at impact.
W4 -- weight of housing plus cap plus cargo
= 19 + 2 + 44 (assuming that the inner container cavity is completely full of fuel plates and cladding which are considered homogenous, thus possessing the properties of zirconium)
= 65 pounds.
Ng = 1768.5 g's (from Section A.5.(b) of this attachment)
P*
c' Tmax, and t(max) can then be found as before.
T (65) 768.5)
PT = 344 +
= 14,713 pounds 14,713 "c *
.150
= 98,087 psi; max"{(
. ) +3341')
t l
= 49,157 psi i
Attachment C to S#87-2738 Page 28 98,087 "t(max)=
+ 49,157
= 98,200 psi.
This is less than the stress limit of 101,730 psi calculated in Section A.S.(c) and it is therefore acceptable.
The maximum shear stress of 49,157 psi is less than the material's shear yield strength of 58,698 psi calculated in Section A.5.(c) and is also acceptable.
(e)
Evaluation of Loading of Plug Internal Threads Due to. Bottom and Bottom Corner _ Drops The screws and the plug are made from different materials. The screws-are made from 17-4 PH (ultimate strength = 140,000 psi) and the plug is made from 304 stainless steel (ultimate strength-
= 75,000 psi). For the case where the ultimate strength of the external threads (screws) in a threaded connection greatly exceeds the ultimate strength of the internal threads (plug), the following is used to determine the shear area:
n -- number of threads per inch
= 20; O (min) -- minimum major diameter of 1/2-20 UNF-2A external s
threads
= 0.4906 in (Reference (7), Table 2.21);
E (max) -- maximum pitch diameter of internal threads n
= 0.4731 in (Reference (7), Table 2.21);
L, -- length of engagement (minimum depth of threads in plug-)
= 0.98 in; AS -- thread shear area
=inL,0(min)[h+0.57735(0(min)-E(""*))
3 3
n (Reference (7),AppendixAS)
=(1)(20)(0.98)(0.4906)[(2)(20)+0.57735(0.4906-0.4731)l
= 1.06 in2 I
f 4
i
Attachment C to 5#87-2738 Page 29 The maximum shear stress is found from:
-PT -- total axial load per screw
= 1076.7 pounds (from Section A.5.(c) of this attachment) x -- shear stress PT 1076.7
- AT " 1.06
= 1016 psi The shear stress is limited to the shear yield strength of the material (t{e)e. The static shear yield strength of the 304 stainless s l plug assembly material is 15,821 psi (0.577 times the compressive yield strength of the material which is specified in Section A.S.(a) of this. attachment).
Since the peak shear stress of 1016 psi is_less than the static shear yield strength of the material of 15,821 psi, the stress is acceptable.
(f) Evaluation of Loading of Plug Internal Threads Due to Top and Top Corner Drops Using the approach of Section A.S.(e), only the applicable component weight and the deceleration developed at impact
. change. W4 is used rather than W3 and therefore PT from Section A.S.(d)istheloadoneachscrew.
PT= 14,713 pounds; I
14.713 1.06 l
13,880 psi;
=
Again, this is acceptable since it is less than the static shear yield strength of the material of 15,821 psi.
I
Attachment C to S#87-2738 Page 30 L(g) Summary of Calculated Component Stresses Drop Calculated Section Orientation Stress (psi)
Limit (psi)
A.5.(a) bottom 6844 27,420 A.S.(c bottom 8492 101,730 A.S. e bottom 1016 15,821 A.5. b top 25,708 27,420 A.S. d) top 98,200 101,730-A.S. f) top 13,880 15,821 As can be seen from this table, no component is stressed above i
yield and therefore the joint would be maintained after an axial drop accident.
j
'6.
Evaluation of Deformation of Plug Rim The following evaluation is provided to confirm that the top of the inner container (i.e., the rim at the upper end of the plug, W 2077456, item 2) would not be deformed to the extent that the test penetration extension on the cap (W 2077456, item 3) would be con-tacted in a 30 foot top or top corner drop accident. This ensures preservation of another portion of the containment boundary of the inner container during application of an axial load.
i l
For this assessment, for conservatism, the inner container will be j
assumed to fall 30 feet onto an unyielding surface (i.e., the l
protection offered by the NRBK-41 cask will not be considered)
{
striking the top of the rim on the plug.
The energy associated with the loaded inner container will be assumed to be dissipated entirely in deformation of the plug rim.
1 The incremental energy absorption technique presented in Section A.1 of this attachment will be used to determine the amount of defor-mation. Material redistribution is represented as shown in Figure 3 based on an original contact area determined as follows:
DO -- outer diameter of plug rim (from W 7593C78) 4.925 in-
=
l 01 -- inner diameter of plug rim (from W 7593C78) 4.52 in;
=
AO -- original contact area f
w/4 (D0' - DI')
=
n/4 (4.925' - 4.52')
=
3.0 in';
=
')
Attachment C to S#87-2738 Page 31 Other variables, as defined in Section A.2 of this attachment unless otherwise indicated, are:
MRD = 1 in. (height of thin-walled portion of plug - W 7593C78) 27.9 x 10' psi (based on the value of Young's modulus at E
=
70*F shown in the NRBK-41 cask SARP and degraded for an elevated temperature of 300*F) o, stat = 27,420 psi (from Section A.S.(a) of this attachment);
y o, stat = 62,500 psi (based on the value for ultimate strength at 70"F u
shown in the NRBK-41 cask SARP and degraded for an elevated temperature of 300*F)'
cu= 0.63 in/in; W
26.5 + 44
=
70.5 pounds (weight of inner container plus cargo);
=
h 360 in;
=
DINC 0.005 in;
=
L = 15.87 in.
Results indicate that the plug would be crushed approximately 0.16 inch. Since the top of the test penetration extension on the cap is at least 0.249 inch below the top of the undeformed plug, the extension would not be contacted in a top drop accident based on the conservative model used. Thus, this portion of the containment boundary of the inner container is also protected in a drop accident.
l l
A--____-_--_________
Attachment C to S#87-2738 Page 32 1
B.
Lateral Loading of the Inner Container f
1 1.
Energy Absorption As was done for the axial and corner drop evaluations presented in I
Section A of this attachment, the deceleration loading developed 1
during a 30-foot side drop of the NRBK-41 cask is determined using the incremental energy absorption technique shown in Section A.I.
The deceleration loading is then applied to the inner container, and the
)
ability of the inner container joint to be maintained is evaluated.
1 To use the technique of Section A.1, the impact area needs to be specified(seestepA.1.(a).5). To maximize the g-load, the container will be assumed to be oriented so that one of the trunnions impacts the unyielding surface in combination with the protruding portion of i
the base plates.
The bolts securing the skid to the cask would fail in this orientation causing the skid to be removed from the cask so that the skid is not considered in evaluating a side drop.
For conservatism, no credit is taken for dissipation of energy related to skid assembly deformation.
At the trunnion assembly, the four-inch diameter trunnion boss (see Figure 7) would crush to absorb a portion of the kinetic energy of the I
package.
Impact would occur on the trunnion plate with material redistributed along the length of the boss. Therefore, the length of the column being crushed will be considered to be the length of the boss (3.12 inches) minus the thickness of the trunnion plate (1 inch) or 2.12 inches (= L = MRD). The cross-sectional area of the four-inch diamater cylincirical trunnion boss is 12.57 in2 At he base pi&te, the combined thickness of the base plate and the supplementary base plate is two inches.
The width of the base plate surface parallel to the trunnion plate is 16 inches. Thus, the base plate contact area is 32 in2, and the total contact area (A0) for the trunnion and the base plates is 44.57 in2 The crush area is determined in the same manner as for a flat top drop (Section A.2 of this attachment). Thus, Ac -- crush area
l
Attachment C to S#87-2738 Page 33 2.
Orop Analyses TRUNNION BOSS TRUNNION PLATE r'
,s f
g
\\% '/
O O
_ q _
i i
1 A
nh /
l Figure 7 - Cask Trunnion Configuration The remainder of the variables (defined as in Section A.2) are as follows:
E = 28 x 10" psi; o, stat = 35,000 psi; y
o, stat = 75,000 psi; u
l l
u_____---__-_-
I Attachment C to S#87-2738 Page 34 u = 0.63 in/in; c
W = 9010 pounds; h = 360 in;.
DINC = 0.025 in.
The results are that a maximum deceleration (Ng) of 600 g's is developed.over a cumulative crush depth of 0.72 in.
~
3.
Assessment of Inner Container Adequacy The inner container is designed to' fit snugly within the cavity of the NRBK-41 cask. The inside radius of the cask cavity is nominally 2.50
. inches..The outer radius of the inner container is 2.475 inches.
To verify the integrity of the inner container, the following is provided to demonstrate that not even the walls of the inner container would be permanently deformed in a side drop accident.
The maximum stress at the point of contact between the inner container and the cask cavity can be found using the following from Reference (9), Table 33 Case 2c for a cylinder in a cylindrical socket.
.Wj weight of inner container plus cargo 26.5 + 44
=
70.5 pounds
=
L --
length of inner container housing 14.37 in
=
P total load Wj x Ng = (70.5) (600)
=
42,300 pounds
=
l load per linear inch p
P 42,300 E" 14.37
)
2944 lbs/in
-)
=
D diameter of the cask cavity i --
5.00 in
=
D diameter of the inner container 2 --
4.95 in i
=
j l
{
tL_________---._----.--
1
Attachment C to S#87-2738 Page 35 KD --
constant defined for Case 2c, Table 33, Reference (9) 0 D 3
2 0-0 3
2 (5.00)(4.95)
(5.00 - 4.95) 495 in
=
The cask cavity is fabricated from 304L stainless steel. The inner container is made from 304 stainless stea1.
E i --
Young's modulus of 304L stainless steel material (cask cavity) at 300*
27x10[ psi (basedonthevalueat70*FshownintheNRBK-41
=
cask SARP and degraded for the elevated temperature)
E Young's modulus of 304 stainless steel material (inner 2 --
container) at 300*F E
=
i vi --
Poisson's ratio for 304L stainless steel material (cask cavity) 0.3
=
v2 --
Poisson's ratio for 304 stainless steel material (inner container)
=
vi for the case where E
=E t
2 and vi = v2 = 0.3, Reference (9) provides the following equation for calculating the maximum stress:
c --
maximum stress
.591([0E)1/2 =0.591((2944) x10')j/2 i
=
7489 psi
=
Since the maximum stress of the most severely loaded portion of the inner container is less than the static yield strength of the material of 27,420 psi (from Section A.S.(a) of this attachment), the inner container will undergo no permanent deformation in a side drop accident of the NRBK-41 cask.
i
_____._______o
Attachment C to S#87-2738 Page 36 C.. Pressure Increase in a 1475'F Fire Accident.
The following analyses demonstrate the ability of the MIN-41 inner container to withstand the pressure-increase associated with subjecting the NRBK-41 cask to the hypothetical 1475'F fire accident.-
The peak temperature experienced by the MIN-41. inner container and its
-cargo in a fire accident'would be 580*F. This peak temperature was obtained from the TRUMP results for the fire accident which are discussed' in Attachment A.
Since users are prohibited from shipping liquids in the inner container by a restriction in the Certificate of Compliance for the NR8K-41 cask (see Attachment-E), the peak total pressure:within the MIN-41 during a 1475'F fire would normally be due to the constant volume of heated air within the container.
1 To find the pressure due to the heated air inside the container, the air i
is treated as a constant mass and-volume, initially at room temperature i
.(70'F = 530*R) and raised to a maximum temperature equal to the temper-ature'.of the container and its contents during the fire (580*F = 1040*R).
Using the ideal gas-law, the known conditions at 70'F can be equated to conditions at 580*F to determine the unknown pressure. For constant volume, the gas law:
P V PV 3 1 2p T
T 1
2 reduces to P
P 3
p 5"5' Therefore,
= (l4*7 1040) = 28.8 psia.
p2=
1 1
To allow for possible additional pressurization resulting from the effect of the elevated temperature on the cargo, a pressure limit of 114 psia will.be imposed on the inner container and its contents. This permits pressure generation'of 85.2 psia in addition to that resulting from the heated air in the container. Thus, PT (total permissible pressure) =
114 psia.
At the closure joint, the force exerted due to this pressure is determined as follows:
1 Dc -- diameter of the c-ring groove in the container housing
= 4.08 in; v
A
f Attachment C to S#87-2738 Page 37 Ac -- area of this section
= (w/4) D
= (w/4) (4.08)'
c
= 13.07 in';
F -- force exerted at the joint due to the increased pressure
= (P ) (A ) = (114) (13.07)
T c
= 1490 pounds.
At assembly, the joint is preloaded by an amount which is greater than the load required to compress the c-ring. The total required preload is equal to the maximum possible load to compress the c-ring, 1260 pounds as deter-mined by the manufacturer, plus load F.
Application of the additional preload will ensure that exposure of the NRBK-41 cask to a fire would not result in separation of the inner container joint as a result of the increase in pressure within the container.
Therefore, for assembly, a preload of 2750 pounds will be established.
The inner container shell is evaluated for its ability to withstand the higher pressure by representing it as a thin-walled cylindrical pressure vessel with both ends capped. A wall thickness of 0.13 inches, which is the local thickness of the shell in the flats region shown in Section A-4 of W 2077445, is used conservatively to represent the shell thickness over its entire length.
CaseIcofTable29ofReference(9)isusedto evaluate the shell for the increase in pressure.
R -- container inner radius
= 2.22 in; t -- wall thickness
= 2.35 - 2.22
= 0.13 in; q -- pressure
=PT (from above)
= 114.0 psi; a -- hoop stress "aR " (114) J2.22) t
- 0. : 3
= 1947 psi i
{
The inner container shell is fabricated from 304 stainless steel material i
which has a yield strength of 18,000 psi at 600*F. Since the hoop stress is significantly less than the yield strength of the material, there would l
be no permanent deformation of the shell due to a pressure increase associated with the hypothetical 1475'F fire accident.
l
Attachment C to S#87-2738 Page 38 References
-(1) WAPD-TM-753, Dynamic Properties of Materials, May 1969 (2) WAPD-RE0(C)-302, A1W-3 Core Barrel and Thermal Shield Disposal Container Safety Analysis Report For Packaging. Bettis Atomic Power Laboratory, April 1982 (3) WAPD-RE0(C)-448, A1W-3 Holddown/ Support Barrel and Shroud Disposal' Container Safety Analysis Report For Packaging, Bettis Atomic Power Laboratory, November 1982 i
(4) Marks' Handbook, McGraw-Hill Book Company, Inc., 4th Edition, 1941 (5) Design of Machine Members, Venton L. Doughtie and Alex Vallance, McGraw-Hill, Fourth Edition, 1964 (6) Machinery's Handbook, Industrial Press, Inc., N.Y., N.Y., 1971 (7) Screw-Thread Standards for Federal Services, Handbook H28 (1969) - Part 1 (8) Nuclear Engineering Handbook, H. Etherington, McGraw-Hill, First Edition, 1958 (9) Formulas for Stress and Strain, Raymond J. Roark and Warren C. Young, McGraw-Hill, 5th edition, 1975 (10)
IMPAC2 - A Lumped Mass Nor'inear Spring Computer Program to Analyze Type B Shipping Container impact Problems, J. Counts and J. B. Payne, Informal Report No. LA-6643-MS, Los Alamos Scientific Laboratory, Los Alamos, New Mexico, January 1977 i
(11) WAPD-RE0(C)-122, S3G Core Basket Disposal Container Safety Analysis l
Report for Packaging, Bettis Atomic Power Laboratory, June 1980 (12) NRBK-41 Cask Safety Analysis Report for Packaging transmitted by WAPD-CL-1794 dated March 25, 1968, as amended (13) WAPD-LP(CES)SE-170,LWBRSeedSafetyAnalysisReportforPackaging i
.y Attachment D to Sf87-2738 Page 1-i' I'
CRITICALITY ANALYSES.0F AN NRBK-41 CASK FOR NORMAL CONDITIONS 0F TRANSPORT AND HYPOTHETICAL ACCIDENT CONDITIONS 4
m_____________.____
Attachment D to S#87-2738 Page 2 Criticality Safety Analysis. NRBK-41 Cask Table of Contents Paae 1.0 Summary and Conclusions 3
2.0 NRBK-41 Fuel Packaae 4
3.0 Calculational Method 5
3.1 RCP01 Computer Program 5
3.2 Nuclide Cross Section Data 5
4.0 Analytical Model and Assumptions 6
4.1 NRBK-41 Cask Model and Assumptions 6
4.2 Undamaged NRBK-41 Cask Array Model 7
4.3 Damaged NRBK-41 Cask Array Model 8
4.4 Nuclide Inventories 8
.5.0 Criticality Results 8
5.1 Undamaged Cask Array:
Normal Shipment Conditions 8
5.2 Damaged Cask Array:
Accident Conditions 9
5.3 H/X Parameter and Fuel Height Variations 11 6.0 Criticality Benchmark Results 12 7.0 Minimum Marain to Criticality 14 8.0 References 14 Table 1 NRBK-41 Cask and IN-41 Container 15 Table 2 NRBK-41 Undamaged Cask Array 16 Table 3 NRBK-41 Damaged Cask Array 17
' Table 4 Criticality Results for Undamaged Arrays 18 Figure 1 IN-41 Container 19 Figure 2 MIN-41 Container' 20 Figure 3 IN-41 Container in the NRBK-41 Cask 21 Figure 4 RCP01 Representation of Undamaged Array _
22 Figure 5 RCP01 Representation of Damaged Array 23 Figure 6 RCP01 Keff vs. Size of Cask Array 24 i
Attachment D to S#87-2738 Page 3
)
l Criticality Analysis for NRBK-41 Cask 1.0 Summary and Conclusions Criticality calculations and analyses arg3gerfggged to ggmit the transport of irradiated or unirradiated U,
U and Pu fuel specimens in the lead-lined NRBK-41 shipping casks.
The NRBK-41 ship-ping casks that are analyzed comprise two similar types:
(1) the un-modified NRBK-41 cask and IN-41 insert and (2) the modified NRBK-41 cask
)
j and modified IN-41 or MIN-41 insert. The analyses are completed in ac-j cordance with those Code of Federal Regulations 10CFR71 requirements j
which allow shipment of Fissile Class I packages.
{
Standards for Fissile Class I packages, specified in Section 71.57(a) of 10CFR71, require that any number of undamaged packages would be sub-critical in any arrangement and with optimum interspersed hydrogenous moderation unless there is a greater amount of interspersed moderation in the packaging, in which case the greater amount may be assumed for this determination.
The results of the analyses, contained in this at-tachment, show under normal' undamaged conditions of transport that the maximum Keff NR which are each fully loaded with 350 equivalent gramsis<0.91andsubcri V optimally moderated and arranged in a three-dimensional most-reactive array.
It is concluded that any number of loaded NRBK-41 casks are considered nuclearly safe satisfying Fissile Class I package criteria for normal, undamaged shipment conditions.
For damaged conditions, the standards for Fissile Class I packages iden-tified in Section 71.57(b) of 10CFR71 require that two hundred and fifty (250), if each package were subjected to the tests specified in Section 71.73 (Hypothetical Accident Conditions), would be subcritical if stacked together in any arrangement, closely reflected on all sides of the stack by water, and with optimum interspersed hydrogenous modera-tion. The analyses performed herein show that for damaged conditions involving the fuel loaded casks previously described for the undamaged case that the maximum Keff is < 0.95 and subcritical for any number up to 8000 casks or thirty-two (32) times the 10CFR71 requirement.
Since the assembly of 8000 loaded casks is considered an implausible event then it is concluded, based on the results of this analysis, that any number of NRBK-41 casks in transport or staged in a stacked array are considered nuclearly safe for damaged as well as undamaged conditions satisfying the Fissile Class I package criteria of 10CFR71.
)
The details regarding the calculational method and model and the results used to support the conclusions for Fissile Class I application for the i
NRBK-41 shipping cask are provided in subsequent sections.
s
Attachment D to S#87-2738 Page 4 2.0 NRBK-41 Fuel Packaoe The criticality analysis for the NRBQ1 cask is based on a fuel loading which is equivalent t 2g0gramg3gf U per cask container.
N equ h-alence parameters for U and Pu g sses are 1.4 and 1.{ h is 2
respectively,inthat1/1.g3gramof U or W.6 gram of equivalent to 1.0 gram of 233,. gPu as well as with gtion is tested U
is equivalency form by calculating cases with 0
U fissile materials.
In these analyses, all cases except one use for conservatism an un-modified NRBK-41 package, i.e., a cask minus the thermal shield and sup-l plementary base plate and with an IN-41 inner container. Omission of the modifications is conservative since the stainless steel added as part of the modification to the NRBK-41 cask is a thermal neutron ab-sorber which would reduce the keff.
In addition, the modified IN-41 containerorMIN-41inserthasasglercavityvolumethantheIN-41 container so that the hydrogen to U ratio for the same loading is reduced more from the optimum condition.
Although additional calcula-tions for the modified package are not necessary, one case is run to verify that the keff is lower.
Figure 1 shows the dimensions of the IN-41 container and the cask which make up the unmodified package.
The dimensions were extracted from the drawings used by the vendor in the fabrication of the container and the cask. The dimensional tolerances, shown in Figure 1, were applied in the calculational representation of the container and the cask, in a manner described in Section 4, that would yield conservative results.
Although the fuel containers are sealed to prevent in-leakage, water moderation is allowed and as near optimum H/X conditions as can be es-tablished, within the dimensional limitations of the container and cask cavity, are factored into the analyses.
For the modified NRBK-41 package, a 0.25-inch cylindrical stainless steel thermal shield is added around the cask with a.062-inch gap separating the thermal shield and cask outer surface.
A one-inch thick stainless steel plate with a 0.12-inch recess in its top surface is added to the cask's base plate.
The modified IN-41 container, desig-nated the MIN-41 insert, is reduced in cavity volume, tapered over the upper quarter of its length to a smaller diameter at the top, and has slightly thicker walls than the IN-41 container.
Figure 2 provides dimensions of the modified cask and container used in the analyses.
Ac-tual dimensions and tolerances for the modified cask are found on Wes-tinghouse drawing 1755E01. Dimensions for the MIN-41 insert are on Wes-tinghouse drawing 2D77456 (Enclosure 2 to this letter). Water modera-tion and H/X conditions were represented in the same manner as for the IN-41 container and unmodified cask.
l
Attachment D to Sf87-2738 Page 5, 3.0 Calculational Method 3.1 RCP01 Comouter Proaram The criticality calculations for the NRBK-41 casks were per-formed using the Monte Carlo computer program RCP01 described in Reference 1.
RCP01 performs neutron reaction rate calcula-tions in three dimensions using explicitly detailed energy and spatial description of all materials within its solution space.
An initial source shape estimate is introduced to start an iterative process using Monte Carlo techniques that results in converged source and neutron flux distributions consistent with the Keft or multiplication constant for the system being analyzed.
The calculational uncertainties are determined from the iterative procedure in terms of 95% confidence intervals.
The program has been widely used at the Bettis Laboratory to calculate benchmark experiments and has been accepted as an analytical standard.
3.2 Nuclide Cross Section Data Thenuclygcrgggsectiondataprincipallyusedforhydrogen,
- oxygen, U,
U, stainless steel and concrete are based on proven data used in the Bettis nuclear design models as well as in the analyses of critical experiments.
The data are obtained from the Bettis XAP library system described in Reference 2.
The individual nuclide cross section data sets are organized into multi-group sets containing 57 fast energy groups and 25 thermal energy groups.
The identities of the nuclide cross section libraries used in thi33$"*N3b8 *
- kb*Pu as w The fission spectra used for U,
U and the thermal hydrogen scattering kernel are identified also in Table 1.
The plutonium cross section data are based on ENDF/B-IV evalua-tions and are identical to those used in the analyses of criti-cal assemblies described in Reference 3.
l The fast cross section data used for lead are ENDF/IV data, and the thermal cross section data are widely accepted as the best data available.
been measured experimentally and calculated for aTheeffectofaleadv 2
ector has 2
assembly, as reported in Reference 4.
While this experiment was not calculated with RCP01, it was concluded from this study that the RCP01 program can predict reactivity increases due to lead versus H O reflection and that these predictions are con-2 sistent with the experimental data.
An additional evaluation of the cross section set for lead has been conducted for the NRBK-41 cask analysis to assess the sensitivity of the Keff results to changes in the groupwise inelastic scattering data.
l
Attachment'D to S#87-2738 Page 6 s
The scattering cross sections were adjusted upward to their maximum values where' applicable to account for the variances 3
reported for.the' experimental cross section measurements of 1
lead. This approach is considered very conservative in that the inelastic scattering cross sections are adjusted al1~in the same direction in order to assess the reactivity effect due to' the enhanced scatter of neutrons from the lead. reflector into f
the fuel source. regions. As shown in Section 5.2, the results
.l of these changes are negligible within the statistical uncer-
'taintiestand provide additional confidence in the reliability and acceptability of the lead cross section data being used.
.4.0-Analytical Model and Assumptions 4.1 NRBK-41 Cask Model and Assumptions The RCP01 calculational model of the IN-41 container and the -
unmodified NRBK-41. lead-lined cask is shown-in Figure.3.
The exact representation is modified for the calculation to account for the dimensional tolerances in a manner that would yield-conservative results.
The model or unit cell thus established provides the basis for subsequent modeling of the arrays.
The following modeling conditions and assumptions are established for the unit cell or initial ' assembly which is comprised-of the IN-41 container and unmodified NRBK-41 cask:
235 233 (1)
The fuel 1ging is 350 grams 0 or 250 grams 0 or 220 grams Pu or any combination thereof that satisfies the following relationship:
235U grams /350 + 2330 grams /250 + 239Pu grams / 220 s 1 (2)
Any replacement of steel or air by lead within that al-lowed by the dimensional tolerances of the unit cell is considered conservative.
(3)
The container as well as the cask cavity are flooded with water, and water moderation of the fuel is allowed in the analysis as a given condition.
The maximum density water 0
at 39.2 F is used.
(4)
The nominal dimensions of the IN-41 container restrict the H/X to a value which is considered below optimum for minimum critical mass / volume conditions.
Consequently, any enlargement of the fueled region within the allowable tolerances or elimination of the insert itself would al-low more moderation and an increase in H/X towards op-timum. However, to ensure that optimum moderation condi-tions are established for full height fuel loaded condi-tions, partial height conditions with corresponding 1
i
..m
l l
l Attachment D to 4
S#87-2738 I
l Page 7 f
\\
.l
. variance in the H/X ratios are also calculated and the l
results presented in Section 5.3.
1 (5)
Fission products, present in irradiated fuel materials, are omitted for conservatism.
1 (6)
Dimensional tolerances are applied as follows:
(a) The I.D. of the IN-41 container is maximized to op-timize H/X.
(b) The I.D. of the cask inner liner is maximized to op-timize H/X for cases with the insert omitted as pointed out in Section 4.2.
l (c) The 0.D. of the cask inner liner is minimized and l
the I.D. of the cask outer liner is maximized to i
maximize lead reflector material.
(d) The thickness of the cask outer liner is minimized to reduce steel as well as to minimize the center-to-center distances of the units in an array.
I (8)
AK anaiysis,al of 0.98, is conservatively assumed in this ritic and is justified based on the benchmark cal-culations discussed in Section 6.0 for large geometry, j
low leakage critical configurations similar to the arrays calculated for this analyses.
4.2 Undamaced NRBK-41 Cask Array Model There are presently only six (6) NRBK-41 shipping casks in ex-istence.
Nevertheless, very large arrays of the unmodified casks stacked in a three-dimensional cubic configuration are q
analyzed to demonstrate compliance with the 10CFR71 require-ments. Arrays of 2x2x2(8), 6x6x6(196),10x10x10(1000), and 20x20x20(8000) casks are evaluated.
lhe RCP01 representation of an X,Y slice of the 10x10x10 array is shown in Figure 4, as an example.
The finite arrays are reflected on five (5) sides by a thick region (>1.5 feet) of water.
The sixth side or bot-tom of the array is reflected by concrete which is a better reflector than water.
The arrays are analyzed for water or air in the interstitial spaces to determine maximum reactivity coupling between the loaded casks. Moreover, the arrays are analyzed a!so with and without the fuel container inserts in the casks to evaluate as an accident condition the omission of the insert and the op-timization of the fuel and moderator over the entire cask cavity.
Omission of the fuel container insert eliminates the
Attachment D to S#87-2738 Page 8 need to evaluate other type insert materials, e.g., Zr or A1, used in cask shipments and as the results will show in Section 5.1 also adds more conservatism to the analyses. Also, a 10x10x10 undamaged array of modified NRBX-41 casks with MIN-41 inner containers was analyzed with air in the interstitial spaces for comparison to the same case with the array of un-modified casks and IN-41 containers.
4.3 Damaaed NRBK-41 Cask Array Model Figure 5 shows the RCP01 representation of the 10x10x10(1000) array of damaged casks.
The following conditions and assump-tions, which are considered to conservatively represent poten-tial damage in an accident, were applied in analyzing the damaged cask array:
(1)
The array of casks is completely deformed such that the cask outer steel liners are lost and the inner containers are essentially immersed in solid lead.
(2)
The center-to-center distance for the casks is reduced to account for the reduction in the array volume upon defor-l mation.
l (3)
For non-fire conditions, the lead fills the interstitial spaces and is conserved for all casks in the array.
(4)
For hypothetical fire conditions, a 5% reduction in lead volume is assumed for each cask in the array resulting in a closer center-to-center distance.
(5)
Full water moderation of the fuel in each container is l
maintained and the array is reflected on five sides with water and the sixth side with concrete.
4.4 Nuclide Inventories The atomic densities used for materials in the calculations of the normal, undamaged arrays are shown in Table 1.
5.0 Criticality Results 5.1 Undamaaed Cask Arrav:
Normal Shioment Conditions Table 2 provides a summary of the Keff results obtained in analyzingtheNRBK-41casksforundamagedarraycondigns.
Each cask in the array contains 350 equivalent grams U which is optimally water moderated. The array is reflected by water i
and concrete.
Graphic plots of the various sized arrays are shown in Figure 6.
The four curves were developed to highlight
4 i
' Attachment D to S#87-2738 Page 9 the effects on'the: array K-effective.in changing the intersti-tial environment as well as omission of the container inserts.
TheresultsshowthatthemaximumK,fker(Case 16)'isobtained for the largest ~ array (8000 casks), w ein the interstitial spaces contain air;and the container inserts.are omitted.
The maximum K ff obtained for the analyses.of, the undamaged arrays is. 0.91,and includes a-1.0% A O reactivity bias to account for j
variances associated with the lead inelastic scattering cross-section measurements.
The 1.0% op bias for the lead cross-section data is justified based on the results shown in Table 3, Cases 7 and 9.
The plots.in Figure 6 demonstrate that water in the intersti-tial spaces l essentially." poisons" the~ array such that the K-effective remains well subcritical and relatively constant as
-the array size increases. However, when the interstitial i
spaces contain air, reactivity coupling of the individual units increases such that as the array size increases the K f in-Nevertheless, based on the slope of the plo!,f creases.
in excess of 1,000,000 casks would be needed to approach a criti-cal condition.
This is considered an implausible result.
i
}ggresultsforCases15and17inTable2showthatyggnthe U is replaced by the equivalent fissile amount of U, the K ff3 for the 10x10x10 arrays are essentially the same within t$e-statisticaluncertainties.233This confirms that the equiv-alency parameter of 1.4 times U mass is valid to the NRBK-41 cask geometry and material. environment used for this analysis.
As expected, the array of modified NRBK-41 casks represented in Case 18 shows a lower kept than the array of unmodified casks in Case 7 because of the increased steel volume associated with the thermal shields of the modified casks and the smaller diameter MIN-41 inserts. Thus the modified NRBK-41 cask with j
the MIN-41 insert is even more,nuclearly safe than the un-j i
modified cask with the IN 41 container.
{
Consequently, the undamaged array results demonstrate that any f
number of unmodified or modified NRBK-41 casks arranged in a-stacked three-dimensional most-reactive array reflected by
.{
L water is well subcritical thus satisfying the 10CFR71 Section i
71.57(a) requirements allowing the shipment of Fissile Class I packages under normal conditions.
l 5.2 Damaced Cask Arrav: Accident Conditions L
Table 3 provides a summary of the Keff results obtained in analyzing the unmodified NRBK-41 gks gU, ordamagg9 array tions and with casks loaded with U,
Pu fissile l
b
1 Attachment D to Sf87-2738 i
Page 10
]
material.
No cases were run for the modified NRBK-41 casks as j
a result of the conclusions reached in evaluating the undamaged j
casks in Section 5.1.
The maximum Keff,. Case 4, is obtaihed for the largest arggg j
(8000 casks) wherein each cask is loaded with 350 grams V,
3 optimally moderated and the insert is omitted. The maxK l
eff is
< 0.95 for this case and includes a 1% Aj0 reactivity bias as-sociated with the variances in the lead inelastic scattering cross-section measurements.
Cases 5, 6 and 7 were run to demonstrate that replaced by the equivalent fissile amount of 23f*o"r$$9 U
Pu, the l
K-effectivesarethesap$5w nsa cal uncedaindes.
gsconfirmsthatthe U equivalency parameters of 1.4 times U mass and 1.6 times 239Pu mass are valid for the NRBK-41 cask, the modified NRBK-41' cask and insert applications.
Case 8 was run to test the change in K ff incurred as a result ofa5%reductionintheleadvolumeofeachcaskandoptimal rearrangement due to heat and melting from a hypothetical fire accident.
As shown in Table 3,2ge test case was run for a 10x10x10 cask array containing Pu fuel.
The result indi-cated a well shutdown condition for these extreme accident con-ditions.
Case 9 represents the condition wherein the lead inelastic scattering cross-section data were adjusted upward to factor in the variances associated with the experimental measurements.
Comparison of Case 9 to Case 7 indicates that the only dif-ference in the calculations is the lead cross-section data sets used.
The Keff results for these two cases are essentially the same within statistical uncertainties. However, based on these results a positive 1% reactivity bias is added to the maximum calculated Keff to account for the lead cross section data un-certainties Case 10wasrunfor$9c mparison to Case 9 to demonstrate that 2
a 5 gram change in Pu content of each cask yielded no noticeable change in Keff within the statistical uncertainties of the calculation Comparison of results for Cases 13 through 16 in Table 2 for the undamaged array and for Cases 1 through 4 in Table 3 for the damaged array demonstrate that with everything being the same regarding the fuel loading and H/X ratio that as the array becomes very large, the damaged array is more reactive than the undamaged configuration.
Attachment D to S#87-2738 Page 11 i
It is judged that cask arrays of sizes larger than 8000 units l
are considered implausible.
Consequently, based on the results i
provided, it is concluded that any number of unmodified NRBK-41
(
casksandmogiedNRBK-41caskseachloadedwith350equiv-l alent grams V, optimally moderated, stacked together in the most reactive arrangement and subjected to hypothetical damage conditions will be subcritical satisfying the Fissile Class I package criteria of 10CFR71.
I 5.3 H/X Parameter and Fuel Heiaht Variations One of the assumptions used in the accident analysis is that the homogenization of the fuel and moderator over the entire volume of the container or the cask cavity is the most reactive fuel configuration because this geometry affords optimization of the hydrogen to fuel (H/X) ratio.
Since the height (H) of the fuel region is approximately three times greater than its diameter (D), and H/X is close to optimum at full height load-ing, a reduction in the fuel height and shift of H/D towards unity could conceivably provide a more reactive geometry than the full height loading.
i Consequently, an auxiliary set of calculations was performed to evaluate variations in H/X and height / diameter.
The results of these calculations are presented in Table 4.
Because precision is necessary to obtain useful results for this parameter evaluation, the majority of these RCP calculations use a much larger number of neutron histories than the calculations per-formed in Sections 5.1 and 5.2.
Many of the calculations were repeated with a different sequence of random numbers, and the results were weighted by the number of iterations to give an average keff and average 95% C.I. error.
The basic geometry used for these calculations is that of Case 15 in Table 2, i.e., a 10x10x10 array of undamaged casks in air with no insert.
The fuel / moderator mixture height was set at 1.0, 0.7, 0.4 or 0.3125 times full cask cavity height, with 0.3125 times H representing a height to diameter ratio of unity. The remaining volume in the cask was filled with either air or water and identical cases run for each fissile material configuration to assess which condition was more reactive.
The fuel materials consiggged ggge the ggge as those used in the accident analysis:
U, U and Pu.
235 Cases lA through 5B compare U loaded at fractional heights of 1.0, 0.7 and 0.4; Cases 2A through 5B compare the effects of g vs. water as the volume filler.
The results show that for U the full height loading is the most reactive configura-tion, and that for partial height loadings air is the more reactive filler material.
Based on these results, only air was
Attachment D to-S#87-2738 Page.12 used in subsequent calculations with the 2330 and 239Pu fissile material s..
233 Cases 6 through 9 compare 0 loaded at fractional heights of 1.0, 0.7, 0.4 and 0.3125. This series was expected to show an
' initial increase in koff.with lowering fuel height since the full' height-H/X of 539 represents;a slightly' overmoderated con-dition.
The results show that there is a small initial in-o crease in k fr with lowering ~ fuel height, however, k drops l
rapidlywit0thecontinuedloweringoffuelheight.,ff 1
i J
Cates 10,-11 and 12 were low neutron history runs with 239Pu at j
fractional heights of 1.0, 0.7 and 0.4.
The initial H/X of.624 j
is far.below optimum (-900), so.it was expected based on the previous calculations that k ff would strictly decrease.with decreasingfuel, height,andIheresultsagreewithexpecta-tions.
H The. consistent trend shown in this fuel height-and H/X~
parameter' study for.the NRBK-41 cask geometry is that the nega-tive reactivity incurred in changing the H/X from a. condition of full height (near optimum H/X) to the partial height condi-tions analyzed is of a larger absolute value than the positive reactivity incurred from changing the height / diameter ratio from full height. to the optimum condition of H/D at unity or 0.3125 times the full height condition.
In the cases in which 2330 is loaded into the cask cavity, the full height loading is a slightly overmoderated condition. An initial decrease in fuel height therefore has positive reac-tivity contributions from both H/X and height / diameter changes.
However, the results shown by Cases 6A thorugh 78 for this con-dition indicate that the reactivity increase is slight as the k rfs agree within the statistical uncertainties. As fuel e
height is lowered further and H/X passes through optimum to an undermoderated condition, k is seen'to decrease. Con-sequently, it is concluded Ikat for cases involving fuel load-ings in the more restrictive geometry of the NRBK-41 inserts, lowering the fuel height will decrease the k rr because the e
starting condition at full height for any of the fissile materials in this analysis is undermoderated.
This affirms the assumption that the full height feel loading is the most reat-tive configuration for the geometr es associated with the NRBK-41 cask inserts'as well as the modified NRBK-41 cask insert or MIN-41 container.
I 6.0
[riticality Benchmark Results Criticality benchmark experiments similar to the large array of NRBK-41 casks analyzed for this evaluation do not exist. However, the results
4 l
Attachment D to S#87-2738 Page 13 reported in Reference (5) for experiments involving 233U as well as 2350 critical assemblies and comparisons to calculated representations showed T'
that RCP01 underestimates the k by 1 to 2% for large, low leakage as-sembliesandoverestimatesthek,ff ff by 1 to 2% for small, high leakage units. The total neutron leakage of the assemblies ranged from.014 to 0.47.
The neutron leakages associated with the large arrays of NRBK-41 casks analyzed are less than 0.01.
Consequently, a kcritic 1 of 0.98 is assumedforthisanalysistoaccountforthegeometricalddferences that would potentially exist between large critical assemblies and the RCP01 representations.
a Any uncertainties in keffectives resulting from the differences in the nuclear cross data are considered very small, less than 1% based on the criticality benchmark experiments identified below and comparisons made to RCP01 calculated representations:
(1) The analyseg3gf 32 gggogeneous aqueous cMcal expeHments Gat use either U or U in a single spherical of cylindrical con-tainer (Reference 5).
(2) 'The analysis of plutonium aqueous homogeneous systems in both a
cylindrical and spherical geometry (Reference 3).
233 (3) The analyses of 002 - Th02 critical assemblies associated with the LWBR program (References 6 and 7).
The first two of these sets are based on ENDF/B-IV nuclear data while the third set is based on the LWBR nuclear calculational model described in Reference '8.
All the RCP01 calculated k ff results for these criti-cals were greater than 0.99.
Since the fuei and water nuclear data used in the NRBK-41 cask analyses differ very little from or are the same as the nuclear data used in the analyses of the criticals identified above, a reactivity bias of no more than 1% op is assigned to the calculated ke fe to account for the uncertainties in the nuclear cross section sefs,ctbeer than the lead cross section data used in this analysis.
o Based on the evaluation of the lead cross section set, showa in Section 5.1, an additional reactivity bias of 1% Ap is included.
In summary, a kc itical of 0.98 is assumed to account for the geometri-caldifferencesbetweenexperimentandtheRCP01representationof large, low leakage arrays. Moreover, a conservative 2% Ap reactivity bias is assigned to the calculated maximum k rree to account for un-certainties in the nuclear cross section data usebjf,or this analysis, e
o
Attachment D to S#87-2738 Page 14 7.0 Minimum Marain to Criticality From Table 2, the highest keff result is obtained for the damaged array of 20x20x20 loaded NRBK-41 casks.
The k is 0.892
.037.
The maxi-is 0.949 and includes the 2% 6 p,7freactivity bias for cross sec-mum keff tion uncertainties.
Based on the minimum shutdown formulation shown below and a kc it 0.98, then the minimum margin to criticality or shutdownis3.5xpinamostreactiveanddamagedconfiguration In view of the implausibility of assembling 8000 fuel loaded casks shutdown margin is considered sufficient to affirm the nuclear safety of the NRBK-41 cask packaging and shipment.
Minimum Shutdown Marain Formulation min kminimum kmaximum f) %
=
shutdown crit eff x 100%
kminimum kmaximum crit eff 8.0 References 1
WAPD-TM-1267, "RCP01 - A Monte Carlo Program for Solving Neutron and Proton Transport Problems in Three-Dimensional Geometry with Detailed Energy Description," dated August 1978.
2 WAPD-TM-823(L), "XAP - A Multigroup Cross Section Library System,"
dated February 1971.
3 WAPD-TM-1480, " Monte Carlo Analysis of Pu-H 0 and U0 Pu0 -H 0 2 2 Critical Assemblies with ENDF/B-IV," dated pril 198 4
Nuclear Science and Engineering 44, "The Reactivity Effect of Replacing a Water Reflector with Lead," pp. 413-422 dated 1971.
5 WAPD-TM-1299, " Analysis of Homogeneous 2330 and 2350 Critical As-semblies with ENDF/B-IV Data," dated October 1977.
6 WAPD-TM-1117, "BMU Series of 2330 Fueled Critical Experiments,"
dated January 1975.
7 WAPD-TM-1101,"2330 0xide-Thorium 0xide Detailed Cell Critical Ex-periments," dated October 1974.
8 WAPD-TM-1314, "The Calculational Model Used in the Analysis of Nuclear Performance of the Light Water Breeder Reactor (LWBR),"
dated August 1978.
Attachment D to S#87-2738 Page 15 Table 1. NRBK-41 Cask and IN-41 Container l
XAP Library Nuclides, Atomic Densities and Fission Soectra used in RCP01 calculations Isotope /
Fast Thermal Atomic Density
- Material Group ID Group ID (Atoms /b-cm)
A.
Resonance Nuclides U-233 U233-14 U233LWB2B 1.748 x 10'6 4
U-234 U234-5 U234-5 3.145 x 10 4 U-235 U235-59 U235-59 2.447 x 10~7 U-236 U236-11 U236-11 4.591 x 10 6 U-238.
U238-15 0238-15 2.142 x 10-Pu-239 PU239-6 PU239-6 1.480 x 10-4 Pu-240 PU240-6 PU240-6 4.741 x 10-6 Pu-241 PU241-6 PU241-6 2.294 x 10-7 B.
Smooth Nuclides Hydrogen H-22 H-22 6.686 x 10~2 Oxygen 0-28 0-28 3.343 x 10-2 Stainless Steel 304 SS-8 SS-8 8.72 x 10-2 2
Lead Pb-4 Pb-3 3.306 x 10 2 Concrete CONCRT-2 CONCRT-2 5.817 x l 1.0x10g' Air (Nitrogen) N-1 N-1 C.
Fission Spectra U-233:
SPEC (29)
U-235:
SPEC (30) i Pu-239:
SPEC (31)
D.
Thermal Scattering Kernel for Hydrogen bound in H 0:
1002H073 (Haywood Kernel) 2
- Atomic densities shown here for fuel isotopes are based on concentrations smeared over IN-41 container volume.
The atomic densit when used in composition with CONCRT-2, are 2.418 x 10'{es of H-22 and 0 28, and 1.407 x 10-5, respectively.
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Attachment D - to S#87-2738 Page 18 Table 4 Criticality Results for 10x10x10 Undamaged Arrays 235, 2330 and 239Pu of NRBK-41 casks with varied active fuel heiahts for 0
Fuel & Moderator (I)
Material Above(2)
Case Fuel Fractional Height H/X Fuel & Moderctor keff + 95% Cl. Avg. Eeff + 95%
C.
lA 350g 235U 1.0 385
.8527 +.00841
.8588 1 0072).8558 +.0095 1B per cask 1.0 385 2A 350g 235U 0.7 270 Air
.8457 +.0100'2
.8503 1 0080 y.8480 +.0109 2B per cask 0.7 270 Air 3A 350g 2350 0.7 270 Water
.8303 +.00891.8300 ;..105
.8296 1 0090 ;
3B per cask 0.7 270 Water' 2350 0.4 154 Air
.7846+.0106}.778910125 4A 350g
.7730 1 0120 4B per cask 0.4 154 Air 5A 350g 2350 0.4 154 Water
.7587 +.0072)
SB per cask 0.4 154 Water
.7575 ~.0085 3 7581 +.0034 6A 250g 233U 1.0 539
.8375 3 0139}.8366
.8356 +.0063 6B per cask 1.0 539 7A 250g 233U 0.7 377 Air
.8495 1 0090)2.8512
.8530 +.0109 7B per cask 0.7 377 Air 1
8A 250g 2330 0.4 216 Air
.8126 +.01022.81171 0H4
.8109 3 0096$
80 per cask 0.4 216 Air 9A 250g 233U 0.3125 168 Air
.7639 +.01067
.76141 0113).7627 +.0113 98 per cask 0.3125 168 Air 218.75g (39 u 2 P 1.0 624 Air
.873 +.065 10 per cask 3)
~
11 218.759 239 u 0.7 437 Air
.845 +.038 P
per cask 12 218.75g 239Pu 0.4 250 Air
.763 +.029
~
per cask Notes:
(1) Cask cavity full height is 16 inches.
(2) The remainder of the cask cavity volume not occupied by the fuel / moderator (3)350 grams 23qcupiedbythematerialshown.
mixture is q U divided by equivalence parameter 1.6 = 218.75 grams 239Pu.
l
FIGURE 1 'to S#87-2738 IN_41 CONTAINER Page 19 9
i SIDE VIEW 304 stainless steel 1.D.
4.50+.000
.014
- 15.85 +.03 14.06 +.03 x
P dj+.000-H
.010 Dimensions are in inches NRBK-41 SHIPPING CASK 1
SIDE VIEW l.25 +.03 STEEL i
LEAD 11.00 +.00
.03 STEEL d2 +.06 T_
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CASK
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Attachment D to S#87-2738 Figure 2 90 tilfi-41 Con tainer
~ " ~
a
.- 4. 44 5
- All dimensions are in inches.
13.629 12.774
.L h-4.94
--+l Modified NRBK-41 Shipping Cask with Tnermal Shield and Baseplate q p. 0.062 air gap
-4 k-0.25 thermal shield 1
u T
Original NRBK-41 Cask i
0.41 dimensions are unchanged 35.13 0.62 m m
~
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s Attachment D to 5887-2738 Page 21
- /
FIGtil:[ 3 1
RCP R[PRI SIN 1 A110f!,10P VI[l! tir:11 C(Lt
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1N111 AL ASSEMBLY 1
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Attachment D to Sf87-2738 Page 22 l
Figure 4 RCP01 Representation of 10x10x10 Undamaged Array of NRBK-41 Casks; X, Y Slice 7soo.a3 i
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Attachment D to S#87-2738 Page: 23
' Figures RCP01 Representation of 10x10x10 Damaged Array of NRBK-41 Casks; X, Y Slice -
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. Attachment E'to:
'Sf87-2738 Page 1 l
I I
L.
PROPOSED, REVISED CERTIFICATE OF COMPLIANCE FOR THE~
NRBK-41 CASK i:
i,
1 DOE F 58231 U.S. DEPARTMENT OF ENERGY (5 85)
CERTIFICATE OF COMPLIANCE OMB Approesi (Forrnerly FV.618)
For Radioactive Materials Packages No 1910 2000 1a. Certificate Numoer 1b Revision No ic. Package Identification No.
Id. Page No.
- 10. Total No. Pages USA /5814-1/B( )F(DOE-NR) 3(Proposed)
l
- 2. PREAM BLE 2a Thts certificate is issued under the authority of 49CFR Part 173.7 (d).
2b.
The packaging and contents cescribed in item 5 below, meets the safety stan6,as se' forth in subpart E," Package Approval Standards" and subpart F, " Package and Special Form Tests" Title 10, Code of Feoeral Regulations, Part 71.
2c.
This certificate does not relieve the consignor from compliance with any requirement of the regulations of the U.S. Departrrent of Transportation or other applicable regulatory agencies, including the government of any country through or into which the package will be transported.
- 3. This certi4cate is issued on the basis of a safety analysis report of the package design or application-(1) Prepared by (Name and address T (2) Title and identification of report or application:
(3)Date:
March 11. 1968 Battelle Memorial Institute
" Safety Analysis Report for Columbus, Ohio Radioactive Materials Shipping For: Bettis Atomic Power Lab Cask #NRBK-41", as amended P. O. Box 79 Wott Mifflin. Da-1R177-nn74
- 4. CONDITIONS This certificate is conditional upon the fu filling of the applicable Operational and Quality Assurance requirements of 49CFR parts 100199 and 10CFR Part 71. and the conditions specified in item 5 below.
i 5 Description of Packaging and Authorized Contents. Model Number Fissile Class. Other Conditions, and
References:
EBK-41 RADI0 ACTIVE MATERIALS SHIPPING CASKS This revision is applicable to NRBK-41 casks which have been modified to include thermal shields in accordance with Westinghouse Drawing 1755E01.
For casks which do not have thermal shields, Revision 2 of this Certificate of Compliance remains in effect until either:
(1) the casks are modified, or (2) the applicable expiration date for Revision 2, April 30, 1988, occurs.
a.
Description of Packaaina Top-loading, cylindrical, lead-shielded, 304L stainless steel clad cask for the shipment of irradiated test specimens.
Cask has a one-inch thick stainless steel plate welded to the bottom. A second one-inch thick stainless steel plate with a 1/8-inch deep, 25.5-inch diameter recess is welded to the first plate to provide a thermal break for the bottom surface of the cask.
In addition, the cask has bolted-on, all-welded, 48-inch square "I" beam skids utilized to distribute the cask load. The cavity of the cask is formed by 1/4-inch thick stainless steel surrounded by poured lead shielding with a 1/2-inch thick, stainless steel, cylindrical sheet outer shell. Also, the cask has a seal-welded,1/4-inch thick, stainless steel outer thermal shield which provides a 1/16-inch air gap between the outer surface of 6a. Date of issuance.
l 6b. Expiration Date: ~ February 1, 1991 I
f FOR THE U S DEPARTMENT OF ENERGY
- 70. AddresslotDOEIssvong Officel 7b Signature. Name. and Titte lot DOE Appronng Ottocoat) l Naval Reactors U. S. Department of Energy C. H. Schmitt Washington, D.C. 20585 Deputy Director, Naval Reactors L
- _ _ _ - _ _ _ _. Certificate of Compliance USA /5814-1/B( )F (DOE-NR),
f Rev. 3 (Pr' posed) l 5.a.
Description of Packaaina (Continued) the cask outer shell and the inside surface of the thermal shield. The cask may be wrapped in polyvinyl chloride when shipment is made in a closed vehicle.
The physical characteristics of the cask are as follows:
Outer Cavity Approximate Thickness Dimensions Dimensions Loaded of Lead Ref Drawing Model #
M Heiaht & Lenath Weiaht (Lbs)
Shieldina (W)
HRBK-41 27.16" 46.98" 5"
16" 9000 10" 1755E01 b.
Authorized Contents The authorized contents of the package consist of fissile and radioactive materials which must be contained in sealed inner containers.
An inner container is required for all shipments, both fissile and non-fissile.
There is only one inner container authorized for use with the NRBK-41 cask.
This inner container is designated the
" MIN-41".
Inner Container Description and Restrictions The housing of the " MIN-41" container is a seamless stainless steel pipe, 4.95 inches in diameter (0.25 in, nominal wall thickness over most of its length increasing to 0.537 in. in the closure region) and 15.87 inches in length.
The bottom of the container is formed by a 0.38 inch thick plate welded inside the l
end of the housing.
The top closure consists of:
(1) a c-ring which is installed in a step in the housing, (2) a flat plate (cap) which is used to compress the c-ring and which incorporates a standpipe enclosing a leak test fitting, (3) a plug assembly which mates with the upper end of the housing by j
means of a breech lock arrangement, and (4) eight screws which are screwed through the plug assembly and tightened against the cap to compress the c-ring and secure the plug assembly within the housing. The " MIN-41" container is fabricated in accordance with Westinghouse Drawing 2D77456.
The " MIN-41" container must be dry loaded and may not be utilized to ship liquids.
j The shipper is responsible for ensuring that there are no adverse chemical reactions between the specimens and the inner container.
c.
Fissile Class Except as exempted by 10CFR71.53, casks containing fissile materials (see Item 5.d.(1) below) shall be shipped as Fissile Class I.
d.
Other Conditions (Restrictions)
(1) The fissile content of the NRBK-41 cask shall be limited to a maximum of 350 l
235 235 equivalent grams of U
, where the number of equivalent grams of U is determined by the equation: 1 x grams of U235 + 1.4 x grams of U233
+ 1.6 x grams of plutonium.
_ ______-_-- _ _ _ _ - Certificate of Compliance l
Rev. 3 (Proposed)
'5.d.
Other Conditions (Restrictions) - (Continued)
(2) The decay heat load of the specimens being shipped must be limited to 900 BTU /hr for an exclusive use shipment or 250 BTV/hr for a non-exclusive use shipment. Adherence to the applicable limit will ensure that temperatures on the exterior surfaces of the NRBK-41 cask will be within limits.
(3) Dose rates on the surface of the loaded cask must be limited to 200 mrem /hr.
When the cask is to be wrapped in polyvinyl chloride, dose rate measurements must be taken prior to wrapping the cask.
(4) The contents must be limited such that at a temperature of 580 F, the total internal pressurization of the MIN-41 inner container, including any contribution due to pressure-generating decomposition of the cargo, would not exceed 114 psia.
(5) All shipments must be dry (i.e. shipment of liquids is not permitted).
]
(6) All shipments must be contained in sealed inner containers. The inner l
container designated for use with the NRBX-41 cask is the " MIN-41" container.
l (7)
The cask may be wrapped in polyvinyl chloride if the shipment is to be made in i
a closed vehicle.
(8) Hydrogenous moderation of the package contents is permitted.
l (9)
The inner container must be tested for leaktightness within one year prior to use to a minimum sensitivity of 10-7 atm-cm /sec.
3 (9)
Prior to each shipment, the inner container must be leak tested after assembly to a minimum sensitivity of 10-3 atm-cm/sec.
j 3
i e.
References I
l None f.
Additional Information The Division of Materials Licensing, in their letter to Naval Reactors of April 4,1969, concurred that the NRBK-41 cask met the requirements of 10CFR71.
l l
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