ML22158A273

ATR SAR Complete Rev. 17_Part2

ML22158A273
Person / Time
Site:	07109330
Issue date:	05/31/2022
From:	US Dept of Energy (DOE)
To:	Storage and Transportation Licensing Branch
P SAVEROT NMSS/DFM/FFLB 3014157505
Shared Package
ML22158A270	List: ML22158A271 ML22158A272 ML22158A273 ML22158A274 ML22158A275 ML22158A276 ML22158A277 ML22158A278
References
Download: ML22158A273 (173)

Text

Docket No. 71-9330 ATR FFSC Safety Analysis Report Rev. 17, May 2022 2.12.1-19 Figure 2.12.1 CN1-1 Impact Damage Figure 2.12.1 CN1-1 Impact on Closure Handle

Docket No. 71-9330 ATR FFSC Safety Analysis Report Rev. 17, May 2022 2.12.1-20 Figure 2.12.1 CD1-1 Drop Orientation Figure 2.12.1 CD1-1 Impact Side Stiffening rib locations

Docket No. 71-9330 ATR FFSC Safety Analysis Report Rev. 17, May 2022 2.12.1-21 Figure 2.12.1 Opening of CTU Following CD1-1 Figure 2.12.1 Inspection of Payload Following CD1-1

Docket No. 71-9330 ATR FFSC Safety Analysis Report Rev. 17, May 2022 2.12.1-22 Figure 2.12.1 Inspection of Fuel Element Following CD1-1 Figure 2.12.1 Inspection of Closure Assembly Following CD1-1

Docket No. 71-9330 ATR FFSC Safety Analysis Report Rev. 17, May 2022 2.12.1-23 Figure 2.12.1 CD2.A-1 Figure 2.12.1 Index Lug Near Closure End, CD2.A-1

Docket No. 71-9330 ATR FFSC Safety Analysis Report Rev. 17, May 2022 2.12.1-24 Figure 2.12.1 Index Lug Near Bottom End, CD2.A-1 Figure 2.12.1 View of Closure Following CD2.A-1 Areas of pinching between body and closure Impact side Impact side

Docket No. 71-9330 ATR FFSC Safety Analysis Report Rev. 17, May 2022 2.12.1-25 Figure 2.12.1 CD2.B-1 Drop Orientation Figure 2.12.1 CTU Position Following CD2.B-1 Drop

Docket No. 71-9330 ATR FFSC Safety Analysis Report Rev. 17, May 2022 2.12.1-26 Figure 2.12.1 Index Lug Near Bottom End, CD2.B-1 Figure 2.12.1 CTU in Chiller Unit

Docket No. 71-9330 ATR FFSC Safety Analysis Report Rev. 17, May 2022 2.12.1-27 Figure 2.12.1 CD3-1 Drop Orientation Figure 2.12.1 CTU Following CD3-1 Impact Impact surface

Docket No. 71-9330 ATR FFSC Safety Analysis Report Rev. 17, May 2022 2.12.1-28 Figure 2.12.1 Deformation Near Closure End Following CD3-1 Figure 2.12.1 View of Closure Following CD3-1 Sheared locking pin Previously bent locking pin

Docket No. 71-9330 ATR FFSC Safety Analysis Report Rev. 17, May 2022 2.12.1-29 Figure 2.12.1 CD4-1 Drop Orientation Figure 2.12.1 View of Impact End Following CD4-1

Docket No. 71-9330 ATR FFSC Safety Analysis Report Rev. 17, May 2022 2.12.1-30 Figure 2.12.1 View of Side Bowing Following CD4-1 Figure 2.12.1 CP3-1 Drop Orientation - Front

Docket No. 71-9330 ATR FFSC Safety Analysis Report Rev. 17, May 2022 2.12.1-31 Figure 2.12.1 CP3-1 Drop Orientation - Front Figure 2.12.1 CTU Following CP3-1 Impact Minor deformation on rib from impact

Docket No. 71-9330 ATR FFSC Safety Analysis Report Rev. 17, May 2022 2.12.1-32 Figure 2.12.1 CD5-1 Drop Orientation Figure 2.12.1 CTU Following CD5-1 Impact

Docket No. 71-9330 ATR FFSC Safety Analysis Report Rev. 17, May 2022 2.12.1-33 Figure 2.12.1 CD5-1 Impact Damage on Bottom 180º Side Figure 2.12.1 CD5-1 Impact Damage on Closure End Impact corner

Docket No. 71-9330 ATR FFSC Safety Analysis Report Rev. 17, May 2022 2.12.1-34 Figure 2.12.1 CD5-1 Impact Damage on Closure Area Figure 2.12.1 CD2.C-1 Drop Orientation Impact corner

Docket No. 71-9330 ATR FFSC Safety Analysis Report Rev. 17, May 2022 2.12.1-35 Figure 2.12.1 Side View of CTU Following CD2.C-1 Drop Figure 2.12.1 Index Lug Near Closure End, CD2.C-1 Index lug pressed flush

Docket No. 71-9330 ATR FFSC Safety Analysis Report Rev. 17, May 2022 2.12.1-36 Figure 2.12.1 Cracked Weld Under Index Lug, CD2.C-1 Figure 2.12.1 CP2-1 Drop Orientation

Docket No. 71-9330 ATR FFSC Safety Analysis Report Rev. 17, May 2022 2.12.1-37 Figure 2.12.1 CTU Following CP2-1 Impact Figure 2.12.1 CP2-1 Impact Damage Impact area

Docket No. 71-9330 ATR FFSC Safety Analysis Report Rev. 17, May 2022 2.12.1-38 Figure 2.12.1 CP1-1 Drop Orientation Figure 2.12.1 CTU Following CP1-1 Impact

Docket No. 71-9330 ATR FFSC Safety Analysis Report Rev. 17, May 2022 2.12.1-39 Figure 2.12.1 CP1-1 Impact Damage (Shown Index Lugs Down)

Figure 2.12.1 Attempted Closure Removal Deformed TID post

Docket No. 71-9330 ATR FFSC Safety Analysis Report Rev. 17, May 2022 2.12.1-40 Figure 2.12.1 Exposure of Thermal Shield Figure 2.12.1 Insulation After Removal of Thermal Shield

Docket No. 71-9330 ATR FFSC Safety Analysis Report Rev. 17, May 2022 2.12.1-41 Figure 2.12.1 Middle Insulation After Removal of Thermal Shield Figure 2.12.1 Bottom End Plate Condition

Docket No. 71-9330 ATR FFSC Safety Analysis Report Rev. 17, May 2022 2.12.1-42 Figure 2.12.1 View of Inner Tube at Closure End Figure 2.12.1 Inner Tube Deformation at Closure End Flattening of FHE endplate Inward deformation of inner pipe

Docket No. 71-9330 ATR FFSC Safety Analysis Report Rev. 17, May 2022 2.12.1-43 Figure 2.12.1 End View (Bottom) of Opened CTU Figure 2.12.1 Removal of ATR Fuel Element Broken end box of fuel element Pieces of neoprene from FHE

Docket No. 71-9330 ATR FFSC Safety Analysis Report Rev. 17, May 2022 2.12.1-44 Figure 2.12.1 Fuel Handling Enclosure Deformation Figure 2.12.1 ATR Fuel Element Inspection FHE deformation greatest near closure end

Docket No. 71-9330 ATR FFSC Safety Analysis Report Rev. 17, May 2022 2.12.1-45 Figure 2.12.1 ATR Fuel Element at Head End Figure 2.12.1 ATR Fuel Element Damage at Bottom End

Docket No. 71-9330 ATR FFSC Safety Analysis Report Rev. 17, May 2022 2.12.1-46 Figure 2.12.1 Top View ATR Fuel Element at Bottom End Figure 2.12.1 ATR Fuel Element Fuel Plates Left Side

Docket No. 71-9330 ATR FFSC Safety Analysis Report Rev. 17, May 2022 2.12.1-47 Figure 2.12.1 ATR Fuel Element Fuel Plates Right Side

Docket No. 71-9330 ATR FFSC Safety Analysis Report Rev. 17, May 2022 2.12.1-48 This page intentionally left blank.

Docket No. 71-9330 ATR FFSC Safety Analysis Report Rev. 17, May 2022 2.12.2-1 2.12.2 Certification Tests on CTU-2 This report describes the methods and results of a series of tests performed on the Advanced Test Reactor (ATR) Fresh Fuel Shipping Container (FFSC) transportation package, shown in Figure 2.12.2-1. The objective of testing was to conduct drop tests in accordance with the requirements of 10 CFR 71, §71.71 Normal Conditions of Transport (NCT), and §71.73 Hypothetical Accident Conditions (HAC). This test was primarily directed at verification of the loose fuel plate basket structural integrity and the performance of the package insulation. The package and ATR fuel element payload performance are supported by the tests described in Section 2.12.1, Certification Tests on CTU-1.

Testing was performed at HiLine Engineering in Richland, Washington on May 17, 2007. Color photographs and videos were taken to document the test events and results.

2.12.2.1 Overview There are three primary objectives for the certification test program:

1. To demonstrate that, after a worst-case series of HAC free drops, the package maintains containment of radioactive contents.

2. To demonstrate that, after a worst-case series of HAC free drops, geometry of both the fuel and package are controlled as necessary to maintain subcriticality.

3. To demonstrate that, after the free drops, the package retains the thermal protection necessary to maintain the fuel below its melting point during the thermal evaluation.

Several orientations were tested to ensure that the worst-case series of free and puncture drop events had been considered. Post-impact examination demonstrated that the package sufficiently met the design objectives. The specific objectives of this test were to demonstrate:

Any displacement of package insulation and/or thermal shields are bounded in the thermal analysis, Reconfiguration of the loose fuel plate basket and/or loose fuel plate payload is bounded in the criticality analysis.

2.12.2.2 Pretest Measurements and Inspections The ATR FFSC packaging (serial number CTU1), loose fuel plate basket (serial number 1),

and simulated ATR loose fuel plates were received at HiLine. The packaging and payload are identified as ATR FFSC Certification Test Unit CTU-2. The components arrived fully constructed and ready for testing.

The ATR loose fuel plates were simulated. The payload was comprised of a combination of 2-and 4-inch wide,.06-inch thick, 5052H32 aluminum flat plates. All plates were 49.5-inches long. There were 15, 2-inch wide plates and 10, 4-inch wide plates making up a total payload weight of 20.7 lbs.

The CTU was dimensionally inspected to the drawings at the fabricator and the fabrication records forwarded to PacTec. A Certificate of Compliance was issued by the fabricator of the CTUs documenting compliance with the fabrication drawings. Minor discrepancies

Docket No. 71-9330 ATR FFSC Safety Analysis Report Rev. 17, May 2022 2.12.2-2 between the drawings and CTUs were identified and independently evaluated. The evaluations concluded that the discrepancies were minor and would not significantly affect the CTU during testing.

There were five fabrication deviations associated with the S/N CTU1 package fabrication:

The 3/8-16 UNC index lug screws were obtained without specified ASTM F-879 certifications.

The #10-24 UNC closure handle screws were obtained without specified ASTM F-879 certifications.

Chemical overtesting of the package body closure plate material identified a manganese content 0.02% above the ASTM A479 maximum allowable.

The tap failed when tapping one of the four #10-24 tapped holes for the closure handle screws. As a result, one of the four tapped holes had full threads to a depth of

.44-inches rather than the specified.5-inches.

The handle width is specified to be 7.5 +/-.3-inches. When measured in the free state (not secured to the closure), the handle width was undersized by approximately 0.1-inches.

Other deviations relative to the CTU are the absence of the stainless nameplate and the use of temporary rigging attachments. These items are also insignificant relative to the weight of the CTU and their impact upon the drop tests.

2.12.2.2.1 Component Weights Component weights were measured and recorded as shown in Table 2.12.2-1.

2.12.2.2.2 Drop Test Pad Measurement and Description The drop pad consists of a 7-foot square x 5-foot thick concrete block covered with a 6-foot square x 2.5-inch thick steel plate. The estimated weight of the pad is greater than 44,000 lbs.

Thus the test pad was qualified as an essentially unyielding surface for the approximately 300 lb CTU.

2.12.2.2.3 Equipment and Instruments Instrumentation used for the component weights and drop tests is given in Table 2.12.2-2.

Calibrated test and measurement equipment used were the weight scale and temperature meter. Those two instruments were calibrated in accordance with HiLine procedures. It is noted that the HiLine calibration procedures require National Institute of Standards and Technology (NIST) traceability and that the HiLine records adequately demonstrated that the calibrations were NIST traceable.

A plumb bob with a stretch resistant string was used to determine the appropriate drop height. HiLine project personnel under the supervision of PacTec personnel measured the plumb bob and string using steel tape measures. The angle of the CTU prior to each drop was measured using a mechanical inclinometer.

One low speed digital video camera was used to record the drop events. In addition, color photographs were taken to document the testing.

Docket No. 71-9330 ATR FFSC Safety Analysis Report Rev. 17, May 2022 2.12.2-3 2.12.2.3 Summary of Tests and Results 2.12.2.3.1 Initial Conditions All three HAC drops, CD1-2, CD3-2, and CD4-2, were performed at ambient temperature.

Ambient temperature and the package surface temperature was recorded before and after each drop. After each drop the closure was removed and the basket inspected. The basket was reassembled (the basket screws tightened to the finger tight condition) and the package re-closed for the following test. One tie wrap (securing the loose fuel plate payload) failed in the CD1-2 test and the second tie wrap failed in the CD3-2 test. Neither of the two tie wraps were replaced between tests.

2.12.2.3.2 Summary of Testing Table 2.12.2-3 identifies the testing performed on the ATR FFSC CTU.

2.12.2.4 Certification Tests 2.12.2.4.1 Drop Tests The three CTU-2 HAC drop tests were performed to augment the CTU-1 tests for the package, and to demonstrate acceptable performance of the loose fuel plate basket payload. In CTU-1, the package was subjected to end drops on both the closure and the bottom ends of the package.

CTU-2 restricted the end drop test to just the bottom end to properly assess axial insulation displacement.

There were no NCT or puncture bar tests performed on the package, since CTU-1 adequately demonstrates acceptable package performance under those conditions. The two side drops subjected the loose fuel plate basket and simulated fuel to worst case impact conditions with the basket oriented perpendicular and parallel to the target surface.

The test identification numbering reflects the same drop orientation as performed in CTU-1. For example, CD3-2 is the same orientation as the third HAC drop in CTU-1, test CD3-1. The -2 identifies this drop as a CTU-2 test.

2.12.2.4.1.1 CD1-2 -Flat (pocket side down) Side HAC Drop The CTU was fitted with swivel lift eyes, and the lift eyes were threaded into the package lift points. This configuration oriented the package such that the package pocket side impacted the target surface. Slings were used to rig the CTU from the swivel lift eyes to the crane remote release hook. Figure 2.12.2-5 illustrates the drop orientation. Initial conditions were as follows:

Ambient temperature:

73 ºF Avg. surface temperature:

78 ºF Time:

10:04 a.m. 5/17/2007 Drop height:

30 ft Following impact, the CTU bounced slightly and landed on the impact side. There was minor visible exterior damage, principally scuff marks, resulting from the drop. Close examination of

Docket No. 71-9330 ATR FFSC Safety Analysis Report Rev. 17, May 2022 2.12.2-4 the package, on the impacted surface side, reveals minor distortion of the outer shell localized at the stiffening ribs. Figures 2.12.2-6 and 2.12.2-7 show the CTU prior to and following the drop.

There was no bowing or other significant visible deformation. There was no visible deformation or rotation of the closure, and the locking pins condition and function were unaffected by the drop.

The basket was not affected by the drop, however the finger operated screws securing the two basket halves were loosened approximately one turn. One fuel tie wrap was broken but the simulated loose fuel plates were not damaged. The simulated fuel plates were replaced in the basket without installing new tie wraps, and the basket closure screws again tightened to the finger tight condition.

2.12.2.4.1.2 CD3 Flat Side HAC Drop (90º from CD1-2)

Following the CD1-2 drop, lift points were welded to the package to enable a side drop rotated 90º from CD1-2 (Figure 2.12.2-8):

Ambient temperature:

78 ºF Avg. surface temperature:

85 ºF Time:

10:50 a.m. 5/17/2007 Drop height:

30 ft The CTU rebounded from the drop pad approximately 1 ft following the 30 ft drop and came to rest on its side (rotated 90º from the drop orientation). As with the CD1-2 event, the outer shell exhibited minor deformation at the stiffening rib locations (reference Figure 2.12.2-9). There was no visible deformation or rotation of the closure, and the locking pins were undamaged and in good working order.

The closure was opened and the basket removed following the drop. The basket exhibited no signs of any deformation but the finger tightened basket screws were loosened approximately 1 turn by the drop.

The basket was opened and it was discovered that the second plastic tie wrap was broken (Figure 2.12.2-10). The simulated fuel plates were found to exhibit no significant damage. The simulated fuel plates were replaced in the basket without installing new tie wraps, and the basket closure screws again tightened to the finger tight condition.

Docket No. 71-9330 ATR FFSC Safety Analysis Report Rev. 17, May 2022 2.12.2-5 2.12.2.4.1.3 CD4 CG over Bottom End (Vertical)

Following CD3-2, the temporary rigging attachments were removed and the CTU rigged for CD4-2 by lifting the package from the closure handle (Figure 2.12.2-11). Initial conditions were recorded as follows:

Ambient temperature:

88 ºF Avg. surface temperature:

90 ºF Time:

11:20 a.m. 5/17/2007 Drop height:

30 ft The CTU appeared to impact slightly off of true vertical; impacting near one corner of the package. This impact dented the lift point feature inward approximately 1/2-inch, and on one adjacent side, bulged out the square outer tube surface by approximately 1/2-inch. Following impact, the CTU rebounded vertically approximately 2-feet, tipped over, and landed on the CD3-2 impact side. There was no overall bowing or of the package or other significant visible deformation. There was no visible deformation or rotation of the closure. Figure 2.12.2-12 shows the bottom end of the CTU following the drop.

There was no visible damage to the closure or the locking pins. The closure was removed and the basket extracted following CD4-2. Damaged to the basket was limited to a small dent at the end of the basket that was situated closest to the package bottom. Upon destructive examination of the package, it was discovered that the weld between the package inner shell and the component at the bottom of the payload cavity had intruded into the payload cavity in a localized area (Figure 2.12.2-13). When the package impacted in CD4-2, the basket was partially supported by that weld bead. The end plate of the basket was slightly deformed (Figure 2.12.2-

14) as the basket seated on the bottom of the package payload cavity. The damage was minor and did not impair the ability of the basket to retain the fuel plates.

The simulated fuel plates experienced localized deformation at the end of the basket closest to the package bottom (Figure 2.12.2-15 and Figure 2.12.2-16). Above this area the simulated fuel plates were not deformed.

2.12.2.5 Post-test Disassembly and Inspection The final acceptance criteria for the ATR FFSC package lies with the criticality evaluation. Any increase in reactivity of the contents resulting from the certification tests must not exceed the allowable as defined in the criticality evaluation. The inspections required to support determination of compliance with the acceptance criteria are identified as follows:

Inspect the outer shell to verify the thermal performance of the package is unimpaired by the free drop events. The thermal analysis assumes that the outer shell is intact such that there is no significant communication between the environment and the outer/inner shell annular space during the thermal event.

Inspect the insulation to verify compliance with the assumptions of the thermal analysis.

Docket No. 71-9330 ATR FFSC Safety Analysis Report Rev. 17, May 2022 2.12.2-6 Inspect the overall package to verify that the package geometry remains within the criticality analyses assumptions.

Inspect the simulated fuel plate payload to verify that the fuel geometry remains within the assumptions of the criticality analyses.

Any deviation of the test results from these acceptance criteria must be reconciled with the criticality evaluation.

2.12.2.5.1 CTU Inspection The CTU-2 was disassembled and inspected on May 17, 2007. Prior to disassembly the exterior dimensions were recorded for comparison to the pre-test condition. Table 2.12.2-4 lists the measured dimensions and Figure 2.12.2-17 identifies the location of the identified measurements.

The closure handle was unaffected by the first two drops. In the CD4-2 drop, the handle was dented when it was struck by the rigging shackle. During the CD4-2 CG over bottom (vertical)

HAC drop, the outer wall bulged out at the bottom end of the package and caused the width of the package to increase from 8 inches to approximately 8 5/8 inches in that area.

The CTU was disassembled systematically by cutting away the outer layers of the packaging using an abrasive saw. The destructive examination was necessary due to the required inspection of the interior insulation. The package was cut lengthwise along two opposite corners and at the ends to expose the thermal shield.

The stainless steel thermal shields were all intact (Figure 2.12.2-18 through Figure 2.12.2-20).

There was minor deformation of the thermal shields at the interface to the stiffening rib. This deformation resulted from the CD4-2 drop and caused the thermal shields to buckle one end and pull away from the stiffening rib at the other end. Figure 2.12.2-21 is typical of this condition.

The gap between the thermal shield and the stiffening rib, where the shield pulls away from the rib, is less than 1/16-inch.

Following documentation of the thermal shields the shields were removed to enable examination of the insulation. For reference purposes the ribs are labeled 1 through 3 (Figure 2.12.2-22).

The number 1 rib is closest to the bottom end of the package.

As can be seen in Figure 2.12.2-23 through Figure 2.12.2-26 the largest gap occurred at the closure end of the package. The gap ranges from 1-inch to 1 3/4 inches at that location. At the rib 3 and rib 2 locations the gap ranged from 1-to 1 1/2-inches. At the rib 3 location the gap ranged from 1/2-to 1-inch. All gaps are within the 1.85-inch gap assumed in the thermal analysis.

Following thermal shield and insulation removal an abrasive saw was used to separate the bottom end plate from the inner tube. Figure 2.12.2-13 illustrates the condition of the end plate.

The endplate showed no drop related deformation and there were no visual indications of broken welds or other damage near the end plate. Using a lathe, the bottom end plate was cut from the insulation pocket to determine the extent of possible insulation compression in the insulation pocket (Figure 2.12.2-27). There was no indication of compression in that region and it was determined that there was no need to open the closure insulation pocket.

Docket No. 71-9330 ATR FFSC Safety Analysis Report Rev. 17, May 2022 2.12.2-7 The inner tube was inspected and, in general, showed no signs of buckling or large deformations.

A minor deformation occurred near the bottom end of the package (Figure 2.12.2-28 and Figure 2.12.2-29) corresponding to the same area of deformation as the outer shell. The tube was bent in that area yielding a slight outward bulge of about 1/16-inch and, closer to the weld between the inner shell and the package bottom, an inward deformation of approximately 1/4-inch. These deformations were localized and did not impair free movement of the basket in the payload cavity. There were no weld failures.

The closure assembly remained fully functional throughout the test series. The only damage to the closure was the handle deformation caused by the rigging shackle. The locking pins and the engagement lugs showed no signs of any deformation. The closure could be freely removed and installed through the tests.

In conclusion, CTU-2 satisfied the acceptance criteria of preventing loss or dispersal of the contents, the outer shell remained intact, the insulation remained within the assumptions of the thermal analysis, and the package and fuel geometry remained greatly unchanged. The deformations of the package and condition of the ATR loose fuel plates were evaluated, against both the criticality evaluation and thermal analysis, and determined to be within the bounds of the assumptions and conditions used to ensure safety.

Docket No. 71-9330 ATR FFSC Safety Analysis Report Rev. 17, May 2022 2.12.2-8 Table 2.12.2 Component Weights Component Weight (lbs)

Body Assembly 224.1 Closure Assembly 8.9 Loose Plate Fuel Basket 29.9 Simulated Fuel Plate Weight 20.7 Package (fully loaded) 283.6 Table 2.12.2 Instrumentation for Drop Tests Item Description Model Serial Number Calibration Due Date Comments Drop Height Indicators N/A N/A N/A String plumb bobs made specifically for this testing. The length was established using a metal tape measure.

Tape Measure N/A N/A N/A 35-ft. steel tape Mechanical inclinometer N/A N/A N/A Used to identify CTU orientation Weight Scale Ohaus, Model CD11 0042508-6BD 7/19/2007 Used to measure weights of CTU components. The scale calibration documents included NIST traceable records.

Temperature meter Carson, Model 4085 41372269 3/1/2008 Handheld temperature reader for measuring ambient temperature and CTU surface temperature.

Meter calibration documents included NIST traceable records.

Docket No. 71-9330 ATR FFSC Safety Analysis Report Rev. 17, May 2022 2.12.2-9 Table 2.12.2.3 - Summary of Testing Test No.

Test Description Comments CD1-2 Flat side drop, pocket side down.

Fuel plates oriented perpendicular to target (see Figure 2.12.2-3).

Flat side drop from 30-feet. No visible damage to package. Both closure locking pins remained in the locked position. Closure could be freely opened and payload extracted. The eight hand tightened screws securing the basket halves together were loose (approximately one turn). No visible damage to basket or simulated fuel plates.

CD3-2 Flat side drop, pockets and index lugs on side. Fuel plates oriented parallel to target (see Figure 2.12.2-4).

Flat side drop from 30-feet. No visible damage to package. Both closure locking pins remained in the locked position. Closure could be freely opened and payload extracted. The eight hand tightened screws securing the basket were loose (approximately one turn).

The plastic wire ties securing the fuel bundle failed as shown in Figure 2.12.2-10. No significant deformation was observed in the fuel plates.

CD4-2 CG over bottom end (vertical)

Vertical end drop from 30-feet; bottom end of package impacting the target. Both closure locking pins remained in the locked position.

Closure could be freely opened and payload extracted. The eight hand tightened screws securing the basket were loose (approximately one turn).

The bottom end of the package was deformed on two surfaces (Figure 2.12.2-12). The surface with the threaded hole was dented inward and the adjacent surface 90º apart was bulged outward.

The surface of the basket end plate contacting the bottom of the package was slightly dented.

The simulated fuel plates were deformed at the bottom end of the basket (Figure 2.12.2-15 and Figure 2.12.2-16).

Table 2.12.2 Package Length Measurements Test ID 1

2 3

4 5

6 7

8 Pre-Test (in.)

72 7/16 72 1/2 72 7/16 72 1/2 72 7/16 72 7/16 72 7/16 72 1/2 CD1-2 (in.)

72 7/16 72 1/2 72 7/16 72 1/2 72 7/16 72 7/16 72 7/16 72 7/16 CD3-2 (in.)

72 7/16 72 1/2 72 7/16 72 1/2 72 7/16 72 7/16 72 7/16 72 7/16 CD4-2 (in.)

72 7/16 72 1/2 72 3/8 72 7/16 72 5/16 72 5/16 72 3/16 72 3/8

Docket No. 71-9330 ATR FFSC Safety Analysis Report Rev. 17, May 2022 2.12.2-10 Figure 2.12.2 ATR FFSC CTU-2 (CTU-2 uses package S/N CTU1)

Figure 2.12.2 Loose Fuel Plate Basket and Simulated Fuel Plates

Docket No. 71-9330 ATR FFSC Safety Analysis Report Rev. 17, May 2022 2.12.2-11 Figure 2.12.2 Basket Orientation in CD1-2 Figure 2.12.2 Basket Orientation in CD3-2

Docket No. 71-9330 ATR FFSC Safety Analysis Report Rev. 17, May 2022 2.12.2-12 Figure 2.12.2 CD1-2 Drop Orientation Figure 2.12.2 CTU Following CD1-2 Impact (impact side facing up)

Docket No. 71-9330 ATR FFSC Safety Analysis Report Rev. 17, May 2022 2.12.2-13 Figure 2.12.2 CD1-2, Extracting Basket Following Drop Figure 2.12.2 CD3-2 Drop Orientation

Docket No. 71-9330 ATR FFSC Safety Analysis Report Rev. 17, May 2022 2.12.2-14 Figure 2.12.2 CD3-2 Deformation at Stiffening Rib Location Figure 2.12.2 CD3 Failed tie wraps

Docket No. 71-9330 ATR FFSC Safety Analysis Report Rev. 17, May 2022 2.12.2-15 Figure 2.12.2 CD4 Drop Orientation Figure 2.12.2 CD4-2 Impact Damage to Package

Docket No. 71-9330 ATR FFSC Safety Analysis Report Rev. 17, May 2022 2.12.2-16 Figure 2.12.2 Weld bead protruding into package payload cavity (inner shell has been removed in this photo)

Figure 2.12.2 Dented area - basket end plate

Docket No. 71-9330 ATR FFSC Safety Analysis Report Rev. 17, May 2022 2.12.2-17 Figure 2.12.2 CD4-2 Impact Damage to Simulated Fuel Plates Figure 2.12.2 CD4-2 Impact Damage to Simulated Fuel Plates (close up view)

Docket No. 71-9330 ATR FFSC Safety Analysis Report Rev. 17, May 2022 2.12.2-18 Figure 2.12.2 CTU Measurement Locations Figure 2.12.2 Thermal Shield Condition, View 1

Docket No. 71-9330 ATR FFSC Safety Analysis Report Rev. 17, May 2022 2.12.2-19 Figure 2.12.2 Thermal Shield Condition, View 2 Figure 2.12.2 Thermal Shield Condition, View 3

Docket No. 71-9330 ATR FFSC Safety Analysis Report Rev. 17, May 2022 2.12.2-20 Figure 2.12.2 Thermal Shields at Interface to Stiffening Rib Figure 2.12.2 Exposed Insulation - Overview

Docket No. 71-9330 ATR FFSC Safety Analysis Report Rev. 17, May 2022 2.12.2-21 Figure 2.12.2 Insulation Gap at Package Closure End Figure 2.12.2 Insulation Gap at Rib #3

Docket No. 71-9330 ATR FFSC Safety Analysis Report Rev. 17, May 2022 2.12.2-22 Figure 2.12.2 Insulation Gap at Rib #2 Figure 2.12.2 Insulation Gap at Rib #1 (nearest impact)

Docket No. 71-9330 ATR FFSC Safety Analysis Report Rev. 17, May 2022 2.12.2-23 Figure 2.12.2 End Plate Insulation Condition Figure 2.12.2 Tube to Bottom End Plate - View 1

Docket No. 71-9330 ATR FFSC Safety Analysis Report Rev. 17, May 2022 2.12.2-24 Figure2.12.2 Tube to Bottom End Plate - View 2

Docket No. 71-9330 ATR FFSC Safety Analysis Report Rev. 17, May 2022 2.12.3-1 2.12.3 Structural Evaluation for MIT and MURR Fuel The ATR FFSC may be utilized to transport a MIT fuel assembly or a MURR fuel assembly.

Both of these fuels are high-enriched aluminum-clad uranium aluminide plate type fuel elements or they may be low enriched aluminum-clad U-Mo alloy fuel elements, similar to the ATR fuel evaluated in this chapter. Since no MIT or MURR fuel elements were included in the drop tests, the following evaluation conservatively estimates a degree of failure and movement of the MIT and MURR Fuel Handling Enclosures (FHE) to develop a worst case pitch expansion of the corresponding fuel elements for evaluation in Section 6.10, Appendix B: Criticality Analysis for MIT and MURR Fuel. By conservatively bounding potential damage and evaluating the exceptional worst case pitch expansion of the MIT and MURR fuel elements the ATR FFSC complies with the performance requirements of 10 CFR §71.

2.12.3.1 Structural Design Discussion A comparison is provided to highlight the similarities and differences between the MIT and MURR designs and the physically tested ATR design. Through this comparison, it is expected that both NCT and HAC testing would result in similar results for the MIT and MURR fuel elements. Similar to the ATR LFPB, the MIT and MURR FHEs are designed to restrict postulated fuel element pitch expansion under the HAC conditions.

The results of NCT conditions on the MIT and MURR payload are assumed to be equivalent to the ATR payload; i.e. there is no damage to the FHE or fuel element under NCT.

For conservatism in evaluating the HAC conditions, the MIT and MURR FHE damage postulated exceeds the results obtained during testing of the ATR payloads. The MIT and MURR FHEs are assumed to separate (fail) and spread apart to permit a worst case reactivity configuration of the fuel elements. The individual fuel plates of the fuel elements are assumed to spread apart uniformly to fill the resulting space.

2.12.3.1.1 Fuel Elements The ATR FFSC packaging is not modified for the use of the MIT and MURR fuel elements. The MIT and MURR FHE are used in place of the ATR FHE or the LFPB within the ATR FFSC packaging. Similar to the ATR FHE and LFPB, the MIT and MURR FHEs are principally fabricated of aluminum construction and secured with stainless steel locking pins.

The MIT and MURR fuel elements are very similar to the ATR fuel element in design, materials, and fabrication. The weight of the fuel elements are 10 lb, 15 lb, and 25 lb, for the MIT, MURR, and ATR fuel elements respectively. All three fuel elements are fabricated of the same fuel type, aluminum-clad uranium aluminide fuel plates, with all fuel plates swaged into the side plates, and include cast or wrought aluminum end boxes. As such, the structural performance of the MIT and MURR fuel types are anticipated to behave very similarly to the ATR fuel element.

Table 2.12.3-1 compares the three fuel element design dimensions. Figure 2.12.3-1 compares the three fuel elements in their overall length and fuel plate length in inches. In this figure, the inside dimension identifies the fuel plate length.

For comparative purposes, an approximate moment of inertia is calculated for all three fuel elements using AutoCAD. The results are presented in Figure 2.12.3-2. The values were

Docket No. 71-9330 ATR FFSC Safety Analysis Report Rev. 17, May 2022 2.12.3-2 determined by taking a cross section of the fuel plate region and selecting the solid boundaries to compute the moments of inertia about the identified axes.

The comparison of the moments of inertia demonstrates that the three fuel elements are similar in stiffness and expected to perform in a similar fashion during NCT and HAC drop events. The length and weight of the fuel elements is clearly bounded by the ATR fuel element. The materials of construction and fabrication techniques are the same for each fuel type. The relatively minor dimensional changes of the ATR fuel element plates as a consequence of the testing identified in Section 2.6, Normal Conditions of Transport, and Section 2.7, Hypothetical Accident Conditions, further justifies the similar performance of the MIT and MURR fuel elements.

Table 2.12.3-1 -Fuel Element Design Component MIT MURR ATR Approximate Weight, lbs 10 15 25 Number of Fuel Plates 15 24 19 Nominal Plate Spacing, in.

.08

.08

.08 Fuel Plate Length, in.

23.00 25.50 49.50 Fuel Plate Thickness, in.

.08

.05

.05,.08,.10 Approximate Fuel Plate Width, in.

2.5 2.0 - 4.3 2.0 - 3.9

Docket No. 71-9330 ATR FFSC Safety Analysis Report Rev. 17, May 2022 2.12.3-3 Figure 2.12.3 MIT, MURR, and ATR Fuel Elements ATR MURR MIT

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Docket No. 71-9330 ATR FFSC Safety Analysis Report Rev. 17, May 2022 2.12.3-5 IX-X = 6.4 in4 IY-Y = 3.8 in4 IX-X = 5.4 in4 IX-X = 11.0 in4 IA-A = 7.7 in4 IY-Y = 10.5 in4 IY-Y = 15.1 in4 IB-B = 2.6 in4 Figure 2.12.3 Fuel Element Moments of Inertia

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Docket No. 71-9330 ATR FFSC Safety Analysis Report Rev. 17, May 2022 2.12.3-7 2.12.3.1.2 Fuel Handling Enclosures The MIT FHE incorporates two end spacers and a two-piece machined aluminum enclosure to protect the MIT fuel element from damage during loading and unloading operations. The enclosure halves are identical segments machined from 6061 aluminum plate. Neoprene rub strips are used to cushion the contact points between the fuel element and enclosure. The end spacers are also fabricated of 6061 aluminum. The end spacers lock the enclosure halves together and are secured using stainless steel ball lock pins. The end spacers also prevent axial movement since the MIT fuel element is much shorter than the package cavity. The weight of the MIT FHE is 25 lb. Figure 2.1-3 illustrates the assembly view of the MIT FHE.

The MURR FHE is designed in the same manner as the MIT FHE. The weight of the MURR FHE is 30 lb. Figure 2.1-4 illustrates the assembly view of the MIT FHE.

The MIT and MURR FHE design is similar to the 30-lb LFPB in that it utilizes machined enclosure halve segments to encase the payload. The use of the enclosure halves makes the MIT and MURR FHEs more robust than the ATR FHE, which weighs 15 lb. The wall thickness of the enclosure halves is 0.19 in compared to the 0.09 in thick sheet used in the ATR FHE. For comparison, the typical machined wall thickness of the LFPB is also 0.19 in thick. The weight of the enclosures and fuel elements are 35 lb, 45 lb, 40 lb, and 50 lb for the MIT payload, MURR payload, ATR payload, and LFPB payload respectively.

Based on the similarity in design and function, the structural and thermal performance of the MIT and MURR FHEs is anticipated to be similar to the physical testing performed using the ATR FHE and LFPB.

2.12.3.1.3 Loose Fuel Plates MIT and MURR loose fuel plates are not evaluated for use within the LFPB.

2.12.3.2 Allowable Damage For HAC tests the MIT and MURR fuel elements are anticipated to perform in a similar manner to the ATR fuel element based on the comparable designs and assembly techniques. To conservatively encompass potential damage, the FHE halves are considered to separate while each half is sized at the extreme tolerances to encourage the maximum space around each fuel element. Based on the maximum space developed by the separated FHE, the fuel element plates separate to create a more reactive configuration for the fuel. The proposed pitch expansion greatly exceeds the results of the physical testing performed on the ATR fuel element.

Axial movement of the fuel element within the package inner tube, which occurs by hypothetical neglect of the FHE end spacers, has no adverse effect on the performance of the ATR FFSC.

Energy dissipated by failure of the spacers would result in lowering the HAC loads to the MIT and MURR elements. However, the structural tests identified that the ATR fuel element survives the impact loads with damage that has no impact on reactivity. The MURR and MIT fuel elements are of similar materials and of similar construction to the ATR fuel elements.

Assuming the spacers to fail with no energy absorption, the impact velocities of the MURR and MIT FHEs on the end fitting of the package would be nearly identical. It is therefore concluded that the damage to MURR and MIT fuel elements is bounded by the damage sustained by the ATR fuel element in the structural tests. However, for conservatism, the fuel plate pitch of the

Docket No. 71-9330 ATR FFSC Safety Analysis Report Rev. 17, May 2022 2.12.3-8 MURR and MIT elements is set to the condition that results in the worst case reactivity under the volumetric constraints presented by the FHEs.

The HAC criticality array model is a 5x5x1 array of packages and all fuel elements are positioned at the same axial location. The FHE end spacers are conservatively neglected and modeled as water. Axial shifting of fuel elements from the modeled configuration would result in a less reactive condition; therefore, failure of the FHE end spacers is not a criticality concern.

For the thermal evaluation, the position of the MIT or MURR fuel element is naturally bounded by the ATR fuel element since its length extends to each end of the package.

The modeled separation of the FHE halves inside the inner tube of the package is determined by using the maximum inner diameter of the packages inner tube and the minimum outer radius of each FHE half as illustrated in Figures 2.12.3-3 and 2.12.3-4. The FHE cavity dimensions are expanded using the maximum tolerance of the parts. Note that this is only hypothetically possible, since this causes the corners of the FHE for both the MIT and MURR to exceed the point of interference with the inner tube wall.

The dimensions for the criticality model of the MIT FHE are determined in the following manner:

Package inner tube maximum inside diameter: Diameter is specified as 6.0 in. OD X 0.12 in. wall thickness +/- 0.030 in. OD and +/- 10% thickness (per drawing 60501-10 and ASTM A269). Resulting maximum ID is 5.814 in.

Minimum outside radius of the FHE half: Radius is specified as 2.8 in +/- 0.2 (per drawing 60501-40). Resulting minimum radius is 2.6 in.

Minimum wall thickness of the FHE half: Wall is specified as 0.19 in +/- 0.06 (per drawing 60501-40). Resulting minimum thickness is 0.13 in.

Maximum cavity height of the FHE half: Wall height specified as 2.82 in +/- 0.06 (per drawing 60501-40). Resulting maximum height is 2.88 in. (which is greater than the 2.6 maximum radius).

Maximum cavity width of the FHE half: Wall width specified as 1.62 in +/- 0.06 (per drawing 60501-40). Resulting maximum width is 1.68 in.

The dimensions for the criticality model of the MURR FHE are determined in the following manner:

Package inner tube maximum inside diameter: Diameter is specified as 6.0 in. OD X 0.12 in. wall thickness +/- 0.030 in. OD and +/- 10% thickness (per drawing 60501-10 and ASTM A269). Resulting maximum ID is 5.814 in.

Minimum outside radius of the FHE half: Radius is specified as 2.8 in +/- 0.2 (per drawing 60501-50). Resulting minimum radius is 2.6 in.

Minimum wall thickness of the FHE half: Wall is specified as 0.19 in +/- 0.06 (per drawing 60501-50). Resulting minimum thickness is 0.13 in.

Maximum cavity height of the FHE half: Wall height specified as 2.00 in +/- 0.06 (per drawing 60501-50). Resulting maximum height is 2.06 in.

Maximum cavity width of the FHE half: Wall width specified as 1.85 in +/-.06 (per drawing 60501-50). Resulting maximum width is 1.91 in.

Docket No. 71-9330 ATR FFSC Safety Analysis Report Rev. 17, May 2022 2.12.3-9 The thermal evaluation in Section 3.6, Thermal Evaluation for MIT, MURR, and Small Quantity Payloads, makes the following conservative assumptions to bound damage to the fuel elements and FHEs as a result of NCT and HAC events.

Idealized contact between the FHE and the package inner tube. The majority of the heat input to the fuel element comes from the radial direction rather than the axial direction.

By maximizing the contact, the greatest heat is transferred. Deformation of the payload would have the effect of reducing the contact area, and therefore reducing the conductive heat input.

Axial movement of the fuel element, as a result of deformation of the FHE end spacers has a negligible effect. The majority of the heat input to the fuel element comes from the radial direction rather than the axial direction (ends). As the fuel element moves closer to the ends of the package the heat input rises. However, the heat input from either end of the package is negligible compared to the heat input received axially from the sides.

Furthermore, any credible axial distance of the MIT and MURR fuel elements to the end of the package is bounded by the ATR fuel element.

The criticality evaluation in Section 6.10, Appendix B: Criticality Analysis for MIT and MURR Fuel, makes the following conservative assumptions to bound damage to the fuel element as a result of HAC events.

Neglecting the function of the end spacers, the two halves are pushed apart to the maximum extent to maximize the available space for pitch expansion.

Although it is not feasible in actual practice to push the FHEs to the center of the array if the two FHE halves are already pushed apart, both the MIT and MURR models are shifted by 0.307-in towards the center of the array.

Fuel element end boxes are not modeled. For criticality purposes, any amount of damage to the end boxes is acceptable.

Note that the MIT and MURR FHEs are sliced off in the corners because such a translation is not possible without interference.

Due to the conservative assumptions utilized for the thermal and criticality evaluations, the allowable damage to the FHEs is considered severe and therefore far exceeding the physical testing results performed using the ATR fuel element and LFPB payloads covered in Section 2.12.1, Certification Tests on CTU-1, and Section 2.12.2, Certification Tests on CTU-2.

For containment purposes, the MIT and MURR fuel element plates must remain intact to prevent the fuel meat from within the fuel plate from exiting the package. The MIT and MURR fuel elements are fully supported over the length of the fuel plates by the FHE enclosure halves. The enclosure halves are specifically designed to fully support each fuel element and minimize any deformation or change in the fuel plate geometry. By design the MIT and MURR FHEs are more robust (thicker side walls) than the ATR FHE and therefore provide better support compared to the testing performed using the ATR fuel element and ATR FHE.

Docket No. 71-9330 ATR FFSC Safety Analysis Report Rev. 17, May 2022 2.12.3-10 Nominal MIT FHE Dimensions Maximum Tolerances Incorporated to Separate FHE Halves Figure 2.12.3 MIT FHE Damage Hypothetical Interference ATR FFSC Inner Tube MIT FHE MIT FHE ATR FFSC Inner Tube

Docket No. 71-9330 ATR FFSC Safety Analysis Report Rev. 17, May 2022 2.12.3-11 Nominal MURR FHE Dimensions Maximum Tolerances Incorporated to Separate FHE Halves Figure 2.12.3 MURR FHE Damage Hypothetical Interference ATR FFSC Inner Tube MURR FHE MURR FHE ATR FFSC Inner Tube

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Docket No. 71-9330 ATR FFSC Safety Analysis Report Rev. 17, May 2022 2.12.4-1 2.12.4 Finite Element Analysis This appendix provides a summary of the information presented in the finite element analysis calculation, reference [1].

2.12.4.1 References

1. Spears, R. E., Selected Sequential Drop Analyses for the ATR FFSC with Heavier LEU Fuel Elements, INL/RPT-22-67296, ECAR-5644, Revision 2, May 2022.

2. Abaqus Version 2021.HF6, Dassault Systemes Simulia Corp., 2021.

3. Snow, S. D., Software Validation Report for Abaqus Standard and Explicit Version 2021.HF6 for Structural Analyses, ECAR-5544, Revision 0, July 2021.

2.12.4.2 Introduction The primary method of demonstration of the ATR FFSC is certification test, as documented in Appendix 2.12.1, Certification Tests on CTU-1 and Appendix 2.12.2, Certification Tests on CTU-2. Finite element analysis is used to evaluate the effect of the increased weight of the ATR LEU and MURR LEU fuel elements on the following:

• Structural integrity of the ATR LEU fuel element in the worst-case HAC free drop impact. The structural integrity of the ATR LEU fuel element is an important factor in the criticality evaluation.

• Structural integrity of the package closure.

The Finite Element Analysis (FEA) simulations performed in this appendix are based on a fine-mesh model using an explicit dynamics formulation appropriate for impact simulations. The models is benchmarked, and several orientations are chosen for analysis to determine the worst-case performance of the ATR FFSC package. Since the MURR LEU fuel element, contained in the MURR FHE, has the same weight as the ATR LEU fuel element contained in the ATR FHE, only the ATR LEU fuel element is modeled. Further, since the MURR criticality evaluation does not rely on a fully intact element, it is not necessary to demonstrate MURR LEU fuel element structural performance.

The FEA model is subjected to simulated free drop impacts representing the 4-ft NCT free drop, the 30-ft HAC free drop, and the 40-inch puncture drop. The results of the FEA simulations demonstrate that the ATR LEU fuel element remains intact, with the fuel plates remaining attached to the fuel elements side plates, essentially in the as-fabricated configuration (not including the end boxes, which have no safety function), and the package closure retains the contents within the confines of the ATR FFSC package.

This appendix is a summary of the analysis. Full details of any aspect of the analysis may be found in [1].

2.12.4.3 Design Input 2.12.4.3.1 Geometry The FEA model includes the ATR FFSC packaging, the ATR FHE, and the ATR LEU fuel element. The model of the ATR FFSC packaging and FHE are built using the drawings in

Docket No. 71-9330 ATR FFSC Safety Analysis Report Rev. 17, May 2022 2.12.4-2 Appendix 1.3.2, Packaging General Arrangement Drawings. The ATR LEU fuel element, also known as the LOWE fuel element, is modeled using the Idaho National Laboratory drawings listed in Section 4.0 of [1].

2.12.4.3.2 Material Properties Material properties are generally chosen as the minimum published properties from the associated ASTM or other governing specification. For the benchmark case, properties used were the actual values recorded in the fabrication data for the certification test units. This was done in order to enhance the accuracy of the benchmark result. In some cases, material properties are varied to obtain a more conservative result. This is done to try to focus more of the free drop energy in a particular feature. For example, in the free drop labeled CD5, whose purpose is to demonstrate the retention of the closure, the package body was given maximum properties to maximize the impact acceleration, and the closure given minimum properties, to increase the deformation and potential failure strain of the closures bayonet lugs. In the

analysis,

• Minimum material properties means the corresponding ASTM specification minimum allowable values,

• Actual material properties means the values recorded for the certification test unit,

• Maximum material properties, since ASTM does not specify maximum allowable values, means the actual properties for the certification test unit.

The material model used in the analysis is elastic-plastic bilinear using Von Mises stress with isotropic strain hardening.

For an element having material properties with failure defined, the element will only provide stress and strain results when its strains are less than the failure strain. When failure strain is reached, the element is removed from the mesh display and ceases to resist deformation.

Further details on material properties may be found in Section B-2.0 of [1].

2.12.4.3.3 Model Details FEA model components are made using a variety of element types suitable to the shape, location, and performance of the component. The general contact coefficient of friction is 0.1, which is relatively low in order to avoid absorbing drop energy in friction. Details of model construction, material application, and mesh density are provided in Section B-3.0 of [1].

The construction of the fuel element depends on the connection of the fuel plates to the side plates. In fabrication practice, this is done by placing each fuel plate into a corresponding groove in the side plate, and roll-swaging the side plate material until it grips the fuel plate. The specification for the fuel element requires that the force required to pull the fuel plate out of the side plate groove after being swaged in place is a minimum of 150 lb per linear inch of fuel plate.

The method of reproducing this in the FEA model is discussed in Section B-2.2 of [1]. In summary, a swage beam is applied to the model in such a way that thermal contraction of the swage beams is used to plastically compress the side plates onto the fuel plates. A coefficient of friction of 1.2 is used, along with the plastic compression force, to produce the joint between the fuel plates and the side plates. The thermal contraction is applied with sufficient magnitude that the swage beams fail, and take no further role in the drop events. The process is calibrated to

Docket No. 71-9330 ATR FFSC Safety Analysis Report Rev. 17, May 2022 2.12.4-3 ensure that the joints are not too strong, and do not produce a pull-out force greater than 150 lb/in.

All free drops were made onto an unyielding surface. The puncture bar is a six-inch diameter mild steel bar with a conservatively sharp edge radius and material properties of steel.

2.12.4.4 Free Drop and Puncture Orientations The free drop and puncture drop orientations are based on the drops used in the initial certification testing described in Appendix 2.12.1, Certification Tests on CTU 1, and Appendix 2.12.2, Certification Tests on CTU-2, and are shown in Table 2.12.4-1, which is extracted from Table 2 of [1]. With only one exception, the drop names are the same in this analysis as in the original certification testing.

Each test in the first column is a HAC free drop from 30 ft, and represents a sequence of NCT 4-ft free drop (column headed preceded by), the HAC free drop in the first column, and HAC 40-in puncture drop (column headed followed by). A NCT free drop is included only in those cases where the damage is expected to be cumulative, such as for the CG over corner deformation. Similarly, a puncture drop is included only when the damage expected from puncture is expected to be cumulative.

All cases except the benchmark case use the weight of package components from Table 2.1-1 of this SAR document, and an ATR LEU fuel element weight of 44 lb. The benchmark case uses the actual weight of the certification test unit (CTU-1) components from Table 2.12.1-1 and the actual weight of the dropped (HEU) fuel of 22.1 lb.

The choice of these orientations is based on the following considerations:

• CD5 (as benchmark): The ATR FFSC package is very strong and deforms very little, even in the 30-ft free drop. The CD5 impact provided the only measurable deformation from the test data. It is preceded by the same orientation in the corresponding NCT (4-ft) drop CN1, since that was the actual test sequence used.

• CD3, Flat side: This is an impact on the largest area provided by the package, and due to its stiffness, should provide one of the largest impact magnitudes. It will subject the fuel to a very strong lateral load which could deform the fuel element or dislodge the fuel plates from the side plates. It is followed by puncture drop CP1. CP1 is a vertical drop centered on the closure.

• CD4, Bottom down: This is a vertical, bottom-down orientation that will subject the fuel element to a very strong axial load which could dislodge the fuel plates from the side plates. It is followed by puncture drop CP1 on the closure.

• CD5, CG over corner: this applies the only load on the bayonet closure lugs which would have any chance of shearing them off. It is preceded by the 4-ft free drop in the same orientation, CN1, and followed by puncture drop CP1.

• CD-New, rotated 10°: This orientation was added to test the performance of the locking pins of the closure, similar to what occurred in certification testing when the package didnt land flat on a side. The aim is to achieve a high rotational acceleration about the package long axis, which requires the locking pins to accelerate the closure rotationally.

A 10° angle provides enough space to allow the package to rotate, while at the same time maintaining a large tangential component causing package rotation. If the angle were less than 10°, the package would not have time to rotate before impact of the

Docket No. 71-9330 ATR FFSC Safety Analysis Report Rev. 17, May 2022 2.12.4-4 opposite edge; if it were larger, more of the impact force would be directed radially toward the package axis without causing rotation. CD-New is followed by puncture drop CP3, an oblique impact on the closure in an effort to rotate it farther.

• CD5 (soft): Using minimum material properties, this orientation will provide a maximum deformation of the impacted end.

A slapdown orientation is not included because the CD3 (flat side) and CD4 (bottom end) drops are considered bounding for fuel integrity. The ATR FFSC package body has three intermediate thick steel plates separating the inner round tube from the outer square tube (SAR drawing 60501-10, item 8), and additionally each end consists of a thick steel plate. Thus, in a flat side drop such as CD3, impact is supported by essentially five relatively rigid steel plates, whereas in a slapdown drop (in which the primary end and the secondary end of the package strike the surface alone in sequence), only the plate at one end is striking at one time. In addition, the bayonet closure, having no attachment bolts or seals, is not subject to significant damage from lateral forces.

2.12.4.5 Results All model runs were made using Abaqus explicit dynamics software, Version 2021.HF6 [2] and validated per [3]. A summary of the results of the six analysis sequences is given in Table 2.12.4-2, extracted from Table 5 of [1]. Initial conditions for each drop in the sequence is discussed in Section B-2.3 of [1]. The method of obtaining sequential impacts is discussed in Section B-2.4 of [1]. Impact magnitudes, filtered using a Butterworth filter with a cutoff frequency of 1,000 Hz, are provided in Table 6 of [1]. All results are discussed in detail in Section B-4.0 of [1].

2.12.4.5.1 CD5 Benchmark Case As stated in Table 2.12.4-1, the model was run with the CG over closure-end corner using the actual certification test component weights and an ATR fuel element weight of 22.1 lb. The results are documented in Section B-4.1 of [1]. The sequence consisted of a 4-ft NCT free drop (CN1) in the CG over corner orientation, followed by a 30-ft HAC free drop in the same orientation (CD5). The results are compared to the certification drop test results, including Figure 2.12.4-1 (certification test results taken from Figure 2.12.1-33) and Figure 2.12.4-2, representing the FEA model results. The deformation of the impacting corner is compared in the table below:

Certification Test, inches (Table 2.12.1-3)

FEA Model, inches CN1 (NCT 4-ft free drop)

~1/8 (0.13) 0.18 CD5 (HAC, 30-ft free drop, including cumulative deformation from CN1)

~5/8 (0.63) 0.83 These results show good agreement between the physical test and the model. Of note, the energy i.e., the drop height, in both cases is identical. Thus, where the deformation is greater, the material strain must be greater. Since material strain is a failure criterion (if the strain reaches a limiting strain in a model element, the element is removed from the model), the FEA model

Docket No. 71-9330 ATR FFSC Safety Analysis Report Rev. 17, May 2022 2.12.4-5 results may be viewed as conservative. For the ATR fuel element, no deformation measurements were made during the certification test. However, a photo of the impacted end of the fuel element was made (after all testing, see Figure 2.12.1-58), which may be compared to the view of the FEA model fuel after the benchmark test, Figure 2.12.4-3. As expected, for both cases, the end box structures have failed and most of the end box material is separated from the fuel element. The remaining fuel element structures (fuel plates and side plates) are intact in both cases, and have similar non-significant deformation damage. Thus, the FEA model is considered acceptable for further analysis.

2.12.4.5.2 CD3 Case The CD3 drop is a HAC free drop from 30 ft on one of the packages flat sides, followed by puncture drop CP1. The purpose of this case is to demonstrate the integrity of the LEU fuel elements structure under the bounding lateral impact. The weight of the fuel element is 44 lb.

As shown in Table 2.12.4-1, the strength of the packaging components is set at the maximum values, and the strength of the fuel element set at the ASTM minimum values.

As expected, the fuel element end box on the closure end experienced failure, but only during the second (puncture) drop. As shown in Figure 2.12.4-4, the fuel element retains its initial configuration, and deformation is limited to the ends of the fuel plates.

2.12.4.5.3 CD4 Case The CD4 drop is a HAC free drop from 30 ft on the closed end opposite the closure, followed by puncture drop CP1. The purpose of this case is to demonstrate the integrity of the LEU fuel elements structure under the bounding axial impact. The weight of the fuel element is 44 lb. As shown in Table 2.12.4-1, the strength of the packaging components is set at the maximum values, and the strength of the fuel element set at the ASTM minimum values.

As expected, the fuel element end boxes on each end experienced failure, one during the free drop and one during the puncture drop on the closure end. As shown in Figure 2.12.4-5, the fuel element retains its initial configuration, and deformation is limited to the ends of the fuel plates.

2.12.4.5.4 CD5 Case This case is essentially identical to the benchmark case, except the weight of the LEU fuel is 44 lb, the material properties are conservatively adjusted, and puncture CP1 is added as a third impact. The purpose of this drop is to demonstrate the retention of the package closure, and thus, the retention of the fissile payload. To that end, the packaging body materials are given maximum material properties in order to maximize the impact acceleration, while the closure lid is given the ASTM minimum properties.

After the third (puncture) event, two of the four bayonet lugs on the closure are sheared off, but two are remaining with relatively little damage, see Figure 2.12.4-6 and Figure 2.12.4-7. One of the two locking pins is bent, but both remain competent to prevent rotation of the closure (see Figure 2.12.4-8). Damage to the package body and to the fuel are similar to the benchmark case and are not bounding. Thus the package closure is retained, maintaining confinement of the payload.

Docket No. 71-9330 ATR FFSC Safety Analysis Report Rev. 17, May 2022 2.12.4-6 2.12.4.5.5 CD-New Case This case, depicted in Figure 2.12.4-9, is focused on the performance of the locking pins, as well as imparting another large lateral impact to the fuel. It consists of a HAC, 30-ft horizontal drop on one long edge, with the package rotated slightly, at 10° away from the surface. This impact imparts a strong rotation to the package body, a rotation which must be transferred to the closure through the locking pins. The free drop is followed by puncture CP3. The weight of the LEU fuel is 44 lb. As shown in Table 2.12.4-1, the material properties of the packaging are given maximum values, except for the locking pins, which are given ASTM minimum values. The fuel element is given minimum material properties.

The model is oriented so that at impact, the locking pin on the top is vertically oriented and disengages through inertia, leaving the lower pin to transfer rotation to the closure. In the puncture impact, both pins are engaged, and one of the ribs in the closure is oriented to fully engage the puncture bar, maximizing the effectiveness of the puncture bar impact (and thus conservatively avoiding a glancing blow that would impart less energy to the closure).

After the 30-ft free drop impact, the single locking pin that is engaged shows some plastic strain but no failed elements, see Figure 2.12.4-10. After the puncture impact, both pins show some plastic strain and the first pin from the free drop impact shows some failed elements. However, the pins remain competent to prevent rotation of the closure. The packaging components and the fuel element do not show significant damage from the sequence. Thus, the closure is secure from rotation and the payload is retained.

2.12.4.5.6 CD5 (soft)

This case is essentially identical to the benchmark case, except the weight of the LEU fuel is 44 lb, and the material properties are conservatively adjusted to maximize the deformation at the closure end. All packaging components and the fuel element are given ASTM minimum material properties.

Figure 2.12.4-11 shows the deformation of the package body, which is similar to the benchmark and CD5 case. The closure and the locking pins show plastic deformation but no significant failure. Thus, the payload is retained. As expected, the fuel element end box on the closure end experiences failure. As shown in Figure 2.12.4-12, the fuel element retains its initial configuration, and deformation is limited to the ends of the fuel plates.

2.12.4.6 Summary The ATR FFSC package was initially licensed utilizing certification drop tests on a prototypic package, and demonstrated its ability to confine the payload and the ability of the ATR HEU fuel element to remain intact. FEA model simulations show that under a number of conservatively configured drop orientations, including sequences of NCT 4-ft, HAC 30-ft, and HAC puncture drops, the ATR FFSC retains its ability to confine the ATR LEU and MURR LEU payloads. In addition, the ability of the ATR LEU fuel element to remain structurally intact, i.e., the fuel plates attached to the side plates, with end box damage, is demonstrated.

Docket No. 71-9330 ATR FFSC Safety Analysis Report Rev. 17, May 2022 2.12.4-7 Table 2.12.4 Free Drop and Puncture Orientations Test Orientation To Show:

Material Properties Preceded by:

Followed by:

CD5 (as benchmark)

CG over corner (Fig. 2.7-3)

Model is reliable Package: actual Fuel: actual or min CN1 (Fig. 2.7-3)

None CD3 Flat side (Fig. 2.7-1)

Integrity of fuel (swage)

Package: max Fuel: min None CP1 (Fig. 2.7-4)

CD4 Bottom down (Fig. 2.7-2)

Integrity of fuel (swage)

Package: max Fuel: min None CP1 (Fig. 2.7-4)

CD5 CG over corner (Fig. 2.7-3)

Retention of closure Square tube: max Closure: min Fuel: min CN1 (Fig. 2.7-3)

CP1 (Fig. 2.7-4)

CD-New 10° rotated (Fig. 2.12.4-1)

Integrity of closure pins Package: max Pins: min Fuel: min None CP3 (Fig. 2.7-6)

CD5 (soft)

CG over corner (Fig. 2.7-3)

Maximum deformation Package: min Fuel: min CN1 (Fig. 2.7-3)

None Note: all items in the first column (Test) are free drops from a height of 30 ft; all of the items in the next-to-last column (Preceded by) are free drops from a height of 4 ft; and all of the items in the last column (Followed by) are puncture drops from a height of 40 inches.

Docket No. 71-9330 ATR FFSC Safety Analysis Report Rev. 17, May 2022 2.12.4-8 Table 2.12.4 Summary of FEA Results Section of [1]

Model Test Sequence Purpose Results B-4.1 C.G. over top corner drop, including CN1 (four ft) and CD5 (30 ft) free drops Benchmark model The accumulated deformation of the top corner in physical test was approx. 5/8 inches. In model run, 0.833 inches, showing acceptable model performance.

B-4.2 Flat side drop (CD3, 30 ft) followed by closure end puncture (CP1)

Demonstrate integrity of fuel plate swage Swage integrity is maintained in test sequence.

B-4.3 Flat bottom end drop (CD4, 30 ft) followed by closure end puncture (CP1)

Demonstrate integrity of fuel plate swage Swage integrity is maintained in test sequence.

B-4.4 C.G. over top corner drop, including CN1 (four ft) followed by CD5 (30 ft) concluding with closure end puncture (CP1)

Demonstrate retention of closure Two of four bayonet lugs sheared off in puncture impact, two other bayonet lugs retained. Closure is retained.

B-4.5 Rotated side drop (CD-New, 30 ft) followed by oblique closure end puncture (CP3)

Demonstrate integrity of pins Pins damaged but remain functional; closure cannot rotate.

B-4.6 C.G. over top corner drop using minimum material properties, including CN1 (four ft) and CD5 (30 ft) free drops Demonstrate maximum package deformations Closure bayonets and pins remain functional, fuel plate swages remain intact.

Docket No. 71-9330 ATR FFSC Safety Analysis Report Rev. 17, May 2022 2.12.4-9 Figure 2.12.4 CTU Following CD5-1 Impact (Figure B-4.1.1-5 [1])

Figure 2.12.4 FFSC Plastic Equivalent Strain After the Second Impact for the Benchmark Test (Figure B-4.1.2-11 [1])

Docket No. 71-9330 ATR FFSC Safety Analysis Report Rev. 17, May 2022 2.12.4-10 Figure 2.12.4 Fuel Without the Failed End Box Displayed Plastic Equivalent Strain After the Second Impact for the Benchmark Test (Figure B-4.1.2-15 [1])

Figure 2.12.4 Plastic Equivalent Strain in the Vicinity of the Swage After the Second Impact for the CD3 Test (Figure B-4.2-7 [1])

Docket No. 71-9330 ATR FFSC Safety Analysis Report Rev. 17, May 2022 2.12.4-11 Figure 2.12.4 Plastic Equivalent Strain in the Vicinity of the Swage After the Second Impact for the CD4 Test (Figure B-4.3-7 [1])

Figure 2.12.4 Lid Plastic Equivalent Strain After the Third Impact for the CD5 Test (Figure B-4.4-7 [1])

Sheared Bayonets

Docket No. 71-9330 ATR FFSC Safety Analysis Report Rev. 17, May 2022 2.12.4-12 Figure 2.12.4 Lid Body Plastic Equivalent Strain After the Third Impact for the CD5 Test (Figure B-4.4-8 [1])

Figure 2.12.4 Pins Plastic Equivalent Strain After the Third Impact for the CD5 Test (Figure B-4.4-9 [1])

Failed elements behind the two elements

Docket No. 71-9330 ATR FFSC Safety Analysis Report Rev. 17, May 2022 2.12.4-13 Figure 2.12.4 Orientation of Model Run CD-New Figure 2.12.4 Pins Plastic Equivalent Strain After the First Impact for the CD-New (10° Rotated) Test (Figure B-4.5-7 [1])

Docket No. 71-9330 ATR FFSC Safety Analysis Report Rev. 17, May 2022 2.12.4-14 Figure 2.12.4 FFSC Body Plastic Equivalent Strain After the Second Impact for the CD5 (Soft) Test (Figure B-4.6-6 [1])

Figure 2.12.4 Fuel Element Plastic Equivalent Strain in the Vicinity of the Swage After the Second Impact for the CD5 (Soft) Test (Figure B-4.6-11 [1])

Docket No. 71-9330 ATR FFSC Safety Analysis Report Rev. 17, May 2022 3-1 3.0 THERMAL EVALUATION This chapter identifies and describes the principal thermal design aspects of the ATR FFSC.

Further, this chapter presents the evaluations that demonstrate the thermal safety of the ATR FFSC package1 and compliance with the thermal requirements of 10 CFR 712 when transporting a payload consisting of an assembled, unirradiated fuel element or a payload of loose, unirradiated fuel plates. The payloads are summarized in Table 1.1-1 and described in Section 1.2.2, Contents.

Specifically, all package components are shown to remain within their respective temperature limits under the normal conditions of transport (NCT). Further, per 10 CFR §71.43(g), the maximum temperature of the accessible package surfaces is demonstrated to be less than 122 °F for the maximum decay heat loading, an ambient temperature of 100 F, and no insolation.

Finally, the ATR FFSC package is shown to retain sufficient thermal protection following the HAC free and puncture drop scenarios to maintain all package component temperatures within their respective short term limits during the regulatory fire event and subsequent package cool-down.

The analysis in the main body of Chapter 3 pertains only to the ATR HEU fuel element and ATR loose plate basket. The analysis for MIT, MURR, RINSC, Cobra, small quantity payloads, and low enriched uranium fuel elements is contained in Section 3.6, Thermal Evaluation for MIT, MURR, Cobra, Small Quantity Payloads, and LEU Fuel Elements.

3.1 Description of Thermal Design The ATR FFSC package, illustrated in Figure 1.2-1 through Figure 1.2-5 from Section 1.0, General Information, consists of three basic components: 1) a Body assembly, 2) a Closure assembly, and 3) either a Fuel Handling Enclosure (FHE) or a Loose Fuel Plate Basket (LFPB).

The FHE is configured to house an assembled ATR fuel element, while the LFPB is configured to house loose ATR fuel element plates. The maximum gross weight of the package loaded with an ATR LEU fuel element and FHE is 299 pounds. The maximum gross weight of the package loaded with a LFPB containing its maximum payload is 290 pounds.

The ATR FFSC is designed as a Type AF packaging for transportation of an ATR fuel element or a bundle of loose ATR fuel element plates. The packaging is rectangular in shape and is intended to be transported in racks of multiple packages by highway truck. Since the payload generates essentially no decay heat, the worst case thermal conditions will occur with an individual package fully exposed to ambient conditions. The package performance when configured in a rack of multiple packages will be bounded by that seen for an individual package.

1 In the remainder of this chapter, the term packaging refers to the assembly of components necessary to ensure compliance with the regulatory requirements, but does not include the payload. The term package includes both the packaging components and the payload of ATR fuel.

2 Title 10, Code of Federal Regulations, Part 71 (10 CFR 71), Packaging and Transportation of Radioactive Material, 01-01-21 Edition.

Docket No. 71-9330 ATR FFSC Safety Analysis Report Rev. 17, May 2022 3-2 configured in a rack of multiple packages will be bounded by that seen for an individual package.

The principal components of the packaging are shown in Figure 1.2-1 and described in more detail below. With the exception of minor components, all steel used in the ATR FFSC packaging is Type 304 stainless steel. Components are joined using full-thickness fillet welds and full and partial penetration groove welds.

3.1.1 Design Features The primary heat transfer mechanisms within the ATR FFSC are conduction and radiation, while the principal heat transfer from the exterior of the packaging is via convection and radiation to the ambient environment. The Body and Closure assemblies serve as the primary impact and thermal protection for the FHE or the LFPB and their enclosed payloads of an ATR fuel element or loose fuel plates. The FHE and LFPB provide additional thermal shielding of their enclosed payloads during the transient HAC event.

There is no pressure relief system included in the ATR FFSC packaging design. The portions of the packaging that are not directly vented to atmosphere do not contain out-gassing materials.

The package insulation is the only non-metallic component located in the enclosed volumes of the package and it is fabricated of a ceramic fiber. The Closure assembly is not equipped with either seals or gaskets so that potential out-gassing of the neoprene material used in ATR fuel tray and the plastic bag material used as a protective sleeve for the fuel element will readily vent without significant pressure build-up in the payload cavity.

The principal thermal design features of each package component are described in the following paragraphs.

3.1.1.1 ATR FFSC Body The ATR FFSC body is a stainless steel weldment that is approximately 73 inches long and 8 inches square and weighs about 230 lbs (empty). It consists of two nested shells; the outer shell is fabricated of a square stainless steel tube with a 3/16 inch wall thickness, while the inner shell is fabricated from a 6 inch diameter, 0.120 inch wall, stainless steel tube. Three, 1-inch thick stiffening plates (i.e., ribs) are secured to the inner shell by fillet welds at four equally spaced intervals. The ribs are not mechanically attached to the outer shell. Instead, a nominal 0.06 inch air gap exists between the ribs and the outer shell, with a larger nominal gap existing at the corners of the ribs. These design features help to thermally isolate the inner shell from the outer shell during the HAC event.

Further thermal isolation of the inner shell is provided by ceramic fiber thermal insulation which is wrapped around the inner shell between the ribs and by the 28 gauge stainless steel sheet used as a jacket material over the insulation. The insulation is applied in two 0.5-inch thick layers in order to permit over-lapping joints between the layers and prevents direct line-of-sight between the inner shell and the jacket should the insulation shift under normal or accident conditions.

The stainless steel jacket maintains the insulation around the inner shell and provides a relatively low emissivity barrier to radiative heat exchange between the insulation and the outer sleeve.

The insulation jacket is pre-formed to the design shape and dimensions prior to installation. As such, the potential for inadvertent compression of the insulation during installation is minimized.

Docket No. 71-9330 ATR FFSC Safety Analysis Report Rev. 17, May 2022 3-3 Once assembled, the inner shell, ribs, and the jacketed insulation wrap are slid as a single unit into the outer shell and secured to closure plates at both ends by welding. Thermal insulation is built into the bottom end closure plate of the packaging, while the ATR FFSC closure (see below) provides thermal insulation at the top end closure.

Cross-sectional views showing key elements of the ATR FFSC body are provided in Figure 1.2-2 and Figure 1.2-3. Figure 1.2-2 illustrates a cross sectional view at the top end closure of the package and 1.2-3 presents a similar cross sectional view of the package at the bottom end closure.

3.1.1.2 ATR FFSC Closure The ATR FFSC closure engages with the body using a bayonet style engagement via four uniformly spaced lugs on the closure that engage with four slots in the mating body feature. The closure incorporates 1 inch of ceramic fiber thermal insulation to provide thermal protection and is designed to permit gas to easily vent through the interface between the closure and the body.

The closure weighs approximately 10 pounds and is equipped with a handle to facilitate use with gloved hands.

A cross sectional view of the ATR FFSC closure is illustrated in Figure 1.2-4.

3.1.1.3 Fuel Handling Enclosure (FHE)

The Fuel Handling Enclosure (FHE) is a hinged, aluminum weldment used to protect an ATR fuel element from damage during loading and unloading operations. It is fabricated of thin wall (i.e., 0.09 inch thick) 5052-H32 aluminum sheet and features a hinged lid and neoprene rub strips to minimize fretting of the fuel element side plates where they contact the FHE. The surface of the FHE is neither anodized nor coated, but is left as an unfinished aluminum sheet.

Figure 1.2-1 presents an illustration of the FHE. A polyethylene bag is used as a protective sleeve over the ATR fuel element.

3.1.1.4 ATR FFSC Loose Fuel Plate Basket (LFPB)

The Loose Fuel Plate Basket (LFPB) serves to maintain the fuel plates within a defined dimensional envelope during transport. The four identical machined segments are machined from a billet of 6061-T651 aluminum and are joined by threaded fasteners (see Figure 1.2-16).

A variable number of ATR fuel plates may be housed in the basket, with the maximum payload weight being limited to 20 lbs. or less. The empty weight of the loose fuel plate basket is approximately 30 lbs. Like the FHE, the surface of the LFPB is neither anodized nor coated, but is left with its as machined finish.

3.1.2 Contents Decay Heat The ATR FFSC is designed as a Type AF packaging for transportation of an unirradiated ATR fuel element or a bundle of loose, unirradiated ATR fuel plates. The decay heat associated with unirradiated ATR fuel is negligible. Therefore, no special devices or features are needed or

Docket No. 71-9330 ATR FFSC Safety Analysis Report Rev. 17, May 2022 3-4 utilized in the ATR FFSC packaging to dissipate the decay heat. Section 1.2.2, Contents, provides additional details.

3.1.3 Summary Tables of Temperatures Table 3.1-1 provides a summary of the package component temperatures under normal and accident conditions. The temperatures for normal conditions are based on an analytical model of the ATR FFSC package for extended operation with an ambient temperature of 100°F and a diurnal cycle for the insolation loading. The temperatures for accident conditions are based on an analytical model of the ATR FFSC package with the worst-case, hypothetical pre-fire damage as predicted based on drop tests using full-scale certification test units (CTUs).

The results for NCT conditions demonstrate that significant thermal margin exists for all package components. This is to be expected since the only significant thermal loads on the package arise from insolation and ambient temperature changes. The payload dissipates essentially zero decay heat. Further, the evaluations for NCT demonstrate that the package skin temperature will be below the maximum temperature of 122°F permitted by 10 CFR §71.43(g) for accessible surface temperature in a nonexclusive use shipment when transported in a 100°F environment with no insolation.

The results for HAC conditions also demonstrate that the design of the ATR FFSC package provides sufficient thermal protection to yield component temperatures that are significantly below the acceptable limits defined for each component. While the neoprene rubber and polyethylene plastic material used to protect the ATR fuel element from damage are expected to reach a sufficient temperature level during the HAC fire event to induce some level of thermal degradation (i.e., melting, charring, the chemical breakdown of the materials into 2 or more substances, etc.), the loss of these components is not critical to the safety of the package.

Further, the potential combustion of these materials will be restricted due to the lack of available oxygen to the point that any potential temperature rise will be insignificant. See Sections 3.2.2, Technical Specifications of Components, 3.4.3.1, Maximum HAC Temperatures, and 3.5.3, Thermal Decomposition/Combustion of Package Organics, for more discussion.

3.1.4 Summary Tables of Maximum Pressures Table 3.1-2 presents a summary of the maximum pressures achieved under NCT and HAC conditions. Since the ATR FFSC package is a vented package, both the maximum normal operating pressure (MNOP) and the maximum pressure developed within the payload compartment under the HAC condition are 0 psig.

Although the volume between the outer and inner shells is sealed, it does not contain organic or other materials that may outgas or thermally degrade. Therefore, the maximum pressure that may develop within the space will be limited to that achieved due to ideal gas expansion. The maximum pressure rise under NCT will be less than 4 psig, while the pressure rise under HAC conditions will be 39 psig.

Docket No. 71-9330 ATR FFSC Safety Analysis Report Rev. 17, May 2022 3-5 Table 3.1 Maximum Temperatures for NCT and HAC Conditions Location / Component NCT Hot Conditions Accident Conditions Maximum Allowable Normal Accident ATR Fuel Element Fuel Plate 147°F 730°F 400°F 1,100°F ATR Fuel Element Side Plate 148°F 827°F 400°F 1,100°F Neoprene Rub Strips/Polyethylene Bag 151°F 1,017°F 225°F N/A Fuel Handling Enclosure (FHE) 151°F 1,017°F 400°F 1,100°F Loose Fuel Plate Basket (LFPB) 151°F 746°F 400°F 1,100°F Inner Shell 157°F 1,422°F 800°F 2,700°F Ceramic Fiber Insulation, Body

- Maximum

- Average 185°F 151°F 1,460°F 1,220°F 2,300°F 2,300°F 2,300°F 2,300°F Ceramic Fiber Insulation, Closure

- Maximum

- Average 145°F 144°F 1,418°F 1,297°F 2,300°F 2,300°F 2,300°F 2,300°F Closure 145°F 1,445°F 800°F 2,700°F Outer Shell 186°F 1,471°F 800°F 2,700°F Table Notes:

Maximum allowable temperatures are defined in Section 3.2.2, Technical Specifications of Components.

Component temperature assumed to be equal to that of the FHE.

Table 3.1 Summary of Maximum Pressures Condition Fuel Cavity Pressure Outer/Inner Shell Cavity Pressure NCT Hot 0 psi gauge 4 psi gauge HAC Hot 0 psi gauge 39 psi gauge

Docket No. 71-9330 ATR FFSC Safety Analysis Report Rev. 17, May 2022 3-6 3.2 Material Properties and Component Specifications The ATR FFSC is fabricated primarily of Type 304 stainless steel, 5052-H32 and 6061-T651 aluminum, ceramic fiber insulation, and neoprene rubber. The payload materials include 6061-T6 and/or 6061-0 aluminum, uranium aluminide (UAlx), and uranium-molybdenum (i.e., U-10Mo in a foil coated with thin zirconium interlayers). A polyethylene plastic bag is used as a protective sleeve over the fuel element.

3.2.1 Material Properties Table 3.2-1 presents the thermal properties for Type 304 stainless steel and 5052-H32 aluminum from Table TCD of the ASME Boiler and Pressure Vessel Code3. Since the HAC analysis requires thermal properties in excess of the maximum temperature point of 400F provided in Table TCD for 5052-H32 aluminum, the property values for 1100F (i.e., the approximate melting point for aluminum) are assumed to be the same as those at 400F. This approach is appropriate for estimating the temperature rise within the fuel basket during the HAC event since the thermal conductivity of aluminum alloys tends to decrease with temperature while the specific heat tends to increase. The density values listed in the table are taken from an on-line database4. Properties between the tabulated values are calculated via linear interpolation within the heat transfer code.

Table 3.2-2 presents the thermal properties for the ATR fuel element. For analysis purposes, the material used for the side plates, covers, and fuel cladding are assumed to be 6061-0 aluminum. The thermal properties for the fuel plates are determined as a composite of the cladding and the fuel core materials based on the geometry data for the ATR fuel element5 and the thermal properties for the ATR fuel element materials6. The details of the computed values are presented in Section 3.5.2.4, Determination of Composite Thermal Properties for ATR Fuel Plates. For simplicity and given the low sensitivity to temperature, a conservatively high, fixed thermal conductivity value is used for the fuel plates in order to maximize the heat transfer into the fuel components during the HAC event.

The specific heat values are computed as a function of temperature to more accurately capture the change in thermal mass for the fuel plates during the HAC event.

The thermal properties for the non-metallic materials used in the ATR FFSC are presented in Table 3.2-3. The thermal properties for neoprene rubber are based on the Polymer Data Handbook7, while the thermal properties for the ceramic fiber insulation are based on the Unifrax Durablanket S insulation product8 with a nominal density of 6 lb/ft3. The thermal properties are for the uncompressed material in both cases. Although the package design requires that the insulation blanket be compressed by up to 20% at the quadrant points, ignoring 3 American Society of Mechanical Engineers (ASME) Boiler and Pressure Vessel Code,Section II, Materials, Part D -

Properties, Table TCD, Material Group J, 2001 Edition, 2002 and 2003 Addenda, New York 4 Matweb, Online Material Data Sheets, www.matweb.com.

5 ATR Mark VII Fuel Element Assembly, INEEL Drawing No. DWG-405400.

6 Thermophysical And Mechanical Properties Of ATR Core Materials, Report No. PG-T-91-031, August 1991, EG&G Idaho, Inc.

7 Polymer Data Handbook, Oxford University Press, Inc., 1999.

8 Unifrax DuraBlanket S ceramic fiber insulation, Unifrax Corporation, Niagara Falls, NY.

Docket No. 71-9330 ATR FFSC Safety Analysis Report Rev. 17, May 2022 3-7 the compression for the purposes of the thermal modeling and using the thermal properties for the uncompressed material at all locations provides a conservative estimate of the packages performance under the HAC condition. This conclusion arises from the fact that the insulations thermal conductivity decreases with density for temperatures above approximately 500°F (see Table 3.2-3). For example, the thermal conductivity of 8 pcf insulation at 1000°F and 1400°F is 0.0814 and 0.1340 Btu/hr-ft-°F, respectively, versus the 0.0958 and 0.1614 Btu/hr-ft-°F values for 6 pcf insulation at the same temperatures. While compression will increase conductivity below 500°F, ignoring the effects of compression for NCT conditions has an insignificant effect since the peak package temperatures occur in the vicinity of the ribs and are therefore unaffected by a local increase in the thermal conductivity of the insulation. Further, large thermal margins exist for the NCT conditions.

The thermal properties for air presented in Table 3.2-4 are derived from curve fits9. Because the thermal conductivity of air varies significantly with temperature, the computer model calculates the thermal conductivity across thin air filled gaps as a function of the mean gap temperature.

All void spaces within the ATR FFSC package are assumed to be filled with air at atmospheric pressure.

Table 3.2-5 and Table 3.2-6 present the assumed emissivity () for each radiating surface and the solar absorptivity () value for the exterior surface. The emissivity of as-received Type 304 stainless steel has been measured10 as 0.25 to 0.28, while the emissivity of weathered Type 304 stainless steel has been measured11 from 0.46 to 0.50. For the purpose of this analysis, an emissivity of 0.30 is assumed for the emittance from all interior radiating stainless steel surfaces, while the emissivity for the exterior surfaces of the package is assumed to be 0.45. The solar absorptivity of Type 304 stainless steel is approximately 0.5212. Under HAC conditions, the outside of the package is assumed to attain an emissivity of 0.8 in compliance with 10 CFR §71.73(c)(4) and to have a solar absorptivity of 0.9 to account for the possible accumulation of soot.

The 5052-H32 aluminum sheet used to fabricate the FHE will be left with a plain finish while the 6061-T651 billets used to fabricate the Loose Fuel Plate Basket will have a machined surface.

The emissivity for either type of finish can be expected to be low (i.e., 0.10 or lower)12 however, for conservatism, an emissivity of 0.2512 representative of a heavily oxidized surface is assumed for this evaluation. The 6061-0 aluminum used for the ATR fuel components are assumed to have a surface coating of boehmite (Al2O3H2O). A 25 m boehmite film will exhibit a surface emissivity of approximately 0.9213. While a fresh fuel element may have a lower surface emissivity, the use of the higher value will provide a conservative estimate of the temperatures achieved during the HAC event.

9 Rohsenow, Hartnett, and Cho, Handbook of Heat Transfer, 3rd edition, McGraw-Hill Publishers, 1998.

10 Frank, R. C., and W. L. Plagemann, Emissivity Testing of Metal Specimens. Boeing Analytical Engineering coordination sheet No. 2-3623-2-RF-C86-349, August 21, 1986. Testing accomplished in support of the TRUPACT-II design program.

11 "Emissivity Measurements of 304 Stainless Steel", Azzazy, M., prepared for Southern California Edison, September 6, 2000, Transnuclear File No. SCE-01.0100.

12 G. G. Gubareff, J. E. Janssen, and R. H. Torborg, Thermal Radiation Properties Survey, 2nd Edition, Honeywell Research Center, 1960.

13 Heat Transfer in Window Frames with Internal Cavities, PhD Thesis for Arild Gustavsen, Norwegian University of Science and Technology, Trondheim, Norway, September 2001.

Docket No. 71-9330 ATR FFSC Safety Analysis Report Rev. 17, May 2022 3-8 The ceramic fiber insulation has a surface emissivity of approximately 0.9012 based on a combination of the material type and surface roughness. The same emissivity is assumed for the neoprene rubber.

3.2.2 Technical Specifications of Components The materials used in the ATR FFSC that are considered temperature sensitive are the aluminum used for the FHE, the LFPB, and the ATR fuel, the neoprene rubber, and the polyethylene wrap used as a protective sleeve around the ATR fuel element. Of these materials, only the aluminum used for the ATR fuel element is considered critical to the safety of the package. The other materials either have temperature limits above the maximum expected temperatures or are not considered essential to the function of the package.

Type 304 stainless steel has a melting point above 2,700F4, but in compliance with the ASME B&PV Code14, its allowable temperature is limited to 800°F if used for structural purposes.

However, the ASME temperature limit generally applies only to conditions where the materials structural properties are relied on for loads postulated to occur in the respective operating mode or load combination (such as the NCT and HAC free drops). Since the package is vented to atmosphere, no critical structural condition exists following the HAC free drop events and, as such, the appropriate upper temperature limit is 800°F for normal conditions and 2,700F for accident conditions Aluminum (5052-H32, 6061-0/6061-T6) has a melting point of approximately 1,100F4 however for strength purposes the normal operational temperature should be limited to 400F3.

The ceramic fiber insulation has a manufacturers recommended continuous use temperature limit of 2,300°F8. There is no lower temperature limit.

The polyethylene plastic wrap used as a protective sleeve around the ATR fuel element has a melting temperature of approximately 225 to 250F4. For the purposes of this analysis, the lower limit of 225F is used. As a thermoplastic, the polyethylene wrap will melt and sag onto the fuel element when exposed to temperatures in excess of 250F. Further heating could lead to charring (i.e., oxidation in the absence of open combustion) and then thermal decomposition into its volatile components. Thermal decomposition will begin at approximately 750F. Unpiloted, spontaneous ignition could occur at temperatures of approximately 650F15 or higher. The plastic wrap is approximately 7 inches wide (when pressed flat), 67.5 inches long, and weights approximately 3 oz. As demonstrated in Section 3.5.3, Thermal Decomposition/Combustion of Package Organics, the available oxygen in the package is sufficient for consumption of less than 1% of the polyethylene. Loss of the plastic wrap is of no consequence to the thermal safety of the ATR FFSC since its effect on conductive and radiative heat transfer is negligible.

The neoprene rub strips used to minimize fretting of the fuel element side plates have a continuous temperature rating of 200 to 250F and a short term (i.e., 0.5 hour5.787037e-5 days <br />0.00139 hours <br />8.267196e-6 weeks <br />1.9025e-6 months <br /> or less) temperature 14 American Society of Mechanical Engineers (ASME) Boiler & Pressure Vessel Code,Section III, Rules for Construction of Nuclear Facility Components, Division 1, Subsection NB, Class 1 Components, & Subsection NG, Core Support Structures, 2001 Edition, 2002 Addendum.

15 Troitzsch, J., Plastics Flammability Handbook, 2nd Edition, Oxford University Press, New York, 1990.

Docket No. 71-9330 ATR FFSC Safety Analysis Report Rev. 17, May 2022 3-9 limit of approximately 525F 16. For the purposes of this analysis, a limit of 225F is used for NCT conditions, while a peak temperature of 525F is assumed for HAC conditions before thermal degradation begins. Since neoprene is a thermoset polymer, it will not melt, but decompose into volatiles as it degrades. The same limitation on oxygen affecting the combustion of polyethylene also affects neoprene. As discussed in Section 3.5.3, Thermal Decomposition/Combustion of Package Organics, the thermal damage expected for the neoprene material is expected to be limited to potential de-bonding from the FHE surfaces and a very limited thermal decomposition.

Loss of the neoprene rub strips is of no consequence to the thermal safety of the ATR FFSC.

The minimum allowable service temperature for all ATR FFSC components is below -40 F.

16 Parker O-Ring Handbook, ORD 5700/USA, 2001, www.parker.com.

Docket No. 71-9330 ATR FFSC Safety Analysis Report Rev. 17, May 2022 3-10 Table 3.2 Thermal Properties of Package Metallic Materials Material Temperature

(ºF)

Thermal Conductivity (Btu/hr-ft-ºF)

Specific Heat (Btu/lbm-ºF)

Density (lbm/in3)

Stainless Steel Type 304 70 8.6 0.114 0.289 100 8.7 0.115 200 9.3 0.119 300 9.8 0.123 400 10.4 0.126 500 10.9 0.128 600 11.3 0.130 700 11.8 0.132 800 12.2 0.133 1000 13.2 0.136 1200 14.0 0.138 1400 14.9 0.141 1500 15.3 0.142 Aluminum Type 5052-H32 70 79.6 0.214 0.097 100 80.8 0.216 150 82.7 0.219 200 84.4 0.222 250 85.9 0.225 300 87.2 0.227 350 88.4 0.229 400 89.6 0.232 1100 89.6 0.232 Notes:

Values for 1100°F are assumed equal to values at 400°F.

Docket No. 71-9330 ATR FFSC Safety Analysis Report Rev. 17, May 2022 3-11 Table 3.2 Thermal Properties of ATR Fuel Materials Material Temperature

(ºF)

Thermal Conductivity (Btu/hr-ft-ºF)

Specific Heat (Btu/lbm-ºF)

Density (lbm/in3)

Aluminum Type 6061-0 32 102.3 0.0976 62 0.214 80 104.0 170 107.5 260 109.2 0.225 350 109.8 440 110.4 0.236 530 110.4 620 109.8 0.247 710 108.6 800 106.9 0.258 890 105.2 980 103.4 0.269 1080 101.1 0.275 ATR Fuel Plate 1 80 60.5 0.177 0.114 800 0.213 ATR Fuel Plates 2 and 18 80 78.5 0.189 0.108 800 0.228 ATR Fuel Plates 3,4,16 &

17 80 76.2 0.182 0.112 800 0.220 ATR Fuel Plates 5 to 15 80 74.6 0.176 0.115 800 0.212 ATR Fuel Plate 19 80 54.5 0.173 0.115 800 0.209 Notes:

Values determined based on composite value of aluminum cladding and fuel core material (see Appendix 3.5.2.4). Thermal conductivity value is valid for axial and circumferential heat transfer within fuel plate.

Docket No. 71-9330 ATR FFSC Safety Analysis Report Rev. 17, May 2022 3-12 Table 3.2 Thermal Properties of Non-Metallic Materials Material Temperature

(ºF)

Thermal Conductivity (Btu/hr-ft-ºF)

Specific Heat (Btu/lbm-ºF)

Density (lbm/ft3)

Comments Neoprene 0.11 0.52 76.8 Ceramic Fiber Insulation 70 0.0196 0.28 6

200 0.0238 400 0.0343 600 0.0499 800 0.0703 1000 0.0958 1200 0.1262 1400 0.1614 1600 0.2017 Ceramic Fiber Insulation 70 0.0300 0.28 8

200 0.0313 400 0.0369 600 0.0463 800 0.0620 1000 0.0814 1200 0.1053 1400 0.1340 1600 0.1669 Notes:

Conductivity value represents uncompressed neoprene.

Conductivity values are for uncompressed insulation. Compression of the material will increase the thermal conductivity for temperatures below approximately 500°F where conduction dominates and decrease the thermal conductivity for temperatures above 500°F where heat transfer via radiation dominates.

8 pcf ceramic fiber insulation is not used in the ATR FFSC Package. Data is provided for comparison purposes to demonstrate the effect of insulation compression on thermal conductivity.

Docket No. 71-9330 ATR FFSC Safety Analysis Report Rev. 17, May 2022 3-13 Table 3.2 Thermal Properties of Air Temperature

(ºF)

Density lbm/in3) 1 Specific Heat (Btu/lbm-ºF)

Dynamic Viscosity (lbm /ft-hr)

Thermal Conductivity (Btu/hr-ft-ºF)

Prandtl Number 2 Coef. Of Thermal Exp.

(ºR-1) 3

-40 Use Ideal Gas Law w/

Molecular wt

= 28.966 0.240 0.03673 0.0121 Compute as Pr = cp k Compute as

(ºF+459.67) 0 0.240 0.03953 0.0131 50 0.240 0.04288 0.0143 100 0.241 0.04607 0.0155 200 0.242 0.05207 0.0178 300 0.243 0.05764 0.0199 400 0.245 0.06286 0.0220 500 0.248 0.06778 0.0240 600 0.251 0.07242 0.0259 700 0.253 0.07680 0.0278 800 0.256 0.08098 0.0297 900 0.259 0.08500 0.0315 1000 0.262 0.08887 0.0333 1200 0.269 0.09620 0.0366 1400 0.274 0.10306 0.0398 1500 0.277 0.10633 0.0412 Table Notes:

1) Density computed from ideal gas law as = PM/RT, where R= 1545.35 ft-lbf/lb-mole-R, T= temperature in °R, P= pressure in lbf/ft2, and M= molecular weight of air. For example, at 100°F and atmospheric pressure of 14.69lbf/in2, = (14.69*144 in2/ft2*28.966 lbm/lb-mole)/1545.35*(100+459.67) = 0.071 lbm/ft3 = 4.099x10-5 lbm/in3.

2) Prandtl number computed as Pr = cp k, where cp = specific heat, = dynamic viscosity, and k = thermal conductivity. For example, at 100°F, Pr = 0.241*0.04607/0.0155 = 0.72.

3) Coefficient of thermal expansion is computed as the inverse of the absolute temperature. For example, at 100°F, 100+459.67) = 0.00179.

Docket No. 71-9330 ATR FFSC Safety Analysis Report Rev. 17, May 2022 3-14 Table 3.2 NCT Thermal Radiative Properties Material Assumed Conditions Assumed Emissivity ()

Absorptivity ()

Outer Shell, Exterior Surfaces (Type 304 Stainless Steel)

Weathered 0.45 0.52 Outer Shell, Interior Surface and Inner Shell (Type 304 Stainless Steel)

As-Received 0.3 Ceramic Fiber Insulation &

Neoprene 0.90 Fuel Handling Enclosure and Loose Fuel Plate Basket (6061-T651 &5052-H32 Aluminum)

Oxidized 0.25 ATR Fuel Side Plates and Fuel Cladding (6061-0 Aluminum)

Boehmite film 0.92 Ambient Environment 1.00 N/A Table 3.2 HAC Thermal Radiative Properties Material Assumed Conditions Assumed Emissivity ()

Absorptivity ()

Outer Shell, Exterior Surfaces (Type 304 Stainless Steel)

Sooted/Oxidized 0.80 0.90 Outer Shell, Interior Surface and Inner Shell (Type 304 Stainless Steel)

Slightly Oxidized 0.45 Ceramic Fiber Insulation &

Neoprene 0.90 Fuel Handling Enclosure and Loose Fuel Plate Basket (6061-T651 &5052-H32 Aluminum)

Oxidized 0.25 ATR Fuel Side Plates and Fuel Cladding (6061-0 Aluminum)

Boehmite film 0.92 Ambient Environment 1.00 N/A

Docket No. 71-9330 ATR FFSC Safety Analysis Report Rev. 17, May 2022 3-15 3.3 Thermal Evaluation for Normal Conditions of Transport This section presents the thermal evaluation of the ATR FFSC for normal conditions of transport (NCT). Under NCT, the package will be transported horizontally. This establishes the orientation of the exterior surfaces of the package for determining the free convection heat transfer coefficients and insolation loading. While the package would normally be transported in tiered stacks of multiple packages, the evaluation for NCT is conservatively based on a single, isolated package since this approach will yield the bounding maximum and minimum temperatures achieved by any of the packages. Further, the surface of the transport trailer is conservatively assumed to prevent heat exchange between the package and the ambient. Thus, the bottom of the ATR FFSC is conservatively treated as an adiabatic surface.

The details of the thermal modeling used to simulate the ATR FFSC package under NCT conditions are provided in Appendix 3.5.2, Analytical Thermal Model.

3.3.1 Heat and Cold 3.3.1.1 Maximum Temperatures The maximum temperature distribution for the ATR FFSC occurs with a diurnal cycle for insolation loading and an ambient air temperature of 100°F, per 10 CFR §71.71(c)(1). The evaluation of this condition is conducted as a transient using the thermal model of an undamaged ATR FFSC described in Appendix 3.5.2.1, Description of Thermal Model for NCT Conditions.

Figure 3.3-1 and Figure 3.3-2 illustrate the expected heat-up transient for an ATR FFSC loaded with an ATR fuel element. The transient analysis assumes a uniform temperature condition of 70ºF for all components prior to loading and exposure to the specified NCT condition at time = 0. The figures demonstrate that the ATR FFSC package will respond rapidly to changes in the level of insolation and will reach it peak temperatures within the first day or two after loading. Table 3.3-1 presents the maximum temperatures reached for various components of the package. As seen from the table, all components are within in their respective temperature limits. Figure 3.3-3 illustrates the predicted temperature distribution within the ATR FFSC package at the time of peak temperature.

The maximum temperature distribution for the ATR FFSC without insolation loads occurs with an ambient air temperature of 100°F. Since the package payload dissipates essentially zero watts of decay heat, the thermal analysis of this condition represents a trivial case and no thermal calculations are performed. Instead, it is assumed that all package components achieve the 100°F temperature under steady-state conditions. The resulting 100°F package skin temperature is below the maximum temperature of 122°F permitted by 10 CFR §71.43(g) for accessible surface temperature in a nonexclusive use shipment.

3.3.1.2 Minimum Temperatures The minimum temperature distribution for the ATR FFSC occurs with a zero decay heat load and an ambient air temperature of -40°F per 10 CFR §71.71(c)(2). The thermal analysis of this condition also represents a trivial case and no thermal calculations are performed. Instead, it is assumed that all package components achieve the -40°F temperature under steady-state conditions.

Docket No. 71-9330 ATR FFSC Safety Analysis Report Rev. 17, May 2022 3-16 As discussed in Section 3.2.2, Technical Specifications of Components, the -40°F temperature is within the allowable operating temperature range for all ATR FFSC package components.

3.3.2 Maximum Normal Operating Pressure The payload cavity of the ATR FFSC is vented to the atmosphere. As such, the maximum normal operating pressure (MNOP) for the package is 0 psig.

While the volume between the outer and inner shells is sealed, it does not contain organic or other materials that may outgas or thermally degrade. Therefore, the maximum pressure that may develop within the space will be limited to that achieved due to ideal gas expansion.

Assuming a temperature of 70°F at the time of assembly and a maximum operating temperature of 190°F (based on the outer shell temperature, see Table 3.3-1, conservatively rounded up), the maximum pressure rise within the sealed volume will be less than 4 psi.

Docket No. 71-9330 ATR FFSC Safety Analysis Report Rev. 17, May 2022 3-17 Table 3.3 Maximum Package NCT Temperatures Location / Component NCT Hot Conditions Maximum Allowable ATR Fuel Element Fuel Plate 147°F 400°F ATR Fuel Element Side Plate 148°F 400°F Neoprene Rub Strips/Polyethylene Bag 151°F 225°F Fuel Handling Enclosure (FHE) 151°F 400°F Loose Fuel Plate Basket (LFPB) 151°F 400°F Inner Shell 157°F 800°F Ceramic Fiber Insulation, Body

- Maximum

- Average 185°F 151°F 2,300°F 2,300°F Ceramic Fiber Insulation, Closure

- Maximum

- Average 145°F 144°F 2,300°F 2,300°F Closure 145°F 800°F Outer Shell 186°F 800°F Table Notes:

The maximum allowable temperatures under NCT conditions are provided in Section 3.2.2, Technical Specifications of Components.

Component temperature assumed to be equal to that of the FHE.

Docket No. 71-9330 ATR FFSC Safety Analysis Report Rev. 17, May 2022 3-18 60 80 100 120 140 160 180 200 0

12 24 36 48 60 72 84 96 Time - Hours Temperature - F Outer Shell - Max.

Outer Shell - Avg.

Insulation Jacket - Max.

Insulation - Avg.

Inner Shell - Max.

Fuel Handling Enclosure (FHE) - Max.

Figure 3.3 ATR FFSC Package Heat-up, NCT Hot Conditions 60 80 100 120 140 160 0

24 48 72 96 Time - Hours Temperature - F ATR Side Plates - Max.

ATR Side Plates - Avg.

ATR Fuel Plates - Max.

ATR Fuel Plates - Avg.

Inner Shell - Max.

Fuel Handling Enclosure (FHE) - Max.

Figure 3.3 ATR Fuel Element Heat-up, NCT Hot Conditions

Docket No. 71-9330 ATR FFSC Safety Analysis Report Rev. 17, May 2022 3-19 Figure 3.3 Package NCT Temperature Distribution

Docket No. 71-9330 ATR FFSC Safety Analysis Report Rev. 17, May 2022 3-20 3.4 Thermal Evaluation for Hypothetical Accident Conditions This section presents the thermal evaluation of the ATR FFSC package under the hypothetical accident condition (HAC) specified in 10 CFR §71.73(c)(4) based on an analytical thermal model of the ATR FFSC. The analytical model for HAC is a modified version of the quarter symmetry NCT model described in Section 3.5.2.1, Description of Thermal Model for NCT Conditions, with the principal model modifications consisting of simulating the expected package damage resulting from the drop events that are assumed to precede the HAC fire and changing the package surface emissivities to reflect the assumed presence of soot and/or surface oxidization.

Physical testing using full scale certified test units (CTUs) is used to establish the expected level of damage sustained by the ATR FFSC package from the 10 CFR 71.73 prescribed free and puncture drops that are assumed to precede the HAC fire event. Appendix 2.12.1, Certification Tests on CTU-1 and Appendix 2.12.2, Certification Tests on CTU-2 provide the configuration and initial conditions of the test articles, the test facilities and instrumentation used, and the test results. Section 3.5.2.2, Description of Thermal Model for HAC Conditions, provides an overview of the test results, the rationale for selecting the worst-case damage scenario, and the details of the thermal modeling used to simulate the package conditions during the HAC fire event.

3.4.1 Initial Conditions The initial conditions assumed for the package prior to the HAC event are described below in terms of the modifications made to the NCT thermal model to simulate the assumed package conditions prior to and during the HAC event. These modifications are:

Simulated the worst-case damage arising from the postulated HAC free and puncture drops as described in Section 3.5.2.2, Description of Thermal Model for HAC Conditions, Assume an initial, uniform temperature distribution of 100ºF based on a zero decay heat package at steady-state conditions with a 100ºF ambient with no insolation.

This assumption complies with the requirement of 10 CFR §71.73(b)2 and NUREG-160917, Increased the emissivity of the external surfaces from 0.45 to 0.8 to account for possible soot accumulation on the surfaces, per 10 CFR §71.73(c)(4),

Increased the emissivity of the interior surfaces of the outer shell from 0.30 to 0.45 to account for possible oxidization of the surfaces during the HAC event, Following the free and puncture bar drops, the ATR FFSC package is assumed come to rest in a horizontal position prior to the initiation of the fire event. Since the package geometry is essentially symmetrical about its axial axis, there are no significant thermal differences whether the 17 NUREG-1609, Standard Review Plan for Transportation Packages for Radioactive Material, §3.5.5.1, U.S.

Regulatory Commission, Office of Nuclear Materials Safety and Standards, March 1999.

Docket No. 71-9330 ATR FFSC Safety Analysis Report Rev. 17, May 2022 3-21 package is right-side up, up-side down, or even on its end. The potential for the ATR fuel element payload being re-positioned depending upon the package orientation is not significant to the peak temperatures developed under HAC conditions given the modeling approach used to compute the heat transfer from the inner shell to the ATR fuel element. Therefore, the peak package temperatures predicted under this evaluation are representative of those achieved for any package orientation.

3.4.2 Fire Test Conditions The fire test conditions analyzed to address the 10 CFR §71.73(c) requirements are as follows:

The initial ambient conditions are assumed to be 100ºF ambient with no insolation, At time = 0, a fully engulfing fire environment consisting of a 1,475ºF ambient with an emissivity of 1.0 is used to simulate the hydrocarbon fuel/air fire event.

The assumption of a flame emissivity of 1.0 bounds the minimum average flame emissivity coefficient of 0.9 specified by 10 CFR §71.73(c)(4),

The convection heat transfer coefficients between the package and the ambient during the 30-minute fire event are based on an average gas velocity18 of 10 m/sec. Following the 30-minute fire event the convection coefficients are based on still air, The ambient condition of 100ºF with insolation is assumed following the 30-minute fire event. Since a diurnal cycle is used for insolation, the evaluation assumes that the 30-minute fire begins at noon so as to maximize the insolation heating during the post-fire cool down period. A solar absorptivity of 0.9 is assumed for the exterior surfaces to account for potential soot accumulation on the package surfaces.

The transient analysis is continued for 11.5 hours5.787037e-5 days <br />0.00139 hours <br />8.267196e-6 weeks <br />1.9025e-6 months <br /> after the end of the 30-minute fire to ensure that the peak package temperatures are captured.

3.4.3 Maximum Temperatures and Pressure 3.4.3.1 Maximum HAC Temperatures The outer shell and the ceramic fiber insulation provide thermal protection to the ATR FFSC package during the HAC fire event. The level of thermal protection can be seen via the thermal response curves presented in Figure 3.4-1 and Figure 3.4-2. As illustrated in the figures, while the exterior of the package quickly rises to nearly the temperature of the fire, the heat flow to the FHE and its enclosed ATR fuel element payload is sufficiently restricted that the maximum temperatures of both the FHE and the ATR fuel element are well below the melting point of aluminum. This result occurs despite the conservative assumption of direct contact between the FHE and the inner shell at 3 locations (e.g., the equivalent of four locations for a full model).

18 Schneider, M.E and Kent, L.A., Measurements Of Gas Velocities And Temperatures In A Large Open Pool Fire, Heat and Mass Transfer in Fire - HTD Vol. 73, 1987, ASME, New York, NY.

Docket No. 71-9330 ATR FFSC Safety Analysis Report Rev. 17, May 2022 3-22 This level of thermal protections is further illustrated by the perspective views presented in Figure 3.4-3 and Figure 3.4-4 of the temperature distribution in the ATR FFSC package after 30 minutes of exposure to the HAC fire and at the point when the peak ATR fuel element temperature is attained (approximately 22 minutes after the end of the fire). The figures show that the ceramic fiber insulation limits the elevated temperatures resulting from the fire event to regions adjacent to the outer shell. The assumed absence of the ceramic fiber insulation adjacent to the ribs as a result of the pre-fire free drop event can be seen in each figure.

A similar thermal performance is seen for the package when loaded with the Loose Fuel Plate Basket (LFPB). Figure 3.4-5 presents the thermal response curve, while Figure 3.4-6 and Figure 3.4-7 present perspective views of the temperature distribution in the ATR FFSC package after 30 minutes of exposure to the HAC fire and at the point when the peak LFPB temperature is attained (approximately 22 minutes after the end of the fire). A lower maximum temperature is achieved in the LFPB vs. that seen for the FHE because of the higher thermal mass associated with the LFPB. Further, since the LFPB is modeled without its payload of loose fuel plates, these results will bound those seen for a LFPB with a payload.

Table 3.4-1 presents the component temperatures seen prior to the fire, at the end of the 30-minute fire event, and the peak temperature achieved during the entire simulated HAC thermal event. As seen, all temperatures are within their allowable limit. It is expected that the neoprene rub strips and the polyethylene bag used as a protective sleeve for the ATR fuel element will thermally degrade due to the level of temperature achieved. In the case of the polyethylene bag, the bag is expected to melt and sag onto the fuel element when exposed to temperatures in excess of 250F. Further heating will lead to charring and then thermal decomposition into its volatile components. While spontaneous ignition is unexpected under the unpiloted conditions, the effect would be minimal since, per Section 3.5.3, Thermal Decomposition/Combustion of Package Organics, the available oxygen in the package is sufficient for consumption of less than 1% of the polyethylene. As a thermoset polymer, the neoprene is expected to simply decompose into volatiles as it thermally degrades. These components are not critical to the safety of the package and any out-gassing associated with their thermal degradation will not contribute to package pressurization since package is vented.

The results presented above also demonstrate that inclusion of insolation effects prior to the fire would not have affected the safety basis of the design. The low thermal mass of the package effectively mitigates the HAC impact of higher initial component temperatures due to insolation.

As seen from Table 3.3-1, consideration of the maximum insolation loading raises the package component temperatures by approximately 50°F above the initial 100°F level assumed by the HAC evaluation. The thermal response curves presented in Figures 3.4-1 and 3.4-2 demonstrate that the fire condition recovers this 50°F temperature difference for the outer components within the first few seconds of fire exposure. Further, since all package components exhibit thermal margins greater than 50°F as shown in Table 3.4-1, the inclusion of insolation effects prior to the fire event would not have impacted the safety basis for the design.

3.4.3.2 Maximum HAC Pressures The payload cavity of the ATR FFSC is vented to the atmosphere. As such, the maximum pressure achieved under the HAC event will be 0 psig. Section 3.5.3, Thermal

Docket No. 71-9330 ATR FFSC Safety Analysis Report Rev. 17, May 2022 3-23 Decomposition/Combustion of Package Organics, provides the justification for assuming a 0 psig package pressure for the HAC event.

Although the volume between the outer and inner shells is sealed, it does not contain organic or other materials that may outgas or thermally degrade. Assuming a temperature of 70°F at the time of assembly and a maximum temperature of 1,475°F (based on the outer shell temperature, see Table 3.4-1), the maximum pressure rise within the sealed volume due to ideal gas expansion will be less than 39 psig. This level of pressurization will occur for only a few minutes and then quickly reduce as the package cools.

3.4.4 Maximum Thermal Stresses The temperature difference between the inner and outer shells during the HAC event (see the average inner and outer shell temperatures presented in Figure 3.4-1) will result in differential thermal expansion between the shells. This differential thermal expansion is expected to peak at approximately 6 minutes after the initiation of fire exposure when the average outer shell temperature is 1,344°F and the average inner shell temperature is 196°F. Based on the differential thermal expansion for Type 304 stainless steel19 the change in length is computed as:

inches L

T T

L L

DTE IS IS OS OS InnerShell OuterShell 9.0 70 70

where:

OS = 10.7(10-6) in/in/ºF at 1,300 ºF IS = 8.9(10-6) in/in/ºF at 200 ºF TOS = 1,344 ºF TIS = 196 ºF L = 73 inches (conservatively for both shells)

After 6 minutes of exposure to the fire the difference in shell lengths will decrease as the inner shell heats up. The differential expansion will reach 0-inches approximately 6 minutes after the end of the fire event when the inner and outer shells reach thermal equilibrium and then go negative as the outer shell continues to cool faster than the inner shell. The largest negative thermal differential expansion achieved is approximately 0.22-inches.

The result of this variation in differential thermal expansion may take one of three forms:

1) the outer shell buckles outward,

2) the outer shell buckles inward, or

3) the weld attaching the inner shell to either the closure plate or the bottom end plate will fail and permit the outer shell and the affected plate to move freely.

While in reality, a square tube is likely to buckle inward on two of the four faces and outward on the remaining two faces simultaneously, the two buckling modes are treated independently for the purposes of this evaluation. The possibility of the outer shell buckling outwards is the assumption upon which the thermal modeling presented in Section 3.5.2.2, Description of 19 American Society of Mechanical Engineers (ASME) Boiler and Pressure Vessel Code,Section II, Materials, Part D - Properties, 2001 Edition, 2002 and 2003 Addenda, New York,Table TE-1, Group 3. Coefficient B = 8.9x10-6 inches/inch/°F at 200°F and 10.7x10-6 inches/inch/°F at 1,300°F.

Docket No. 71-9330 ATR FFSC Safety Analysis Report Rev. 17, May 2022 3-24 Thermal Model for HAC Conditions is based. This mode is seen as likely given the level of metal softening that will occur with the outer shell quickly reaching over 1,200°F and the expected pressurization of the void space between the inner and outer shells. Buckling the outer shell in this fashion will act to lower the rate of inward heat transfer. As such, ignoring the outer shells displacement due to differential thermal expansion, as assumed by the HAC thermal modeling, yields conservatively high package temperatures.

The second possibility is that the outer shell buckles inward under the differential thermal expansion. Should this occur, the maximum deflection would be 0.9-inches/2 = 0.45-inches assuming a zero length deflection and only one buckle along the length of the outer shell. In reality, the actual deflection would measure perhaps 0.33-inches after properly accounting for the curvature in the buckled section. Since this level of deflection would still leave 0.5-inches or more of insulation separating the inner shell from the outer shell, no significant impact on the predicted peak HAC temperatures will occur.

The final possibility which the differential thermal expansion may manifest itself is in the failure of the one of the welds attaching the inner shell to the closure and bottom end plates. If this occurs, besides releasing any potential pressure buildup in the void between the inner and outer shells, the outer shell and the associated end plate will extend away from the inner shell at the point of the weld failure. The size of the gap will maximize at about 0.9-inches and then decrease. Since the insulation jacket is cut out to fit around the hardware used to index the packages to one another, the insulation jacket and the underlying insulation will be pulled in the same direction as the outer shell, thus preventing the creation of a gap between the interface of the insulation wrap and the end plate. Even if such a gap would occur, no direct exposure of cavity within the inner shell to the outer shell surfaces will result since the closure plugs at each end of the package are longer than the predicted movement under differential thermal expansion.

Instead, the likely and worst case scenario is that the movement of the outer shell, the insulation jacket, and the insulation will create a gap of approximately 0.9-inches at the interface between the first support rib and the insulation. Combining this gap with an insulation shift of up to 1.75-inches at this same locations due to a pre-fire, 30-foot end drop (see Section 3.5.2.2, Description of Thermal Model for HAC Conditions) could result in a scenario where there is a 0.9-inch gap between the support rib and the insulation jacket and up to a 0.9 + 1.75 = 2.65-inch gap between the support rib and the end of the insulation wrap. A sensitivity thermal analysis of this geometry showed that the peak inner shell temperature reported in Table 3.4-1 remained bounding, while the maximum temperature of the ATR fuel element increased by less than 25ºF.

This modest change in temperature occurs because there is little difference in temperature between the outer shell and the stainless steel insulation wrap. Since this level of temperature increase is well within the thermal margins apparent from Table 3.4-1, the potential thermal impact due to the package geometry displacement under differential thermal expansion is seen as being not significant to the safety of the package.

3.4.5 Accident Conditions for Air Transport of Fissile Material 10 CFR §71.55(f) requires that the package be subcritical subsequent to the application of a series of accident condition tests, including a thermal test. A criticality analysis of the worst-case geometric configuration of the packaging and contents materials is performed in Section 6.7, Fissile Material Packages for Air Transport, which considers the presence of all of the

Docket No. 71-9330 ATR FFSC Safety Analysis Report Rev. 17, May 2022 3-25 moderating and reflecting material in the package. The tendency of the fire event to decrease the availability of moderating material due to combustion is conservatively neglected. Thus, the effects of the fire test of 10 CFR §71.55(f)(1)(iv) do not need to be specifically evaluated.

Table 3.4 HAC Temperatures Location / Component Pre-fire End of Fire Peak Maximum Allowable ATR Fuel Element Fuel Plate 100°F 586°F 730°F 1,100°F ATR Fuel Element Side Plate 100°F 676°F 827°F 1,100°F Neoprene Rub Strips/

Polyethylene Bag 100°F 1,016°F 1,017°F N/A Fuel Handling Enclosure (FHE) 100°F 1,016°F 1,017°F 1,100°F Loose Fuel Plate Basket (LFPB) 100°F 584°F 746°F 1,100°F Inner Shell 100°F 1,422°F 1,422°F 2,700°F Ceramic Fiber Insulation, Body

- Maximum

- Average 100°F 100°F 1,460°F 1,220°F 1,460°F 1,220°F 2,300°F 2,300°F Ceramic Fiber Insulation, Closure

- Maximum

- Average 100°F 100°F 1,418°F 1,297°F 1,418°F 1,297°F 2,300°F 2,300°F Closure 100°F 1,445°F 1,445°F 2,700°F Outer Shell 100°F 1,471°F 1,471°F 2,700°F Table Notes:

The maximum allowable temperatures under HAC conditions are provided in Section 3.2.2, Technical Specifications of Components.

Docket No. 71-9330 ATR FFSC Safety Analysis Report Rev. 17, May 2022 3-26 50 150 250 350 450 550 650 750 850 950 1050 1150 1250 1350 1450 1550

-30 30 90 150 210 270 330 390 450 510 570 630 690 Temperature - F Time - Minutes Outer Shell - Max.

Outer Shell - Avg.

Insulation Jacket - Max.

Insulation - Avg.

Inner Shell - Max.

Fuel Handling Enclosure (FHE) - Max.

Figure 3.4 ATR FFSC Package Thermal Response to HAC Event 50 150 250 350 450 550 650 750 850 950 1050 1150 1250 1350 1450 1550

-30 30 90 150 210 270 330 390 450 510 570 630 690 Temperature - F Time - Minutes ATR Side Plates - Max.

ATR Side Plates - Avg.

ATR Fuel Plates - Max.

ATR Fuel Plates - Avg.

Inner Shell - Max.

Fuel Handling Enclosure (FHE) - Max.

Figure 3.4 ATR Fuel Element Thermal Response to HAC Event

Docket No. 71-9330 ATR FFSC Safety Analysis Report Rev. 17, May 2022 3-27 (Note: the positive x-axis is oriented towards the top of the package and the positive z-axis towards the package closure end)

Figure 3.4-3 -Temperature Distribution at End of HAC 30-Minute Fire (Note: the positive x-axis is oriented towards the top of the package and the positive z-axis towards the package closure end)

Figure 3.4-4 -Temperature Distribution at Peak ATR Fuel Element Temperature

Docket No. 71-9330 ATR FFSC Safety Analysis Report Rev. 17, May 2022 3-28 50 150 250 350 450 550 650 750 850 950 1050 1150 1250 1350 1450 1550

-30 30 90 150 210 270 330 390 450 510 570 630 690 Temperature - F Time - Minutes Outer Shell - Max.

Outer Shell - Avg.

Insulation Jacket - Max.

Insulation - Avg.

Inner Shell - Max.

Loose Fuel Plate Basket - Max.

Figure 3.4 ATR FFSC Package with LFPB Thermal Response to HAC Event (Note: the positive x-axis is oriented towards the top of the package and the positive z-axis towards the package closure end)

Figure 3.4 FFSC-LFPB Temperature Distribution at End of HAC Fire

Docket No. 71-9330 ATR FFSC Safety Analysis Report Rev. 17, May 2022 3-29 (Note: the positive x-axis is oriented towards the top of the package and the positive z-axis towards the package closure end)

Figure 3.4 FFSC-LFPB Temperature Distribution at Peak LFPB Temperature

Docket No. 71-9330 ATR FFSC Safety Analysis Report Rev. 17, May 2022 3-30 3.5 Appendices 3.5.1 Computer Analysis Results 3.5.2 Analytical Thermal Model 3.5.3 Combustion/Decomposition of Package Organics

Docket No. 71-9330 ATR FFSC Safety Analysis Report Rev. 17, May 2022 3-31 3.5.1 Computer Analysis Results Due to the size and number of the output files associated with each analyzed condition, results from the computer analysis are provided on a CD-ROM.

3.5.2 Analytical Thermal Model The analytical thermal model of the ATR FFSC package was developed for use with the Thermal Desktop 20 and SINDA/FLUINT21 computer programs. These programs are designed to function together to build, exercise, and post-process a thermal model. The Thermal Desktop computer program is used to provide graphical input and output display function, as well as computing the radiation exchange conductors for the defined geometry and optical properties.

Thermal Desktop is designed to run as an AutoCAD application. As such, all of the CAD tools available for generating geometry within AutoCAD can be used for generating a thermal model. In addition, the use of the AutoCAD layers tool presents a convenient means of segregating the thermal model into its various elements.

The SINDA/FLUINT computer program is a general purpose code that handles problems defined in finite difference (i.e., lumped parameter) and/or finite element terms and can be used to compute the steady-state and transient behavior of the modeled system. Although the code can be used to solve any physical problem governed by diffusion-type equations, specialized functions used to address the physics of heat transfer and fluid flow make the code primarily a thermal code.

The SINDA/FLUINT and Thermal Desktop computer programs have been validated for safety basis calculations for nuclear related projects22,23.

Together, the Thermal Desktop and SINDA/FLUINT codes provide the capability to simulate steady-state and transient temperatures using temperature dependent material properties and heat transfer via conduction, convection, and radiation. Complex algorithms may be programmed into the solution process for the purposes of computing heat transfer coefficients as a function of the local geometry, gas thermal properties as a function of species content, temperature, and pressure, or, for example, to estimate the effects of buoyancy driven heat transfer as a function of density differences and flow geometry.

3.5.2.1 Description of Thermal Model for NCT Conditions A 3-dimensional, one-quarter symmetry thermal model of the ATR FFSC is used for the NCT evaluation. The model simulates one-quarter of the package, extending from the closure to the 20 Thermal Desktop, Versions 4.8 and 5.1, Cullimore & Ring Technologies, Inc., Littleton, CO, 2005/2007.

21 SINDA/FLUINT, Systems Improved Numerical Differencing Analyzer and Fluid Integrator, Versions 4.8 and 5.1, Cullimore & Ring Technologies, Inc., Littleton, CO, 2005/2007.

22 Software Validation Test Report for Thermal Desktop and SINDA/FLUINT, Versions 4.8 and 5.1, Packaging Technology, Inc., File No. TR-VV-05-001, Rev. 1 and Rev. 2.

23 Thermal Desktop and SINDA/FLUINT Testing and Acceptance Report, Version 5.1, AREVA Federal Services, LLC, File No. AFS-TR-VV-006, Rev. 0.

Docket No. 71-9330 ATR FFSC Safety Analysis Report Rev. 17, May 2022 3-32 axial centerline of the package. Symmetry conditions are assumed about the packages vertical axis and at the axial centerline. This modeling choice captures the full height of the package components and allows the incorporation of the varying insolation loads that will occur at the top and sides of the package. Program features within the Thermal Desktop computer program automatically compute the various areas, lengths, thermal conductors, and view factors involved in determining the individual elements that make up the thermal model of the complete assembly.

Figure 3.5-1 and Figure 3.5-2 illustrate the solid and hidden line views of the package thermal model. The model simulates one-half of the closure end half of the package (i.e.,

symmetry is assumed about the packages vertical plane) and extends approximately 36.5 inches in the axial direction (e.g., from closure to the mid-point of the center support rib). As seen from the figure, the modeling captures the various components of the packaging, including the index lug and mating pocket used to align the stacked packages, the recessed exterior surface area of the package closure, the FHE, and the ATR fuel element. Also captured, but not easily seen in the figure due to the scale of the figures, are the nineteen (19) individual fuel plates that comprise the ATR fuel element.

The model is composed of solid and plate type elements representing the various package components. Thermal communication between the various components is via conduction, radiation, and surface-to-surface contact. Since the ATR FFSC Package dissipates essentially no decay heat, the peak temperatures internal to the package are driven by the external heating occurring during NCT and HAC conditions. While the potential for developing convective flows within the air filled cavity between the outer shell and the insulation jacket is small due to the cavity dimensions, if convective heat transfer was to develop it could raise the peak temperatures developed under either NCT or HAC conditions since it would reduce the thermal resistance to heat flowing inward from the outer shell. To address this possibility, the thermal conductivity associated with the air overpack nodes in the lower quadrant of the package are increased by a factor of 2 from that for conduction as a means of simulating the type of enhanced heat transfer that convection would cause. The affected nodes are limited to those in the lower quadrant of the package since, in the assumed horizontal orientation of the package under both NCT and HAC conditions, the buoyancy forces associated with convection will tend to drive the flow in this portion of the package in a circular motion, but would only produce a stratified temperature layer in the upper quadrant. However, since subsequent examination of the temperature distribution at the end of the fire event showed no discernible difference in the insulation jacket temperature between the upper and lower quadrants, it is concluded that the heat transfer within these cavities is dominated by radiation and conduction and the potential for convective heat transfer can be ignored. Despite this conclusion, the factor of 2 has been retained in the models as a conservatism.

A total of approximately 8,050 nodes, 2,800 planar elements, and 3,700 solid elements are used to simulate the modeled components. In addition, one boundary node is used to represent the ambient environment for convection purposes and a second boundary node is used to represent the ambient temperature for the purpose of radiation heat transfer.

Figure 3.5-3 and Figure 3.5-4 illustrate the quarter symmetry thermal models of the FHE and the ATR fuel element. The FHE thermal model uses planar elements to represent the 0.09 inch thick sides of the enclosure, while solid elements are used to represent the 0.25 inch thick end cap.

Docket No. 71-9330 ATR FFSC Safety Analysis Report Rev. 17, May 2022 3-33 Heat transfer between the FHE and the inner shell of the package is modeled as a combination of radiation and conduction across the air-filled void space, as well as via direct contact along 3 edges of the FHE. The contact conductance simulates the physical contact between an impact deformed FHE and the inner shell. Figure 3.5-5 illustrates a cross-section through the combined modeling for the inner shell, the FHE, and the ATR fuel element. The left side of the figure illustrates the placement of the thermal nodes (indicated by the small circles) used to simulate each of the components, the use of curved elements to represent the 19 fuel plates, and the assumed points of direct contact between the FHE and the inner shell. The right side of the figure includes depiction of the solid elements that are used to simulate the air voids in and around the FHE. The heat transfer between the FHE and the ATR fuel element is computed as conductance through the 0.125 inch thick neoprene rub strips (see Figure 3.5-5) and radiation and conductance through the air voids.

The heat transfer due to direct contact conservatively assumes the FHE has been deformed as a result of the HAC drop event to create flat areas measuring 0.5 inches wide at the lower 2 points of contact, 0.75 inches wide at the top, and extending the entire length of the FHE.

Although this type of damage would only occur for the HAC condition (if it occurs at all), it is conservatively assumed for the NCT modeling as well. A conservatively high contact conductance9 of 1 Btu/min-in2-ºF is assumed.

A detailed model of the ATR fuel element is used to simulate the heat transfer within the fuel element and between the fuel element and the FHE. The detailed thermal model, illustrated in Figure 3.5-4 and Figure 3.5-5, includes a separate representation of each composite fuel plate, the side plates (including the cutouts), and the upper end box casting. Heat transfer between the individual fuel plates is simulated via conduction and radiation across the air space separating the plates. The curvature and separation distance between the plates is based on the information presented in Section 3.5.2.4, Determination of Composite Thermal Properties for ATR Fuel Plates. Each quarter segment of the fuel plates is represented by four thermal nodes in the circumferential direction and 16 nodes along its length.

The thermal modeling for the Loose Fuel Plate Basket uses the same model for the ATR FFSC, but replaces the thermal modeling of the FHE and the ATR fuel element with the thermal modeling for the Loose Fuel Plate Basket depicted in Figure 3.5-6. Approximately 500 nodes, 280 planar elements, and 530 solids are used to simulate the basket. Since the payload for the basket may contain a variable number and size of fuel plates, the thermal modeling is based on an empty basket. This approach is conservative since the addition of a payload will serve to increase the thermal mass of the basket and, thus, reduce its temperature rise under the transient conditions associated with the HAC event. Since the unirradiated fuel plates have essentially zero decay heat, there will be no temperature rise between the loose fuel plates and the basket. As such, modeling of the loose fuel plate payload is both unnecessary and conservative for the purposes of this evaluation.

The heat transfer from the exterior surfaces of the ATR FFSC is modeled as a combination of convection and radiation exchange. Appendix 3.5.2.3, Convection Coefficient Calculation, presents the methodology used to compute the convection coefficients from the various surfaces.

The radiation exchange is computed using a Monte Carlo, ray tracing technique and includes the affect of reflection and/or transmission, according to the optical properties assigned to each surface (see Section 3.2.1, Material Properties).

Docket No. 71-9330 ATR FFSC Safety Analysis Report Rev. 17, May 2022 3-34 In addition, heating of the exterior surfaces due to solar insolation is assumed using a diurnal cycle. A sine wave model is used to simulate the variation in the applied insolation on the surfaces of the package over a 24-hour period, except that when the sine function is negative, the insolation level is set to zero. The timing of the sine wave is set to achieve its peak at 12 pm and peak value of the curve is adjusted to ensure that the total energy delivered matched the regulatory values. As such, the total energy delivered in one day by the sine wave solar model is given by:

6 hr

18 hr

t Q peak sin t

12 hr

2

d 24 hr

Q peak

Using the expression above for the peak rate of insolation, the peak rates for top and side insolation may be calculated as follows:

2 2

top 2

top in min Btu 0447

.0 in hr Btu 68

.2 Q

hr 24 cm cal 800 Q

2 2

sside 2

side in min Btu 0112

.0 in hr Btu 67

.0 Q

hr 24 cm cal 200 Q

Conversion factors of 1 cal/cm2-hr = 0.0256 Btu/hr-in2 are used in the above calculations. These peak rates are multiplied by the sine function and the solar absorptivity for Type 304 stainless steel (i.e., 0.52) to create the top and side insolation values as a function of time of day.

3.5.2.2 Description of Thermal Model for HAC Conditions The thermal evaluations for the hypothetical accident condition (HAC) are conducted using an analytical thermal model of the ATR FFSC. The HAC thermal model is a modified version of the quarter symmetry NCT model described in Section 3.5.2.1, Description of Thermal Model for NCT Conditions, with the principal model modifications consisting of simulating the expected package damage resulting from the drop events that are assumed to precede the HAC fire and changing the package surface emissivities to reflect the assumed presence of soot and/or surface oxidization.

Physical testing using full scale certified test units (CTUs) is used to establish the expected level of damage sustained by the ATR FFSC package from the 10 CFR 71.73 prescribed free and puncture drops that are assumed to precede the HAC fire event. Appendix 2.12.1, Certification Tests on CTU-1 and Appendix 2.12.2, Certification Tests on CTU-2 document the configuration and initial conditions of the test articles, the test facilities, the instrumentation used, and the test results. The drop tests covered a range of hypothetical free drop orientations and puncture bar drops. The results from both sets of CTU drop tests showed the following:

Docket No. 71-9330 ATR FFSC Safety Analysis Report Rev. 17, May 2022 3-35

1) The worst case physical damage to the exterior of the package occurs from a CG over corner drop. The resulting damage (depicted in Figure 3.5-7) is thermally insignificant in that there is no breach in the outer shell and the compaction of the underlying insulation is minor and offset by an increase in the gap between the outer shell and the insulation in other areas.

2) The oblique, CG over side puncture bar drop caused a 0.5 inch indentation to the side of the package at the center of the impact region and less near the edges. No tearing of the outer shell occurred.

3) The end drops caused the ceramic fiber insulation to slide axially between each set of ribs, as depicted in Figure 3.5-9. The amount of re-positioning varied from approximately 1 to 1.75 inches and results in the compression of the insulation in the axial direction by 6 to 10%. No compression or shifting of the insulation in the radial direction was noted from the drop tests. While the insulation jacket showed some crimping at the edges, it was essentially undamaged.

Based on the above observations, the NCT was modified for the HAC evaluations via the following steps:

1) A 1.85 inch long segment of insulation was removed between each set of ribs. This degree of insulation re-positioning/compression conservatively bounds the maximum observed distance of 1.75 inches. Heat transfer across the vacated segments of insulation is then computed as radiation and conduction across an air filled space. Figure 3.5-10 illustrates the change made to the NCT thermal model to capture the expected insulation re-positioning. The change in the insulations thermal conductivity as a result of the compression is conservatively ignored since thermal conductivity decreases with density at temperatures in excess of approximately 500°F (see Table 3.2-3).

2) All other geometric aspects of the NCT thermal model are assumed to be unchanged for the HAC evaluations since the observed damage to the outer shell resulting from the free and puncture drops has a superficial impact to the thermal protection offered by the ATR FFSC to the HAC fire event.

3) The surface emissivities for the various components of the package are revised as presented in Table 3.2-6 vs. that given in Table 3.2-5.

3.5.2.3 Convection Coefficient Calculation The convective heat transfer coefficient, hc, has a form of:

L k

Nu hc

, where k is the thermal conductivity of the gas at the mean film temperature and L is the characteristic length of the vertical or horizontal surface.

Natural convection from each surface is computed based on semi-empirical relationships using the local Rayleigh number and the characteristic length for the surface. The Rayleigh number is defined as:

Pr

T L

g

Ra 2

3 c

2 L

Docket No. 71-9330 ATR FFSC Safety Analysis Report Rev. 17, May 2022 3-36 where:

gc = gravitational acceleration, 32.174 ft/s2

= coefficient of thermal expansion, °R-1 T = temperature difference, °F

= density of air at the film temperature, lbm/ft3

= dynamic viscosity, lbm/ft-s Pr = Prandtl number = (cp ) / k L = characteristic length, ft k = thermal conductivity at film temperature cp = specific heat, Btu/lbm-hr-°F RaL = Rayleigh #, based on length L Note that k, cp, and are each a function of air temperature as taken from Table 3.2-4. Values for are computed using the ideal gas law, for an ideal gas is simply the inverse of the absolute temperature of the gas, and Pr is computed using the values for k, cp, and from Table 3.2-4. Unit conversion factors are used as required to reconcile the units for the various properties used.

The natural convection from a discrete vertical surface is computed using Equation 6.39 to 6.42 of Rohsenow, et. al.9, which is applicable over the range 1 < Rayleigh number (Ra) < 1012:

4 1

L T

Ra C

Nu

9

/

4 16

/

9 Pr 492

.0 1

671

.0 C

L

)

2.8/Nu ln(1 2.8 Nu T

L

1/3 V

t t

Ra C

Nu

42

.0 0.81 0.22 V

t Pr 0.61 1

Pr 0.13 C

6 1

6 t

6 L

c

)

(Nu

)

(Nu k

L h

Nu

Natural convection from horizontal surfaces is computed from Equations 4.39 and 4.40 of Rohsenow, et. al.9, and Equations 3.34 to 3.36 of Guyer24, where the characteristic dimension (L) is equal to the plate surface area divided by the plate perimeter. For a heated surface facing upwards or a cooled surface facing downwards and Ra > 1:

1/10 10 t

10 L

c

)

(Nu

)

(Nu k

L h

Nu

4 1

L L

Ra C

1.677 1

ln 4.1 Nu

24 Guyer, E.C., Handbook of Applied Thermal Design, McGraw-Hill, Inc., 1989.

Docket No. 71-9330 ATR FFSC Safety Analysis Report Rev. 17, May 2022 3-37

9

/

4 16

/

9 L

Pr 492

.0 1

671

.0 C

3 1

t Ra 14

.0 Nu For a heated surface facing downwards or a cooled surface facing upwards and 103 < Ra < 1010, the correlation is as follows:

T L

2.5/Nu 1

ln 2.5 Nu Nu

1/5 9

/

2 10

/

9 T

Ra Pr 9.1 1

527

.0 Nu

The forced convection coefficients applied during the HAC fire event are computed using the relationships in Table 6-5 of Kreith25 for a flat surface, where the characteristic dimension (L) is equal to the length along the surface and the free stream flow velocity is V. The heat transfer coefficient is computed based on the local Reynolds number, where the Reynolds number is defined as:

L

V ReL

For Reynolds number (Re) < 5x105 and Prandtl number (Pr) > 0.1:

33

.0 5.0 Pr Re 664

.0 L

Nu For Reynolds number (Re) > 5x105 and Prandtl number (Pr) > 0.5:

]

200 23

[Re Pr 036

.0 8.0 33

.0

L Nu Given the turbulent nature of the 30-minute fire event, a characteristic length of 0.25 feet is used for all surfaces to define the probable limited distance for boundary growth. The turbulent heat transfer coefficient relationship used for HAC modeling is a modified version of the Colburn relation recommended by the advisory material for the IAEA (see Advisory Material for the IAEA Regulations for the Safe Transport of Radioactive Material, TS-G-1.1, Rev. 1, International Atomic Energy Agency, 2008). The same advisory material states that "pool fire gas velocities are generally found to be in the range of 5-10 m/s". The above forced convection relationships yields a convection heat transfer rate of approximately 40 W/m2-K, which matches that obtained with the IAEA recommended Colburn relation and conservatively bounds the experimental values in large pool fires.

25 Kreith, Frank, Principles of Heat Transfer, 3rd edition, Harper & Row, 1973.

Docket No. 71-9330 ATR FFSC Safety Analysis Report Rev. 17, May 2022 3-38 3.5.2.4 Determination of Composite Thermal Properties for ATR Fuel Plates The ATR fuel plates are a composite material consisting of a fissile fuel matrix sandwiched within aluminum cladding. For the purposes of this calculation, the fuel composite is treated as a homogenous material with lumped thermal properties as defined below. This modeling approach is justified since the thermal gradient within the fuel element will be very low given that the un-irradiated fuel has essentially no decay heat.

Because of the thinness of the plates, the average conductivity is required only for the axial and circumferential direction.

Conductivity through the plates is not required as this analysis assumes a zero temperature gradient in that direction. Mean density and specific heat values are also defined below.

Circumferential and Axial Conductivity Ignoring the affect of curvature, the heat flow can be written as, y

T k

z x

y T

k z

x y

T k

z x

q

2 2

1 1

where

i ix x

From which, x

k x

k x

k

2 2

1 1

Mean Density The mean density of the fuel plates is computed from:

2 2

1 1

z y

x z

y x

z y

x Mass

, from which we get x

x x

2 2

1 1

Mean Specific Heat In the same manner used to define the mean density, the mean specific heat for the fuel plates is computed as; z

y x

c z

y x

c z

y x

c p

p p

2 2

1 1

2 1

from which we get, x

x c

x c

c p

p p

2 2

1 1

2 1

The thermal properties for the individual plates making up the ATR fuel element are computed using the above approach and thermophysical and geometric data6,5 for the ATR fuel element.

k1 k2 x1 x2 y

Docket No. 71-9330 ATR FFSC Safety Analysis Report Rev. 17, May 2022 3-39 Based on these data sources, the radius of the inner plate is 3.015 inches, while the radius of the outer plate is 5.44 inches. The gap between the plates is 0.078 inches. The thickness of the aluminum cladding is 0.015 inches.

While the thermal properties for the aluminum cladding and the fissile fuel matrix material will vary with temperature, for the purposes of this evaluation, fixed material properties are assumed in order to simplify the calculation. To provide conservatism for this modeling approach conservatively low value is assumed for the specific heat for each component, while a conservatively high thermal conductivity value is used. This methodology will result in over-predicting the temperature rise within the composite material during the HAC fire event.

The thermal properties used in this calculation are:

1) Aluminum cladding thermal conductivity = 191 W/m-K, conservatively high value from

[6], page 18

2) Fissile fuel matrix (UAlx) conductivity:

a. 53 W/m-K, conservatively high based on Table 2.3 from [6], at 300K for fuel plates 1, 2, 18, & 19

b. 43 W/m-K, conservatively high based on Table 2.3 from [6], at 300K for fuel plates 3, 4, 16, & 17

c. 36.1 W/m-K, conservatively high based on Table 2.3 from [6], at 300K for fuel plates 5 to 15

3) Aluminum cladding density = 2702 kg/m3, from [6], page 16

4) Fissile fuel matrix (UAlx) density:

a. 3409 kg/m3, from [6], Table 2.5, for fuel plates 1, 2, 18, & 19

b. 3671 kg/m3, from [6], Table 2.5, for fuel plates 3, 4, 16, & 17

c. 3933 kg/m3, from [6], Table 2.5, for fuel plates 5 to 15

5) Aluminum cladding specific heat = 896 and 1080 J/kg-K, from [6], Table 3.2, at 300 &

600K, respectively

6) Fissile fuel matrix (UAlx) specific heat:

a. 666 & 803 J/kg-K, from [6], Table 2.4, value at 300 & 700K, respectively, for fuel plates 1, 2, 18, & 19

b. 616 & 743 J/kg-K, from [6], Table 2.4, value at 300 & 700K, respectively, for fuel plates 3, 4, 16, & 17

c. 573 & 692 J/kg-K, from [6], Table 2.4, value at 300 & 700K, respectively, fuel plates 5 to 15 Table 3.5-1 presents the composite thermal conductivity, specific heat, and density values for each of the nineteen (19) fuel plates making up the ATR fuel element. These composite values are based on the thermal property values given above and the geometry depicted in Figure 3.5-11.

Docket No. 71-9330 ATR FFSC Safety Analysis Report Rev. 17, May 2022 3-40 Table 3.5 Composite ATR Fuel Plate Thermal Properties Plate Plate Thickness, in UAlx Thickness, in Circumferential Conductivity (W/m-K)

Inner radius, in Outer radius, in Mean radius, in Mean density, kg/m^3 Mean specific heat, J/(kg K) @ 300 K Mean specific heat, J/(kg K) @ 700 K 1

0.08 0.05 104.8 3.015 3.095 3.055 3143.9 740.1 892.3 2

0.05 0.02 135.8 3.173 3.223 3.198 2984.8 790.9 953.5 3

0.05 0.02 131.8 3.301 3.351 3.326 3089.6 762.9 919.8 4

0.05 0.02 131.8 3.429 3.479 3.454 3089.6 762.9 919.8 5

0.05 0.02 129.0 3.557 3.607 3.582 3194.4 736.9 888.9 6

0.05 0.02 129.0 3.685 3.735 3.710 3194.4 736.9 888.9 7

0.05 0.02 129.0 3.813 3.863 3.838 3194.4 736.9 888.9 8

0.05 0.02 129.0 3.941 3.991 3.966 3194.4 736.9 888.9 9

0.05 0.02 129.0 4.069 4.119 4.094 3194.4 736.9 888.9 10 0.05 0.02 129.0 4.197 4.247 4.222 3194.4 736.9 888.9 11 0.05 0.02 129.0 4.325 4.375 4.350 3194.4 736.9 888.9 12 0.05 0.02 129.0 4.453 4.503 4.478 3194.4 736.9 888.9 13 0.05 0.02 129.0 4.581 4.631 4.606 3194.4 736.9 888.9 14 0.05 0.02 129.0 4.709 4.759 4.734 3194.4 736.9 888.9 15 0.05 0.02 129.0 4.837 4.887 4.862 3194.4 736.9 888.9 16 0.05 0.02 131.8 4.965 5.015 4.990 3089.6 762.9 919.8 17 0.05 0.02 131.8 5.093 5.143 5.118 3089.6 762.9 919.8 18 0.05 0.02 135.8 5.221 5.271 5.246 2984.8 790.9 953.5 19 0.1 0.07 94.4 5.349 5.449 5.399 3196.9 724.3 873.2 An average UAlx thickness of 0.020 inches exists for Plates 1 an 19 vs. the 0.05 and 0.07 inches assumed by this analysis based on the assumption of a constant cladding thickness. However, for the purposes of developing composite fuel plate properties for this evaluation, the UAlx thicknesses identified in the table yield conservative bounding thermal property values.

Docket No. 71-9330 ATR FFSC Safety Analysis Report Rev. 17, May 2022 3-41 (Note: the positive x-axis is oriented towards the top of the package and the positive z-axis towards the package closure end)

(Note: the positive x-axis is oriented towards the top of the package and the positive z-axis towards the package closure end)

Figure 3.5 Solid and & Hidden Line Views of Package Quarter Symmetry Thermal Model

Docket No. 71-9330 ATR FFSC Safety Analysis Report Rev. 17, May 2022 3-42 (Note: the positive x-axis is oriented towards the top of the package and the positive z-axis towards the package closure end)

Figure 3.5 Reverse, Hidden Line View of Package Quarter Symmetry Thermal Model (Note: the positive x-axis is oriented towards the top of the package and the positive z-axis towards the package closure end)

Figure 3.5 Reverse, Hidden Line View of FHE Quarter Symmetry Thermal Model

Docket No. 71-9330 ATR FFSC Safety Analysis Report Rev. 17, May 2022 3-43 ATR Fuel Element Modeling, View Along Centerline of Element ATR Fuel Element Modeling, View Along Outside of Element (Note: the positive x-axis is oriented towards the top of the package and the positive z-axis towards the package closure end)

Figure 3.5 Centerline and Side Views of ATR Fuel Element Thermal Model

Docket No. 71-9330 ATR FFSC Safety Analysis Report Rev. 17, May 2022 3-44 Modeling Showing Direct Contact Modeling with Solid Elements for Air Figure 3.5 Thermal Model of ATR Fuel Element and FHE within Inner Shell (Note: the positive x-axis is oriented towards the top of the package and the positive z-axis towards the package closure end)

Figure 3.5 Thermal Model of Loose Fuel Plate Basket (LFPB)

Docket No. 71-9330 ATR FFSC Safety Analysis Report Rev. 17, May 2022 3-45 Figure 3.5 Worst Case Package Damage Arising from Corner Drop Figure 3.5 Worst Case Package Damage Arising from Oblique Puncture Drop Figure 3.5 Insulation Re-positioning Arising from End Drop

Docket No. 71-9330 ATR FFSC Safety Analysis Report Rev. 17, May 2022 3-46 Insulation Modeling for NCT Conditions Insulation Modeling for HAC Conditions Figure 3.5 Thermal Modeling of Insulation Re-positioning for HAC Conditions Figure 3.5 ATR Fuel Element Cross Section

Docket No. 71-9330 ATR FFSC Safety Analysis Report Rev. 17, May 2022 3-47 3.5.3 Thermal Decomposition/Combustion of Package Organics The organic material in the ATR FFSC subject to thermal decomposition and/or combustion is limited to polyethylene, neoprene, and the adhesive used to attach the neoprene, as well as up to 4 kg of cellulosic material. The fuel elements and, optionally, the loose fuel plates are enclosed in polyethylene bags prior to their placement in the FHEs and loose plate basket. The bags serve no safety function beyond providing investment protection of the payload material. Similarly, neoprene (polychloroprene) rub strips are attached via adhesive to the FHEs to provide investment protection against fretting on the elements and loose plates. As such, the loss of the organic material under either NCT or HAC conditions has no safety implication beyond the potential for gas and heat generation. The following sections provide a bounding assessment on the potential safety impact associated with the loss of organic material within the ATR FFSC package.

3.5.3.1 Organic Material Within Package The amount of organic material in the package varies with the payload configuration. While the bounding amount of polyethylene is constant at 200 g, the amount of neoprene varies with payload configuration. The sections below identify the quantity and important thermal properties associated with the organic materials present in the ATR FFSC package.

Polyethylene Properties of polyethylene related to its thermal decomposition/combustion are as follows:

a) chemical formulation7: -[CH2-CH2]n-,

b) heat of combustion (Hc)28: 46,500 kJ/kg, c) oxygen index29,30: 17.4%,

d) melting temperature30: 109-135°C e) temperature for 1% decomposition30: 275°C f) autoignition temperature31: 330 to 410°C Oxygen index (OI) is the minimum oxygen concentration required to support self-sustaining combustion of the polymer. Since piloted conditions do not exist within the ATR FFSC payload cavity, self-sustaining combustion of the polyethylene can't occur when the oxygen concentration drops below 17.4%. Low oxygen concentrations will not only prevent self-sustaining combustion, but will raise the autoignition temperature. Combustion of polyethylene in air is governed by the following equation:

2 2

2 2

2 4

2 11.28N O

H 2

CO 2

)

3.76N O

(

3 H

C

The above equation demonstrates that complete combustion of a mole of polyethylene requires 3 moles of oxygen and, since oxygen constitutes approximately 21% of air, 14.28 moles of air.

28 NUREG-1805, Fire Dynamics Tools, Nuclear Regulatory Commission, Washington, D.C.

29 office.wendallhull.com/matdb/

30 SFPE Handbook of Fire Protection Engineering, 3rd Edition, Section 1, Chapter 7, Table 1.7-4, NFPA, 2003.

31 MSDS for Polyethylene, #1488, prepared by International Programme on Chemical Safety, 2004.

Docket No. 71-9330 ATR FFSC Safety Analysis Report Rev. 17, May 2022 3-48 The total quantity of gas generated is 15.28 moles, or an increase of 7% over the original gas quantity existing before combustion. Per SAR section 1.2.2, the amount of polyethylene in the package is limited to 200 g or less. Based on a molecular weight of approximately 28 g/g-mole of polyethylene, the 7.14 g-moles of polyethylene represented by the 200 g would require 7.14 x 14.28 = 102.0 g-moles of air for complete combustion.

Neoprene Properties of neoprene (polychloroprene) related to its thermal decomposition/combustion are as follows:

a) chemical formulation7: -[CH2-Cl-C=CH-CH2]n-,

b) heat of combustion (Hc)28: 10,300 kJ/kg, c) oxygen index29,32: 32-35% at one atmosphere, d) melting temperature: N/A - thermoset material e) temperature for initial decomposition30: 342°C f) autoignition temperature32: >380°C in a 21% oxygen concentration environment.

As a thermoset plastic, uncontrolled heating of neoprene will result in reaching the decomposition temperature before the melting point is obtained. The high oxygen index demonstrates why neoprene can't support combustion without an external ignition source. The typical adhesives33 used to bond the rub strips to the FHEs consist of principally of solvents that outgas during the curing process. The non-volatile components consist of polymers, including polychloroprene, and cure and vulcanization agents. As a result, the cured adhesive layer exhibits properties33 similar to neoprene.

Combustion of neoprene in air is governed by the following equation:

2 2

2 2

2 5

4 19.74N Cl O

H 5.2 CO 4

)

3.76N O

(

25

.5 Cl H

C

From the above equation, complete combustion of a mole of neoprene is seen to require 5.25 moles of oxygen and, since oxygen constitutes approximately 21% of air, 24.99 moles of air.

The total quantity of gas generated is 27.24 moles, or an increase of 9% over the original gas quantity existing before combustion.

Based on the surface area of rub strips depicted on each SAR drawing, a thickness of 0.125 in, and an adhesive layer thickness of 2 mils, the total quantity of neoprene and neoprene like material used in each FHE is summarized in Table 3.5-2. With a molecular weight of approximately 88.5 g/g-mole of neoprene, the 4.62 g-moles of neoprene represented by the minimum 409 g of neoprene contained within the 60501-40 FHE assembly would require 4.62 x 24.99 = 115.5 g-moles of air for complete combustion.

The same limitation on package oxygen that prevents significant combustion of polyethylene will also prevent combustion of the neoprene. This same conclusion applies to the cellulosic material. Further, given the higher oxygen index and autoignition temperature for neoprene versus polyethylene, there is a low probability any neoprene material will be involved in combustion. Instead, it is expected that the only damage to be incurred by the neoprene will be a 32 Safe Use of Oxygen and Oxygen Systems, ASTM, 2nd Edition.

33 Product and MSDS sheets for 3MTM Spray 80 Neoprene Contact Adhesive or 3MTM Scotch-Weld Neoprene Contact Adhesive 1357.

Docket No. 71-9330 ATR FFSC Safety Analysis Report Rev. 17, May 2022 3-49 de-bonding from the FHE surfaces and a small amount of thermal decomposition. Since thermal decomposition is an endothermic process, the loss of the material will act to lower the temperatures predicted within the FHE.

3.5.3.2 Air Quantity Within Package Since the ATR FFSC payload cavity is not sealed, the quantity of gas filling the cavity volume will vary with time as a function of the cavity's bulk gas temperature, the thermal decomposition of the enclosed organic material, and diffusion of gas through the package closure gaps. The following sections address these various mechanisms affecting the air/oxygen content within the package.

Potential Combustion Due to Resident Air Quantity The ATR FFSC payload cavity has a length of 67.88 in and a diameter of 5.76 in. The gross cavity volume is 1768.8 in3. The ATR fuel element and the ATR FHE have volumes of approximately 155 and 223 in3, based on weights of 25 and 15 lbs, respectively, and mean densities of 0.112 and 0.097 lbs/ in3, respectively. The net cavity space is therefore approximately 1,391 in3 (22.8 liters). Table 3.5-3 summarizes the net cavity volume existing for all payload configurations. As seen from the table, only the MIT FHE (SAR drawing 60501-40) loaded with a MIT fuel element results in a larger net cavity volume than the ATR FHE (SAR drawing 60501-30) loaded with an ATR fuel element. Given the substantially higher HAC temperature predicted for the ATR FHE (see Section 3.4, Thermal Evaluation for Hypothetical Accident Conditions) versus that for the MIT FHE (see Section 3.6, Thermal Evaluation for MIT, MURR, Cobra, Small Quantity, and LEU Payloads) and the larger quantity of neoprene used (see Table 3.5-2), the ATR FHE is the appropriate payload configuration for assessing the thermal safety related to the organic material in the package.

At 100°F, approximately 0.9 g-moles of air are required to fill a 1,391 in3 (22.8 liters) cavity space to a pressure of 14.7 psia, while at 626°F (330°C, i.e., the lower autoignition temperature for polyethylene), the quantity of air required to fill the cavity space drops to approximately 0.5 g-mole. As such, the resident air quantity in the payload cavity is sufficient to support combustion of less than 0.5% of the polyethylene (i.e., 0.5 g-mole/102 g-mole air per 200 g polyethylene). The potential heat release from this quantity of polyethylene is: 0.5% x 200 g x 46,500 kJ/kg = 46,500 J = 44 Btu. Based on a combined ATR payload weight of 40 lbs and a specific heat of approximately 0.2 Btu/lbm-°F34, the net increase in the mean payload temperature would be less than 6°F even if this heat release occurred instantaneously. The use of the combined payload weight for this calculation is appropriate since the combustion occurs in the vapor space and not on a surface. Further, combustion of the limiting 0.5% of the polyethylene can neither occur instantaneously nor in only one concentrated area since the availability of the oxygen within the cavity will be rate limited by the diffusion process from reaching the potential site(s) of polyethylene combustion. In fact, oxygen diffusion will also prevent the entire resident oxygen quantity from being consumed. As such, the estimated 6°F rise in payload temperature is highly conservative.

34 Approximate specific heat of ATR fuel plates per Table 3.2-2

Docket No. 71-9330 ATR FFSC Safety Analysis Report Rev. 17, May 2022 3-50 Given the lower heat of combustion of neoprene versus that for polyethylene and the greater air quantity needed for complete combustion, the potential temperature rise from the combustion of polyethylene bounds that for neoprene by a factor of over 3.

Potential Combustion Due to Air Induced Via Pressure Forces Once the residual air existing in the payload cavity prior to the start of the HAC event is consumed, further combustion will require additional air to enter the cavity via pressure and diffusion forces. The pressure forces will arise due to the balance between ideal gas expansion/contraction and gas generation within the package versus the pressure resistance associated with gas flow through the gaps around the package closure. Heatup of the package during the 30-minute fire event will result in elevated cavity pressure and a continuous outflow of gas from the cavity. This gas flow will switch to an inflow condition once the peak bulk gas temperature is reached and the package begins to cool down.

While an accurate estimate of the gas flow due to pressure forces requires a detailed modeling of the flow paths and resistance factors, a bounding estimate on the rate of gas flow into the package due to pressure differential can be made by assuming zero vent resistance and zero internal gas generation. These assumptions assure that the minimum gas quantity is achieved at the point where packaging cooling begins, thus maximizing the potential for the reverse gas flow necessary to restore atmospheric pressure within the package.

Assuming that the bulk average gas temperature within the package is represented by the mean of the average temperatures over the length of the package's inner shell and the FHE, the cavity gas quantity within the package can be estimated as a function of time during the HAC transient.

Figure 3.5-12 presents the predicted package gas quantity for the HAC transient depicted in Figure 3.4-1 for the ATR fuel element. As seen, the package gas quantity rapidly drops during the 30-minute fire event as the cavity gas expands under HAC heating. Shortly after the cessation of the fire event, the package begins to cool and the gas flow switches to an inflow.

However, due to the rate of package cooldown, greater than 10 hours1.157407e-4 days <br />0.00278 hours <br />1.653439e-5 weeks <br />3.805e-6 months <br /> are required to restore the approximately 0.5 g-moles of gas expelled during package heatup. The calculated reverse gas flow peaks at 0.0025 g-moles per minute. The potential polyethylene combustion supported by this flow rate is 0.0025 g-moles per minute x 200 g polyethylene per 102 g-mole air x 46,500 kJ/kg = 228.1 J/min = 0.22 Btu/min. Clearly this flow rate is too low to permit any significant rate of combustion, especially when considering the facts that the reverse gas flowrate decreases rapidly from this peak rate and that accounting for flow resistance through the vent geometry will reduce this potential heat gain even more.

When the above discussion is added to the fact that the oxygen concentration at the start of the inflow condition will be well below the oxygen index of 17.5% required to support combustion, the fact that oxygen diffusion within the package will extend the time for the entering air to reach the site of elevated polyethylene temperatures, and as seen in Figure 3.5-12, that the package temperatures will fall below the lower autoignition temperature for polyethylene after 90 minutes, it is reasonable to conclude that the contribution to package heatup from airflow due to pressure differential is essentially zero.

Docket No. 71-9330 ATR FFSC Safety Analysis Report Rev. 17, May 2022 3-51 Potential Combustion Due to Air Induced Via Diffusion Beside pressure differential, the other force available to drive oxygen inflow to the package cavity is diffusion. Assuming that the oxygen inside the package cavity is consumed as fast as it enters, the rate of oxygen diffusion can be determined via Fick's first law or:

y w

D J

where: J = mass flux of oxygen per area, g/cm2 D = diffusion coefficient of oxygen in nitrogen, cm2/sec

= density of air, g/cm3 y

w

= change in mass fraction of oxygen over diffusion path While diffusion of oxygen in nitrogen is used to reflect that fact that the environment within the payload cavity is assumed to be oxygen depleted, in reality there is little difference between diffusion in air or nitrogen. The diffusion coefficient is a function of temperature and pressure.

Diffusion increases with increasing temperature since the molecules move rapidly and decreases with increasing pressure since higher fluid density increases the number of molecules per unit volume, increasing the number of collisions, thus slowing the speed of transport. The diffusion coefficient for oxygen in air at 1 atm and 25°C is 0.206 cm2/sec35. Since the fluid pressure is assumed to remain near atmospheric throughout the HAC event, there is no need to adjust the diffusion coefficient for pressure effects. However, the temperature of the fluid both within and exterior to the package will increase significantly during the HAC transient, thus necessitating an adjustment36 in the diffusion coefficient via:

N O

D, 2

N O

N O

3 N

O P

1 M

1 M

1 T

0.0018583 D

where: D = diffusion coefficient of oxygen in air, cm2/sec T = temperature, K M = molecular mass of oxygen and nitrogen P = pressure, atm D,O-N = collision integral for molecular diffusion of oxygen in nitrogen O-N = collision diameter, Angstroms From Table E.1 and the equations provided in Transport Phenomena36, MO = 31.999, MN =

28.013, O = 3.433, N = 3.667, O/ = 113, and N/ = 99.8. O-N = 0.5x(3.433 + 3.667) = 3.55.

O-N/ = (113 x 99.8)0.5 = 106.2. Assuming the maximum flame temperature of 1475°F (1075K),

the dimensionless temperature is T/O-N = 1075/106.2 = 10.1. From Table E.236, D,O-N =

0.741. Thus, DO-N at a pressure of 1 atm and 1475°F is 1.815 cm2/sec.

35 CRC Handbook of Engineering Tables, Dorf, R. editor, CRC Press LLC, 2004.

36 Transport Phenomena, 2nd Ed., Eqn 17.3-12 and Appendix E, Bird, R., Stewart, W, and Lightfoot, E., John Wiley

& Sons, Inc., 2002.

Docket No. 71-9330 ATR FFSC Safety Analysis Report Rev. 17, May 2022 3-52 The bayonet type closure plug for the ATR FFSC package results in a labyrinth like leakage path (see Figure 3.5-13). To conservatively bound the available leakage area for air exchange via diffusion, the closure plug geometry can be simplified as simply the barrel portion (i.e., flow path over segment A-B, Figure 3.5-13). Per the Table 3.5-4, the maximum diffusion area represented by this flow path is 1.71 in2 (11 cm2). Based on the derived diffusion coefficient, an air density of 0.000325 g/cm3at 1475°F (1075K), and a total diffusion path length of 2.5 in (6.4 cm, i.e., the total length of the closure plug), the maximum diffusion rate during the 30-minute fire event is calculated as:

2 2

3 cm 11 4.6 0

21

.0

/sec cm 815

.1 g/cm

.000325 0

Area J

J x Area = 0.00021 g/sec = 0.0004 g-mole/min Following the fire event, the ambient temperature will drop to 100°F and the ambient density will rise to 0.001128 g/cm3. The diffusion coefficient for oxygen in air at 1 atm and 25°C is 0.206 cm2/sec35, or approximately 11% of the diffusion coefficient determined for the fire conditions. The net effect of the higher density and lower diffusion coefficient is a diffusion rate of 0.00008 g/sec, or 38% of the rate determined at fire conditions.

Based on the 0.22 Btu/min temperature rise determined in the previous section for the 0.0026 g-mole/min oxygen flow associated with the pressure differential, the 0.0004 g-mole/min oxygen diffusion rate would generate a maximum 0.03 Btu/min temperature rise, dropping to less than 0.013 Btu/min following the end of the fire event. Since accounting for the diffusion resistance within the payload cavity will reduce the potential heat generation rate even more, a reasonable conclusion is that the contribution to package heatup from oxygen diffusion can be ignored.

3.5.3.3 Pressure Loss Across Closure Leakage Path The ATR FFSC package is not sealed, but uses a bayonet type closure plug that results in a labyrinth like leakage path, see Figure 3.5-13. The size of the various pathways illustrated in the figure are listed in Table 3.5-4. The maximum pressure rise within the package is associated with the minimum flow area and the maximum gas generation and thermal expansion, with the total pressure loss estimated from a summation of the individual pressure losses associated with each portion of the flow path. Normalizing the individual pressure losses to the flow velocity in the A-B channel allows direct addition of the individual loss coefficients and eases the calculation of the pressure loss based on a single flow velocity. The normalizing to flow velocity involves multiplying the calculated loss coefficient by the square of the area ratio.

The entrance loss at the beveled portion of the closure plug can be estimated using a conical inlet with adjoining wall (i.e, Diagram 3-737). Based on a 15° bevel angle on closure plug and L/Dh >

0.6, the total loss coefficient at the entrance is:

13

.0 5.0 2

1

v P

37 Handbook of Hydraulic Resistance, 3rd Ed., Idelchik, I.E., Begell House Publishers, 1996.

Docket No. 71-9330 ATR FFSC Safety Analysis Report Rev. 17, May 2022 3-53 where v is the flow velocity upstream of the inlet area. This value is conservatively increased to 50

.0 1

for a blunt, flush inlet (Diagram 3-137). Since the loss coefficient is based on the flow velocity after entering the gap, no adjustment for flow area in A-B is necessary:

5.0 1

1

K The pressure loss associated with flow in the A-B channel is a function of wall friction losses.

Given the short path length and smooth wall surfaces, the associated pressure loss will be insignificant and can be ignored.

Flow between B-B' can be approximated as a 90-degree turn with sharp corners (Diagram 6-637). Here the rectangular side length ratio (ao/bo) is equal to (5.64 x pi)/((5.76-5.64)/2) = 295.3 and the ratio of cross section areas (b1/b0) is equal to 0.006/0.06 = 0.1 (based on the minimum gap width after the turn). With these values, the loss factor extrapolated from Diagram 6-6 is 1.3 2

. Given uncertainties in the extrapolation, the computed value is doubled to 6.2 for conservatism38. Since the loss coefficient is based on the flow velocity in the gap approaching the turn, no adjustment for flow area in A-B is necessary:

2.6 2

2

K Flow between B'-E can also be approximated as a 90-degree sharp corner turn (Diagram 6-637).

Again, the rectangular side length ratio (ao/bo) is equal to (5.967 x pi)/(0.006 min gap) = 3124 and the ratio of cross section areas (b1/b0) is equal to 0.235/0.006 = 39. With these values, the loss factor can be conservatively estimated from Diagram 6-6 as 55

.0 3

. Converting to the loss coefficient based on the gap area for flow path A-B yields:

20

.9 11

.0 45

.0 2

3 3

K Flow between E-F can also be approximated as a sudden expansion with a discharge to ambient.

A loss factor of 1 is used to account for these losses. Converting to the loss coefficient based on the gap area for flow path A-B yields:

06

.0 77

.1 45

.0 2

4 4

K The parallel flow path to B'-E consisting of B'-C, C-D, and D-E can be conservatively ignored as its inclusion will serve to lower the estimated total pressure loss. Therefore, a bounding estimate of the total loss factor associated with the minimum expected flow path areas is calculated as 16 06

.0 2.9 2.6 5.0 4

3 2

1

K K

K K

The pressure loss for flow through the closure plug leakage path can be computed as a function of velocity and density via cg v

P 2

5.0 16

. Since mass flow is also a function of velocity and 38 This flow loss is a reasonable upper bound given a worst case assumption that the flow comes to a complete stop before the turn and then needs to re-accelerate into the smaller gap. When adjusted for velocity differences, the flow loss under this worst case scenario would be approximately (0.45 in2/0.11 in2)2 x 0.5 = 8.4, where 0.5 is the loss factor associated with a blunt inlet fitting.

Docket No. 71-9330 ATR FFSC Safety Analysis Report Rev. 17, May 2022 3-54

density, Area

v m

, the pressure loss relationship can be re-formulated as a function of mass flow via:

cg m

P 2

Area

/

5.0 16

where Area is the flow area in the path A-B (0.45 in2 minimum) and the density is for the bulk gas temperature. From the data used to develop Figure 3.5-12, the maximum gas flow required to maintain atmospheric pressure within the ATR FFSC cavity due to only ideal gas expansion occurs during package heatup. The peak flowrate of 0.035 g-mole/min occurs approximately 8 minutes after the start of the 30-minute HAC fire and when the bulk gas temperature within the payload cavity has reached 230°F (110°C). Based on a molecular weight of 28.96 g/g-mole for air, the associated mass flow and density are 1.01 g/min (0.00004 lbm/sec) and the gas density is 0.00091 g/cm3 (0.057 lbm/ft3). Substituting these values into the above equation yields:

P = 5(10-6) psi for the conservative assumption of minimum flow areas within all vent gaps. The pressure loss at nominal gap dimensions will be even lower.

This maximum pressure rise due to thermal expansion of the cavity gas is too low to create an issue. Thermal decomposition of polyethylene and neoprene will generate additional gases that would need to be vented. While only a small fraction of the material is expected to be thermally decomposed due to a combination of the temperature levels achieved and the time above the thermal decomposition temperature level, a bounding maximum pressure rise can be estimated assuming the entire inventory of both polyethylene and neoprene decomposes over a 60 minute period. The potential gas quantity associated with the total decomposition of the 200 g of polyethylene is 200 g/(28 g/g-mole) x 2 g-moles H2 per g-mole polyethylene = 14.3 g-moles H2.

Similarly, the 1,926 g of neoprene associated with the SAR 60501-70 FHE assembly will generate 1926 g/(88.5 g/g-mole) x (2 g-moles H2 + 1 g-moles HCl) per g-mole neoprene = 65.3 g-moles H2 and HCl. The combined gas generation rate is therefore (14.3 + 65.3 g-mole)/60 minutes, or 1.33 g-moles/minute. Alternately, considering the entire mass of neoprene and polyethylene to decompose in 60 minutes gives a mass flow rate of (1,926 + 200)/60 = 35.4 g/min, or 0.0013 lbm/s. Conservatively assuming the gas is H2 at a temperature of 1,017 °F from Table 3.1-1, the gas density is 1.1(10-6) lbm/in3. Substituting into the equation above yields:

P = 0.16 psi Four kg of cellulosic material (8.8 lb) would have a mass flow rate of 8.8/602 = 0.0024 lbm/s.

Factoring the neoprene result by the square of the mass flow rate yields:

P = 0.16 x (0.0024/0.0013)2 = 0.55 psi These results are based on three significant conservatisms:

Minimum flow area in all gaps 100% of the mass decomposes to gas (no residue)

Density of gas minimized (high-temperature hydrogen)

Docket No. 71-9330 ATR FFSC Safety Analysis Report Rev. 17, May 2022 3-55 As such, the assumption of a 0 psig pressure throughout the HAC event is valid for the purposes of determining the safety basis of the design.

Based on the level of and type of damage noted in Appendix 2.12.1, Certification Tests on CTU-1 and Appendix 2.12.2, Certification Tests on CTU-2, no change to the net vent areas based on the assumed minimum gaps is expected. Thus the above conclusions remain valid for the damaged package configuration as well.

Docket No. 71-9330 ATR FFSC Safety Analysis Report Rev. 17, May 2022 3-56 Table 3.5 Neoprene Quantity Per Assembly SAR Drawing Neoprene Surface Area, in2 Neoprene Volume, in3 Neoprene Adhesive Volume, in3 Neoprene Quantity, g 60501-10 N/A N/A N/A N/A 60501-20 N/A N/A N/A N/A 60501-30 475 59 1.0 1210 g 60501-40 162 20 0.3 409 g 60501-50 266 33 0.5 676 g 60501-60 547 68 1.1 1393 g 60501-70 748 94 1.5 1926 g Notes: Based on density of 1.23 g/cm3 (76.8 lb/ft3 per Table 3.2-3)

Table 3.5 Net Cavity Volume vs. Payload Assembly SAR Drawing Gross Cavity Volume, in3 FHE Volume, in3 Payload Volume, in3 Net Cavity Volume, in3 Comments 60501-20 1768.8 307.4 168.1 1293.3 ATR Loose Plate FHE 60501-30 154.6 223.2 1390.9 ATR FHE - Design basis selection due to combination of net cavity size and peak HAC temperature for FHE 60501-40 256.1 88.5 1424.1 MIT FHE 60501-50 307.4 126.1 1335.4 MURR FHE 60501-60 286.9 142.9 1339.0 RINSC FHE 60501-70 307.4 168.1 1293.3 Small Quantity FHE Notes: Based on 30 lb weight and density of 0.0976 in3 per Tables 2.1-1 and 3.2-2 Based on 20 lb weight and density of 0.112 in3 per Tables 2.1-1 and 3.2-2 Based on 15 lb weight and density of 0.097 in3 per Tables 2.1-1 and 3.2-1 Based on 25 lb weight and density of 0.112 in3 per Tables 2.1-1 and 3.2-2 Based on 25 lb weight and density of 0.0976 in3 per Tables 2.1-1 and 3.2-2 Based on 10 lb weight and density of 0.113 in3 per Tables 2.1-1 and 3.6-4 Based on 30 lb weight and density of 0.0976 in3 per Tables 2.1-1 and 3.2-2 Based on 15 lb weight and density of 0.119 in3 per Tables 2.1-1 and 3.6-4 Based on 28 lb weight and density of 0.0976 in3 per Tables 2.1-1 and 3.2-2 Based on 17 lb weight and density of 0.119 in3 per Tables 2.1-1 and 3.6-4

Docket No. 71-9330 ATR FFSC Safety Analysis Report Rev. 17, May 2022 3-57 Table 3.5 Closure Leakage Path Areas Flow Path Inner/Outer Diameter, in Gap Width/Length, in Flow Path Area, in2 A to B 5.64 + 0.01 5.76 + 0.06 1.69 Max: 1.71 Min: 0.45 B to B' 5.70 (mean) 0.006 to 0.03 Max: 0.54 Min: 0.11 B' to C 5.967 (min) 0.006 to 0.03 Max: 0.56 Min: 0.11 C to D 6.38 + 0.02 6.44 + 0.01 0.281 Max: 0.91 Min: 0.30 D to E 6.21 (mean) 0.006 to 0.03 Max: 0.59 Min: 0.12 B' to E 5.967 + 0.01 6.44 + 0.01 0.281 Max: 1.92 Min: 1.77 E to F 5.967 + 0.01 6.44 + 0.01 0.56 Max: 1.92 Min: 1.77 Notes: Tolerance from ASTM A269 Based on bayonet tab of width of 0.25 in. centered in slot width of 0.281 in., and tolerances of +0.01 on both parts.

Based on 40% of gross area accounting for area of bayonet tabs and ignoring additional smaller gaps 0.0 0.2 0.4 0.6 0.8 1.0 1.2 1.4 50 150 250 350 450 550 650 750 850 950 1050 1150 1250 1350 1450

-30 30 90 150 210 270 330 390 450 510 570 630 690 Gas Quntity - g/moles Temperature - F Time - Minutes Inner Shell - Max.

Fuel Handling Enclosure (FHE) - Max.

Bulk Gas Temperature Cavity Gas Quantity (Free Vent) - g/mole Gas Flow Into Package Lower Autoignition Temperature for Polyethylene Figure 3.5 Free Vent Gas Flow During HAC Transient

Docket No. 71-9330 ATR FFSC Safety Analysis Report Rev. 17, May 2022 3-58 a) Package Closure b) Enlarged Flow Paths at Package Closure Figure 3.5 Free Vent Gas Flow Path

Docket No. 71-9330 ATR FFSC Safety Analysis Report Rev. 17, May 2022 3-59 3.6 Thermal Evaluation for MIT, MURR, Cobra, Small Quantity, and LEU Payloads This section identifies and describes the principal thermal design aspects of the ATR FFSC for the transport of one assembled fuel element of the type: MIT HEU, MURR HEU, Cobra, ATR LEU, MIT LEU or DDE, MURR LEU or DDE, NBSR DDE, or small quantity payloads as described in Section 1.2.2.4, Small Quantity Payload. The evaluation presented herein demonstrates that the thermal performance of the ATR FFSC when transporting these payloads is bounded by the temperatures reported for the transport of the ATR fuel element payload.

Specifically, the evaluations presented herein demonstrate the thermal safety of the ATR FFSC package1 complies with the thermal requirements of 10 CFR 712 when transporting a payload consisting of either an assembled, unirradiated fuel element, or loose, unirradiated fuel plates, or other small quantity payloads as described in Section 1.2.2, Contents.

All package components are shown to remain within their respective temperature limits under the normal conditions of transport (NCT). Further, per 10 CFR §71.43(g), the maximum temperature of the accessible package surfaces is demonstrated to be less than 122 °F for the maximum decay heat loading, an ambient temperature of 100 F, and no insolation. Finally, the ATR FFSC package is shown to retain sufficient thermal protection following the HAC free and puncture drop scenarios to maintain all package component temperatures within their respective short term limits during the regulatory fire event and subsequent package cool-down.

3.6.1 Description of Thermal Design The ATR FFSC package, as described and illustrated in Chapter 1.0, General Information, consists of three basic components: 1) a Body assembly, 2) a Closure assembly, and 3) either a Fuel Handling Enclosure (FHE) or a Loose Fuel Plate Basket (LFPB). The FHE is configured to house one assembled fuel element or DDE, while the LFPB is configured to house loose fuel element plates. The maximum gross weight of the fully loaded package is approximately 299 lbs.

The ATR FFSC is designed as a Type AF packaging. The packaging is rectangular in shape and is intended to be transported in racks of multiple packages by highway truck. Since the payload generates essentially no decay heat, the worst case thermal conditions will occur with an individual package fully exposed to ambient conditions. The package performance when configured in a rack of multiple packages will be bounded by that seen for an individual package.

The thermal design aspects of the principal components of the packaging are described in more detail in Section 3.1, Description of Thermal Design. The paragraphs below present the thermal design features of the various payloads and their associated FHEs.

3.6.2 Fuel Handling Enclosures Fuel handling enclosures are used with all payloads (except the NBSR DDE), and associated loose fuel element plates. Except as noted, the FHE are machined, two-piece aluminum

Docket No. 71-9330 ATR FFSC Safety Analysis Report Rev. 17, May 2022 3-60 enclosures used to protect the fuel element from damage during loading and unloading operations. The FHE consist of two identical segments machined from 6061 aluminum plate or bar stock. The FHE features neoprene rub strips to further protect the fuel. The FHE is neither anodized nor coated, but is left as unfinished aluminum. Spacer weldments on either end of the enclosure halves are used to position and support the FHE within the ATR FFSC cavity. The spacers are also fabricated of 6061 aluminum. A polyethylene bag may be used as a protective sleeve over the fuel elements. The following table presents a directory of figure depictions of the FHE and fuel elements, and design weights of the FHE. Note, the MIT, MURR, and Cobra loose fuel element plates are shipped in the Small Quantity FHE (SQFHE). Loose plates may be shipped with kraft paper and adhesive tape for property protection, and aluminum or cellulosic dunnage, as described in Section 1.2.1.1.8, Small Quantity Payload FHE. The NBSR DDE does not utilize a FHE.

Fuel Exploded View of FHE Fuel Element Figure FHE Design Weight, lb ATR LEU Figure 1.2-1 Figure 1.2-12 15 MIT HEU, LEU, DDE Figure 1.2-6 Figure 1.2-13 25 MURR HEU, LEU, DDE Figure 1.2-7 Figure 1.2-14 30 RINSC Figure 1.2-8 Figure 1.2-15 28 Small Quantity Figure 1.2-9 Not shown in figures 30 Cobra Figure 1.2-10 Figure 1.2-17 28 3.6.3 Contents Decay Heat The ATR FFSC is designed as a Type AF packaging for transportation of an unirradiated fuel elements or a bundle of loose, unirradiated fuel plates. The decay heat associated with un-irradiated fuel is negligible. Therefore, no special devices or features are needed or utilized in the ATR FFSC packaging to dissipate the decay heat. Section 1.2.2, Contents, provides additional details regarding the potential contents of the ATR FFSC.

3.6.4 Summary Tables of Temperatures Table 3.6-1 provides a summary of the maximum package component temperatures achieved under NCT and HAC conditions for either the MIT HEU or MURR HEU fuel element payloads.

These temperatures are either bounded by or similar to those reported in Table 3.1-1 for the transport of the ATR HEU fuel element payload. Those values unbounded by the values found in Table 3.6-1 remain well below the maximum allowable temperatures. Based on the results for the MURR fuel element, the maximum temperatures achieved under NCT and HAC conditions for the Cobra fuel element, small quantity payloads (including the RINSC fuel element), and LEU fuel elements and DDEs are shown by qualitative analysis below to also be bounded by the results presented in Table 3.1-1.

Docket No. 71-9330 ATR FFSC Safety Analysis Report Rev. 17, May 2022 3-61 The MIT HEU and MURR HEU payload temperatures for NCT are based on an analytical model of the ATR FFSC package under extended operation with an ambient temperature of 100°F and a diurnal cycle for the insolation loading. The temperatures for HAC are based on an analytical model of the ATR FFSC package with the worst-case, hypothetical pre-fire damage as predicted based on drop tests using full-scale certification test units (CTUs). The ATR FFSC with the Cobra fuel element, small quantity payloads, or LEU fuel elements and DDEs was not specifically modeled as part of this evaluation. Instead, their thermal performance is estimated using a qualitative approach based on the thermal characteristics of the other payloads and their associated thermal performance.

The MIT HEU and MURR HEU payload results for NCT demonstrate that significant thermal margin exists for all package components. This is expected since the only significant thermal loads on the package arise from insolation and ambient temperature changes. The payload dissipates essentially zero decay heat. Further, the evaluations for NCT demonstrate that the package skin temperature will be below the maximum temperature of 122°F permitted by 10 CFR

§71.43(g) for accessible surface temperature in an nonexclusive use shipment when transported in a 100°F environment with no insolation. Given the significant thermal margin existing for the other payloads and the similar materials of fabrication, the Cobra fuel element, small quantity payloads, and LEU fuel elements and DDEs are also predicted to exhibit large thermal margins.

The MIT HEU and MURR HEU payload results for HAC conditions demonstrate that the design of the ATR FFSC package provides sufficient thermal protection to yield component temperatures that are significantly below the acceptable limits defined for each component.

While the neoprene rubber and polyethylene plastic material used to protect the fuel element from damage are expected to reach a sufficient temperature level during the HAC fire event to induce thermal decomposition, the loss of these components is not critical to the safety of the package. As demonstrated in Section 3.5.3, Thermal Decomposition/Combustion of Package Organics, the available oxygen in the package, plus that which may enter the package under pressure differential and gas diffusion forces, is insufficient to result in any significant heat generation due to combustion. Given the similar materials of fabrication and equivalent thermal mass as the MURR payload, the Cobra fuel element, small quantity payloads, and LEU fuel elements and DDEs are also predicted to exhibit large thermal margins under HAC conditions.

3.6.5 Summary Tables of Maximum Pressures Table 3.6-2 presents a summary of the maximum pressures achieved under NCT and HAC conditions. Since the ATR FFSC package is a vented package, both the maximum normal operating pressure (MNOP) and the maximum pressure developed within the payload compartment under the HAC condition are 0 psig. Section 3.5.3, Thermal Decomposition/Combustion of Package Organics, provides the justification for assuming a 0 psig package pressure for the HAC event.

Although the volume between the outer and inner shells is sealed, it does not contain organic or other materials that may outgas or thermally decompose. Therefore, the maximum pressure that may develop within the space will be limited to that achieved due to ideal gas expansion. The maximum pressure rise under NCT will be less than 4 psig, while the pressure rise under HAC conditions will be 39 psig.

Docket No. 71-9330 ATR FFSC Safety Analysis Report Rev. 17, May 2022 3-62 In the remainder of this appendix, the use of the terms MIT and MURR, unless distinguished by the term LEU, shall be taken to mean MIT HEU and MURR HEU, respectively.

Table 3.6 Maximum Temperatures for NCT and HAC Conditions Location / Component NCT Hot Conditions Accident Conditions Maximum Allowable Normal Accident Fuel Element Fuel Plate 143°F 640°F 400°F 1,100°F Fuel Element Side Plate 143°F 644°F 400°F 1,100°F Neoprene Rub Strips/Polyethylene Bag 143°F 710°F 225°F N/A Fuel Handling Enclosure (FHE) 143°F 710°F 400°F 1,100°F Inner Shell 157°F 1,417°F 800°F 2,700°F Ceramic Fiber Insulation, Body

- Maximum

- Average 184°F 149°F 1,462°F 1,253°F 2,300°F 2,300°F 2,300°F 2,300°F Ceramic Fiber Insulation, Closure

- Maximum

- Average 145°F 143°F 1,402°F 1,236°F 2,300°F 2,300°F 2,300°F 2,300°F Closure 145°F 1,439°F 800°F 2,700°F Outer Shell 184°F 1,475°F 800°F 2,700°F Table Notes:

Maximum allowable temperatures are defined in Section 3.2.2, Technical Specifications of Components.

Component temperature assumed to be equal to that of the FHE.

Table 3.6 Summary of Maximum Pressures Condition Fuel Cavity Pressure Outer/Inner Shell Cavity Pressure NCT Hot 0 psi gauge 4 psi gauge HAC Hot 0 psi gauge 39 psi gauge

Docket No. 71-9330 ATR FFSC Safety Analysis Report Rev. 17, May 2022 3-63 3.6.6 Material Properties and Component Specifications The ATR FFSC is fabricated primarily of Type 304 stainless steel, 5052-H32 and 6061-T651 aluminum, ceramic fiber insulation, and neoprene rubber. The payload materials include 6061-T6 and/or 6061-0 aluminum, uranium aluminide (UAlx), uranium silicide (U3Si2), and uranium molybdenum (U-10Mo in a foil coated with thin zirconium interlayers). A polyethylene plastic bag is used as a protective sleeve over the fuel element.

3.6.6.1 Material Properties The material specifications for the ATR FFSC package are defined in Section 3.2.1, Material Properties. Table 3.6-3 presents the thermal properties for 6061 aluminum used for the MIT and MURR FHEs, as taken from Table TCD of the ASME Boiler and Pressure Vessel Code3.

Although the design permits a variety of aluminum tempers to be used, a single data set is provided since the material temper has little to no effect on its thermal properties. Further, because the HAC analysis requires thermal properties in excess of the maximum temperature point of 400F provided in Table TCD, the property values at 1100F (i.e., the approximate melting point for aluminum) are assumed to be the same as those at 400F. This approach is appropriate for estimating the temperature rise within the fuel basket during the HAC event since the thermal conductivity of aluminum alloys tends to decrease with temperature while the specific heat tends to increase. The density values listed in the table are taken from an on-line database4. Properties between the tabulated values are calculated via linear interpolation within the heat transfer code.

Table 3.6-4 presents the thermal properties for the MIT and MURR fuel elements. For analysis purposes, the material used for the side plates and end fittings are assumed to be 6061-0 aluminum.

The thermal properties for the fuel plates are determined as a composite of the cladding and the fuel core materials based on the geometry data for the MIT and MURR fuel element39,40 and the thermal properties for the ATR fuel element materials6. This approach is the same as used for the ATR fuel element. The details of the computed values are presented in Appendix 3.6.9.2.3, Determination of Composite Thermal Properties for MIT and MURR Fuel Plates. For simplicity, the thermal properties are assumed to be constant with temperature based on the use of conservatively high thermal conductivity and conservatively low specific heat values. This approach maximizes the heat transfer into the fuel components during the HAC event, while under-estimating the ability of the components to store the heat.

The RINSC fuel elements are fabricated with a nominally 0.020-in thick mixture of uranium silicide (U3Si2) and aluminum powder as the fuel meat and a nominally 0.015-in thick aluminum alloy cladding. The twenty-two (22) flat fuel plates have a 2.8-in width, an overall length of 25-in, and an active fuel region of 22.5 to 24.0-in. These fuel plate meat and cladding thicknesses match those of the interior plates for the ATR fuel element and are similar to those for the MURR fuel plates. The side plates are fabricated of ASTM B 209, aluminum alloy 6061-39 Massachusetts Institute of Technology, Test Research Training Reactor 3 Fuel Plate, EG&G, Idaho, Inc.,

Drawing No. 410368, Rev. A.

40 University of Missouri at Columbia, Test Research Training Reactor 4 MURR Fuel Plate, EG&G, Idaho, Inc.,

Drawing No. 409406, Rev. E.

Docket No. 71-9330 ATR FFSC Safety Analysis Report Rev. 17, May 2022 3-64 T6 and 6061-T651 and are approximately 0.188-in thick. This is similar to the side plate thicknesses of the ATR, MIT, and MURR fuel elements.

The thermal conductivity of the RINSC fuel plates are similar to data obtained in the measurements of the thermal conductivities for the uranium aluminide (UAlx) based fuels41.

Similarly, the thermal mass of the fuel plates are comparable despite the higher density of uranium silicide versus uranium aluminide since the ratio of the specific heats of the two materials is nearly the inverse of the density ratio.

The Cobra fuel elements are fabricated with a nominally 0.025-in thick mixture of Uranium and either aluminum as UAlx (HEU) or with silicon as U3Si2 (LEU) as the fuel meat and a nominally 0.014-in thick aluminum alloy cladding. The fuel is constructed using six concentric circular layers of fuel plates, divided into three equal segments by radial, aluminum alloy separator plates. The fuel plates are approximately 38 inches long. The entire fuel element, including aluminum alloy end fittings, is approximately 61 inches long. The diameter of the element (outside edge of the separator plates) is approximately 3.25 inches. The remarks above concerning thermal conductivity and thermal mass for the RINSC fuel elements apply to the Cobra fuel elements as well.

The LEU fuel elements and DDEs are described in Section 3.6.9.2.6, Determination of Thermal Properties for LEU Fuel Elements.

The additional small quantity payloads are fabricated as described in Section 1.2.2.4, Small Quantity Payload. Small quantity payloads may be shipped with aluminum or cellulosic dunnage.

The thermal properties for air and for the non-metallic materials used in the ATR FFSC are presented in Section 3.2.1, Material Properties, as is the assumed emissivity () for each radiating surface and the solar absorptivity () value for the exterior surface. The 6061-0 aluminum used for the MIT and MURR fuel components are assumed to have a surface coating of boehmite (Al2O3H2O). A 25 m boehmite film will exhibit a surface emissivity of approximately 0.9213.

While a fresh fuel element may have a lower surface emissivity, the use of the higher value will provide a conservative estimate of the temperatures achieved during the HAC event.

3.6.6.2 Technical Specifications of Components The materials used in the ATR FFSC that are considered temperature sensitive include the aluminum used for the FHEs, the LFPB, and the fuel elements, the neoprene rubber, and the polyethylene wrap used as a protective sleeve around the fuel elements. Of these materials, only the aluminum used for the fuel elements is considered critical to the safety of the package. The other materials either have temperature limits above the maximum expected temperatures or are not considered essential to the function of the package.

Section 3.2.2, Technical Specifications of Components, presents the basis for the temperature limits of the various components. These temperature limits are applicable to this safety evaluation as well.

41 IAEA-TECDOC-643, Research Reactor Core Conversion Guidebook, Volume 4: Fuels (Appendices I-K),

International Atomic Energy Agency, Vienna, Austria.

Docket No. 71-9330 ATR FFSC Safety Analysis Report Rev. 17, May 2022 3-65 Table 3.6 Thermal Properties of Package Metallic Materials Material Temperature

(ºF)

Thermal Conductivity (Btu/hr-ft-ºF)

Specific Heat (Btu/lbm-ºF)

Density (lbm/in3)

Aluminum Type 6061-T651 /

T6511 70 96.1 0.214 0.098 100 96.9 0.216 150 98.0 0.220 200 99.0 0.222 250 99.8 0.224 300 100.6 0.227 350 101.3 0.230 400 101.9 0.231 1100 101.9 0.231 Notes:

Values for 1100°F are assumed equal to values at 400°F.

Docket No. 71-9330 ATR FFSC Safety Analysis Report Rev. 17, May 2022 3-66 Table 3.6 Thermal Properties of MIT and MURR Fuel Materials Material Temperature

(ºF)

Thermal Conductivity (Btu/hr-ft-ºF)

Specific Heat (Btu/lbm-ºF)

Density (lbm/in3)

Aluminum Type 6061-0 32 102.3 0.0976 62 0.214 80 104.0 170 107.5 260 109.2 0.225 350 109.8 440 110.4 0.236 530 110.4 620 109.8 0.247 710 108.6 800 106.9 0.258 890 105.2 980 103.4 0.269 1080 101.1 0.275 MURR Fuel Plate 80 57.9 0.165 0.121 800 0.200 MIT Fuel Plate 80 72.6 0.176 0.115 800 0.212 Notes:

Values determined based on composite value of aluminum cladding and fuel core material (see Appendix 3.5.2.4). Thermal conductivity value is valid for axial and circumferential heat transfer within fuel plate.

Docket No. 71-9330 ATR FFSC Safety Analysis Report Rev. 17, May 2022 3-67 3.6.7 Thermal Evaluation for Normal Conditions of Transport The ATR FFSC with the MIT or MURR fuel element payloads is transported horizontally under normal conditions of transport (NCT). This establishes the orientation of the exterior surfaces of the package for determining the free convection heat transfer coefficients and insolation loading.

While the package would normally be transported in tiered stacks of multiple packages, the evaluation for NCT is conservatively based on a single, isolated package since this approach will yield the bounding maximum and minimum temperatures achieved by any of the packages.

Further, since the surface of the transport trailer is conservatively assumed to prevent heat exchange between the package and the ambient, the bottom of the ATR FFSC is treated as an adiabatic surface.

The details of the thermal modeling used to simulate the ATR FFSC package under NCT conditions are provided in Appendix 3.5.2, Analytical Thermal Model, while details of the thermal modeling of the MIT and MURR FHEs and fuel elements are provided in Appendix 3.6.9.2.1, Description of MIT and MURR Payload Thermal Models for NCT Conditions. The ATR FFSC with Cobra fuel elements or small quantity payloads was not specifically modeled as part of this evaluation. Instead, their thermal performance is estimated using a qualitative approach based on the thermal characteristics of the other payloads and their associated thermal performance. See below for the details of this qualitative basis.

3.6.7.1 Heat and Cold 3.6.7.1.1 Maximum Temperatures The maximum temperature distribution for the ATR FFSC occurs with a diurnal cycle for insolation loading and an ambient air temperature of 100°F, per 10 CFR §71.71(c)(1). The evaluation of this condition is conducted as a transient using the thermal model of an undamaged ATR FFSC described in Appendix 3.6.9.2.1, Description of MIT and MURR Payload Thermal Models for NCT Conditions. Figure 3.6-1 illustrates the expected heat-up transient for an ATR FFSC loaded with a MIT fuel element. The transient analysis assumes a uniform temperature condition of 70ºF for all components prior to loading and exposure to the specified NCT condition at time = 0.

The figures demonstrate that the ATR FFSC package will respond rapidly to changes in the level of insolation and will reach it peak temperatures within the first day or two after loading. The higher thermal mass of the MIT FHE on a unit length basis versus that of the ATR FHE is reflected in the delayed response of the MIT FHE to changes in the inner shell temperature, whereas the ATR FHE was seen in Figure 3.3-1 to respond more rapidly. A similar temperature response curve is seen for the MURR FHE.

Table 3.6-5 presents the maximum temperatures reached for various components of the package.

As seen from the table, all components are within in their respective temperature limits. Figure 3.6-2 illustrates the predicted temperature distribution within the ATR FFSC package with a MIT fuel element payload at the end of the evaluated transient heat up period and near the time of peak temperature. Figure 3.6-3 presents the temperature distribution within the ATR FFSC package with a MURR fuel element payload.

Docket No. 71-9330 ATR FFSC Safety Analysis Report Rev. 17, May 2022 3-68 The maximum temperature distribution for the ATR FFSC without insolation loads occurs with an ambient air temperature of 100°F. Since the package payload dissipates essentially zero watts of decay heat, the thermal analysis of this condition represents a trivial case and no thermal calculations are performed. Instead, it is assumed that all package components achieve the 100°F temperature under steady-state conditions. The resulting 100°F package skin temperature is below the maximum temperature of 122°F permitted by 10 CFR §71.43(g) for accessible surface temperature in a nonexclusive use shipment.

The ATR FFSC with the small quantity payload was not specifically modeled as part of this evaluation. Instead, its thermal performance is estimated using a qualitative approach based on the thermal characteristics of the other payloads and their associated thermal performance. Using this approach, it is estimated that the maximum temperatures attained for the transportation of the small quantity payload within the ATR FFSC will be bounded by that presented for the MURR payload. This conclusion is based on the facts that the combined weight of the small quantity payload and MURR FHE's with their enclosed fuel elements, plates, or foils are similar (see Section 1.2.2.3, MURR HEU Fuel Element, and Section 1.2.2.4, Small Quantity Payload), the FHE's are both fabricated of 6061 aluminum, and the fuel elements have similar thermal properties (see Section 3.6.6.1). This conclusion is further supported by the fact that Table 3.6-5 demonstrates that the MIT and MURR fuel elements produce essentially the same peak NCT temperatures despite their design differences. As such, the limited design differences between the MURR and small quantity payloads will not yield a significant difference in their NCT thermal response.

The ATR FFSC with the RINSC fuel element payload and the Cobra fuel element payload are not specifically modeled as part of this evaluation. Instead, their thermal performance is estimated using a qualitive approach based on the thermal characteristics of the other payloads and their associated thermal performance. (See Section 3.6.9.2.4, Determination of Thermal Properties for RINSC Element, Section 3.6.9.2.5, Determination of Thermal Properties for Cobra Element, and Section 3.6.9.2.6, Determination of Thermal Properties for LEU Fuel Elements for details). Using this approach, it is estimated that the maximum temperatures attained for the transportation of the RINSC, Cobra, and LEU fuel elements and DDEs are considered bounded by the analysis of the MURR payload and no additional analysis is required.

3.6.7.1.2 Minimum Temperatures The minimum temperature distribution for the ATR FFSC occurs with a zero decay heat load and an ambient air temperature of -40°F per 10 CFR §71.71(c)(2). The thermal analysis of this condition also represents a trivial case and no thermal calculations are performed. Instead, it is assumed that all package components achieve the -40°F temperature under steady-state conditions.

As discussed in Section 3.2.2, Technical Specifications of Components, the -40°F temperature is within the allowable operating temperature range for all ATR FFSC package components.

3.6.7.2 Maximum Normal Operating Pressure The payload cavity of the ATR FFSC is vented to the atmosphere. As such, the maximum normal operating pressure (MNOP) for the package is 0 psig.

Docket No. 71-9330 ATR FFSC Safety Analysis Report Rev. 17, May 2022 3-69 While the volume between the outer and inner shells is sealed, it does not contain organic or other materials that may outgas or thermally decompose. Therefore, the maximum pressure that may develop within the space will be limited to that achieved due to ideal gas expansion.

Assuming a temperature of 70°F at the time of assembly and a maximum operating temperature of 190°F (based on the outer shell temperature, see Table 3.6-5, conservatively rounded up), the maximum pressure rise within the sealed volume will be less than 4 psi.

Table 3.6 Maximum Package NCT Temperatures Location / Component MIT Fuel Payload MURR Fuel Payload Maximum Allowable Fuel Element Fuel Plate 143°F 142°F 400°F Fuel Element Side Plate 143°F 142°F 400°F Neoprene Rub Strips/Polyethylene Bag 143°F 142°F 225°F Fuel Handling Enclosure (FHE) 143°F 142°F 400°F Inner Shell 157°F 157°F 800°F Ceramic Fiber Insulation, Body

- Maximum

- Average 184°F 149°F 184°F 148°F 2,300°F 2,300°F Ceramic Fiber Insulation, Closure

- Maximum

- Average 145°F 143°F 145°F 143°F 2,300°F 2,300°F Closure 145°F 145°F 800°F Outer Shell 184°F 184°F 800°F Table Notes:

The maximum allowable temperatures under NCT conditions are provided in Section 3.2.2, Technical Specifications of Components.

Component temperature assumed to be equal to that of the FHE.

Docket No. 71-9330 ATR FFSC Safety Analysis Report Rev. 17, May 2022 3-70 60 80 100 120 140 160 180 200 0

12 24 36 48 60 72 84 96 Time - Hours Temperature - F Outer Shell - Max.

Outer Shell - Avg.

Insulation Jacket - Max.

Insulation - Avg.

Inner Shell - Max.

Fuel Handling Enclosure (FHE) - Max.

Figure 3.6 ATR FFSC Package Heat-up with MIT Payload, NCT Hot Conditions

Docket No. 71-9330 ATR FFSC Safety Analysis Report Rev. 17, May 2022 3-71 Figure 3.6 Package NCT Temperature Distribution for MIT Payload

Docket No. 71-9330 ATR FFSC Safety Analysis Report Rev. 17, May 2022 3-72 Figure 3.6 Package NCT Temperature Distribution for MURR Payload

Docket No. 71-9330 ATR FFSC Safety Analysis Report Rev. 17, May 2022 3-73 3.6.8 Thermal Evaluation for Hypothetical Accident Conditions This section presents the thermal evaluation of the ATR FFSC package under the hypothetical accident condition (HAC) specified in 10 CFR §71.73(c)(4)2 based on an analytical thermal model. The analytical model of the ATR FFSC for HAC is a modified version of the quarter symmetry NCT model described in Appendix 3.5.2.1, Description of Thermal Model for NCT Conditions, with the principal model modifications consisting of simulating the expected package damage resulting from the drop events that are assumed to precede the HAC fire and changing the package surface emissivities to reflect the assumed presence of soot and/or surface oxidization. The analytical model of the MIT and MURR fuel elements are the same as those described in Appendix 3.6.9.2.1, Description of MIT and MURR Payload Thermal Models for NCT Conditions. The evaluations of the ATR FFSC with a small quantity payload and RINSC and Cobra payloads under HAC conditions are accomplished using a qualitative approach in the same manner as accomplished for NCT conditions (see Section 3.6.7.1.1, Maximum Temperatures).

Physical testing using full scale certified test units (CTUs) is used to establish the expected level of damage sustained by the ATR FFSC package from the 10 CFR 71.73 prescribed free and puncture drops that are assumed to precede the HAC fire event. Appendix 3.5.2.2, Description of Thermal Model for HAC Conditions, provides an overview of the test results, the rationale for selecting the worst-case damage scenario, and the details of the thermal modeling used to simulate the package conditions during the HAC fire event.

3.6.8.1 Initial Conditions The initial conditions assumed for the package prior to the HAC event are described below in terms of the modifications made to the NCT thermal model to simulate the assumed package conditions prior to and during the HAC event. These modifications are:

Simulated the worst-case damage arising from the postulated HAC free and puncture drops as described in Appendix 3.5.2.2, Description of Thermal Model for HAC Conditions, Assume an initial, uniform temperature distribution of 100ºF based on a zero decay heat package at steady-state conditions with a 100ºF ambient with no insolation.

This assumption complies with the requirement of 10 CFR §71.73(b)2 and NUREG-160917, Increased the emissivity of the external surfaces from 0.45 to 0.8 to account for possible soot accumulation on the surfaces, per 10 CFR §71.73(c)(4),

Increased the emissivity of the interior surfaces of the outer shell from 0.30 to 0.45 to account for possible oxidization of the surfaces during the HAC event, Following the free and puncture bar drops, the ATR FFSC package is assumed come to rest in a horizontal position prior to the initiation of the fire event. Given that the package geometry is essentially symmetrical about its axial axis, there are no significant thermal differences whether the package is right-side up, up-side down, or on its side. None of the payloads are expected to be re-positioned as a result of the pre-fire drop and puncture bar events based on the limited damage seen

Docket No. 71-9330 ATR FFSC Safety Analysis Report Rev. 17, May 2022 3-74 for the ATR FHE as a result of the drop tests conducted on the ATR FFSC presented in Section 2.12.1, Certification Tests on CTU-1, and given the greater robustness of the payload FHEs.

However, even if the end spacers are conservatively assumed to buckle as a result of the HAC drop event, no significant temperature increase will occur since direct contact between the FHE and the closure plug will be prevented and because the average radial heat transfer through the sides of the package does not change significantly as a function of axial position. Therefore, the peak package temperatures predicted under this evaluation based on no payload re-positioning or reconfiguration are representative of those achieved for any package orientation and/or credible re-positioning of the enclosed payloads.

3.6.8.2 Fire Test Conditions The fire test conditions analyzed to address the 10 CFR §71.73(c) requirements are as follows:

The initial ambient conditions are assumed to be 100ºF ambient with no insolation, At time = 0, a fully engulfing fire environment consisting of a 1,475ºF ambient with an emissivity of 1.0 is used to simulate the hydrocarbon fuel/air fire event.

The assumption of a flame emissivity of 1.0 bounds the minimum average flame emissivity coefficient of 0.9 specified by 10 CFR §71.73(c)(4),

The convection heat transfer coefficients between the package and the ambient during the 30-minute fire event are based on an average gas velocity18 of 10 m/sec. Following the 30-minute fire event the convection coefficients are based on still air, The ambient condition of 100ºF with insolation is assumed following the 30-minute fire event. Since a diurnal cycle is used for insolation, the evaluation assumes that the 30-minute fire begins at noon so as to maximize the insolation heating during the post-fire cool down period. A solar absorptivity of 0.9 is assumed for the exterior surfaces to account for potential soot accumulation on the package surfaces.

The transient analysis is continued for 11.5 hours5.787037e-5 days <br />0.00139 hours <br />8.267196e-6 weeks <br />1.9025e-6 months <br /> after the end of the 30-minute fire to ensure that the peak package temperatures are captured.

3.6.8.3 Maximum Temperatures and Pressure 3.6.8.3.1 Maximum HAC Temperatures The thermal performance of the ATR FFSC package loaded with a MIT fuel element payload is summarized in Table 3.6-6, while Table 3.6-7 presents a summation of the results with a MURR fuel element payload. With the exception of the neoprene rub strips and the polyethylene bag used as a protective sleeve around the fuel elements, all other components of the package are seen to remain well below their allowable short term temperature limits. As with the ATR payload, the thermal decomposition of the neoprene strips and polyethylene bag will not impact the safety of the package and any associated out-gassing will not contribute to package pressurization since the package is vented. As demonstrated in Section 3.5.3, Thermal

Docket No. 71-9330 ATR FFSC Safety Analysis Report Rev. 17, May 2022 3-75 Decomposition/Combustion of Package Organics, the available oxygen in the package is sufficient for consumption of less than 1% of the polyethylene and the quantity of air that enters the cavity under pressure differential and gas diffusion forces is insignificant. The discussion in Section 3.5.3 also provides validation of a 0 psig package pressure for the HAC event.

The outer shell and the ceramic fiber insulation provide thermal protection to the ATR FFSC package during the HAC fire event. The level of thermal protection can be seen via the thermal response curves presented in Figure 3.6-4 and Figure 3.6-5 for the ATR FFSC package loaded with a MIT and MURR fuel element payload, respectively. As seen from the figures, while the exterior of the package quickly rises to nearly the temperature of the fire, the heat flow to the FHE and its enclosed fuel element payloads is sufficiently restricted to limit the maximum temperatures of both the FHE and the fuel element to well below the melting point of aluminum.

The higher thermal mass of the MIT and MURR FHEs in comparison with that of the ATR FHE is reflected in their correspondingly slower heat up and longer cool down during the fire transient when compared to that see in Figure 3.4-1 for the ATR FHE. The higher temperature reached by the MURR FHE versus that seen for the MIT FHE is due to the conservative assumption of direct contact between the FHE and the inner shell along two line locations for the MURR FHE versus one line location for the MIT FHE. Similarly, the difference in the shape of the FHE temperature response curve seen for the MIT FHE between 30 minutes and 60 minutes versus that seen for the MURR FHE for the same time period is related to the fact that the top end of the shorter MIT FHE lies below one of the packages support ribs while the top of the MURR FHE is adjacent to it (see Figures 3.6-6 and 3.6-7).

Although the peak temperature achieved by the MURR FHE is about 20ºF hotter than that achieved by the MIT FHE, the peak temperatures reached by the MIT and MURR fuel elements are approximately the same. This results from a combination of the higher thermal mass and greater separation distance between the end of the fuel element and the start fuel plates associated with the MURR fuel element versus that for the MIT fuel element.

The results demonstrate that thermal performance is similar to that achieved with the transport of a LFPB payload (see Section 3.4.3, Maximum Temperature and Pressure) due to the fact that these FHE have a thermal mass similar to that of the LFPB. The result of the higher thermal mass is that the MIT and MURR FHEs have a peak temperature that is approximately 300ºF cooler than that seen for the ATR FHE and the enclosed fuel elements reach peak temperatures that are 90 to 180ºF cooler than that seen for the ATR fuel element. The thermal performance of the ATR FFSC packaging with either the MIT or MURR payload is similar to that seen for the ATR payload.

The results presented above also demonstrate that inclusion of insolation effects prior to the fire would not have affected the safety basis of the design. As documented in Section 3.4.3.1, Maximum HAC Temperatures, consideration of the maximum insolation loading raises the package component temperatures by approximately 50°F above the initial 100°F level assumed by the HAC evaluation. Since all package components exhibit thermal margins significantly greater than 50°F, the inclusion of insolation effects prior to the fire event would not have impacted the safety basis for the design.

As with the evaluation for NCT, the thermal performance of the ATR FFSC with the small quantity payload, RINSC, and Cobra fuel elements under HAC conditions was not specifically modeled as part of this evaluation. Instead, based on the similarity between the MURR and small

Docket No. 71-9330 ATR FFSC Safety Analysis Report Rev. 17, May 2022 3-76 quantity payloads, the thermal performance is qualitatively estimated to be bounded by that presented for the MURR payload. Since the combined weight of the small quantity payload and MURR FHE's with their enclosed fuel elements, plates, or foils are similar (see Section 1.2.2.3, MURR HEU Fuel Element, and Section 1.2.2.4, Small Quantity Payload) and the thermal mass of the two payloads are similar, the transient response of the small quantity payload can be expected to be similar to that presented for the MURR payload. This conclusion is further supported by the fact that Table 3.6-6 and Table 3.6-7 show that similar transient results occur for the MIT and MURR fuel element payloads despite their design differences. As such, the limited design differences between the MURR and small quantity payloads will not yield a significant difference in their HAC thermal response. This same logic applies to the RINSC, Cobra, and LEU fuel elements as further discussed in Section 3.6.9.2.4, Determination of Thermal Properties for RINSC Element, Section 3.6.9.2.5, Determination of Thermal Properties for Cobra Element, and Section 3.6.9.2.6, Determination of Thermal Properties for LEU Fuel Elements. Additionally, the SQFHE thermal response without its small quantity payload is expected to be similar with the conservative ATR LFPB thermal response. The empty SQFHE and LFPB are constructed of similar materials and have the same thermal mass of 30 lbs. The LFPB thermal evaluation is conservatively performed without its loose fuel plate payload, see Sections 3.4.3.1 and 3.5.2.1 for discussion of the LFPB thermal evaluation. Therefore, use of the SQFHE for any payload amount up to the maximum loaded SQFHE weight of 50 lbs is bounded by the thermal response of the LFPB evaluation. The addition of any small quantity payload mass to the SQFHE will increase the thermal mass and thereby increase the conservatism of the thermal response with respect to the empty LFPB thermal evaluation results.

3.6.8.3.2 Maximum HAC Pressures The payload cavity of the ATR FFSC is vented to the atmosphere. As such, the maximum pressure achieved under the HAC event will be 0 psig. Section 3.5.3, Thermal Decomposition/Combustion of Package Organics, provides the justification for assuming a 0-psig package pressure for the HAC event.

Although the volume between the outer and inner shells is sealed, it does not contain organic or other materials that may outgas or thermally decompose. Assuming a temperature of 70°F at the time of assembly and a maximum temperature of 1,475°F (based on the outer shell temperature, see Table 3.6-6), the maximum pressure rise within the sealed volume due to ideal gas expansion will be less than 39 psig. This level of pressurization will occur for only a few minutes and then quickly reduce as the package cools.

3.6.8.4 Maximum Thermal Stresses The ATR FFSC package is fabricated principally of sheet metal and relatively thin structural steel shapes. As such, the thermal stresses developed within each component during the HAC fire event will be low and not significant to the safety of the package.

The temperature difference that exists between the inner and outer shells during the HAC event (see the average inner and outer shell temperatures presented in Figure 3.6-4) will result in differential thermal expansion between the shells. The thermal impact related to the potential package geometry displacement due to this differential thermal expansion was evaluated in

Docket No. 71-9330 ATR FFSC Safety Analysis Report Rev. 17, May 2022 3-77 Section 3.4.4, Maximum Thermal Stresses, and found not to be significant to the safety of the package.

Table 3.6 HAC Temperatures with MIT Payload Location / Component Pre-fire End of Fire Peak Maximum Allowable MIT Fuel Element Fuel Plate 100°F 345°F 640°F 1,100°F MIT Fuel Element Side Plate 100°F 346°F 643°F 1,100°F Neoprene Rub Strips/Polyethylene Bag 100°F 599°F 690°F N/A Fuel Handling Enclosure (FHE) 100°F 599°F 690°F 1,100°F Inner Shell 100°F 1,417°F 1,417°F 2,700°F Ceramic Fiber Insulation, Body

- Maximum

- Average 100°F 100°F 1,462°F 1,253°F 1,462°F 1,253°F 2,300°F 2,300°F Ceramic Fiber Insulation, Closure

- Maximum

- Average 100°F 100°F 1,401°F 1,233°F 1,401°F 1,233°F 2,300°F 2,300°F Closure 100°F 1,439°F 1,439°F 2,700°F Outer Shell 100°F 1,475°F 1,475°F 2,700°F Table Notes:

The maximum allowable temperatures under HAC conditions are provided in Section 3.2.2, Technical Specifications of Components.

Docket No. 71-9330 ATR FFSC Safety Analysis Report Rev. 17, May 2022 3-78 Table 3.6 HAC Temperatures with MURR Payload Location / Component Pre-fire End of Fire Peak Maximum Allowable MURR Fuel Element Fuel Plate 100°F 371°F 636°F 1,100°F MURR Fuel Element Side Plate 100°F 380°F 644°F 1,100°F Neoprene Rub Strips/Polyethylene Bag 100°F 648°F 710°F N/A Fuel Handling Enclosure (FHE) 100°F 648°F 710°F 1,100°F Inner Shell 100°F 1,417°F 1,417°F 2,700°F Ceramic Fiber Insulation, Body

- Maximum

- Average 100°F 100°F 1,462°F 1,222°F 1,462°F 1,222°F 2,300°F 2,300°F Ceramic Fiber Insulation, Closure

- Maximum

- Average 100°F 100°F 1,402°F 1,236°F 1,402°F 1,236°F 2,300°F 2,300°F Closure 100°F 1,439°F 1,439°F 2,700°F Outer Shell 100°F 1,475°F 1,475°F 2,700°F Table Notes:

The maximum allowable temperatures under HAC conditions are provided in Section 3.2.2, Technical Specifications of Components.

Docket No. 71-9330 ATR FFSC Safety Analysis Report Rev. 17, May 2022 3-79 50 150 250 350 450 550 650 750 850 950 1050 1150 1250 1350 1450 1550

-30 30 90 150 210 270 330 390 450 510 570 630 690 Temperature - F Time - Minutes Outer Shell - Max.

Outer Shell - Avg.

Insulation Jacket - Max.

Insulation - Avg.

Inner Shell - Max.

Fuel Handling Enclosure (FHE) - Max.

Figure 3.6 ATR FFSC Package Thermal Response to HAC Event with MIT Payload 50 150 250 350 450 550 650 750 850 950 1050 1150 1250 1350 1450 1550

-30 30 90 150 210 270 330 390 450 510 570 630 690 Temperature - F Time - Minutes Outer Shell - Max.

Outer Shell - Avg.

Insulation Jacket - Max.

Insulation - Avg.

Inner Shell - Max.

Fuel Handling Enclosure (FHE) - Max.

Figure 3.6 ATR FFSC Package Thermal Response to HAC Event with MURR Payload

Docket No. 71-9330 ATR FFSC Safety Analysis Report Rev. 17, May 2022 3-80 (Note: the positive x-axis is oriented towards the top of the package and the positive z-axis towards the package closure end)

Figure 3.6-6 -Temperature Distribution at Time of Peak MIT Fuel Element Temperature

Docket No. 71-9330 ATR FFSC Safety Analysis Report Rev. 17, May 2022 3-81 (Note: the positive x-axis is oriented towards the top of the package and the positive z-axis towards the package closure end)

Figure 3.6-7 -Temperature Distribution at Time of Peak MURR Fuel Element Temperature

Docket No. 71-9330 ATR FFSC Safety Analysis Report Rev. 17, May 2022 3-82 3.6.9 Appendices 3.6.9.1 Computer Analysis Results 3.6.9.2 Analytical Thermal Model

Docket No. 71-9330 ATR FFSC Safety Analysis Report Rev. 17, May 2022 3-83 3.6.9.1 Computer Analysis Results Due to the size and number of the output files associated with each analyzed condition, results from the computer analysis are provided on a CD-ROM.

Docket No. 71-9330 ATR FFSC Safety Analysis Report Rev. 17, May 2022 3-84 3.6.9.2 Analytical Thermal Model The analytical thermal model of the ATR FFSC package and the MIT and MURR fuel element payloads were developed for use with the Thermal Desktop20 and SINDA/FLUINT21 computer programs. These programs are designed to function together to build, exercise, and post-process a thermal model. Appendix 3.5.2, Analytical Thermal Model, provides an overview of the capability and functionality of these programs. The SINDA/FLUINT and Thermal Desktop computer programs have been validated for safety basis calculations for nuclear related projects22. The ATR FFSC with the small quantity payload was not specifically modeled as part of this evaluation. Instead, its thermal performance is estimated using a qualitative approach based on the thermal characteristics of the other payloads and their associated thermal performance.

3.6.9.2.1 Description of MIT and MURR Payload Thermal Models for NCT Conditions A 3-dimensional, one-quarter symmetry thermal model of the ATR FFSC is used for the NCT evaluation. The model simulates one-quarter of the package, extending from the closure to the axial centerline of the package. Symmetry conditions are assumed about the packages vertical axis and at the axial centerline. This modeling choice captures the full height of the package components and allows the incorporation of the varying insolation loads that will occur at the top and sides of the package. Program features within the Thermal Desktop computer program automatically compute the various areas, lengths, thermal conductors, and view factors involved in determining the individual elements that make up the thermal model of the complete assembly. Details of the thermal modeling of the ATR FFSC packaging are provided in Appendix 3.5.2.1, Description of Thermal Model for NCT Conditions.

A detailed model of the MIT and MURR fuel elements are used to simulate the heat transfer within the fuel elements and between the fuel element and their associated FHEs and spacer weldments. The detailed thermal models, illustrated in Figure 3.6-8 to Figure 3.6-13, include a separate representation of each composite fuel plate, the side plates, and the end fittings. Heat transfer between the individual fuel plates is simulated via conduction and radiation across the air space separating the plates. The curvature and separation distance between the plates is based on the information presented in Appendix 3.6.9.2.3, Determination of Composite Thermal Properties for MIT and MURR Fuel Plates.

The thermal modeling for the MIT fuel element and FHE is similar to that described for the ATR fuel element payload. Figure 3.6-8 illustrates the quarter symmetry thermal model of the MIT FHE and one of the two spacer weldments. The FHE thermal model uses planar elements to represent the 0.19 inch thick sides of the enclosure and the 0.25 inch thick elements of the spacer weldment. Solid elements are used to represent the ends of the FHE. Heat transfer between the FHE and the inner shell of the package is modeled as a combination of radiation and conduction across the air-filled void space, as well as via direct contact along 1 edge of the FHE. The contact conductance simulates a conservative idealized physical contact (i.e., a flat, smooth interface and that the FHE is oriented within the package such that the edge is aligned with the vertical axis of the package) between the FHE and the inner shell. Due to the robustness of the MIT FHE, no change to the direct contact between the FHE and the inner shell conservatively assumed for the NCT condition is expected as a result of the HAC drop event.

Docket No. 71-9330 ATR FFSC Safety Analysis Report Rev. 17, May 2022 3-85 Figure 3.6-9 illustrates a cross-section through the combined modeling for the inner shell, the FHE, and the MIT fuel element. The left side of the figure illustrates the placement of the thermal nodes (indicated by the small circles) used to simulate each of the components, the use of planar elements to represent the 15 fuel plates, and the assumed points of direct contact between the FHE and the inner shell. The right side of the figure includes depiction of the solid elements that are used to simulate the air voids around the FHE. The heat transfer between the FHE and the MIT fuel element is computed as conductance through the 0.125 inch thick neoprene rub strips and radiation and conductance through the air voids.

Figure 3.6-10 illustrates a side and end view of the thermal model of the MIT fuel element as it would be for a complete fuel element. Approximately 1,140 nodes, 350 planar elements, and 445 solids are used to represent the quarter symmetry thermal model of the MIT fuel element, FHE, and the spacer weldment.

The thermal modeling for the MURR fuel element and FHE is similar to that described above for the MIT fuel element payload. Figure 3.6-11 illustrates the quarter symmetry thermal model of the MURR FHE and one of the two spacer weldments. The FHE thermal model uses planar elements to represent the 0.19 inch thick sides of the enclosure and the 0.25 inch thick elements of the spacer weldment. Solid elements are used to represent the ends of the FHE. Heat transfer between the FHE and the inner shell of the package is modeled as a combination of radiation and conduction across the air-filled void space, as well as via direct contact along 2 edges of the FHE. The contact conductance simulates a conservative idealized physical contact (i.e., a flat, smooth interface and an alignment that places 2 edges of the FHE in contact) between the FHE and the inner shell. Due to the robustness of the MURR FHE, no change to the direct contact between the FHE and the inner shell conservatively assumed for the NCT condition is expected as a result of the HAC drop event.

Figure 3.6-12 illustrates a cross-section through the combined modeling for the inner shell, the FHE, and the MURR fuel element. The left side of the figure illustrates the placement of the thermal nodes (indicated by the small circles) used to simulate each of the components, the use of curved, planar elements to represent the 24 fuel plates, and the assumed points of direct contact between the FHE and the inner shell. The right side of the figure includes depiction of the solid elements that are used to simulate the air voids around the FHE. The heat transfer between the FHE and the MURR fuel element is computed as conductance through the 0.125 inch thick neoprene rub strips and radiation and conductance through the air voids.

Figure 3.6-13 illustrates a side and end view of the quarter symmetry thermal modeling used for the MURR fuel element. Approximately 1,400 nodes, 700 planar elements, and 340 solids are used to represent the quarter symmetry thermal model of the MURR fuel element, FHE, and the spacer weldment.

The heat transfer from the exterior surfaces of the ATR FFSC is modeled in the same manner as that used for the evaluation of the ATR fuel element payload and assumes a combination of convection and radiation exchange. Appendix 3.5.2.3, Convection Coefficient Calculation, presents the methodology used to compute the convection coefficients from the various surfaces.

The radiation exchange is computed using a Monte Carlo, ray tracing technique and includes the affect of reflection and/or transmission, according to the optical properties assigned to each surface (see Section 3.2.1, Material Properties).

Docket No. 71-9330 ATR FFSC Safety Analysis Report Rev. 17, May 2022 3-86 In addition, heating of the exterior surfaces due to solar insolation is assumed using a diurnal cycle. The methodology used to simulate and apply the insolation loading is described in Appendix 3.5.2.1, Description of Thermal Model for NCT Conditions.

3.6.9.2.2 Description of Thermal Model for HAC Conditions The thermal evaluations for the hypothetical accident condition (HAC) are conducted in the same manner and using the same methodology as that described in Appendix 3.6.9.2.1, Description of MIT and MURR Payload Thermal Models for NCT Conditions. No change to the geometry or position of the MIT and MURR fuel element payloads are expected as a result of the drop and puncture bar events that are assumed to precede the HAC fire event.

3.6.9.2.3 Determination of Composite Thermal Properties for MIT and MURR Fuel Plates The MIT and MURR fuel plates are a composite material consisting of a fissile fuel matrix sandwiched within aluminum cladding. For the purposes of this calculation, the fuel composite is treated as a homogenous material with lumped thermal properties. The methodology used to compute the composite thermal properties for each fuel element is the same as that described in Appendix 3.5.2.4, Determination of Composite Thermal Properties for ATR Fuel Plates.

Each MIT element contains up to 515 g U-235, enriched up to 94 wt.%, which equates to a density of approximately 1.5 g U/cc in the fuel matrix. The thermal properties for the individual plates making up the MIT fuel element are computed using the approach used with the ATR Fuel Plates and the geometric data39,42 for the MIT fuel element. Each of the fifteen (15) fuel plates contained in the MIT fuel element has a thickness of 0.08 inches and a width of 2.526 inches.

The nominal gap between the plates is 0.078 inches. Since the aluminum cladding contains 110 grooves on each side of the plate, the effective thickness of the cladding is reduced from 0.025 inches to 0.02 inches. Table 3.6-8 presents the composite thermal conductivity, specific heat, and density values for the fuel plates. These composite values are based on the described geometry of the fuel plates and the same thermophysical data6 used for the ATR fuel plates.

The thermal properties for the MIT element used are:

1) Aluminum cladding thermal conductivity = 191 W/m-K, conservatively high value from

[6], page 18

2) Fissile fuel matrix (UAlx) conductivity = 38.5 W/m-K, conservatively high based on Table 2.3 from [6] at 300K for 1.5 g U/cc

3) Aluminum cladding density = 2702 kg/m3, from [6], page 16

4) Fissile fuel matrix (UAlx) density = 3846 kg/m3, from [6], Table 2.5 for 1.5 g U/cc

5) Aluminum cladding specific heat = 896 & 1080 J/kg-K, from [6], Table 3.2 at 300 &

700K, respectively

6) Fissile fuel matrix (UAlx) specific heat = 587 & 709 J/kg-K, from [6], Table 2.4, value at 300 & 700K, respectively, for 1.5 g U/cc 42 Massachusetts Institute of Technology, Test Research Training Reactor 3 Welded Fuel Element Assembly, EG&G Idaho, Inc. Drawing No. DWG-419486, Rev. A.

Docket No. 71-9330 ATR FFSC Safety Analysis Report Rev. 17, May 2022 3-87 Each MURR element contains up to 785 g U-235, enriched up to 94 wt.%, which equates to a density of approximately 1.44 g U/cc in the fuel matrix. The thermal properties for the individual plates making up the MURR fuel element are also computed using the approach used with the ATR Fuel Plates and the geometric data40,43 for the MURR fuel element. Due to the curved geometry of the twenty-four (24) fuel plates contained in the MURR fuel element, each plate has a different geometry. The inner plate has an inner radius of 2.77 inches and an arc length of 1.993 inches, while the outer plate has an inner radius of 5.76 inches and an arc length of 4.342 inches. The nominal gap between the plates is 0.08 inches. The thickness of the aluminum cladding is 0.01 inches. Table 3.6-9 presents the composite thermal conductivity, specific heat, and density values for the twenty four (24) fuel plates making up the MURR fuel element. These composite values are based on the described geometry of the fuel plates and the same thermophysical data6 used for the ATR fuel plates.

The thermal properties for the MURR fuel element used in this calculation are:

1) Aluminum cladding thermal conductivity = 191 W/m-K, conservatively high value from

[6], page 18

2) Fissile fuel matrix (UAlx) conductivity = 39.8 W/m-K, conservatively high based on Table 2.3 from [6], at 300K for 1.44 g U/cc

3) Aluminum cladding density = 2702 kg/m3, from [6], page 16

4) Fissile fuel matrix (UAlx) density = 3793 kg/m3, from [6], Table 2.5 for 1.44 g U/cc

5) Aluminum cladding specific heat = 896 & 1080 J/kg-K, from [6], Table 3.2, at 300 &

700K, respectively

6) Fissile fuel matrix (UAlx) specific heat = 596 & 719 J/kg-K, from [6], Table2.4, value at 300 & 700K, respectively, for 1.44 g U/cc 3.6.9.2.4 Determination of Thermal Properties for RINSC Element The RINSC fuel elements are fabricated with a nominally 0.020-in thick mixture of uranium silicide (U3Si2) and aluminum powder as the fuel meat and a nominally 0.015-in thick aluminum alloy cladding. The twenty-two (22) flat fuel plates have a 2.8-in width, an overall length of 25-in, and an active fuel region of 22.5 to 24.0-in. The fuel plate meat and cladding thicknesses match those of the interior plates for the ATR fuel element and are similar to those for the MURR fuel plates. The side plates are fabricated of ASTM B 209, aluminum alloy 6061-T6 and 6061-T651 and are approximately 0.188-in thick. This is similar to the side plate thicknesses of the ATR, MITR, and MURR fuel elements.

The thermal conductivity of the RINSC fuel plates41 are similar to data obtained in the measurements of the thermal conductivities for the uranium aluminide (UAlx) based fuels6.

Similarly, the thermal mass of the fuel plates are comparable despite the higher density of uranium silicide versus uranium aluminide since the ratio of the specific heats of the two materials is nearly the inverse of the density ratio.

43 University of Missouri at Columbia, MURR UAlx Fuel Element Assembly, EG&G Idaho, Inc. Drawing No.

DWG-409407, Rev. N.

Docket No. 71-9330 ATR FFSC Safety Analysis Report Rev. 17, May 2022 3-88 The ATR FFSC with the RINSC fuel element payload is not specifically modeled as part of this evaluation. Instead, its thermal performance is estimated using a qualitative approach based on the maximum temperatures attained for the transportation of the MURR fuel element within the ATR FFSC. This conclusion is based on the facts that the combined weight of the RINSC and MURR FHE's with their enclosed fuel elements are the same, the FHE's are both fabricated of 6061 aluminum, and the fuel elements have similar thermal properties (see above). This conclusion is further supported by the fact that the MIT and MURR fuel elements produce essentially the same peak temperatures despite their design differences. As such, the limited design differences between the MURR and RINSC payloads will not yield a significant difference in their thermal response.

3.6.9.2.5 Determination of Thermal Properties for Cobra Element As with the MURR and MIT fuel elements, the temperature of the Cobra fuel element will vary based on the heat that flows into the ATR FFSC package from insolation (NCT) and the hypothetical fire (HAC). The temperature of the fuel element will depend primarily on the nature of the dominant resistances in the heat path between the fuel and the environment, and on the heat capacity or thermal mass of the package. A comparison can be made between the Cobra fuel element case and the MURR case, whose resulting temperatures are given in Table 3.6-5 (NCT) and Table 3.6-7 (HAC). The dominant resistances consist of the non-metallic links in the heat path from the outside to the inside (such as air gaps and rubber); the resistance through the metallic elements is comparatively negligible (such as steel and aluminum), and will be neglected in what follows. The non-metallic elements in the heat path are:

The insulation between the inner and outer shells of the package, based on the thermal conductivity, thickness, and area of the insulation The air gap between the inner shell and the FHE, based on the radiative heat transfer properties and the area The contact conductance of the FHE resting on the inner shell The neoprene rubber between the FHE and the fuel element The air gap between the FHE and the fuel element The contact conductance of the fuel element resting on the rubber Because the Cobra fuel element is transported within a FHE having a design very similar to that of the MURR fuel element, the dominant heat transfer resistances will be very similar, including the same number and approximate size of air gaps, emissivities, contact conductances, and rubber thickness. The thermal behavior in transient heat transfer also depends on the thermal mass of the components. From Table 2.1-1, the weight of the MURR fuel element and its FHE are 15 lb and 30 lb, respectively, for a total of 45 lb; and the weight of the Cobra fuel element and its FHE are 20 lb and 28 lb, respectively, for a total of 48 lb. Since the heat capacity of all aluminum alloys is very similar, and since the total weight of each fuel element plus FHE is essentially the same, the thermal mass will be essentially the same. Thus, the Cobra fuel element case will have essentially the same thermal behavior in NCT and HAC to the MURR fuel

Docket No. 71-9330 ATR FFSC Safety Analysis Report Rev. 17, May 2022 3-89 element case. In addition, the temperatures calculated for the MURR case show significant margins to the limiting temperatures for the various components.

3.6.9.2.6 Determination of Thermal Properties for LEU Fuel Elements The full size prototype LEU fuel elements and DDEs to be transported in the ATR FFSC are:

ATR LEU fuel element MIT LEU fuel element and DDE MURR LEU fuel element and DDE NBSR DDE Each of the LEU fuel elements has the same design as the corresponding HEU fuel element, including bounding dimensions, except for a different number of fuel plates in some cases. The DDEs have the same core design as the corresponding LEU fuel element, differing only in the end structures. The NBSR DDE does not have a corresponding LEU fuel element used in the ATR FFSC, but its materials, construction, and weight are similar to those of the other LEU fuel elements and DDEs.

The U-10Mo fuel meat is an alloy of nominally 10-wt% molybdenum with uranium which is enriched up to 20%. The meat material is clad with a nominal 0.001-inch thick layer of zirconium and with 6061 aluminum alloy. The ATR LEU fuel element and the MIT LEU and MURR LEU fuel elements and DDEs are transported within the same FHE as is used for the corresponding HEU fuel elements. The NBSR DDE is transported with cellulosic protective material (such as cardboard).

The LEU fuel elements and DDEs are not explicitly modeled for this evaluation, because they are considered to be bounded by the ATR, MIT, and MURR HEU fuel elements which are modeled in this Appendix.

Because all of the payload fuel elements have negligible decay heat, the temperature under NCT will be completely determined by solar loading, and the thermal characteristics of the fuel elements or FHEs will have little influence over the temperature. Under HAC, the heat input is dominated by the hypothetical fire. Because the thermal mass dominates the heat transfer relations, the effect of any differences in the overall conductivity of the fuel elements is negligible. The higher thermal mass of the LEU fuel elements and DDEs serves to damp the thermal transient response in the model compared to the HEU case. The LEU fuel elements are heavier than the corresponding HEU fuel elements by factors of 1.8 (ATR), 1.9 (MIT), and 1.9 (MURR). The DDEs all have weights comparable to the LEU prototypic fuel elements. Thus, all of the LEU payloads are significantly heavier than the current HEU payloads. This will result in a lower peak temperature for the LEU fuel elements and DDEs under HAC.

The thermal decomposition of the cellulosic material surrounding the NBSR DDE will not have a significant effect on the payload temperature, due to the limited supply of oxygen in the package during the fire, as discussed in Section 3.5.3, Thermal Decomposition/Combustion of Package Organics. In addition, the loss of these components is not critical to the safety of the package. Furthermore it is noted that the margin of safety against melting of the fuel plates is over 300 °F for the ATR HEU element (see Table 3.1-1) and over 400 °F for the MIT/MURR

Docket No. 71-9330 ATR FFSC Safety Analysis Report Rev. 17, May 2022 3-90 fuel plate (see Table 3.6-1). Thus, the thermal integrity of the LEU fuel elements will be maintained through the HAC fire event.

Docket No. 71-9330 ATR FFSC Safety Analysis Report Rev. 17, May 2022 3-91 Table 3.6 Composite MIT Fuel Plate Thermal Properties Plate Plate Thickness, in UAlx Thickness, in Axial and Circumferential Conductivity (W/m-K)

Plate Width, in Mean density, kg/m^3 Mean specific heat, J/(kg K) @ 300 K Mean specific heat, J/(kg K) @ 700 K 1 to 15 0.08*

0.03 125.6 2.314 3192.3 736.5 888.4

• - mean plate thickness estimated at 0.07 inches after allowance for ribbing Table 3.6 Composite MURR Fuel Plate Thermal Properties Plate Plate Thickness, in UAlx Thickness, in Axial and Circumferential Conductivity (W/m-K)

Inner radius, in Outer radius, in Plate Arc Length, in Mean density, kg/m^3 Mean specific heat, J/(kg K) @ 300K Mean specific heat, J/(kg K) @ 700 K 1

0.05 0.03 100.3 2.77 2.82 1.993 3288.8 692.6 835.2 2

0.05 0.03 100.3 2.9 2.95 2.095 3288.8 692.6 835.2 3

0.05 0.03 100.3 3.03 3.08 2.197 3288.8 692.6 835.2 4

0.05 0.03 100.3 3.16 3.21 2.300 3288.8 692.6 835.2 5

0.05 0.03 100.3 3.29 3.34 2.402 3288.8 692.6 835.2 6

0.05 0.03 100.3 3.42 3.47 2.504 3288.8 692.6 835.2 7

0.05 0.03 100.3 3.55 3.6 2.606 3288.8 692.6 835.2 8

0.05 0.03 100.3 3.68 3.73 2.708 3288.8 692.6 835.2 9

0.05 0.03 100.3 3.81 3.86 2.810 3288.8 692.6 835.2 10 0.05 0.03 100.3 3.94 3.99 2.912 3288.8 692.6 835.2 11 0.05 0.03 100.3 4.07 4.12 3.014 3288.8 692.6 835.2 12 0.05 0.03 100.3 4.2 4.25 3.116 3288.8 692.6 835.2 13 0.05 0.03 100.3 4.33 4.38 3.218 3288.8 692.6 835.2 14 0.05 0.03 100.3 4.46 4.51 3.321 3288.8 692.6 835.2 15 0.05 0.03 100.3 4.59 4.64 3.423 3288.8 692.6 835.2 16 0.05 0.03 100.3 4.72 4.77 3.525 3288.8 692.6 835.2 17 0.05 0.03 100.3 4.85 4.9 3.627 3288.8 692.6 835.2 18 0.05 0.03 100.3 4.98 5.03 3.729 3288.8 692.6 835.2 19 0.05 0.03 100.3 5.11 5.16 3.831 3288.8 692.6 835.2 20 0.05 0.03 100.3 5.24 5.29 3.933 3288.8 692.6 835.2 21 0.05 0.03 100.3 5.37 5.42 4.035 3288.8 692.6 835.2 22 0.05 0.03 100.3 5.5 5.55 4.137 3288.8 692.6 835.2 23 0.05 0.03 100.3 5.63 5.68 4.239 3288.8 692.6 835.2 24 0.05 0.03 100.3 5.76 5.81 4.342 3288.8 692.6 835.2

Docket No. 71-9330 ATR FFSC Safety Analysis Report Rev. 17, May 2022 3-92 (Note: the positive x-axis is oriented towards the top of the package and the positive z-axis towards the package closure end)

Figure 3.6 Hidden Line View of MIT FHE and Spacer Quarter Symmetry Thermal Model Modeling Showing Direct Contact Modeling with Solid Elements for Air Figure 3.6 Thermal Model of MIT Fuel Element and FHE within Inner Shell

Docket No. 71-9330 ATR FFSC Safety Analysis Report Rev. 17, May 2022 3-93 MIT Fuel Element Model, Side View of Full Element MIT Fuel Element Model, End View of Full Element Figure 3.6 Side and End Views of MIT Fuel Element Thermal Model