TN 68 TRANSPORT PACKAGING APPENDIX 2.10.1 TABLE OF …Shock Loads (Cask Horizontal) 2.10.1-60 Local...

TN 68 TRANSPORT PACKAGING

APPENDIX 2.10.1

TABLE OF CONTENTS

Page

2.10.1 STRUCTURAL EVALUATION OF CASK BODY

2.10.1.1 Introduction ...................................................................................... 2.10.1-1

2.10.1.2 ANSYS M odel ................................................................................. 2.10.1-1

2.10.1.3 Axisymmetric Loadings .................................................................. 2-10.1-3

2.10.1.4 Asymmetric Loading ........................................................................ 2.10.1-8

2.10.1.5 Transport Shock Loading ............................................................... 2.10.1-18

2.10.1.6 Transport Vibration Loading .......................................................... 2.10.1-19

2.10.1.7 Transport Tie-Down Loading ......................................................... 2.10.1-20

2.10.1.8 6 G Lifting ...................................................................................... 2.10.1-20

2.10.1.9 Summary of Individual Load Cases ............................................... 2.10.1-21

2.10.1.10 Trunnion Local Stress Analysis ..................................................... 2.10.1-22

2.10.1.11 References ...................................................................................... 2.10.1-24

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LIST OF TABLES

2.10.1-4A Individual Load Cases for TN-68 Cask Body Analysis 2.10.1-4B Trunnion Loads 2.10.1-1 Bolt Pre-Load and Gasket Seating Pressure 2.10.1-2 Bolt Pre-Load and Gasket Seating Pressure 2.10.1-3 Fabrication Stresses 2.10.1-4 Fabrication Stresses 2.10.1-5 Internal Pressure (100 psi) 2.10.1-6 Internal Pressure (100 psi) 2.10.1-7 External Pressure (25 psi) 2.10.1-8 External Pressure (25 psi) 2.10.1-9 End Drop on Bottom - Rear Impact Limiter (1 G) 2.10.1-10 End Drop on Bottom - Rear Impact Limiter (1 G) 2.10.1-11 End Drop on Lid - Front Impact Limiter (I G) 2.10.1-12 End Drop on Lid - Front Impact Limiter (IG) 2.10.1-13 Thermal Stress at 1000 F Hot Environment 2.10.1-14 Thermal Stress at 100' F Hot Environment 2.10.1-15 Thermal Stress at -20* F Environmental Conditions 2.10.1-16 Thermal Stress at -20' F Environmental Conditions 2.10.1-17 Thermal Stress at -40' F Environmental Conditions 2.10.1-18 Thermal Stress at -40' F Environmental Conditions 2.10.1-19 Transport Shock Load/Truck,

Horizontal Cask Supported at Rear Trunnions and Front Saddle, (Stress at 1800 Orientation)

2.10.1-20 Transport Shock Load / Truck, Horizontal Cask Supported at Rear Trunnions and Front Saddle, (Stress at 1800 Orientation)

2.10.1-21 1G Down Load, Horizontal Cask Supported at Rear Trunnions and Front Saddle, (Stress at 180' Orientation)

2.10.1-22 1G Down Load Horizontal Cask Supported at Rear Trunnions and Front Saddle, (Stress at 1800 Orientation)

2.10.1-23 Side Drop (1G), (Stress at 900 Orientation) 2.10.1-24 Side Drop (IG), (Stress at 90' Orientation) 2.10.1-25 Side Drop (IG), (Stress at Impact Side 180" Orientation) 2.10.1-26 Side Drop (IG), (Stress at Impact Side 1800 Orientation) 2.10.1-27 C.G. Over Bottom Comer Drop - Rear Impact Limiter (1G),

(Stress at 900 Orientation) 2.10.1-28 C.G. Over Bottom Comer Drop - Rear Impact Limiter (IG),

(Stress at 90' Orientation)

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LIST OF TABLES (Continued)

2.10.1-29 C.G. Over Bottom Comer Drop - Rear Impact Limiter (IG),

(Stress at Impact Side, 1800 Orientation) 2.10.1-30 C.G. Over Bottom Comer Drop - Rear Impact Limiter (IG),

(Stress at Impact Side, 1800 Orientation) 2.10.1-31 C.G. Over Lid Comer Drop - Front Impact Limiter (1G),

(Stress at 90' Orientation) 2.10.1-32 C.G. Over Lid Comer Drop - Front Impact Limiter (IG),

(Stress at 900 Orientation) 2.10.1-33 C.G. Over Lid Comer Drop - Front Impact Limiter (IG),

(Stress at Impact Side, 180' Orientation) 2.10.1-34 C.G. Over Lid Comer Drop - Front Impact Limiter (IG),

(Stress at Impact Side, 180' Orientation)

2.10.1-35 150 Slap Down Drop - Second Impact (Gnornial = 42.22, Grotational = 77.78),

(Stress at 900 Orientation) 2.10.1-36 150 Slap Down Drop - Second Impact

(Gnormal = 42.22, Grotationai = 77.78), (Stress at 90' Orientation)

2.10.1-37 150 Slap Down Drop - Second Impact (Gno.ma = 42.22, Grotational = 77.78),

(Stress at Impact Side, 1800 Orientation)

2.10.1-38 150 Slap Down Drop - Second Impact

(Gnorml = 42.22, Grotationa = 77.78), (Stress at Impact Side, 1800 Orientation)

2.10.1-39 Fire Accident 2.10.1-40 Fire Accident 2.10.1-41 1 g Longitudinal Loading, Horizontal Cask Held at Rear Trunnions and

Front Saddle 2.10.1-42 1 g Longitudinal Loading, Horizontal Cask Held at Rear Trunnions and

Front Saddle 2.10.1-43 Transport Truck Vibration Load (0.3g Long., 0.3g Lat., 0.6g Vert.),

Horizontal Cask Held at Rear Trunnions and Front Saddle

2.10.1-44 Transport Truck Vibration Load (0.3g Long., 0.3g Lat., 0.6g Vert.),


2.10.1-45 Transport Tie-Down Load (10g Long., 5g Lat., 2g Vert.),


2.10.1-46 Transport Tie-Down Load (log Long., 5g Lat., 2g Vert.),


2.10.1-47 6g Lifting Load on Trunnion (Cask Vertical) 2.10.1-48 6g Lifting Load on Trunnion (Cask Vertical)

2.10.1-49 Local Stresses at Rear Trunnion / Cask Body Interface with lg Down

Load (Cask Horizontal)

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LIST OF TABLES (Continued)

2.10.1-50 Local Stresses at Rear Trunnion / Cask Body Interface with Ig Down Load (Cask Horizontal)

2.10.1-51 Local Stresses at Rear Trunnion / Cask Body Interface with Truck Shock Loads (Cask Horizontal)

2.10.1-52 Local Stresses at Rear Trunnion / Cask Body Interface with Truck Shock Loads (Cask Horizontal)

2.10.1-53 Local Stresses at Rear Trunnion / Cask Body Interface with Vibration Loads (Cask Horizontal)

2.10.1-54 Local Stresses at Rear Trunnion / Cask Body Interface with Vibration Loads (Cask Horizontal)

2.10.1-55 Local Stresses at Rear Trunnion / Cask Body Interface with Tie-Down Loads (Cask Horizontal)

2.10.1-56 Local Stresses at Rear Trunnion / Cask Body Interface with Tie-Down Loads (Cask Horizontal)

2.10.1-57 Local Stresses at Rear Trunnion / Cask Body Interface with 6g Lifting Load (Cask Vertical)

2.10.1-58 Local Stresses at Rear Trunnion / Cask Body Interface with 6g Lifting Load (Cask Vertical)

2.10.1-59 Local Stresses at Rear Trunnion / Cask Body Interface with Rail Car Shock Loads (Cask Horizontal)

2.10.1-60 Local Stresses at Rear Trunnion / Cask Body Interface with Rail Car Shock Loads (Cask Horizontal)

2.10.1-61 Transport Rail Car Shock Load (4.7g All Directions) Horizontal Cask Held at Rear Trunnions and Front Saddle

2.10.1-62 Transport Rail Car Shock Load (4.7g All Directions) Horizontal Cask Held at Rear Trunnions and Front Saddle

2.10.1-63 Bijlaard Computation Sheet

2.10.1-iv Rev. 0 4/99

LIST OF FIGURES

2.10.1-1 2.10.1-2 2.10.1-3 2.10.1-4 2.10.1-5 2.10.1-6 2.10.1-7 2.10.1-8 2.10.1-9 2.10.1-10 2.10.1-11 2.10.1-12 2.10.1-13 2.10.1-14 2.10.1-15 2.10.1-16 2.10.1-17 2.10.1-18 2.10.1-19 2.10.1-20 2.10.1-21

2.10.1-22 2.10.1-23 2.10.1-24

2 Rev. 0 4/99

Cask Body Key Dimensions Cask Body Finite Element Model Cask Body Bottom Corner Cask Body Top Comer Cask Lid to Shield Plate Connection Bolt Preload and Seal Reaction Design Internal Pressure (100 psi)

External Pressure Loading (25 psi)

Impact at Bottom End Load Distribution Impact at Lid End Load Distribution

1G Longitudinal Supported by Two Rear Trunnions

Fourier Coefficients for 1 G Lateral

Location of C.G., Trunnions, Saddle, and Impact Limiters

1G Vertical Down, Supported by Two Rear Trunnions and Front Saddle

Side Drop Load Distribution C.G. Over Bottom Corner Drop Load Distribution

C.G. Over Lid End Corner Drop Load Distribution Transverse g Load Combinations

Combined g Load Used for 30 Foot 15' Slap Down Structural Analysis

150 Slap Down (Second Impact at Lid End) Load Distribution

15' Slap Down (Second Impact at Lid End) Rotational Quasistatic

Equilibrium Standard Reporting Locations for Cask Body Weld Stress Locations Lifting: 6g Vertical Up

2.10.1-v

THIS PAGE IS INTENTIONALLY LEFT BLANK.

Rev. 0 4i99

APPENDIX 2.10.1

STRUCTURAL EVALUATION OF CASK BODY

2.10.1.1 Introduction

This appendix presents the structural analyses of the TN-68 cask body including the cylindrical

shell assembly and bottom assembly, the lid, and the local stress at the trunnion / cask body

interface. The specific methods, models and assumptions used to analyze the cask body for the

various individual loading conditions specified in 10CFR71.71(') and 10CFR71.73(2) are

described. Stress results are reported at selected locations for each load case. Maximum stresses

from this appendix are evaluated in Sections 2.6 and 2.7 of Chapter Two where the load

combinations outlined in Regulatory Guide 7.0(3) are performed and the results evaluated against

the ASME Code(4) and Regulatory Guide 7.6(5) design criteria described in Section 2.1.2.

The TN-68 cask body structural analyses generally use static or quasistatic linear elastic methods

so that combinations of loads can be examined by superimposing the results from individual

loads. The stresses and deformations due to the applied loads are generally determined using the

ANSYS( 6) computer program.

The detailed calculations for the lid bolts are presented in Appendix 2.10.2. The calculations for

the outer shell are reported in Appendix 2.10.3. Stress evaluations of the lifting and tie-down

devices are described in Section 2.6 of Chapter 2. The evaluation of the cask body under the 40

inch puncture event is described in Section 2.7.2.

The two analysis methods described in this appendix and used to evaluate the cask body for the

individual loading conditions are:

"° ANSYS Analysis - Axisymmetric and Asymmetric Loads

"0 Bijlaard Trunnion Local Stress Analysis

The Bijlaard(7) trunnion local stress analysis is performed to determine the local stresses in the

gamma shield shell at locations that correspond to stress reporting locations selected for the

ANSYS analyses. This permits the localized trunnion induced stresses to be easily combined

with stresses obtained from appropriate ANSYS load cases. The method of combining stress

results from individual load cases to evaluate the required load combinations is discussed in

Section 2.6 of Chapter Two for normal conditions of transport and Section 2.7 for hypothetical

accident conditions.

2.10.1.2 ANSYS Analysis

Geometry Description

The cask body as shown in Figure 2.10.1-1 consists of:

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1. A 1 1/2 in. thick inner vessel with a welded flat bottom, a flange welded at the top, and a lid bolted to the flange by 48, 1.875 in. diameter high strength bolts and sealed with two metallic o-rings. This is the containment vessel, the primary containment boundary of the cask.

2. A thick cylindrical vessel with a welded flat bottom surrounding the containment. This vessel and a steel disk welded to the lid inner surface provide the gamma shielding.

The lid and the flange are carbon steel forgings as are the gamma shielding components. The cask body is designed as a Class 1 component in accordance with the rules of the ASME Code.

ANSYS Cask Model

A two dimensional ANSYS model is used to evaluate the stresses in the cask body due to the individual load cases. The finite elements used in the model are the axisymmetric shell element, SHELL 61, and the axisymmetric harmonic element, PLANE 25. Both of these elements consider axisymmetric and non-axisymmetric loadings.

The cylindrical containment shell and bottom are modeled using SHELL 61 elements. The remainder of the cask body is modeled with PLANE 25 elements except for the lid bolts which are modelled with the two dimensional elastic beam, BEAM 3. The finite element model of the cask body is shown in Figure 2.10.1-2.

Figure 2.10.1-3 shows an enlarged view of the bottom comer with the weld joining the gamma shielding flat bottom to cylinder simulated by coupling nodes 236-107 and 280-108.

The weld connecting the gamma shielding cylinder to the containment flange is simulated by coupling nodes 63-328 and 64-329 as shown in Figure 2.10.1-4. The gamma shielding is heated prior to assembly with the containment shell and flange for ease of installation. During cooldown, a gap may result between the flange and the gamma shield shell. The gap is filled with shim plates made from SA-516, Grade 70 plate. The plates are fit between the gamma shield shell and the flange behind the weld. These shim plates are not modeled. The weld between the gamma shield and the flange is not affected by the shims. Also shown in this figure are the lid bolts connecting the lid to the containment flange. The connection is simulated by coupling nodes 505, 506 and 507 of the bolts to the corresponding nodes 81, 74, and 67 of the flange; and nodes 501, 502 and 503 of the bolts to the corresponding nodes 438, 439, and 440 of the lid. In this manner the threaded portion of the bolt is fixed to the flange while the bolt head is fixed to the top surface of the lid. In order to prevent the lid from moving into the flange, nodes 79 and 395 are also coupled in the axial or Y direction. The enlarged view in Figure 2.10.1-5 shows the coupling of nodes 394-383 and 395-384 which simulates the weld connecting the containment lid to the gamma shielding disk.

The pairs of nodes listed above, with the exception of nodes 79-395, are coupled in the X, Y and Z directions. The coupling of nodes 79-395 is in the Y direction only. Nodes 80-396 and 82-398 are also coupled in Y-direction. This is accomplished using constraint equations. The reactions

Rev. 0 4/992.10.1-2

at these nodes are monitored during the analysis to insure that tensile forces between the flange

and the lid are not developed.

Appropriate boundary conditions are applied to prevent rigid body motion and to show that the

system of forces applied to the cask in each of the individual load cases is in equilibrium.

Generally a node at the center of the vessel bottom is held in all directions and all nodes at the

center line are held in the X direction. Node 78 (Figure 2.10.1-4) is held in the Z direction to

avoid rigid body motion.

Loading Conditions

The loading conditions analyzed simulate or represent various effects due to the normal

conditions of transport and hypothetical accident conditions specified in 1 OCFR7 1. These

individual loading conditions, called load cases, are superimposed or combined as specified in

Regulatory Guide 7.8 in Sections 2.6 and 2.7 of this SAR.

2.10.1.3 Axisvmmetric Loadings

The following individual axisymmetric load cases analyzed using this ANSYS model are

described in this section.

(1) Bolt preload and lid seating pressure

"(2) Internal pressure loading

(3) External pressure loading

(4) 30 foot end drop on bottom (Rear Impact Limiter)

(5) 30 foot end drop on lid (Front Impact Limiter)

(6) Thermal stresses for hot environment at 100°F ambient temperature

(7) Thermal stresses for cold environment at -20'F ambient temperature

(8) Thermal stresses for cold test at -40'F ambient temperature

(9) Thermal accident condition

(10) Fabrication stress

(11) 1G in the longitudinal direction with the cask axis horizontal, held at the rear trunnions and

supported at the front saddle

Rev. 0 4/992.10.1-3

Since the individual load cases are linearly elastic, their results can be scaled and superimposed as required in order to perform the normal and hypothetical accident condition load combinations. The load combination approaches for these cases are described in Sections 2.6 and 2.7. The magnitudes of the loads used in each individual load case analysis are computed as described in the following paragraphs:

Detailed stresses and displacements in the ANSYS model of the cask body are obtained and stored (on magnetic tape) for every node location for each individual load case. These stored results are postprocessed to printout the stresses at the 39 standard locations on the cask body structure shown in Figures 2.10.1-22 and 2.10.1-23. The locations selected as shown in Figures 2.10.1-22 and 2.10.1-23 are key points that, when carefully studied, indicate the behavior of the entire structure. The maximum stress may occur at a different location for each individual load.

1. Bolt Preload and Lid Seating Pressure

A lid bolt preload corresponding to 86,000 psi direct stress (actual stress is 56,000 psi) in the bolt shank is simulated by specifying an initial strain in the elements representing the bolts. A portion of this strain becomes elastic preload strain in the bolts, and a portion becomes strain in the clamped parts. The required initial strain value of 0.00331 in/in (in the bolts) was determined by trial and error.

The selected bolt preload is sufficient to insure a full seating of the metallic seals under a maximum design internal pressure of 100 psig. The metallic seal seating load is 1,399 lb./in./seal(8 ) or 2,798 lb./in. for 2 seals. This load is simulated by applying a pressure of 3,498 psi on an annular ring on both the containment lid and flange surfaces as shown in Figure 2.10.1-6. The stress intensities from the ANSYS run at the selected locations of the containment vessel and gamma shield are presented in Tables 2.10.1 -1 and 2.

2. Internal Pressure Loading

An internal pressure of 100 psig is applied to the cavity surface as shown in Figure 2.10.1-7. The stress intensities from the ANSYS run at the selected locations of the containment vessel and gamma shield are presented in Tables 2.10.1-5 and 6.

3. External Pressure Loading

An external pressure of 25 psig is applied to the outer surface of the cask body as shown in Figure 2.10.1-8. The stress intensities from the ANSYS run at the selected locations of the containment vessel and gamma shield are presented in Tables 2.10.1-7 and 8.

4. 30 Foot End Drop on Bottom (Rear Impact Limiter)

The dynamic analysis described in Appendix 2.10.8 determined the inertial loads on the TN-68 packaging for a 30 foot end drop onto an unyielding surface. That analysis concluded that the maximum axial deceleration is 66 G for this case. A static elastic analysis of the cask is performed for a unit load (1 G) with inertial forces balanced by the

Rev. 0 4/992.10.1-4

impact force. The results of this elastic analysis will be ratioed up for the actual G load in

the load - combination runs. Since the payload or cargo and the impact limiters are not

included in the model, their loading effects are simulated as distributed pressures applied

on the cask at the appropriate locations. The contacting impact limiter force on the cask is

applied as the reaction pressure on the bottom required to balance the inertial forces of the

system. Thus, the cask body vessel is in equilibrium under the applied forces. The system

of forces on the cask body is presented on Figure 2.10.1-9.

Following is the derivation of the inertia load (pressure) magnitudes for the ANSYS model

run:

"O Weight of Cask: 162,760 lb.

"O Weight of Front Impact Limiter (with spacer): 16,500 lb.

"O Weight of Rear Impact Limiter: 15,450 lb.

"O Weight of Internals: 77,240 lb.

F = 162,760 lb. resultant of all of the distributed inertia forces acting on the cask body.

This force is actually distributed throughout the model and is simulated by applying 1G

acceleration to the finite element model in the longitudinal direction.

Pressure due to internals, P, = 77,240/(mr x 35.52) = 19.509 psi

Pressure due to front impact limiter, PF = 16,500/(rt x 42.252) = 2.942 psi

Reaction pressure due to cask body, internals and front impact limiter,

PR = (162,760 + 77,240 + 16,500)/( 7t x 42.252) = 45.739 psi

The stress intensities from the ANSYS run at the selected locations of the containment

vessel and gamma shield are presented in Tables 2.10.1-9 and 10.

5. 30 Foot End Drop on Lid (Front Impact Limiter)

An analysis similar to that for the 30 foot free drop on the bottom is performed for the 30

foot drop on the lid. The same inertial forces (1 G) are used for the lid or front impact case

as for the rear impact case. The results of this elastic analysis will be ratioed up for the

actual G load in the load - combination runs. The system of forces on the cask body is

presented on Figure 2.10.1-10, and the derivation of the magnitudes follows:

"O Weight of Cask: 162,760 lb.

"o Weight of Front Impact Limiter (with spacer): 16,500 lb.

"O Weight of Rear Impact Limiter: 15,450 lb.

" Weight of Internals: 77,240 lb.

Rev. 0 4/992.10.1-5

F = 162,760 lb. resultant of all of the distributed inertia forces acting on the cask body. This force is actually distributed throughout the model and is also simulated by applying 1G acceleration to the finite element model in the longitudinal direction.

Pressure due to internals, P1 = 77,240/(it x 34.182) = 21.045 psi

Pressure due to rear impact limiter, PR = 15,450/(7t x 42.252) = 2.755 psi

Reaction pressure due to cask body, internals and front impact limiter,

PF = (162,760 +77,240 + 15,450)/( It x 42.252) = 45.5515 psi

The reaction forces are slightly different for the drop on the lid end compared to the drop on the bottom end because of the slight geometry and configuration differences of the model at each end.

The stress intensities from the ANSYS run at the selected locations of the containment vessel and gamma shield are presented in Tables 2.10.1-11 and 12.

6. Thermal Stress for Hot Environment Condition at 100°F Ambient Temperature

The thermal analysis of the cask body is described in Chapter Three. The thermal model is used to obtain the steady state metal temperatures in the cask body for the normal condition which includes 1000 F daily averaged ambient air temperature, maximum decay heat and maximum solar heat loading. The thermal stress evaluations were conservatively based on an outside gamma shield temperature of 301'F and an inside cavity temperature of 256°F (temperature differential of 45°F). The actual temperature difference is less than 10'F from the thermal analysis presented in Chapter 3. Therefore the thermal stresses calculated here are conservative. It is assumed that there is a stress free state at 70'F uniform temperature.

The stress intensities from the ANSYS run at the selected locations of the containment vessel and garmma shield are presented in Tables 2.10.1-13 and 14.

7. Thermal Stresses for Cold Environment Condition at -20'F Ambient Temperature

A uniform temperature of -20'F is input to the structural model to calculate the stresses due to differential thermal expansion between the two shells of the cask. This differential expansion is entirely due to the difference in expansion coefficients of the construction materials of the cask assembly.


Rev. 0 4/992.10.1-6

8. Thermal Stresses for Cold Test at -40°F Ambient Temperature

A uniform temperature of -40°F is input to the structural model to calculate the stresses

due to differential thermal expansion between the two shells of the cask. Stresses occur

due to the difference in coefficients of expansion of the cask construction materials.



9. Thermal Accident Condition

An ANSYS transient thermal analysis of the cask for the 30 minute thermal accident is

reported in Chapter Three. The initial condition is steady state at 100°F ambient

conditions with maximum decay heating. The initial steady state condition is followed by

a 0.5 hour severe thermal transient which is then followed by a cool-down period. The

temperatures through the cross section of the cask at the time which result in the

maximum thermal gradient are used for input to the cask model for thermal stress analysis.


vessel and gamma shield are presented in Tables 2.10.1- 39 and 40.

10. Fabrication Stress

The fabrication stresses are computed due to 0.03 in. shrink-fit (diametrically) between

the inner containment cylinder and outer gamma cylinder. This interference results in an

interface pressure of 403.4 psi between the outer surface of the inner shell and inner

surface of the gamma shell. The two-dimensional axisymmetric finite-element model

was modified by removing all the couplings between the inner and outer cylinders. A run

was made by applying a pressure of 403.4 psi to the outer surface of inner containment

vessel and the inner surface of outer gamma shield cylinder.



11. 1G Unit Loading (Longitudinal)

This analysis is conducted to evaluate the stresses while the cask axis is horizontal and

the rear of the cask is supported at the trunnions and the front of the cask is support at the

saddle. The loading used is the unit (IG) longitudinal deceleration resulting from

braking.

The results of this elastic analysis will be ratioed up for the actual G load (2.3G) in the

load- combination runs. Since the internals and impact limiters are not included in the

model , their loading effects are simulated as distributed pressure applied on the vessel at

"the appropriate locations. The system of forces acting on the cask is presented in Figure

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2.10.1-11. The node in the finite element model corresponding to the rear trunnion location is held in the Y- direction (longitudinal direction).

The forces acting on the cask in this case are:

F = 162,760 lb. resultant of all of the distributed inertia forces acting on the cask body. This force is actually distributed throughout the model and is simulated by applying IG acceleration in the longitudinal direction.

Pressure due to internals, P, = 77,240/(0t x 34.182) =21.045 psi

Weight of both impact limiters (including spacer) = 15,450 + 16,500 = 31,950 lb.

Pressure due to both impact limiters, PF+R= 31,950 /(7t x 42.252) = 5.697 psi

Since the two impact limiters are connected by tie rods, the entire load shifts to the rear cask surface.


2.10.1.4 Asymmetric Loading

The asymmetric loadings to the axisymmetric cask body are applied by special ANSYS harmonic elements. Each load acting on the cask is expanded into a Fourier series and is input into ANSYS as a series of load steps. Each load step contains all of the terms from the applied loads having the same mode number. The number of terms in the Fourier series, required to adequately represent a load varies with the type of load (whether it is a concentrated or a distributed load) and the degree of accuracy required. In the particular case where the applied loads are distributed over a large area (i.e., 180 degrees of the cask circumference), three terms of the series are sufficient to represent the desired loading within a few percent. See Figure 2.10.1-12 which shows the Fourier series approximations of cosine functions acting on the arc from 900 to 2700.

The following individual asymmetric load cases analyzed (using the same two-dimensional ANSYS model previously discussed) are described in this section.

(1) 1G downward loading with the cask axis horizontal, supported vertically on front saddle and two rear trunnions.

(2) 30 foot side drop with the cask axis parallel to the target.

(3) 30 foot CG over bottom corner (rear impact limiter) drop.

(4) 30 foot CG over lid corner (front impact limiter) drop.

Rev. 0 4/992.10.1-8

(5) 30 foot 150 slap down (lid end).

Since the individual load cases are linear and elastic, their results can be scaled and

superimposed as required in order to perform the normal and hypothetical accident load

combinations. Figure 2.10.1-13 illustrates the locations of the CG, trunnions, saddle, and impact

limiters. The magnitudes of the loads used in each individual load case analysis are computed in

the following paragraphs:

1. 1G Unit Loading (Vertical Down)

A static elastic analysis of cask is performed for a unit vertical down load (1 G) inertial

forces on horizontal cask supported at front saddle and rear trunnions. The system of

forces acting on the cask is presented in Figure 2.10.1-14. The results of this elastic

analysis will be ratioed up for the actual g load in the load- combination runs. Since the

internals and impact limiter are not included in the model, their loading effects are

simulated as distributed pressure applied on the vessel at the appropriate locations. The

saddle reaction pressure required to balance the inertial forces is applied at the vessel

bottom. The rear trunnion node is supported in the Z direction. The pressures due to

internals, impact limiters and saddle are assumed to vary as a cosine function over 180

degrees.

The forces acting on the cask in this case are:

A. Cask Body Inertia

F = 162,760 lb. downward resultant of the distributed inertia force acting on the mass of

the cask (shown at the C G) and is simulated by applying 1G acceleration in X-direction

(normal direction).

B. Pressure due to internals:

The downward force acting on the lower half of the inside surface of the cavity due to the

internals which are represented as a pressure varying sinusoidally around the bottom half (90' to 270' range) of the cavity as shown below:

3,t/2 WI P1 (cos 0) (L) cos 0 R dO = (PI) (R) (L) (tC/2)

Therefore, P1 = 2W/(ic RL)

The peak pressure of the internals is calculated as follows:

W, = weight of internals = 77,240 lb.

"Internals peak pressure, P1= 2Wi/(7L RL) = 2 x 77,240/(3.142 x 34.75 x 178.5) = 7.927

psi.

Rev. 0 4/992.10.1-9

This peak pressure is assumed to vary as a cosine function over 180 degrees (900 to 2700 range). The computer inputs are based on the pressures calculated by using three terms of the fourier series coefficients, and are:

Term MODE ISYM Fourier Peak pressure Input pressure coefficient (psi) to be used

1 0 1 0.3182 2.522 2 1 1 -0.50 7.927 -3.964 3 2 1 0.2124 1.684

C. Pressures due to Impact Limiters

PR, PF1, and PF2 are downward pressures applied by the rear and front impact limiters to the outer surfaces of the cask. These pressures are assumed to vary sinusoidally around the top half (270 to 90-) range of the surfaces and are calculated as follows:

Rear impact limiter peak pressure, PR = 2 x 15,450/(3.142 x 42.25 x 12.75) = 18.259 psi

The front (lid side) impact limiter weight (16,500 lb.) is divided in the ratio of two lengths; 5.0 in (R = 39.94") and 6.25 in. (R = 42.25").

WFI = 16,500 x 5/11.25 = 7,333 lb. WF2 = 16,500 x 6.25/11.25 = 9,167 lb.

Front impact limiter peak pressure, PF1 = 2 x 7,333/(3.142 x 39.94 x 5.0) = 23.377 psi Front impact limiter peak pressure, PF2= 2 x 9,167/(3.142 x 42.25 x 6.25) = 22.100 psi

D. Reaction forces at the Supports

Total package weight, W = weight of the package = 271,950 lb.

Trunnion reaction, VT = 150,286 lb (conservatively using higher reaction force at trunnion)

FT is the tangential line load simulating the upward reactions at the rear trunnions. The line load, FT, is related to the total reaction force VT (VT = IrFT.).

Saddle reaction, Vs = 271,950 -150,286 = 121,664 lb. (159,444 lb is conservatively used for calculation).

Rev. 0 4/992.10.1-10

The saddle reaction is applied as a pressure applied to the outer surface of the cask, this

pressure is assumed to vary sinusoidally around the bottom half (900 to 2700 range) of the

surface (8.85" width). The peak pressure is calculated as follow:

Peak Pressure, Ps = 2 x 159,444/(3.142 x 42.25 x 8.85) = 271.59 psi



2. 30 Foot Side Drop

A static elastic analysis of the cask is performed for a unit load (1 G) with inertial forces

balanced by impact forces. The results of this elastic analysis are ratioed up for the actual

g load in the load- combination runs. Since the internals and impact limiter are not

included in the model, their loading effects are simulated as distributed pressure applied

on the vessel at the appropriate locations. The impact limiter reaction pressure required

to balance the inertial forces is applied at the vessel bottom. During the side drop, the

pressure at inner surface due to internals and reaction pressure on the outer side by the

impact limiters is assumed to vary as a cosine function over 180 degrees. The system of

forces acting on the cask is presented in Figure 2.10.1-15.

The forces acting in this case are:


F = 162,760 lb. downward resultant of the distributed inertia force acting on the cask

(shown at the CG) and is simulated by applying 1G acceleration in X-direction (normal

direction).

B. Pressure Due to Internals

The downward force acting on the lower half of the inside surface of the cavity due to the

internals which are represented as a pressure varying as a cosine function around the

bottom half (90' to 2700 range) of the cavity as shown below:

W1 =f P1 (cos 0) (L) cos 0 R dO = (PO) (R) (L) (ic/2) I it/2

Therefore, P1 = 2WV/(ic RL)

The peak pressure of the internals is calculated as follows:

WI = weight of internals = 77,240 lb.

Peak pressure, P1 = 2Wi/(nt RL) = 2 x 77,240/(3.142 x 34.75 x 178.5) = 7.927 psi

Rev. 0 4/992.10.1-11

C. Impact Reaction Pressures:

PR, PFI, and P1: are the peak pressures applied by the rear and front impact limiter reactions on the bottom half of the cask body outer surface during impact. These pressures are assumed to vary sinusoidally around the bottom half of the outer surfaces (900 to 2700 range) and are calculated as follows:

Total cask weight, W = 240,000 lb (cask + internals)

Reaction force, Rear (bottom) = 240,000 x 94.88/185.24 = 122,928 lb.

Reaction force, Front (lid) = 240,000 x 90.36/185.24 = 117,072 lb.

Rear Peak Pressure, PR = 2 x 122,928/(3.142 x 42.25x 12.75) = 145.276 psi

The front (lid side) reaction force is divided in the ratio of two lengths; 5.Oin (R = 39.94") and 6.25 in. (R = 42.25").

RFI = 117,072 x 5/11.25 = 52,032 lb. R2 = 117,072 x 6.25/11.25 = 65,040 lb.

Rear Peak Pressure, PFI = 2 x 52,032/(3.142 x 39.94 x 5.0) = 165.872 psi Rear Peak Pressure, Pm = 2 x 65,040/(3.142 x 42.25 x 6.25) = 156.803 psi

The stress intensities from the ANSYS run at the selected locations of the containment vessel and gamma shield are presented in Tables 2.10.1-23 to 26.

3. 30 Foot C.G. Over Bottom Comer Drop (Rear Impact Limiter)

A static elastic analysis of the cask is performed for a unit load (1 G) with inertial forces balanced by impact forces. The results of this elastic analysis are ratioed up for the actual g load in the load- combination runs. Since the internals and impact limiter are not included in the model, their loading effects are simulated as distributed pressure applied on the vessel at the appropriate locations. For CG-over-comer, the cask is inclined at about 60 degrees with the horizontal. All the applied loads and reaction forces are transformed in axial and normal components. The axial pressure components due to internals, top (front) impact limiter and bottom (rear) impact reaction are assumed uniformly distributed. All the normal pressure components are assumed to vary as cosine functions over 180 degrees of the circumference. The system of forces acting on the cask is presented in Figure 2.10.1-16.

Rev. 0 4/992.10.1-12



F = 162,760 lb. downward resultant of the distributed inertia force acting on the cask

(shown at the C G) and is simulated by applying 0.5G in the normal direction and 0.866G

in the longitudinal direction (resulting 1 G in the vertical direction).


The internals inertia loading was applied in two mutually perpendicular directions (one

along the axis of the cask and the other perpendicular to it). The component along the

axial direction (PtA) was distributed uniformly over the inside surface of the bottom plate.

The other component (PIN) was assumed to vary sinusoidally around the lower half (90*

to 2700 range) of the inside surface of the inner shell; the fourier coefficients for this

function are computed in Figure 2.10.1-12.

WI = weight of internals = 77,240 lb.

Axial pressure, PtA = 77,240 sin 600/(3.142 x 35.52) = 16.895 psi

Peak normal pressure, PIN= 2Wtcos 600/(it RL) = 2 x 77,240x 0.5 /(3.142 x 34.75 x

178.5) = 3.964 psi

C. Pressure Due to Top (Front) Impact Limiter

The inertia load of the nonstriking impact limiter was also applied to the cask in two

mutually perpendicular directions. The axial component (PFA) was applied as a uniform

pressure over the outside surface at the interface with the impact limiter on the lid end.

The other component (Pm) was assumed to vary sinusoidally around the upper half of the

outside surface (2700 to 90 range) of the cask.

WF = weight of front impact limiter = 16,500 lb.

Axial pressure, PFA =16,500 sin 600/ (3.142 x 34.942) = 2.8513 psi

Normal pressure is applied on two lengths; 5.0 in (R = 39.94") and 6.25 in. (R = 42.25").

F1 = 16,500 cos 600 x 5/11.25 = 3,667 lb. F2 = 16,500 cos 600 x 6.25/11.25 = 4,583 lb.

Peak Pressure due to F 1, PNI = 2 x 3,667/(3.142 x 39.94 x 5.0) = 11.690 psi

Peak Pressure due to F2, PF2 = 2 x 4,583/(3.142 x 42.25 x 6.25) = 11.049 psi

Rev. 0 4/992.10.1-13

D. Reaction Pressures Due to Bottom (Rear) Impact Limiter

The reaction pressure of the striking impact limiter was also applied to the cask in two mutually perpendicular directions. The axial component (PRA) was applied as a uniform pressure over the outside surface at the interface with the impact limiter on the bottom end. The other component (PRN) was assumed to vary sinusoidally around the lower half of the outside surface (900 to 2700 range) of the cask.

Total weight = 162,760 + 77,240 + 16,500 = 256,500 lb.

Axial pressure, PRA =256,500 sin 600 /(3.142 x 42.252) = 39.6097 psi

Peak normal pressure, PRN= 2W cos 600 /(i RL) =2 x 256,500 x 0.5/(3.142 x 42.25 x 12.75) = 151.556 psi.


4. 30 Foot CG Over Lid Comer Drop (Front Impact Limiter)

A static elastic analysis of the cask is performed for a unit load (1 G) with inertial forces balanced by impact forces. The results of this elastic analysis are ratioed up for the actual G load in the load - combination runs. Since the internals and impact limiter are not included in the model, their loading effects are simulated as distributed pressure applied on the vessel at the appropriate locations. For CG-over-comer, the cask is inclined at about 60 degrees from the horizontal. All the applied loads and reaction forces are transformed in axial and normal components. The axial pressure components due to internals, top impact limiter and bottom impact reaction are assumed uniformly distributed. All the normal pressure components are assumed to have cosine variation over 180 degrees circumference. The system of forces acting on the cask is presented in Figure 2.10.1-17.



F = 162,760 lb. downward resultant of the distributed inertia force acting on the mass of the cask (shown at the CG) and is simulated by applying 0.5G in the normal direction and 0.866G in the longitudinal direction (resulting IG in the vertical direction).

B. Pressure Due to Internals:

The internals inertia loading was applied in two mutually perpendicular directions (one along the axis of the cask and the other perpendicular to it). The component along the axial direction (Pts) was distributed uniformly over the inside surface of the bottom plate.

Rev. 0 4/992.10.1-14

The other component (PIN) was assumed to vary sinusoidally around the lower half of the inside surface (90' to 2700 range) of the inner shell.

W= weight of internals = 77,240 lb.

Axial pressure, PLA =77,240 sin 600 /(3.142 x 35.52) = 16.895 psi

Peak normal pressure, PN= 2W, cos 600 /(ir RL) = 2 x 77,240x0.5 /(3.142 x 34.75 x 178.5) = 3.964 psi.

C. Pressure Due to Rear Impact Limiter:

The inertia load of the nonstriking impact limiter was also applied to the cask in two mutually perpendicular directions. The axial component (PRA) was applied as a uniform pressure over the outside surface at the interface with the impact limiter on the bottom end. The other component (PRN) was assumed to vary sinusoidally around the upper half of the outside surface (2700 to 90' range) of the cask.

WR = weight of rear limiter = 15,450 lb.

Axial pressure, PRA =15,450 sin 60'/(3.142 x 42.252) = 2.386 psi Peak normal pressure, PRN = 2 x 15,450 cos 600/(3.142 x 42.25 x 12.75) = 9.129 psi

D. Reaction Pressures Due to Front Impact Limiter

The reaction pressure of the striking impact limiter was also applied to the cask in two mutually perpendicular directions. The axial component (PFA) was applied as a uniform pressure over the outside surface at the interface with the impact limiter on the lid end. The other component (PFN) was assumed to vary sinusoidally around the lower half of the outside surface (90' to 270' range) of the cask.

Total weight = 162,760 + 77,240 + 15,450=- 255,450 lb.

Axial pressure, PFA = 255,450 sin 600 / (3.142 x 42.252) = 39.4476 psi

Normal pressure is applied on two lengths; 5.0 in (R = 39.94") and 6.25 in. (R = 42.25").

F1 = 255,450 cos 600 x 5/11.25 = 56,767 lb. F2 = 255,450 cos 600 x 6.25/11.25 = 70,958 lb.

Peak Pressure due to F1, PFN1 = 2 x 56,767/(3.142 x 39.94 x 5.0) = 180.966 psi Peak Pressure due to F2, PFN2 = 2 x 70,958/(3.142 x 42.25 x 6.25) = 171.070 psi


Rev. 0 4/992.10.1-15

5. 30 Foot 150 Slat Down (Lid End)

The maximum transverse G loads predicted by the ADOC computer program (Appendix 2.10.8) resulting in the maximum.impact force are based on combining the normal and rotational accelerations. The following table presents the G loads from the ADOC runs at time of second impact.

G Load From ADOC Run - 30 Foot 150 Slap Down (Lid End Impact)

Gaxial Gaxia = -0.045 (<-) Gnormal at CO Gnori• = -25.4(,) Gat front impact limiter Gnormal = -25.4 (,4)

Grotational = 46.8 (j-) Gat rear impact limiter Gnormal = 19 (1M)

These G loads are also depicted in Figure 2.10.1-18. For stress analysis of the cask body, the maximum combined normal and rotational acceleration of 72.2 G is conservatively increased to 120 G. The G loads used in the ANSYS finite element model analysis are listed in the following table and also and also depicted in Figure 2.10.1-19.

G Load used For 30 Foot 150 Slap Down Structural Analysis

Gaxiaj Gaiai= -0.075 (<--)

Gnormal at CG Gnormal = -42.22 (,!) Gat front impact limiter Gnormal = -42.22 (4)

Grotational = 77.78 (f) Gat rear impact limiter Gnormai = 31.86 (1")

The system of forces acting on the cask is presented in Figure 2.10.1-20. The forces acting on the cask in this case are:


The axial (0.075G) and lateral (42.22G) inertia loads of the cask are applied as inertia (body) loads in the finite element model.


Rev. 0 4/992.10.1-16

The cask contents inertia load was again resolved into components acting in two mutually

perpendicular directions. The component in the axial direction was applied as a pressure

acting on the inside surface of the lid. The lateral component was applied as a pressure

acting over the lower half (90' to 2700 range) of the inside surface of the inner shell. In this

case, these pressures are not only varied sinusoidally around the circumference but also

varied linearly with distance from the center of the surface to which they are applied since

the containment vessel is being subjected to rotational as well as axial and lateral

accelerations.

Wi = weight of internals = 77,240 lb.

Peak normal pressure, P1= 2W1 x 120/(ic RL) = 2 x 77,240x 120/(3.142 x 34.75 x 178.5)

= 951.28 psi

P2 = 951.28 x 97.24/120 = 770.85 psi

P 3 = 951.28 x 55.96/120 = 443.61 psi

P4 = 951.28 x 18.72/120 = 148.4 psi

P5 = 951.28 x 7.81/120 = 61.91 psi

P6 = 951.28 x 23.74/120 = 188.19 psi

P1, P 2, P3, and P 4 are acting over the lower half (900 to 2700 range) of the inside surface of

the inner shell. P5 and P 6 are acting over the upper half (2700 to 900 range) of the inside

surface of the inner shell. The computer inputs are based on the pressures calculated by

using three terms of the fourier series coefficients; the fourier coefficients for this

function are computed in Figure 2.10.1-12.

Axial pressure, P9 =77,240 x 0.075 /(3.142 x 35.52) = 1.46 psi

C. Pressure Due to Rear Impact Limiter:

The inertia load of the rear impact limiter is also applied to the cask in two mutually

perpendicular directions. The axial component (P7) is applied as a uniform pressure over

the outside surface at the interface with the impact limiter on the bottom end. The other

component (Ps) is assumed to vary sinusoidally around the upper half of the outside

surface (2700 to 900 range) of the cask.

WR = weight of rear limiter = 15,450 lb.

Axial pressure, P7 =15,450 x 0.075/(3.142 x 42.252) = 0.21psi

Peak normal pressure, P8 = 2 x 15,450 x 42.22/(3.142 x 42.25 x 12.75) = 770.89 psi

Rev. 0 4/992.10.1-17

D. Reaction Pressures Due to Front Impact Limiter

The reaction pressure of the front impact limiter is also applied to the cask in two mutually perpendicular directions. The axial component (PIo) is applied as a uniform pressure over the outside surface at the interface with the impact limiter on the lid end. The other component (P1 ) is assumed to vary sinusoidally around the lower half of the outside surface (90' to 2700 range) of the cask.

Total weight = 162,760 + 77,240 + 15,450=- 255,450 lb.

Axial pressure, P10 = (255,450) x 0.075/ (3.142 x 42.252) = 3.42 psi

Peak normal pressure, P11= 2 (255,450) x 42.22/(3.142 x 42.25 x 11.25) = 4,445.25 psi.

E. Rotational Acceleration

A rotational acceleration, a, is also input to the finite element model to satisfy the rotational quasistatic equilibrium condition (see Figure 2.10.1-21):

ZM=0

The rotational acceleration, a, required to place the model in equilibrium was calculated Based on:

ZM=I, a=Fd, + F2 d2 +F 3 d3 + ........

Where

F1, F2, etc. are the various inertial loadings in the system of forces.

L= Mass moment of inertia I M = Net moment resulting from axial and normal loads on the finite element model.


2.10.1.5 Transport Shock Loading

Transport By Truck

The transport shock loadings used to evaluate the TN-68 transport cask are based on truck bed accelerations in ANSI N 14.23 which are:

Vertical 3.5G Longitudinal 2.3G Lateral 1.6G

Rev. 0 4/992.10.1-18

"1-1 The resultant transverse load is (3.52 + 1.62)12 = 3.85 G

The transport shock load stresses are obtained by multiplying the IG transverse load (Tables 2.10.1-21 and 22) and 1G longitudinal load (Tables 2.10.1-41 and 42) stresses by 3.85 and 2.3 factors, respectively and summing them. This operation is conducted using an ANSYS postprocessor.

The stress intensities from the ANSYS run at the selected locations of the containment vessel and gamma shield are presented in Tables 2.10.1- 19 and 20.

Transport By Rail

The transport shock loadings used to evaluate the TN-68 transport cask are based on NUREG 766510(9) which specifies a maximum inertia loading of 4.7G in each of the three x-y-z coordinate directions:


The resultant transverse load is (4.72 +4.72)112 = 6.65 G

The transport shock load stresses are obtained by multiplying the 1G transverse load (Tables 2.10.1-21 and 22) and 1G longitudinal load (Tables 2.10.1-41 and 42) stresses by 6.65 and 4.7 factors, respectively and summing them. This operation is conducted using an ANSYS postprocessor.


2.10.1.6 Transport Vibration Loading

Transport By Truck

The input loading conditions used to evaluate the TN-68 cask for transport vibration are also obtained from truck bed accelerations in ANSI N14.23. The peak inertia values used are:


The resultant transverse load is (0.62 +0.32)112 =0.67 G

__ The transport vibration load stresses are obtained by multiplying the IG transverse load (Tables 2.10.1-21 and 22) and IG longitudinal load (Tables 2.10.1-41 and 42) stresses by 0.67 and 0.3

Rev. 0 4/992.10.1-19

factors, respectively and summing them. This operation is conducted using an ANSYS postprocessor.


Transport By Rail

The input loading conditions used to evaluate the TN-68 cask for transport vibration are obtained from NUREG 766510. The peak inertia values used are:


These values are less than the truck vibration loadings. Therefore, the stresses obtained from the truck vibration analysis are used to evaluate the combined stresses for transport inertia load combination.

2.10.1.7 Transport Tie-Down Loading

The input loading conditions used to evaluate the TN-68 cask for transport tie-down loadings are obtained from 10CFR71.45. The peak inertia values used are:

Vertical 2 G Longitudinal 10 G Lateral 5 G

The resultant transverse load is (22 + 52)112 =5.4 G

The transport tie-down load stresses are obtained by multiplying the 1G transverse load (Tables 2.10.1-21 and 22) and 1G longitudinal load (Tables 2.10.1-41 and 42) stresses by 5.4 and 10 factors, respectively and summing them. This operation is conducted using an ANSYS postprocessor.


2.10.1.8 6 G Lifting

The cask is oriented vertically in space and held by the 2 top trunnions. The down load is simulated by applying 6G vertical acceleration to the finite element model. Since the resin, outer shell and trunnions are not included in the model, they are accounted for by increasing the density of the gamma shielding. Figure 2.10.1-24 shows the loading conditions.

Rev. 0 4/992.10.1-20

The internals are not included in the model, their loading effects are simulated by a distributed pressure acting on the inside bottom surface of the cask cavity.

The total cask weight (including internals) is replaced by forces applied to the 2 top trunnions so

that the system of forces acting on the cask is again in equilibrium. A cask weight of 240,000 lb.

is used in the calculations (weight of impact limiters is not included). The two trunnion forces

FTR are replaced by a total force:

Fy = 6.0 x (240,000) x 1.15 = 1,656,000 lbs

A 15% additional load is included to cover the dynamic effects of lifting. This force is acting in

the Y direction on the outer surface of the gamma shielding at the trunnion location. The stress intensities from the ANSYS run at the selected locations of the containment vessel

and gamma shield are presented in Tables 2.10.1- 47 and 48.

2.10.1.9 Summary of Individual Load Cases

Detailed stresses and displacements in the ANSYS model of the cask body are obtained and stored (on magnetic tape) for every node location for each individual load case. These stored

results are postprocessed to printout the stresses at the 39 standard locations on the cask body

structure shown in Figures 2.10.1-22 and 2.10.1-23. The locations selected as shown in Figures

2.10.1-22 and 2.10.1-23 are key points that, when carefully studied, indicate the behavior of the

entire structure. The maximum stress may occur at a different location for each individual load.

The individual load cases analyzed are listed in Table 2.10.1-1A. Linear elastic analyses were

performed for all load cases. Stress intensities at nodal locations on the inner and outer surfaces

of each cask body component are reported in Tables 2.10.1-1 to 2.10.1-48, and Tables 2.10.1-61to 2.10.1-62 as listed in Table 2.10.1-lA. There are no specific limits for individual

stress components. The stress components of each load case are combined for the various load combinations as described in Sections 2.6 and 2.7 by factoring and algebraic addition. Then the

stress intensities at each location are determined, classified and compared with the design criteria of Section 2.1.2 in Sections 2.6 and 2.7.

The local stresses at the trunnion locations are not included in these results. The local stresses

are obtained as described below in Section 2.10.1.8 and they are reported in Tables 2.10.1-49

through 2.10.1-60. The method used is the "Bijlaard" analysis using hand calculations rather

than an ANSYS model.

It should be noted that, for the axisymmetric analyses, the stress is constant around the cask at

every location. For asymmetric analyses with significant differences in stress magnitudes on the

extreme opposite sides of the cask, the stresses at locations on different sides of the cask are

reported. In those cases where the cask is supported on the trunnions, the stresses under

transverse loadings (such as gravity) are nearly equal in magnitude on the top and bottom.

Therefore only the tensile results for the bottom side are listed for these cases.

Rev. 0 4/992.10.1-21

In order to check the reasonableness of the finite element model response, some simple closedform calculations are conducted. While these simple results are unlikely to duplicate the complex area of model and complex loading conditions, they can be used to verify the stresses in simple areas away from discontinuities.

"* Internal Pressure (100 psig)

P*rm P Membrane stress intensity in a cylinder - + P

t 2

= [(100 * 38.5/7.5) + (100/2)] = 513 psi

Average of stress intensities at locations 11 and 32 from computer output at Tables 2.10.1-5 and 2.10.1-6 = ½ (586 + 417) = 502 psi. This is close to the hand-calculated stress intensity.

"• Normal Thermal Condition

E'a'AT Thermal stress in a cylinder = 2(1

AT between locations 11 and 32 = 301 - 256 = 45'F Thermal stress = (28.3E6*6.43E-6*45)/(2*0.7) = 5,849 psi

Average stress intensity at locations 11 and 32 from computer output (Tables 2.10.1-13 and 2.10.1-14) = 1/2 (5599+5435) = 5517 psi. This is close to the hand-calculated intensity.

The above comparison indicates that the finite element response to various simple loads is reasonable.

2.10.1.10 Trunnion Local Stress Analysis

This section discusses the analysis performed to calculate the local stresses in the cask body outer shell at the trunnion locations due to the loadings applied through the trunnions. These local effects are not included in the ANSYS stress results tables reported above. The local stresses must be superimposed on the above stress results for the cases where the inertial loadings are reacted at the trunnions. The local stresses are calculated in accordance with the methodology of Reference WRC-107(7 ) which is based on the "Bijlaard" analysis for local stresses in cylindrical shells due to external loadings.

Loading

The Bijlaard analysis was performed for several different trunnion loading conditions to support various structural evaluations. A summary of the load cases considered is provided in Table 2.10.1-lB.

Rev. 0 4/992.10.1-22

Loads used for the "Bijlaard" analysis are obtained from the packaging inertial loadings and are

also listed in Table 2.10.1-lB. The loads are based on a transport weight of 271,950 lb except

for lifting loads which use a weight of 240,000 lb (which does not include the impact limiter

weight).

Method of Analysis

The local stresses induced in the outer shell cylinder by the trunnions are calculated using

"Bijlaard's" method. The neutron shield and thin outer shell are not considered to strengthen

either the trunnions or the gamma shield shell. The trunnion is approximated by an equivalent

attachment so that the curves of the Reference WRC-107 can be used to obtain the necessary

coefficients. These resulting coefficients are inserted into blanks in the column entitled "Read

Curves For," in a standard computation form, a sample of which is attached as Table 2.10.1-59.

The stresses are calculated by performing the indicated multiplication in the column entitled

"Compute Absolute Values of Stress and Enter Result". The resulting stress is inserted into the

stress table at the eight stress locations, i.e., AU, AL, BU, BL, etc. Note that the sign convention

for this table is defined on the figure as if the load directions are as shown. The membrane plus

bending stresses are calculated by completing Table 2.10.1-59.

Model, Boundary Conditions and Assumptions

The cylindrical body is assumed to be a hollow cylinder of infinite length. Since the trunnions

are located away from the ends of the cylinder, this assumption is acceptable because the local

effects are not significantly affected by the end restraints, i.e., lid and bottom. This is

conservative since end restraints would reduce the local bending effects.

Results of Trunnion Local Stress Analysis

Tables 2.10.1-49 to 60 summarize the resulting nodal stress intensity for the various loading

conditions. These local stresses are combined with the finite element results at the same

locations from the individual load cases above and compared with allowables in Section 2.6.

Rev. 0 4/992.10.1-23

2.10.1.11 References

1. 1 OCFR 71.71, Normal Conditions of Transport. 2. 10CFR 71.73, Hypothetical Accident Conditions. 3. Regulatory Guide 7.8, "Load Combinations for the Structural Analysis of Shipping casks

for Radioactive Material". 4. ASME Code Section III, Division 3, Subsection WB, Containment Systems and Transport

Packagings for Spent Nuclear. 5. Regulatory Guide 7.6, "Design Criteria for the Structural Analysis of Shipping Cask

Containment Vessels". 6. ANSYS Users Manual, Rev. 5.5. 7. WRC Bulletin 107, March 1979, "Local Stresses in Spherical and Cylindrical Shells Due

to External Loadings". 8. High Performance Sealing, Metal Seals Helicoflex Catalog, Helicoflex Co., Boonton, N.J.

ET 507 E 5930. 9. NUREG 766510, "Shock and Vibration Environments for Large Shipping Containers on

Rail Cars and Trucks".

Rev. 0 4/992.10.1-24

Table 2.10.1-1A

Individual Load Cases For TN-68 Cask Body Analysis

Load Individual Load Description Stress Result

Case Tables

Number IL - 1 Bolt Preload and Lid Seating Pressure 2.10.1-1, -2

IL - 2 Fabrication Stress 2.10.1-3, -4

IL - 3 Internal Pressure (100 psi) 2.10.1-5, -6

IL - 4 External Pressure (25 psi) 2.10.1-7, -8

IL - 5 End Drop on Bottom -Rear Impact Limiter (MG) 2.10.1-9, -10

IL - 6 End Drop on Lid - Front Impact Limiter (IG) 2.10.1-11, -12

IL - 7 Thermal Stress Due to Hot Environment 2.10.1-13, -14

IL - 8 Thermal Stress Due to -20°F Cold Environment 2.10.1-15, -16

IL - 9 Thermal Stress Due to -40'F Cold Environment 2.10.1-17, -18

IL - 10 Transport Shock Load/Truck (2.3G Long., 1.6G Lat., 2.10.1-19, -20 3.5G Vert.)

IL - 11 Cask Horizontal, IG Down 2.10.1-21, -22

IL - 12 Side Drop (1G) 2.10.1-23, -24, -25, -26

IL - 13 CG Over Comer Drop on Rear Impact Limiter (IG) 2.10.1-27, -28, -29, -30

IL - 14 CG Over Comer Drop on Front Impact Limiter (IG) 2.10.1-3 1, -32, -33, -34

IL - 15 150 Slap Down Second Impact on Lid End (120G) 2.10.1-35, -36, -37, -38

IL - 16 Thermal Stress Due to Fire Accident 2.10.1-39, -40

IL - 17 Cask Horizontal, 1G Longitudinal 2.10.1-41, -42

IL - 18 Transport Vibration Load/Truck (0.3G Long., 0.3G Lat., 2.10.1-43, -44 0.6G Vert.)

IL - 19 Transport Tiedown Load (10G Long., 5G Lat., 2G Vert.) 2.10.1-45, -46

IL - 20 6G on Trunnion Lifting Load (Cask Vertical, 6GUp) 2.10.1-47, -48

IL - 21 Local Stresses at Rear Trunnion/Cask Body Interface 2.10.1-49, -50 with 1G Down - Cask Horizontal

IL - 22 Local Stresses at Rear Trunnion/Cask Body Interface 2.10.1-51, -52

with Shock Loads/Truck - Cask Horizontal IL - 23 Local Stresses at Rear Trunnion/Cask Body Interface 2.10.1-53, -54

with Vibration Loads/Truck - Cask Horizontal IL - 24 Local Stresses at Rear Trunnion/Cask Body Interface 2.10.1-55, -56

with Tiedown Loads - Cask Horizontal IL - 25 Local Stresses at Front Trunnion/Cask Body Interface 2.10.1-57, -58

with 6G up - Cask Vertical IL - 26 Local Stresses at Rear Trunnion/Cask Body Interface 2.10.1-59, -60

with Shock Loads/Rail - Cask Horizontal

IL - 27 Transport Shock Load/Rail (4.7G All Directions) - Cask 2.10.1-61, -62

Horizontal II

Rev. 0 4/99

Table 2.10.1-4B

Trunnion Loads

Trunnion Load Description Inertial Load Max. Each Trunnion Load

Gravity !G Vertical Vc = 67.99 kips

(Cask Horizontal - Rear Trunnion) Mc - 543.22 in-kips

Shock/Truck 2.3G - Longitudinal VL =312.74 kips

(Cask Horizontal - Rear Trunnion) 1.6G - Lateral Vc = 237.96 kips 3.5G - Vertical P = 21.56 kips

ML = 2498.81 in-kips Mc = 1901.27 in-kips

Vibration/Truck 0.3G - Longitudinal VL = 40.79 kips



Tiedown lOG - Longitudinal VL = 1359.75 kips

(Cask Horizontal - Rear Trunnion) 5G - Lateral Vc = 135.98 kips 2G - Vertical P = 679.88 kips


Lifting 6G- vertical VL = 828.00 kips

(Cask vertical - Front Trunnion) ML= 6218.28 in-kips

Shock/Rail 4.7G - Longitudinal VL = 639.08 kips



Vibration/rail 0.19G - Longitudinal VL = 40.79 kips



Rev. 0 4/99

Table 2.10.1-1

Bolt Preload and Lid Gasket Seating Pressure

Location. Nodal Stress Intensity (psi)

1 0

20 Inner Bottom Plate

3 0 4 0

5 12

6 10

7 11

8 11 9 11

10 1 Inner Shell111

12 1

13 10

14 12

15 39

16 24

17 16

18 38

Rev. 0 4/99

Table 2.10.1-2

Bolt Preload and Lid Gasket Seating Pressure

Location Nodal Stress Intensity (psi)

Flange 19 782

20 757

Lid 21 4

22 173

23 1 Outer Bottom Plate 24 1

25 15

26 1

27 3

28 2

29 3

Outer Shell 30 3

31 3

32 3

33 4

34 1

35 49

36 37

37 1 Weld 38 105

39 234

Rev. 0'4/99

Table 2.10.1-3

Fabrication Stresses


1 2355

2 751 Inner Bottom Plate 3 2355

4 751

5 2192

6 1653

7 9931

8 10076

9 9752

10 9747 Inner Shell1199

11 9749

12 9749

13 9747

14 9741

15 11124

16 9928

17 4326

18 5254

Rev. 0 4/99

Table 2.10.1-4

Fabrication Stresses


Flange 19 8507

20 4085

Lid 21 21

22 112


25 306

26 16

27 2978

28 2225

29 3060

Outer Shell 30 2256

31 3052

32 2250

33 3062

34 2258

35 2839

36 2114

37 44

Weld 38 2186

39 557

Rev. 0 4/99

Table 2.10.1-5

Internal Pressure (100 psi)


1 300

2 135 Inner Bottom Plate

3 238

4 73

5 1411

6 732

7 460

8 453

9 598

10 541 Inner Shell

11 586

12 529

13 587

14 532

15 522

16 415

17 636

18 499

Rev. 0 4/99

Table 2.10.1-6

Internal Pressure (100 psi)


Flange 19 1053

20 711

Lid 21 1392

22 2330

23 1032

Outer Bottom Plate 24 1419

25 2532

26 738

27 296

28 297

29 579

Outer Shell 30 440

31 566

32 417

33 568

34 422

35 521

36 310

37 870

Weld 38 778

39 2478

Rev. 0 4/99

Table 2.10.1-7

External Pressure (25 psi)


1 74 2 15

Inner Bottom Plate

3 61

4 31 5 343

6 163

7 122

8 130

9 150 10 150

Inner Shell

11 147 12 147

13 147

14 148

15 128

16 124

17 158

18 140

Rev. 0 4/99

Table 2.10.1-8

External Pressure (25 psi)


Flange 19 279

20 171

Lid 21 349

22 583


25 629

26 186

27 75

28 73

29 146 Outer Shell 30 111

31 143

32 105

33 143

34 107

35 131

36 79

37 220 Weld 38 189

39 620

Rev. 0 4/99

Table 2.10.1-9

End Drop on Bottom - Rear Impact Limiter (1G)


1 144 2 24

Inner Bottom Plate 3 134

4 39

5 312

6 117

7 102

8 108

9 78

10 79 Inner Shell61

12 61

13 44

14 45

15 37 16 34

17 36

18 23

Rev. 0 4/99

Table 2.10.1-10

End Drop on Bottom - Rear Impact Limiter (1G)


Flange 19 74

20 15

Lid 21 93

22 148


25 446

26 173

27 139

28 104


31 54

32 54

33 37

34 38

35 35

36 23

37 131 Weld 38 19

39 170

Rev. 0 4/99

Table 2.10.1-11

End Drop on Lid - Front Impact Limiter (1G)

Location

Inner Bottom Plate

Inner Shell

I

Nodal Stress Intensity (psi) 36

2 3

3 33

4 9

5 92

6 37

7 37

8 39

9 45

10 45

11 62

12 62

13 78

14 79

15 109

16 98

17 116

18

Rev. 0 4/99

1

72

Table 2.10.1-12

End Drop on Lid - Front Impact Limiter (I G)


Flange 19 259

20 55

Lid 21 327

22 441


25 129

26 49

27 40

28 30


31 51

32 51

33 66

34 68

35 118

36 77

37 44 Weld 38 41

39 548

Rev. 0 4/99

Table 2.10.1-13

Thermal Stress at 1000 F Hot Environment


1 2710


3 2747

4 3269

5 7963

6 4626

7 5713

8 5507

9 5794

10 5426 Inner Shell 11 5599

12 5619

13 6116

14 5244

15 5633 16 6314

17 5303

18 6628

Rev. 0 4/99

Table 2.10.1-14

Thermal Stress at 1000 F Hot Environment


Flange 19 1772

20 3376

Lid 21 2924

22 2058


25 7135

26 2556

27 2839

28 6197

29 4527 Outer Shell 30 6967

31 2905

32 5435

33 4099

34 6551

35 2593

36 4981

37 2783 Weld 38 4860

39 3233

Rev. 0 4/99

Table 2.10.1-15

Thermal Stresses at -201F Environmental Conditions


1 1277 2 1183

Inner Bottom Plate

3 1270

4 1184

5 1415

6 1199

7 1259

8 1266

9 1273

10 1272 Inner Shell 11 1289

12 1290

13 1317

14 1296

15 1253 16 1492

17 1237

18 1404

Rev. 0 4/99

Table 2.10.1-16

Thermal Stresses at -20'F Environmental Conditions


Flange 19 1341

20 1254

Lid 21 1034

22 591


25 3503

26 180

27 451

28 370


31 409

32 359

33 379

34 365

35 964

36 229

37 161 Weld 38 2005

39 1352

Rev. 0 4/99

Table 2.10.1-17

Thermal Stresses at -40'F Environmental Conditions


1 1533 2 1445

Inner Bottom Plate

3 1529

4 1445

5 1805

6 1447

7 1563

8 1573

9 1579

10 1578 Inner Shell

11 1595

12 1595

13 1624 14 1599

15 1536

16 1836

17 1525

18 1727

Rev. 0 4/99

Table 2.10.1-18

Thermal Stresses at -40TF Environmental Conditions

Rev. 0 4/99


Flange 19 1627

20 1535

Lid 21 1247

22 691


25 4195

26 205

27 530

28 440


31 460

32 428

33 464

34 439

35 1175

36 286

37 200 Weld 38 2457

39 1680

Table 2.10.1-19

Transport Shock Load/Truck

Horizontal Cask Supported at Rear Trunnions and Front Saddle

(Stress at 1800 Orientation)


1 459 2 407


4 173

5 190

6 432

7 403

8 436

9 427

10 434 Inner Shell11 557

12 609

13 422

14 497

15 1071

16 877

17 1946 18 1310

Rev. 0 4/99

Table 2.10.1-20

Transport Shock Load/Truck Horizontal Cask Supported at Rear Trunnions and Front Saddle

(Stress at 1800 Orientation)


Flange 19 1675

20 1220

Lid 21 734

22 1367


25 664

26 148

27 360

28 823


31 480

32 726

33 347

34 579

35 1153

36 881

37 383 Weld 38 2348

39 899

Rev. 0 4/99

Table 2.10.1-21

1G Down Load

Horizontal Cask Supported at Rear Trunnions and Front Saddle (Stress at 180' Orientation)


1 109 210


4 46

5 41

6 43

78 93

9 55

10 62

11 102

12 115

13 77

14 95

15 232

16 200

17 467

18

Rev. 0 4/99

Inner Shell

83

334

Table 2.10.1-22

1G Down Load Hoiizontal Cask Supported at Rear Trunnions and Front Saddle

(Stress at 180' Orientation)


Flange 19 569

20 350

Lid 21 35

22 55


25 77

26 22

27 88

28 108


31 70

32 133

33 46

34 105

35 238

36 207

37 53 Weld 38 550

39 168

Rev. 0 4/99

Table 2.10.1-23

Side Drop (IG) (Stress at 900 Orientation)

Rev. 0 4/99

Table 2.10.1-24

Side Drop (IG) (Stress at 900 Orientation)


Flange 19 351

20 272

Lid 21 158

22 219


25 109

26 49

27 225

28 143


31 60

32 57

33 111

34 91

35 222

36 144

37 204 Weld 38 215

39 48

Rev. 0 4/99

Table 2.10.1-25

Side Drop (I G)



1 54 2 63

-Inner Bottom Plate

3 113

4 120

5 214

6 214

8 55

9 89

10 103

11 130

12 137

13 76

14 89

15 36

16

17

18

150

166

Rev. 0 4/99

Inner Shell

7 1la

150

Table 2.10.1-26

Side Drop (IG) (Stress at Impact Side, 180* Orientation)


Flange 19 297

20 209

Lid 21 260

22 266


25 153

26 233

27 47

28 134


31 125

32 204

33 107

34 170

35 55

36 178

37 168 Weld 38 224

39 48

Rev. 0 4/99

Table 2.10.1-27

C.G. Over Bottom Comer Drop - Rear Impact Limiter (MG) (Stress at 90' Orientation)

Location

Inner Bottom Plate

Inner Shell

Nodal Stress Intensity (psi)

1 117

2 138

3 90

4 67

5 265

6 124

7 128

8 123

9 101

10 96

11 87

12 77

13 83

14 76

15 101

16 97

17 118

18 114

Rev. 0 4/99

Table 2.10.1-28

C.G. Over Bottom Comer Drop - Rear Impact Limiter (IG) (Stress at 90' Orientation)


Flange 19 231

20 179

Lid 21 104

22 145


25 318

26 174

27 166

28 105


31 76

32 70

33 77

34 65

35 106

36 91

37 146 Weld 38 216

39 128

Rev. 0 4/99

Table 2.10.1-29

C.G. Over Bottom Comer Drop - Rear Impact Limiter (1G) (Stress at Impact Side, 1800 Orientation)


1 351

2 340 Inner Bottom Plate 3 130

4 149

5 233

6 86

Inner Shell

78

9

10

11

12 13

14

15

16

17

18

70

26

29

16

19

8

13

35

40

62

55

Rev. 0 4/99

I

6 1

Table 2.10.1-30

C.G. Over Bottom Comer Drop - Rear Impact Limiter (IG) (Stress at Impact Side, 1800 Orientation)


Flange 19 96

20 57

Lid 21 144

22 119


25 466

26 158

27 101

28 18


31 14

32 54

33 580

34 54

35 26

36 43

37 205 Weld 38 51

39 102

Rev. 0 4/99

Table 2.10.1-31

C.G. Over Lid Comer Drop - Front Impact Limiter (I G) (Stress at 90' Orientation)

-V Nodal Stress Intensity (psi)

1 102


3 64

4 57

5 120

6 75

8 97

9 83

10 75

11 89

12 80

13 112

14 107

15 174

16 153

17 182

18

Rev. 0 4/99

Location

Inner Shell

IV,,+7

146

Table 2.10.1-32

C.G. Over Lid Comer Drop - Front Impact Limiter (MG) (Stress at 90' Orientation)


Flange 19 433

20 171

Lid 21 317

22 550


25 140

26 49

27 123

28 79

29 81

Outer Shell 30 68

31 74

32 68

33 93

34 86

35 169

36 102

37 112 Weld 38 71

39 402

Rev. 0 4/99

Table 2.10.1-33

C.G. Over Lid Comer Drop - Front Impact Limiter (1G)


Location

Inner Bottom Plate

Inner Shell

Rev. 0 4/99


.1 363

2 354

3 147

4 144

5 75

6 59

7 27

8 23

9 8

10 12

11 13

12 12

13 39

14 43

15 123

16 65

17 132

18 147

I

i

Table 2.10.1-34

C.G. Over Lid Comer Drop - Front Impact Limiter (IG) (Stress at Impact Side, 1800 Orientation)


Flange 19 273

20 178

Lid 21 284

22 659


25 164

26 112

27 37

28 60


31 21

32 61

33 13

34 48

35 156

36 73

37 69 Weld 38 237

39 416

Rev. 0 4/99

Table 2.10.1-35

150 Slap Down Drop - Second Impact (Gnomal = 42.22, G ucational = 77.78)

(Stress at 90* Orientation)

Location Nodal Stress Intei (psi)

1 10380 2 15522

Inner Bottom Plate

3 6059

4 5599

5 8313

6 5538

7 9195

8 8629

9 5544

10 5209 Inner Shell

11 3812

12 3293

13 6307

14 5806

15 11602

16 9754

17 13208

18 10034

nsity

Rev. 0 4/99

Table 2.10.1-36

150 Slap Down Drop - Second Impact (Gnom = 42.22, Grotatjona = 77.78) (Stress at 90' Orientation)


Flange 19 13527

20 7415

Lid 21 17266

22 25841


25 7073

26 1280

27 11167

28 7354

29 6668 Outer Shell 30 5985

31 3379

32 3144

33 4328

34 4153

35 10497

36 4415

37 11009 Weld 38 2615

39 4104

Rev. 0 4/99

Table 2.10.1-37

15' Slap Down Drop - Second Impact (Gno.0 l = 42.22, Gottjor0 = 77.78) (Stress at Impact Side, 1800 Orientation)


1 37507


3 14637

4 13805

5 7478

6 8201

7 99

8 1571

9 3571

10 3680 Inner Shell

11 5612

12 5816

13 3366

14 4212

15 3469

16 3198

17 14353

18 15655

Rev. 0 4/99

Table 2.10.1-38

150 Slap Down Drop - Second Impact (Gnomia = 42.22, Grotatjonal = 77.78) (Stress at Impact Side, 180* Orientation)


Flange 19 13527

20 7415

Lid 21 17266

22 25841


25 7073

26 1280

27 11167

28 7354

29 6668 Outer Shell 30 5985

31 3379

32 3144

33 4328

34 4153

35 10497

36 4415

37 11009 Weld 38 2615

39 4104

Rev. 0 4/99

Table 2.10.1-39

Fire Accident

Location

Inner Bottom Plate

Inner Shell


2822

2198

2877

6409

4447

5625

4764

5392

4993

5208

5175

5704

4824

7361

3052

7992

2429

Rev. 0 4/99

1

2

3

4

5

6

7

8

9

10

11

12

13

14

15

16

17

18

Table 2.10.1-40

Fire Accident


Flange 19 7544

20 8626

Lid 21 3393

22 2870


25 6512

26 4836

27 8873

28 17653

29 5253 Outer Shell 30 7506

31 4018

32 6480

33 5248

34 7211

35 2483

36 8882

37 3853 Weld 38 34774

39 4568

Rev. 0 4/99

Table 2.10.1-41

lg Longitudinal Loading Horizontal Cask Held at Rear Trunnions and Front Saddle

T .. n . i�oaai �tress 1flL�flS1LYLocation

Inner Bottom Plate

Inner Shell

I I

1

Nodal Stress Intens(ty (psi)

2 9

3 34

4 4

5 17

6

7

8

9

10

11

12

13

14

15

16

17

18

199

80 118

94 85

75

76 58

59

77

62

75

17

Rev. 0 4/99

• ml,.3y

Table 2.10.1-42

lg Longitudinal Loading Horizontal Cask Held at Rear Trunnions and Front Saddle


Flange 19 233

20 70

Lid 21 335

22 569


25 417

26 72

27 72

28 282

29 128

Outer Shell 30 90

31 92

32 93

33 74

34 77

35 110

36 57

37 85 Weld 38 101

39 640

Rev. 0 4/99

Table 2.10.1-43

Transport Truck Vibration Load (0.3g Long., 0.3g Lat., 0.6g Vert.)



1 76

2 71 Inner Bottom Plate 3 431s

4 30

5 31

6 55

7 66

8 71

9 65

10 67 Inner Shell90

12 99

13 68

14 81 15 179

16 146

17 332

18 227

Rev. 0 4/99

Table 2.10.1-44

Transport Truck Vibration Load (0.3g Long., 0.3g Lat., 0.6g Vert.) Horizontal Cask Held at Rear Trunnions and Front Saddle

Rev. 0 4/99


Flange 19 314

20 219

Lid 21 95

22 181


25 74

26 21

27 61

28 125


31 74

32 117

33 53

34 93

35 190

36 150

37 25 Weld 38 396

39 96

Table 2.10.1-45

Transport Tiedown Load (10g Long., 5g Lat., 2g Vert.)


Location

Inner Bottom Plate

T

Inner Shell

1

2

3

4

5

6

7 8

9

10

11

12

13

14

15

16

17

18

Rev. 0 4/99


576

578

232

378

1955

829

1124

1232

1181

1269

1343

961

1077

2021

1648

3165

1907

Table 2.10.1-46

Transport Tiedown Load (1Og Long., 5g Lat., 2g Vert.) Horizontal Cask Held at Rear Trunnions and Front Saddle


Flange 19 1311

20 1332

Lid 21 3243

22 5770

23 1497

Outer Bottom Plate 24 1493

25 3753

26 666

27 712

28 2889

29 1494

Outer Shell 30 1190

31 1294

32 1646

33 986

34 1333

35 2340

36 1517

37 663

Weld 38 3972

39 5562

Rev. 0 4/99

Table 2.10

6g Lifting Load c (Cask Ver

Location

Inner Bottom Plate

Inner Shell

.1-47

,n Trunnion tical)

Nodal Stress Intensity

(psi)

1 797

2 13

3 700

4 167

5 1961

6 1323

7 501

8 547

9 428

10 432

11 558

12 559

13 693

14 631

15 812

16 669

17 755

18 939

Rev. 0 4/99

Table 2.10.1-48

6g Lifting Load on Trunnion (Cask Vertical)


Flange 19 579

20 791

Lid 21 351

22 625


25 3712

26 1223

27 851

28 591


31 607

32 608

33 800

34 676

35 118

36 1243

37 1216 Weld 38 427

39 846

Rev. 0 4/99

Table 2.10.1-49

Local Stresses at Rear Trunnion / Cask Body Interface with Ig Down Load

(Cask Horizontal)


1

2

Inner Bottom Plate 3

4

5

6

7 704 8 704

9

10 Inner Shell11

12

13

14

15 16

17

18

Rev. 0 4/99

Table 2.10.1-50

Local Stresses at Rear Trunnion / Cask Body Interface with Ig Down Load (Cask Horizontal)


Flange 19

20

Lid 21

22

23 Outer Bottom Plate 24

25

26

27 756

28 756

29

Outer Shell 30

31

32

33

34

35

36

37

Weld 38

39

Rev. 0 4/99

Table 2.10.1-51

Local Stresses at Rear Trunnion / Cask Body Interface with Truck Shock Loads

(Cask Horizontal)


1

2

Inner Bottom Plate

3

4

5

6 7 6385

8 6385

9

10 Inner Shell1

12

13

14

15

16

17

18

Rev. 0 4/99

Table 2.10.1-52

Local Stresses at Rear Trunnion / Cask Body Interface with Truck Shock Loads (Cask Horizontal)


Flange 19

20

Lid 21

22


25

26

27 6385

28 7989

29 Outer Shell 30

31

32

33

34

35

36

37 Weld 38

39

Rev. 0 4/99

Table 2.10.1-53

Local Stresses at Rear Trunnion / Cask Body Interface with Vibration Loads (Cask Horizontal)


1


3

4

5

6

7 1086

8 1086

9

10 Inner Shell 11

12

13

14

15

16

17

18

Rev. 0 4/99

Table 2.10.1-54

Local Stresses at Rear Trunnion / Cask Body Interface with Vibration Loads (Cask Horizontal)


Flange 19

20

Lid 21

22


25

26

27 1086

28 1383

29 Outer Shell 30

31

32

33

34

35

36

37 Weld 38

39

Rev. 0 4/99

Table 2.10.1-55

Local Stresses at Rear Trunnion / Cask Body Interface with Tiedown Loads (Cask Horizontal)


1


3

4

5

6

7 23439

8 23439

9

10 Inner Shell

11

12

13

14

15

16

17

18

Rev. 0 4/99

Table 2.10.1-56

Local Stresses at Rear Trunnion / Cask Body Interface with Tiedown Loads (Cask Horizontal)


Flange 19

20

Lid 21

22


25

26

27 27904

28 27904

29 Outer Shell 30

31

32

33

34

35

36

37 Weld 38

39

Rev. 0 -4/99

Table 2.10.1-57

Local Stresses at Front Trunnion / Cask Body Interface with 6g Lifting Load (Cask Vertical)


1


3

4

5

6

_________________________________________ I.8

9 10

11

12

13

14

15 19705

16 19705

17

18

Rev. 0 4/99

Location

Inner Shell

7

Table 2.10.1-58

Local Stresses at Front Trunnion / Cask Body Interface with 6g Lifting Load (Cask Vertical)


Flange 19

20

Lid 21

22


25

26

27

28

29 Outer Shell 30

31

32

33

34

35 19705

36 19705

37 Weld 38

39

Rev. 0 4/99

Table 2.10.1-59

Local Stresses at Rear Trunnion / Cask Body Interface with Rail Car Shock Load (Cask Horizontal)


1 2

Inner Bottom Plate 3

4

5

6

7 14089 8 14089

9

10 Inner Shell

12

13

14

15

16

17

18

Rev. 0 4/99

Table 2.10.1-60

Local Stresses at Rear Trunnion / Cask Body Interface with Rail Car Shock Load (Cask Horizontal)


Flange 19

20

Lid 21

22


25

26

27 14089

28 18461

29 Outer Shell 30

31

32

33

34

35

36

37 Weld 38

39

Rev. 0 4/99

Table 2.10.1-61

Transport Rail Car Shock Load (4.7g All Directions) Horizontal Cask Held at Rear Trunnions and Front Saddle


1 820 2 704

Inner Bottom Plate

3 476

4 297

5 340

6 890

7 724

8 790

9 806

10 811 Inner Shell

11 1014

12 1105

13 769

14 898

15 1906

16 1560

17 3408

18 2271

Rev. 0 4/99

Table 2.10.1-62

Transport Rail Car Shock Load (4.7g All Directions) Horizontal Cask Held at Rear Trunnions and Front Saddle


Flange 19 2732

20 2066

Lid 21 1504

22 2775


25 1450

26 290

27 631

28 1553


31 896

32 1322

33 653

34 1057

35 2068

36 1550

37 310 Weld 38 4130

39 2006

Rev. 0 4/99

(

Sf APP.LD LOADS

RADIAL LOAD P I. L8.

CIRC. MOMENT Mo w 6,-LR

LONG. MOMENT MLI N-.-.-.. . H-LBL

TORSION MOMENT Mv I - IK-LB.

SHEAR LOAD VS LB.

SJHEAR WAD VI. L...... LL.

* NOTE- ENTER ALL FORCE VALUES IN ACCORDANCE WITH SIGN CONVENTION

L fOMLM I &GEOMETRIC PARAMETERS

VESSELTHICKNESN T L.%-.. _ I. "* . -r

ArlACHNENT RADIUS ft L. . I. O.N"ST5lR.

VESSEL RADIUS R- I - *4

/ (

DL. Au'

COO=4LA EM T

FROM FIG. READ CURVES FOR COMPUTE ANSOLUTE WUES OF STRESS AND SNTES RESULT SAS AL a. . el PT DL

SA RND 4CftTRIN W.J .& I + + +.4 + + SE ATE S•-1 TrE U DV 4 .-- ÷

115Am~~~~~~~ +TESCET A I U

ADD ALGEBRAICALLY FOR SUMMATION Of A STRESSES 17.0

3C EAN TRSDIETfSN 4C a-x*T I +" + + + + +

ADD AL.GEBRAICALL.Y FOR SUMMATION OF STRESSES ý7,

LONGM

PRESSURE STRESS

LONGITUDINAL SENDING STRESS TOTAL MEMBRANE STRESS

TOhU. SUIffCE STRESS

"MCT °*,1 40,•.1•

CIDmAmEmrNTIAL omp OIW I SIME FOR LOCAL SYSESLES

SINCTLISIMOCAL SHELLS5

NOZZLE NO.

PIPING LOAD COD

"* G ANALYSIS POINTe

tz

o CD

o a

CD

0


Rev. 0 4/99

FIGURE 2.10.1-1

CASK BODY KEY DIMENSIONS REV. 0 4/99

lwalcourt

New Stamp

y

-sH'L E T /

7_ FA. 3..

ZIAE Z5

FIGURE 2.10.1-2 CASK BODY FINITE

ELEMENT MODELREV. 0 4/99

x---snow

FIGURE 2.10.1-3 CASK BODY

BOTTOM CORNER REV. 0 4/99

CO..TAINLl,*M FZANGE

LID BO~rS CVO'.ECTING PMII1AY LID -L0 COflTAINYXKT FLANGE

AND WZ-LD ATTACHII4G Gk'.1A SHIELDING TO FAG

REV. 0 4/99

FIGURE 2.10.1-5 . CASK LID TO SHELD PLATE CONNECTION

REV. 0 4/99

II

75.880"

SEAL LOAD F. "

SEAL PRESSURE

ICONTAINMENT FLANGE

LD. xPs 2

D.B.C .

GAMMA SHIELDING

2 x 1399 #/IN - 2798 #/IN AT Day - 72.10"

Tr (72.10)(2798) PS- Tr /4 (72.902- 71.302)"

3498 psi

BOLT PRELOAD STRESS: 86,000 psi

- The actual lid preload stress is 56,000 psi. however, rid bolt prelood corresponding to 86,000 psi is used for all load combinations.

FIGURE 2.10.1-6 BOLT PRELOAD AND

SEAL REACTION

REV. 0 4/99

FIGURE 2.10.1-7 DESIGN INTERNAL

PRESSURE (100 PSIG) REV. 0 4/99

y

L-

FIGURE 2.10.1-8 EXTERNAL PRESSURE

LOADING (25 PSIG) REV. 0 4/99

I

"",7

FRONT IMPACT LIMITER INERTIA

I INTERNALS= WEIGHT OF BASKET. RAILS. HOLDDOWN RING. AND FUEL ASSEMBLIES

INTERNALS INERTIA

pite

PR

REAR IMPACT LIMITER REACTION

FIGURE 2].10-9 IMPACT AT BOTTOM END

LOAD DISTRIBUTION REV. 0 4/99

REAR IMPACT LIMITER INERTIA

FRONT IMPACT LIMITER REACTION

FIGURE 2.10.1-10 IMPACT AT LID END LOAD DISTRIBUTION

r%,l I"- ll! A I ,%l

KI-V. U -4/YY

,

liNTRNL INERTIA DUE TO INERTIA RESULTANT OF FRONT & REAR CASK INERTIA IMPACT LIMITER

p -- - --~ FF+

REACTION FORCE AT REAR TRUNNION

m

0 FIGURE 2.10.1-11I IN IG LONGITUDINAL SUPPORTED (o BY TWO REAR TRUNNIONS

( ,

/~~I ýý/Iz:IzI LzLI a� 4a

140me

COL 1

THETA( 1)

FOURIER COEFFICIENTS

MODE

0.0000 1.0000 1.0000 2.0000 2.0000 3.0000 3.0000 4.0000 4.0000 5.0000 5.0000 6.0000 6.0000 7.0000. 7.0000 8.0000 8.0000 9.0000 9.0000

10.0000 10.0000

COEFF

-0.3182 0.5000 0.0000

-0.2124 0.0000 .0.0000 0.0000 0.0426 0.0000 0.0000 0.0000

-0.0183 0.0000 0.0000 0.0000 0.0103 0.0000 0.0000 0.0000

-0.0066 0.0000

ISYM

1.0000 1.0000

-1.0000 .1.0000 -1.0000 1.0000

-1.0000 1.0000

-1.0000 1.0000

-1.0000 1.0000

-1.0000 1.0000

-1.0000 1.0000

-1.0000 1.0000

-1.0000 1.0000

-1.0000

FOURIER OUTPUT CURVE FOR COS FUNCTION OVER 90-270 DEGREES

FIGURE 2.10.1-12 FOURIER COEFFICIENTS FOR

IG LATERALREV. 0 4/99

iC

0

REV. 0 4/99

(

C, 0io"

I -ME -- I00.51" IUU=MI

C.G. O-' FRONT IMPACT

SADDLE 48.92"

C (

II3*.: II0.63"

96.74"

1 I

(-69.33"

- -N-C.G. OF

REAR IMPACT LIMITER

LTRUNNION

FIGURE 2.10.1-13 LOCATION OF CG. TRUNNIONS. SADDLE. AND IMPACT LIMITERS

T I

C.G.

I

' ' I I I I I I I I

-1

i i1 4 4 4 1lf ~ f

I

°,ý.ýw

N

FRONT IMPACT LIMITER INERTIA t-Pli PF 2

Pr-

/ / / / /I

REAR IMPACT LIMITER INERTIAPo 11"'

Ps SADDLE REACTION FORCE

REACTION FORCE AT REAR TRUNNION

F FIGURE 2.10.1-14 IG VERTICAL DOWN SUPPORTED

BY TWO REAR TRUNNIONS AND FRONT SADDLE

(

ItCASK INERTIA

INTERNALS T I INERTIA t 10%i-

i- W

m 0

0

0O

\

ell

.i

11! I III . LI, I I

[/ / / -, - -//Z///////////

.\l\

(

(

C

INTERNALS INERTIA

.G.

, F CASK IN!

/It�K�K�K7K�7K77K7KIY3X Al'. T

PF I PF 2

FRONT IMPACT LIMITER REACTI

94.88"

RESULTANT FORCE

ERTIA

PR

REAR IMPACT LIMITER REACTION

0.36"

RESULTANT FORCEp Ny

FIGURE 2.10.1-15 SIDE DROP

LOAD DISTRIBUTION

m

0

(C

1

AA I

,

zzzzzzzzzzz4z

FRONT IMPACT LIMITER NORMAL INERTIA A'

/RESULTANT OF CASK INERTIA

/

IMPACT LIMITER

FIGURE 2.10.1-16 C.G. OVER BOTTOM CORNER DROP

LOAD DISTRIBUTIONREV. 0 4/99

REAR IMPACT LIMITER NORMAL INERTIA PRN

RESULTANT OF CASK INERTIA

0

/

FRONT IMPACT LIMITER NORMAL REACTION

FIGURE 2.10.1-17 C.G. OVER LID END CORNER DROP

LOAD DISTRIBUTION

REV. 0 4/99

)Qrl

C.G.

NORMAL ACCELERATION

.+

ROTATIONAL ACCELERATION

93- 91 -19 (NOTE I)

C.G.

91 ' 92 2. 722 (NOTE I) "91 + 92 - 25.4

(NOTE I)

COMBINED ACCELERATION

NOTE I: TRANSVERSE 9 LOADS REPORTED IN TABLE 2.10.8-3

FIGURE 2.10.1-18 TRANSVERSE 9 LOAD COMBINATIONS

REV. 0 4/99

93

91 " v ý T ý ý ý ý ý

P8

Pti

70 m

0

4.. FIGURE 2.10.1-20 "o 150 SLAP DOWN (SECOND IMPACT AT LID END) 0o LOAD DISTRIBUTION

(

C C

(

F y

m X

0

FIGURE 2.10.1-21 15" SLAP DOWN (SECOND IMPACT AT LID END)

0o ROTATIONAL QUASISTATIC EQUILIBRIUM

((

B

FIGURE 2.10.1-22 STANDARD REPORTING

LOCATIONS FOR CASK BODY

REV. 0 4/99

"I

FIGURE 2.10.1-23 WELD STRESS LOCATIONS

REV. 0 4/99

6W x 1.15= 6 x 240,000 x 1.15 = 828,000 LBS. FTR= 2 2

p = 6 x 77,240 x 1.15 = 134.61 psi I x (35.5) 2 FIGURE 2.10.1-24

LIFTING: 6g

VERTICAL UP

REV. 0 4/99


APPENDIX 2.10.2

TABLE OF CONTENTS

Page 2.10.2 LID BOLT ANALYSIS

2.10.2.1 Introduction ...................................................................................... 2.10.2-1 2.10.2.2 Bolt Load Calculations ..................................................................... 2.10.2-2 2.10.2.3 Load Combinations .......................................................................... 2.10.2-8 2.10.2.4 Bolt Stress Calculations ................................................................. 2.10.2-10 2.10.2.5 Analysis Results ............................................................................. 2.10.2-13 2.10.2.6 Fatigue Analysis ............................................................................. 2.10.2-14 2.10.2.7 Minimum Engagement Length for Bolt and Flange ...................... 2.10.2-19 2.10.2.8 Conclusions .................................................................................... 2.10.2-21 2.10.2.9 References ...................................................................................... 2.10.2-22

LIST OF TABLES

2.10.2-1 Design Parameters for Lid Bolt Analysis 2.10.2-2 Bolt Data 2.10.2-3 Allowable Stresses in Closure Bolts for Normal Conditions 2.10.2-4 Allowable Stresses in Closure Bolts for Accident Conditions

LIST OF FIGURES

2.10.2-1 TN-68 Cask Lid Closure Arrangement 2.10.2-2 TN-68 Cask Lid Bolt

Rev. 0 4/992.10.2-i


Rev. 0 4/99

APPENDIX 2.10.2

LID BOLT ANALYSIS


This section evaluates the ability of the cask closure to maintain a leak tight seal under normal and accident conditions. Also evaluated in this section, are the bolt thread and internal thread stresses, and lid bolt fatigue. The stress analysis is performed in accordance with NUREG/CR6007"').

The TN-68 cask lid closure arrangement is shown in Figure 2.10.2-1. The 5.0 inch thick lid is bolted directly to the end of the containment vessel flange by 48 high strength alloy steel 1.875 inch diameter bolts. Close fitting alignment pins ensure that the lid is centered in the vessel.

The lid bolt is shown in Figure 2.10.2-2. The bolt material is SA-540 Gr. B24 class 1 which has a minimum yield strength of 150 ksi at room temperature.

The following ways to minimize bolt forces and bolt failures for shipping casks are taken directly from with NUREG/CR-6007•'), page xiii. All of the following design methods are employed in the TN-68 closure system.

"* Protect closure lid from direct impact to minimize bolt forces generated by free drops. (use impact limiters)

"* Use materials with similar thermal properties for the closure bolts, the lid, and the cask wall to minimize the bolt forces generated by fire accident

"* Apply sufficiently large bolt preload to minimize fatigue and loosening of the bolts by vibration.

"* Lubricate bolt threads to reduce required preload torque and to increase the predictability of the achieved preload.

* Use closure lid design which minimizes the prying actions of applied loads.

"* When choosing a bolt preload, pay special attention to the interactions between the preload and thermal load and between the preload and the prying action.

Rev. 0 4/992.10.2-1

The following evaluations are presented in this section:

"* Lid bolt torque "* Bolt preload "* Gasket seating load "* Pressure load "* Temperature load "* Impact load "* Puncture load "* Thread engagement length evaluation "* Bearing stress "* Load combinations for normal and accident conditions "• Bolt stresses and allowable stresses "* Lid bolt fatigue

The design parameters of the lid closure are summarized in Table 2.10.2-1. The lid bolt data and material allowables are presented in Tables 2.10.2-2 through 2.10.2-4. A maximum temperature of 300'F is used in the lid bolt region during normal and accident conditions. The following load cases are considered in the analysis.

1. Preload + Temperature Load (normal condition) 2. Pressure Load + 1 Foot Drop (normal condition) 3. Pressure + 30 Foot Comer Drop (accident condition) 4. Pressure + Puncture Load (accident condition)

2.10.2.2 Bolt Load Calculations

Symbols and terminology used in this analysis are taken from NUREG/CR-6007(1 ) and are reproduced in Table 2.10.2-1.

2.10.2.2.1 Lid Bolt Torque

The desired maximum preload stress in the lid bolts is 56,000 psi.

For a 1 7/8" - 8UN -2A bolt, the Tensile Stress Area is 2.414 in2 (see Table 2.10.2-2). Therefore,

F, = 56,000 x Stress Area = 56,000 x 2.414 = 135,200 lb.

The torque required to achieve this preload is (Ref. 1, Section 4.0)

Q = KDbF0 = 0.1 (1.875) (135,200) = 25,350 in. lb. = 2,113 ft. lb.

A bolt torque range of 2,050 to 2,100 ft. lb. has been selected. For the minimum torque,

F, = QIKDb = 2,050x 12/(0.1 x 1.875) = 131,200 Ibs, and

Rev: 0 4/992.10.2-2

Preload stress = 131,200/2.414= 54,350 psi.

2.10.2.2.2 Bolt Preload (Ref. 1, Table 4.1)

Fa = QIKDb = 25,350/0.1(1.875) = 135,200 lb.

Residual torsional moment for maximum torque is,

Mtr= 0.5Q =.5(25,350) = 12,680 in. lb.

Residual torsional moment for minimum torque of 2,050 ft. lb. is,

Mtr= 0.5Q =.5(2,050x12) = 12,300 in. lb.

Residual tensile bolt force,

Far= Fa = 135,200 lbs

2.10.2.2.3 Gasket Seating Load (Seal - 2 Helicoflex HND 229 seals, Aluminum Jacket(2•)

The diameter of the inner seal, D1 s, is 71.3 in., and the diameter of the outer seal, Ds, is 72.9 in. The force to seat the seals is 1399 lbs./in (245 N/mm)(2) times the circumference of the seal. Therefore the force required to seat the seals is:

Inner: 7c (71.3) (1399) = 313,400 lbs. Outer: nt (72.9) (1399) = 320,400 lbs.

Total, Fa = 633,800 lbs.

Therefore, the seal seating load is,

FaI48 = 633,800/48 = 13,200 lb.ibolt.

2.10.2.2.4 Pressure Loads (Ref. 1, Table 4.3)

Axial force per bolt due to internal pressure is,

Fa 4 Nb

Rev. 0 4/992.10.2-3

Dig for outer seal (conservative) = 72.9 in. Then,

7-(72.9 2 )(100-0) 4(48) 8,696 lb./bolt.

The fixed edge closure lid force is,

FZ =ODIb (Pli - PIo) _ 75.88(100) = 1,897 lb. in.-. 4 4

The fixed edge closure lid moment is,

My - (P1 i- PIo)D 1 ,b - 100(75.882) = 17,990in. lb. in.32 32 1

The shear bolt force per bolt is,

7.rE It, (Pi- Po =)D11 , •r(27.8x 106 ý5.0X100X75.88)2 17,950 lb./bolt.

2NbEctc(1-NN,,) 2(48X27.8 x 106 X7.5X0.7)

The lid shoulder takes this shear force, so that F, = 0.

2.10.2.2.5 Temperature Loads

The lid bolt material is SA-540 GR.B24 Cl. 1, 2Ni ¾ Cr 1/3 M,. This is Group E in the thermal coefficients of expansion tables in Reference 3. Both the lid and flange are made of SA-350 Gr. LF3, 3 ½ Ni, which is also Group E. Consetuently, the bolts, lid and flange have the same coefficient of thermal expansion (6.78 x 10- in/in-0 F at 3000F). Therefore, heating to the maximum isothermal temperature will not generate bolt stress.

2.10.2.2.6 Impact Loads (Ref. I, Table 4.5)

The non-prying tensile bolt force per bolt, F,, is,

F = 1.34 sin(xi)(DLF)(ai)(W, + W,) _ 1.34 sin(xi)(1.1)(ai)(89,500) = 2,748 Nb= 48748(ai) sin(xi) lb./bolt. N!, 48

Note: W,+ W, is conservatively assumed to be 89,500 lbs. [Actual weight from chapter 2 = 12,289 (lid and lid bolts) + 25,868 (basket, rails, and shims) + 1,408 (hold down ring) + 47,940 (fuel assemblies) = 87,505 lbs.]

Rev. 0 4/992.10.2-4

The shear bolt force is,

, = cos(xi)(ai)(W1 ) - 12,100(ai) cos(xi) - 252. 1(ai) cos(xi) lb./bolt. Nb 48

The lid shoulder during normal and accident condition drops takes shear force. Therefore,

F,=0.

The fixed-edge closure lid force, Ff, is,

1.34 sin(xi)(DLF)(ai)(W1 + W) 1.34 sin(xi)(1.1)(ai)(89,500) - 553.4 sin(xi)(ai) lb./bolt. Ff = Dlb ;- (75.88)

The fixed-edge closure lid moment, Mf, is,

= 1.34 sin(xi)(DLF)(ai)(Wj + W,) _ 1.34 sin(xi)(1.1)(ai)(89,500) = 5,249 sin(xi)(ai) lb./bolt. 81z 8;z

Normal Condition Loads

Since the bolts are protected by the impact limiter during a 90' end drop, the worst case scenario is taken to be roughly a 600 CG over comer drop. From the impact limiter 1 foot normal condition analysis, Section 2.10.8.5, the maximum g load for a 1 foot 60' CG over comer drop is 5.39 g's (5 g vertical, 2 g transverse, Appendix 2.10.8, Table 2.10.8-11). However, for the lid bolt analysis, the following normal condition loading is conservatively assumed.

ai = 6 gs, and xi = 60'

Therefore,

Fa = 2,748x6xsin(60°) = 14,280 lb./bolt, F, = 0 lb./bolt,

Ff= 553.4x6xsin(60°) = 2,876 lb./bolt, and Mf = 5,249x6xsin(60°) = 27,280 lb./bolt.

Accident Condition Loads

The g load resulting from a 30 foot, 60" CG over comer drop is 35.47 gs (33 g vertical, 13 g transverse, Appendix 2.10.8, Table 2.10.8-13). The accident condition g load for lid bolt analysis is conservatively increased to the following value.

ai = 50 gs, and xi = 60'

Rev. 0 4/992.10.2-5

Therefore,

Fa = 2,748x5Oxsin(60') = 119,000 lb./bolt, F, = 0 lb./bolt,

Ff = 553.4x50xsin(60°) = 23,960 lb./bolt, and Mfy= 5,249x5Oxsin(60°) = 227,300 lb./bolt.

2.10.2.2.7 Puncture Loads (Ref. 1, Table 4.7)

The non-prying tensile bolt force per bolt, F,, is,

-sin(xi)Pun Nb

where,

Pun = The smaller of pbttl

LO.677D pAtSuI

T l 0.75zr(6 2)(37,500) = 3.181X 106

-The smaller of 0.67r(6)(9.5)(70,000) = 7.521 x 107

=>pun = 3.18tx10 6 lb.

The puncture force is greatest when xi = 90'. Conservatively neglect the protection provided by the impact limiter. Then,

- sin(xi)3.181 x 106 F, 48 _ 66,270 lb.

48

Since this force is negative (inward acting), the actual resulting bolt force, Fa = 0, because the applied load is supported by the cask wall and not the lid bolts. The shear bolt force is,

cos(90°)Pun lb.bolt. Nb

The lid shoulder during puncture takes shear force. Therefore,

Fs =0.

Rev. 0 4/992.10.2-6

The fixed-edge closure lid force, Ff, is,

- sin(xi)Pun - sin(900)3.181 x 106 Ff = D7•, 7(75.88) =

The fixed-edge closure lid moment, Mf, is,

Mf sin(xi)Pun sin(90°)3.181 x 10'= -253,100 lb./bolt. 4)z 4)z

ITD BOLT INDIVIDUAL LOAD SUMMARY

Non-Prying Torsional Prying Load Applied Tensile Moment, Prying Force, Moment, Case Load Force, Fa M, (in. lb.) Ff (lb.in.-1) Mf

(lb.) (in. lb. in.1 )

Maximum Torque 135,200 12,680 0 0

Preload Residual Minimum

Torque 131,200 12,300 0 0

Gasket Seating Load 13,200 0 0 0

Pressure 100 psig Internal 8,696 0 1,897 17,990

Thermal 300°F 0 0 0 0

1 Foot Normal Condition Drop 14,280 0 2,876 27,280

(6 gs) Impact 30 foot Accident

Condition Drop 119,000 0 23,960 227,300 (80 gs)

Drop on six inch Puncture diameter rod 0 0 -13,340 -253,100

Rev. 0 4/992.10.2-7

2.10.2.3 Load Combinations (Ref. 1, Table 4.9)

A summary of normal and accident condition load combinations is presented in the following table.

LTD BOlT NORMAl AND ACCIDENT LOAD COMBINATIONS

Rev. 0 4/99

Non-Prying Torsional Prying Load Combination Description Tensile Moment, Prying Force, Moment, Case Force, Fa M, (in. lb.) Ff (lb.in.")) Mf

(lb.) (in. lb. in.-') A.

Preload + Maximum 135,200 12,680 0 0 1. Temperature Torque

(Normal B. Condition) Minimum 131,200 12,300 0 0

Torque

2. Pressure + Normal Impact 22,980 0 4,773 45,270 (Normal Condition) Pressure + Accident

Impact 127,700 0 25,860 245,300

(Accident Condition)

Pressure + Puncture (Accident Condition) 8,696 0 -11,440 -235,100

2.10.2-8

LID BOLT NORMAI

Additional Prviny Bolt Force

Since the prying forces applied in load case 4 (pressure + puncture) acts inward, normal to the cask lid, an additional prying bolt force, F,,p, is generated (Ref. 1, Table 2.1). No additional force is generated for the outward loadings however (load cases 1, 2, and 3), because of the gap between the lid and flange at the outer edge. Fap is calculated in the following way.

c1+c2

where,

C1= 1, Ca 2 3(O Db)2 =-NI + (D.

75.7 8 {.27.8 x 106 (9.53) + (79.88- 70)(27.8 x 106)(7.5)3- 5.0

9.88 - 75.88Y 1-0.3 75.88 (48)(1.8752 )(27.8 x 106)

= 6.320,

"B is the non-prying tensile bolt force, and P is the bolt preload. Since F, - 0, F, < P, and therefore B = P. Parameters B, P, Ff, and Mf are quantities per unit length of bolt circle. For the applied inward force,

F"Nb (135,200)(48) P =B = 70 - -r7.8 27,220 lb. in.

Mflb )(75.88)

Mf -- (-235,l00)(48) - -47,3501lb. in.-', and Ff= 0 lb. in."'. ,r(75.88)

Therefore,

2(-47,350)

Fap =48 1+6.320

= 34,530 lb./bolt.

It is observed that the additional tensile bolt force due to prying for the puncture is less than the accident impact force. The puncture is therefore not critical for bolt stress evaluation.

Rev. 0 4/992.10.2-9

Bolt Bending Moment (Ref. 1, Table 2.2)

The maximum bolt bending moment, Mbb, generated by the applied load is evaluated as follows:

Mb D- (rtb Kb1 Mbb •,Nb I-Kb+ K, MMf

The terms Kb and K, are based on geometry and material properties and are defined in reference 1, Table 2.2. By substituting the values given above,

.(Nb Eb I = 1 48 27.8x 106 .18754 4- 6.792 x W05, and ~Lb*ýDib )64) (,5.0,J 75.88 * 64)

K Ett3_ 27.8x 106 (5.03)

31(1 - N )2 (i-; ) i JD23 3[( 1-0.3 2)+ (1 - 0.3y (75.88 )375.88

= 11.29 x 106

Therefore,

Mb =ri75.88 6.792X101 106Mf=021 f Mbb , - 16.792x10 5 +11.29x10 6

For load case 3, Mf= 45,270 in. lb. Substituting this value into the equation above gives,

Mbb = 12,760 in. lb. / bolt.

2.10.2.4 Bolt Stress Calculations (Ref. 1, Table 5.1)

2.10.2.4.1 Average. Tensile Stress

The bolt preload is calculated to withstand the worst case load combination and to maintain a clamping (compressive) force on the closure joint, under both normal and accident conditions. Based upon the load combination results (see Table LID BOLT NORMAL AND ACCIDENT LOAD COMBINATIONS on pg. 2.10.2-8), it is shown that a positive (compressive) load is maintained on the clamped joint for all load combinations. Therefore, in both normal and accident load cases, the maximum non-prying tensile force of 135,200 lb., from load case 1.A. (maximum torque preload + temperature load), is used for bolt stress calculations.

2.10.2-10 i,, n Areax'•..•T. V --rlJJ

Normal Condition:

/F 135,200 Sba = 1.2732- = 1.27323 - 56,000 psi. = 56.0 ksi.

Db, 1.75322

Accident Condition:

135,200 Sb,1 = 1.2732 -56,000 psi. = 56.0 ksi.

2.10.2.4.2 Bending Stress

Normal Condition:

Sbb = 10. 18 6 Mbb = 10.186 12,760 _ 24,120psi.= 24.1 ksi. Dia 1.7532s

2.10.2.4.3 Shear Stress

For both normal and accident conditions, the average shear stress caused by shear bolt force F, is,

Sb = 0.

For normal and accident conditions the maximum shear stress caused by the torsional moment M, is,

M_ 12,680 Sb, = 5.093 = 5.093 1.7532 3 11,980 psi. = 12.0 ksi.

Dba 1.75323

2.10.2.4.4 Maximum Combined Stress Intensity

The maximum combined stress intensity is calculated in the following way (Ref. 1, Table 5.1).

Sbi = [(Sb, + Sbb) 2 + 4(Sb, + Sb,) 2]05

For normal conditions, the combined tension, shear, bending, and residual torsion results in a maximum stress intensity of:

Sbi = [(56,000 + 24,120)2 + 4 (0 + 11,980)2]o.5 = 83,630 psi. = 83.6 ksi.

Rev. 0 4/992.10.2-11

2.10.2.4.5 Stress Ratios

In order to meet the stress ratio requirement, the following relationship must hold for both normal and accident conditions.

R 2 2+R2 < 1

Where R, is the ratio of average tensile stress to allowable average tensile stress, and R, is the ratio of average shear stress to allowable average shear stress.

For normal conditions:

Rt = 56,000/92,400 = 0.606,

R, = 11,980/55,400 = 0.216,

Rt2 + R,2 = (0.606)2 + (0.216)2 = 0.414 < 1.

For accident conditions:

R= 56,000/115,500 = 0.485,

R= 11,980/69,300 = 0.173,

R,2 +R, 2 = (0.485)2 + (0.173)2 = 0.265 < 1.

2.10.2.4.6 Bearing Stress (Under Bolt Head)

The maximum axial force is 135,2001b. A bolt head corresponding to 2 1/4" diameter bolt is used for 1 7/8" diameter shank due to higher bearing load during transport. The total bearing area under the 2 1/4" Hex bolt head is 6.45 in.2 The bearing stress is,

Bearing Stress = 135,200/6.45 = 20,960 psi. = 21.0 ksi.

The allowable bearing; stress on the lid is taken to be the yield stress of the lid material at 300' F. The lid may be manufactured out of SA-350 GR. LF3 or SA-203 GR. E. The yield stress of SA350 GR. LF3 is less than that of SA-203 GR. E (33.2 ksi. < 35.4 ksi.). Therefore the allowable stress is taken to be 33.2 ksi., which is the lesser of the two materials.

Rev. 0 4/992.10.2-12

2.10.2.5 Analysis Results

A summary of the bolt stresses calculated above is presented in the following table:

SUMMARY OF STRESSES AND ALLOWABLES

The calculated bolt stresses are all less than the specified allowable stresses.

Rev. 0 4/99

Normal Condition Accident Condition Stress Type

Stress Allowable Stress Allowable

Average Tensile 56.0 92.4 56.0 115.5

(ksi.)

Shear (ksi) 12.0 55.4 12.0 69.3

Combined S(ksi) 83.6 124.7 Not Required(')

Interaction E.Q. Rt' +R,2 < 1 0.414 1 0.265 1

Bearing (ksi) Allowable (ksi) 21.0 33.2 Not Required~')

(Sy of lid material)

SUMMARY OF STRESSES AND

2.10.2-13

2.10.2.6 Fatigue Analysis

The purpose of the fatigue analysis is to show quantitatively that the fatigue damage to the bolts during normal transport conditions is acceptable. This is done by determining the fatigue damage factor for each normal transport event. For this analysis it is assumed that the bolts are replaced after 150 round trip shipments. The total cumulative damage or fatigue usage for all events was conservatively determined by adding the usage factors for the individual events. The sum of the individual usage factors was checked to make certain that for the 150 round trip shipments of the TN-68 cask, the total usage factor was less than one. The following sequence of events was assumed for the fatigue evaluation.

1. Operating preload 2. Test pressure 3. Road vibration / shock 4. Pressure and temperature fluctuations 5. 1 foot normal condition drop

Since the bolt preload stress applied to the TN-68 cask lid bolts is higher than all of the other normal and accident condition loads, the stress in the bolt will never exceed the bolt preload stress. Consequently, the application and removal of preload is the only real cyclic loading that occurs in the lid bolts. The following analysis is therefore very conservative since it assumes that the damage factor is the sum of all of the individual event damage factors, and not simply the damage factor for bolt preload.

2.10.2.6.1 Operating Preload

Assuming that the bolts are replaced after 150 round trips, the number of preload cycles is two times the number trips or 300 cycles.

The maximum tensile stress due to bolt preload is 56,000 psi (Section 2.10.2.4.1), and the maximum shear stress due to residual bolt torsion is 11,980 psi (Section 2.10.2.4.3). The corresponding stress intensity is then

S.I. = 56,0002 + 4(11,9802) = 60,910 psi.

2.10.2.6.2 Test Pressure

The proof test, according to reference 11, is 1.25 x Design Pressure = 125 psi, and will only be performed once. The test pressure loads are calculated using the pressure loads computed in Section 2.10.2.2.4.

Fa = 8,696xl.25 = 10,870 psi., Fs = Oxl.25 = 0 psi,

Ff= 1,897xl.25 = 2,371 psi., and Mf= 17,990x1.25 = 22,490 psi.

Rev. 0 4/992.10.2-14

The lid bolt diameter is 1.7532 in. Therefore the stress intensity due to the test pressure load is, F7 10,870 Sa=1.2732F. = 1.2732 10 0= 4,503psi.,

Dbaa 1.75322

Sbb = 10.186 Mbb = 10.186 0.2818(22,490) = 11,980 psi,, Dba 1.75323

Since internal pressure causes no bolt torsion, and all shear loads are taken by the lid shoulder,

Sb, = 0, and Sbt = 0.

S.L = Sbi = [(Sba + Sbb) 2 + 4 (Sbs + Sbt)2]°5 = [(4,503 + 11,980)2 + 4(0)2]0.5 = 16,480 psi.

2.10.2.6.3 Vibration / Shock

Since the TN-68 Cask may be shipped either by truck or by rail car, the shock loading for both cases will be considered.

Truck Shock

Truck shock input was obtained from ANSI N 14.23(5 ). This standard specifies shock loads that correspond to normal transport over rough roads or minor accidents such as backing into a loading dock. Since the TN-68 cask will be transported on interstate highways or major good roads, the shock loads will not be applied continuously to the normal transport mode for the package. The fatigue calculation assumes an average trip of 600 miles in 12 hours.

Assume the driver stops and leaves the interstate every 4 hours and assume that one shock could be experienced during each of these stops. The return trip package behavior is assumed to be the same as the "loaded" trip even though the cargo is no longer present. Therefore shock loading occurs 3 (shocks per trip) x 2 (round trip) x 150 shipments = 900 cycles.

Reference 3 specifies a peak shock loading of 2.3 gs in the longitudinal direction. The weight of the lid, lid bolts, and cask internals is conservatively assumed to be 89,500 lbs. [Actual weight (specified in Table 2-6) = 12,289 lb. (lid and lid bolts) + 25,868 lb. (basket, rails, and shims) + 1,408 lb. (hold down ring) + 47,940 lb. (fuel assemblies) = 87,505 lb.] The bolt force due to shock is,

(89,500 lb)(2.3 gs) / (48 bolts)(2.414 in2 per bolt) = 1,777 psi.

Rev. 0 4/992.10.2-15

Rail Car Shock

Again, assume 150 round trip shipments, averaging 600 miles each way. Reference 6 reports that there are roughly 9 shock cycles per 100 miles of rail car transport. Therefore the total number of cycles is 600 (miles) x 2 (round trip) x 150 (shipments) x 0.09 (Shocks per mile) = 16,200 cycles.

Reference 6 specifies a peak shock loading of 4.7 g's in the longitudinal direction for rail car transport. Consequently, the bolt force due to rail car shock is

(89,500 lb)(4.7 gs) / (48 bolts)(2.414 in2 per bolt) = 3,630 psi.

Vibration

Since vibration accelerations are higher on a truck than on a rail car, the truck vibration loads are considered bounding. According to ANSI N14.23(5 ), the peak vibration load at the bed of a truck in the longitudinal direction is 0.3 g's. This results in a stress of 232 psi, which is negligible for a high strength bolt.

2.10.2.6.4 Pressure and Temperature Fluctuations

The lid bolt material is SA-540 GR. B24 CL. 1, 2Ni ½/2Cr 1/3Mo, which is group E in the coefficients of thermal expansion tables in reference 4. The lid and flange are both made of SA-350 Gr. LF3, 3½/2Ni, which is also group E. Therefore the lid bolts and all of the materials it contacts have the same coefficient of thermal expansion (6.78xi0 6 in. in.-:' OF @ 300OF). Consequently thermal load will cause no stress in the lid bolts.

The pressure fluctuation is conservatively assumed to be the maximum operating pressure, 100 psi, which is far greater that the actual operating pressure. Since the stress intensity in the lid bolts is linearly proportional to the internal / external pressure difference, the stress intensity due to a 100 psi. internal load is,

16,480 psi. x I =psi. -13,190 psi. 125psi.

The pressure fluctuation is assumed to occur once per round trip, since there is no cargo, and therefore no pressurization, during the return trip. So the total number of cycles of pressure' fluctuation is 150.

Rev. 0 4/992.10.2-16

2.10.2.6.5 1 Foot Normal Condition Drop

"The normal condition drop consists of a 1 foot drop in an orientation that results in the most damage. For the side drop the resulting shear load is taken entirely by the lid / flange interface. For the end drop, the load is transferred to the cask body via the impact limiters, protecting the bolts. Therefore the worst case scenario is taken to be roughly a 600 CG over comer drop. The

normal condition impact loads calculated in Section 2.10.2.2.6 are,

Fa = 14,280 lb./bolt, F, = 0 lb./bolt,

Ff = 2,876 lb./bolt, and Mf = 27,280 lb./bolt.

Mbb = 0.2818 Mf.

The lid bolt diameter is 1.7532 in. Therefore, from reference 1, we get the following stresses.

Sb0 = 1.2732F-- = 1.2732 14,280 = 5,915psi, D a 1.75322

bb =10. 18 6 Mb-b -=10. 18 6 0.2818(27,280) -14,530 psi, D • 1.7532 2

Since internal pressure causes no bolt torsion, and all shear loads are taken by the lid shoulder,

Sbs = 0, and Sb = 0.

S.1. = Sbi = [(Sba + Sbb) 2 + 4 (Sbs + Sbt)2]0 5 = [(5,915 + 14,530)2 + 4(0)2]0.5 = 20,450 psi.

Conservatively assume that the cask is dropped once per shipment, resulting in 150 normal condition drops before the lid bolts are changed.

Rev. 0 4/992.10.2-17

2.10.2.6.6 Damage Factor Calculation

The following damage factors are computed based on the stresses and cyclic histories described above, a fatigue strength reduction factor, KF, of 4(10), and the fatigue curve shown in Table 1-9.4 of ASME Section III Appendices(7 ).

Stress S.L x KF Cycles Damage Event Intensity (psi.) Sa (psi.) Factor

(psi.) n N

Operating 60,910 243,600 131,500 300 500 0.60 Preload

Test 16,480 65,920 35,600 1 7,000 0.00 Pressure

Truck 1,77-7 7,108 7,671 900 0 0.00 Shock

Rail Car 3,630 14,520 15,670 16,200 2x10 0.08 Shock

Pressure and 13,190 52,760 56,940 150 2,600 0.06 Temperature

1 Foot Normal 20,450 81,800 44,140 150 4,400 0.03

Condition Drop

Y_ 0.77

Here, n is the number of cycles, N is taken from Figure 1-9.4 of reference 7, and Sa is defined in the following way.

If one cycle goes from 0 to +S.L,

Sa = (1/2) x S. X KF x KE.

If one cycle goes from -S.L to + S.L,

Sa = S.L x KF x KE.

Where KE is the correction factor for modulus of elasticity, 30x10 6 / 27.8x 106 1.08 (7).

The above table shows that the total damage factor is less than one. Therefore the TN-68 cask lid bolts will not fail due to fatigue.

Rev. 0 4/992.10.2-18

Minimum Engagement Length for Bolt and Flange

For a 1%- 8UN - 2A bolt, the material is SA - 540 GR. B24 CL. 1, with

S, = 165 ksi, and Sy = 150 ksi (at room temperature)

The flange material is SA - 350 GR. LF3, with

S, = 70 ksi, and Sy = 37.5 ksi (at room temperature)

The minimum engagement length, Le, for the bolt and flange is (Ref. 4, Page 1149),

L, 2A, 3.146K,,.[! +.57735n(E~min -K a

Where,

A, = tensile stress area = 2.414 in.2 , n = number of threads per inch = 8, K ,,= maximum minor diameter of internal threads = 1.765 in., Esai,, = minimum pitch diameter of external threads = 1.7838 in., ---Dmi= minimum major diameter of external threads = 1.8577 in. (Ref. 4, p. 1294)

Substituting the values given above,

Le = 2(2.414) = 1.484 in. 3.146(l.765)[2 +.57735(8)(1.7838-1.765)]

A correction factor to account for different bolt and mating hole materials, J, is computed in the following way.

J = An x SWe. (Ref. 4, p. 1294)

Where, Se is the tensile strength of the external thread material, and Si is the tensile strength of the internal thread material.

As = shear area of external threads = 3.1416 nL, K,,ma [ 1/(2n) + .57735 (Es r- Kn - ,)]

An = shear area of internal threads = 3.1416 nLe Ds mi, [ 1/(2n) + .57735(Dsi., - E,, ,,)]

Rev. 0 4/99

2.10.2.7

2.10.2-19

For the bolt / Helicoil insert connection:

E.,mo = maximum pitch diameter of internal threads = 1.8038 in.(Ref. 4, p. 1294).

Therefore,

S= 3.1416(8)(l.484)(l.765)[1/(2x8) + .57735 (1.7838 - 1.765)] = 4.829 in.2

A, = 3.1416(8)(1.484)(1.8577)[1/(2x8) + .57735 (1.8577 - 1.8038)] = 6.487 in.2

So,

4.829(165.0)0614

6.487(200.0)

Since the tensile strength of the internal threads (Helicoil insert 1 7/8" -8 (4190-30 CN), Ref. 9) is 200 ksi, which is higher than the bolt material tensile strength (165 ksi), the reduction factor, J, decreases the required engagement length. Consequently, the reduction factor is removed from this analysis for conservatism. Therefore, the minimum required engagement lengthbetween bolt and insert, Q = Le = 1484 in.

The actual minimum engagement length = 2.79 in. > 1.484 in.

For bolt / flange connection:

Since the threaded insert has a very high tensile strength (200 ksi.), it does not bound the required engagement length. The following analysis computes the required engagement length based on the interaction between the bolt and the flange, canceling out the effect of the high strength insert.

Since the tap used on the flange for the 1 7/8" -8, 4190-30 CN helicoil insert is a standard 2" 8 pitch tap (Ref. 8), A, for the flange internal threads is calculated as follows.

Dmmrin = 2.000 in., and E, , = 1.9264 in. (Ref. 4)

A= 3.1416(8)(1.484)(2.000)[1/(2x8) + .57735 (2.000 - 1.9264)] = 7.8319 in.2

Therefore,

= 4.829(165.0) 1 45 7.8319(70.0) 3,and

Q = Le J= (1.484)(1.453) = 2.157 in.

The actual minimum engagement length = 2.79 in. > 2.157 in.

Rev. 0 4/992.10.2-20

2.10.2.8 Conclusions

1. Bolt stresses meet the acceptance criteria of NUREG/CR-6007 "Stress Analysis of Closure Bolts for Shipping Casks".

2. A positive (compressive) load is maintained during normal and accident condition loads since bolt preload is higher than all applied loads.

3. If the TN-68 cask lid bolts are replaced after every 150 round trip shipments, they will not fail due to fatigue during transport.

4. The bolt, insert, and flange thread engagement length is acceptable.

Rev. 0 4/992.10.2-21

2.10.2.9 References

1. Stress Analysis of Closure Bolts for Shipping Cask, NUREG/CR-6007, 1992.

2. High Performance Sealing, Metal Seals Helicoflex Catalog, Helicoflex Co., Boonton, N.J., ET 507 E 5930, Catalog Reference HEL 1.

3. American Society of Mechanical Engineers, ASME Boiler and Pressure Vessel Code, Section Hi, Part D, 1995 with 1996 addenda.

4. Machinery Handbook, 21st Ed, Industrial Press, 1979.

5. Draft American Standard Design Basis for Resistance to Shock and Vibration of Radioactive Material Packages Greater than One Ton in Truck Transport, ANSI N14.23,

6. Shock and Vibration Environments for Large Shipping Containers on Rail Cars and Trucks, NUREG 766510.

7. American Society of Mechanical Engineers, ASME Boiler and Pressure Vessel Code Section III, Division 1, Appendix, 1995 with 1996 addenda.

8. Helicoil Catalog, Heli-Coil 8-Pitch Inserts, Bulletin 913B.

9. Design Criteria for the Structural Analysis of Shipping Cask Containment Vessels, U. S. Nuclear Regulatory Commission, Regulatory Guide 7.6, Revision 1, March 1978.

10. Design Criteria for the Structural Analysis of Shipping Cask Containment Vessels, U. S. Nuclear Regulatory Commission, Regulatory Guide 7.6, Revision 1, March 1978.

11. American Society of Mechanical Engineers, ASME Boiler and Pressure Vessel Code Section III, Division 1, Subsection NB 6220, 1995 with 1996 addenda.

Rev. 0 4/992.10.2-22

TABLE 2.10.2-1

DESIGN PARAMETERS FOR LID BOLT ANALYSIS

* Db Nominal diameter of closure bolt; 1.875 in.

* K Nut factor for empirical relation between the applied torque and achieved preload,

used 0.1 for neolube

* Q Applied torque for the preload (in.-lb.)

* Dib Closure lid diameter at bolt circle, 75.88 in.

* Dig Closure lid diameter at the seal (outer) = 72.90 in.

E, Young's modulus of cask flange material, 27.8 x 106 psi.

E1 Young's modulus of lid material, 27.8 x 106 psi.

0 Nb Total number of closure bolts, 48 * N.l Poisson's ratio of closure lid, 0.3 * Pei Inside pressure of cask, 100 psig. * Dio Closure lid diameter at outer edge, 79.88 in.

* Pli Pressure inside the closure lid, 100 psig.

* tc Thickness of cask wall, 6.0 + 1.5 = 7.5 in.

* ti Thickness of lid, 9.5/5.0 in. * lb Thermal coefficient of expansion, bolt material, 6.27 x 10.6 at R.T., 6.78 x 10-6

in. in.-' OF' at 300-F 0 lc Thermal coefficient of expansion, cask 6.27 x 10-6 at R.T., 6.78 x 10-6

in. in.-' OF at 300°F * 11 Thermal coefficient of expansion, lid 6.27 x 10-6 R.T., 6.78 x 10-6 in. in."' OF' at

300OF * Eb Young's modulus of bolt material, 27.8 x 106 psi.

* ai Maximum rigid-body impact acceleration (g) of the cask

* DLF Dynamic load factor to account for any difference between the rigid body

acceleration and the acceleration of the contents and closure lid = 1.1

& Wc weight of contents = 46,920 (fuel) + 30,320(basket)** = 77,240 lbs.

& W, weight of lid = 12,074 lbs., say 12,100 lbs.

0 Wc+W, 77,240 + 12,074 = 89,314 lbs., say 89,500 lbs.

* xi Impact angle between the cask axis and target surface

* Sy1 Yield strength of closure lid material, 37,500 psi.

* S.1 Ultimate strength of closure lid material, 70,000 psi.

* Syb Yield strength of bolt material (see Table 2.10.2-3)

* Sub Ultimate strength of bolt material (see Table 2.10.2-4)

* P1, Pressure outside the lid * Lb Bolt length between the top and bottom surfaces of closure, 5.0 in.

* Pun Maximum impact force that can be generated by the puncture bar during a normal

impact * Dpb Puncture bar diameter, 6 inches as per 10 CFR 71.73 (c) (3)

** Conservatively using higher basket weight for lid bolt analysis.

Rev. 0 4/99

TABLE 2.10.2-2

BOLT DATA (Ref. 1, Table 5.1)

Bolt: 1 7/8".- UN8 - 2A

N: no of threads per inch =8

P: Pitch = 1/8 .125 in.

Db: Nominal Diameter = 1.875 in.

Db,,: Bolt diameter for stress calculations = Db - .9743p = 1.875 - .9743 (.125) = 1.7532 in

Stress Area = 7c/4 (1.753) 2 = 2.414 in2

Rev. 0 4/99

TABLE 2.10.2-3

ALLOWABLE STRESSES IN CLOSURE BOLTS FOR NORMAL CONDITIONS

(MATERIAL: SA-540 Gr. B24 CL.1)

Temperature Yield Stress (1) Normal Condition Allowables (OF) (ksi)

Ftb (2,4) Fb (3.4) S.L (5) (ksi) (ksi) (ksi)

100 150 100.0 60.0 135.0

200 143.4 95.6 57.4 129.1

300 138.6 92.4 55.4 124.7

400 134.4 89.6 53.8 121.0

500 130.2 86.8 52.1 117.2

600 124.2 82.8 49.7 111.8

Notes:

1. Yield stress values are from ASME Code, Section II, Table y-1(3)

2. Allowable Tensile stress, Fib = 2/3 Sy (Ref. 1, Table 6.1)

3. Allowable shear stress, Fvb = 0.4 Sy (Ref. 1, Table 6.1)

4. Tension and shear stresses must be combined using the following interaction equation:

2 2 - + <' 1.0 (Ref. 1)

5. Stress intensity from combined tensile, shear and residual torsion loads, S.I. < 0.9 Sy

(Ref. 1, Table 6.1)

Rev. 0 4/99

TABLE 2.10.2-4

ALLOWABLE STRESSES IN CLOSURE BOLTS FOR ACCIDENT CONDITIONS

(MATERIAL: SA-540 Gr. B24 Cl.1)

Temperature Yield Stress (1) Accident Condition Allowables (OF) (ksi)

0.6 Sy (3) Fib(2,4) Fb (3,4) (ksi) (ksi) (ksi)

100 150.0 90.0 115.5 69.3 200 143.4 86.0 115.5 69.3 300 138.6 83.2 115.5 69.3 400 134.4 80.6 115.5 69.3 500 130.2 78.1 115.5 69.3 600 124.2 74.5 115.5 69.3

Notes:

1. Yield and tensile stress values are from ASME Code, Table Y- 1 (3). Note that S, is 165 ksi at all temperatures of interest.

2. Allowable Tensile stress, Fb = MINIMUM(0.7 S,,, S,), where 0.7 S,, = 0.7 (165) = 115.5 ksi. (Ref. 1, Table 6.3)

3. Allowable shear stress, Fr,, = MINIMUM(0.42 S,, 0.6 S.), where 0.42 S, = 0.42 (165.) = 69.3 ksi. (Ref. 1, Table 6.3)

4. Tension and shear stresses must be combined using the following interaction equation:

2 2

I tb vb

Rev. 0 4/99

REV. 0 4/99

FIGURE 2.10.2-1 TN-68 CASK LID CLOSURE ARRANGEMENT

lwalcourt

New Stamp

FIGURE 2.10.2-2 TN-68 CASK LID BOLT

REV. 0 4/99

lwalcourt

New Stamp


APPENDIX 2.10.3

TABLE OF CONTENTS

Page

2.10.3 STRUCTURAL EVALUATION OF THE OUTER SHELL

2.10.3.1 Introduction ...................................................................................... 2.10.3-1

2.10.3.2 D escription ....................................................................................... 2.10.3-1

2.10.3.3 M aterials Input D ata ......................................................................... 2.10.3-1

2.10.3.4 A pplied Loads .................................................................................. 2.10.3-1

2.10.3.5 M ethod of A nalysis .......................................................................... 2.10.3-2

2.10.3.6 R esults .............................................................................................. 2.10.3-5

2.10.3.7 R eferences ........................................................................................ 2.10.3-6

LIST OF FIGURES

2.10.3-1 Cask Outer Shell and Connection with Cask Body 2.10.3-2 Finite Element Model Outer Shell 2.10.3-3 Finite Element Model Bottom Comer 2.10.3-4 Finite Element Model Top Comer 2.10.3-5 Internal Pressure (25 psi.) 2.10.3-6 25 psi. & 3g Down (Cask Vertical) 2.10.3-7 25 psi. & 15g Down (Cask Vertical) 2.10.3-8 25 psi. + 5.4g Down & lOg Forward Inertia Load 2.10.3-9 25 psi. & lOg Backward Inertia Load 2.10.3-10 25 psi. & 35g Down Inertia Load

Rev. 0 4/992.10.3 -i


Rev. 0 4/99

APPENDIX 2.10.3

"STRUCTURAL EVALUATION OF THE OUTER SHELL


This section presents the structural analysis of the outer shell of the TN-68 package. The outer shell consists of a cylindrical shell section and closure plates at each end which connect the cylinder to the cask body. The normal loads acting on the outer shell are due to internal and external pressure, normal handling/tiedown loads and 1 foot end/side drops. Maximum membrane plus bending stress intensities due to the pressure difference, handling/tiedown loads and 1 foot end/side drop loads are determined. These stresses are compared to the allowable stress limits in Chapter 2 to assure that the design criteria are met.

2.10.3.2 Description

The outer shell is constructed from low-alloy carbon steel and is welded to the outer surface of the cask body gamma shielding. The cylindrical shell section and the closure plates are 0.75 in. thick. Pertinent dimensions are shown in Fig. 2.10.3-1 and Drawing 972-71-2.

2.10.3.3 Materials Input Data

The outer shell cylindrical section and closure plates are SA-516 Gr. 70. The material properties "are taken from the ASME(1) Code, Section II, Part D. The yield strength of the material is also obtained from the code at a temperature of 250'F.

2.10.3.4 Applied Loads

It is assumed that a pressure of 25 psi may be applied to the inside or outside of the outer shell. This bounding assumption envelopes the actual expected pressures described in Chapter 2.

The handling loads acting on the outer shell are a result of lifting. The weight or inertia g load can include all of the weights of the outer shell, neutron resin shield, and aluminum containers. The most severe Normal Service (Design and Level A) Condition load is assumed 3 G inertia load in the vertical lifting orientation, 1 foot end drop and side drop. The shell is also analyzed for 2/5/10 G tiedown loads when the cask is oriented horizontally to ensure it is not damaged during transportation.

* Cask in the Vertical Orientation - Stress due to 25 psi pressure - Stress due to 25 psi pressure and 3G inertia load (lifting) - Stress due to 25 psi pressure and 1 foot end drop (15G)

Rev. 0 4/992.10.3-1

Cask in the Horizontal Orientation - Stress due to 25 psi pressure - Stress due to 25 psi pressure, 2G vertical, 5Glateral accelerations and lOG

longitudinal forward acceleration - Stress due to 25 psi pressure, 2G vertical, 5G lateral accelerations and lOG

longitudinal backward acceleration - Stress due to 25 psi pressure and 1 foot side drop (35G)

2.10.3.5 Method of Analysis

ANSYS Model

A finite element model is built for the structural analysis of the outer shell and closure plates. The outer shell and closure plates are modeled with ANSYS(2) Plane 42 elements. The element used is an axisymmetric element. Double nodes are created at weld locations. The partial penetration welds are simulated by coupling the nodes where welds exist. In finite element model for pressure run, the entire ½ inch cylinder thickness was conservatively reduced to 0.375 inch to simulate the longitudinal partial penetration weld. The basic geometry of the outer shell and weld sizes used for analysis are shown in Figure 2.10.3-1. The finite element model is shown in Figures 2.10.3-2, -3, and -4.

Cask in the Vertical Orientation

0 Stresses due to 25 psi Pressure

An external pressure of 25 psi will not induce any load or stress in the outer shell since it is in contact with and supported by the resin filled aluminum containers.

An internal pressure of 25 psi is used as the maximum pressure acting in the inner surface of the outer shell as shown on Figure 2.10.3-5. The maximum stress intensity for this load case is 4.5 ksi.

• Stress due to 25 psi pressure and 3G inertia load (lifting - cask in the vertical orientation)

The weight of the resin and aluminum containers is modeled as an additional pressure on the bottom inner surface as shown on Figure 2.10.3-6. The effect of the outer shell dead weight is accounted for by using a 3G gravitational load in the longitudinal direction. The maximum stress intensity for this load case is 9.1 ksi.

* Stress due to 25 psi pressure and 1 foot end drop (15G) (cask in the vertical orientation)

The weight of the resin and aluminum containers is modeled as an additional pressure on the bottom inner surface as shown on Figure 2.10.3-7. The effect of the outer shell dead weight is accounted for by using a 15G gravitational load in the longitudinal direction. The maximum stress intensity for this load case is 27.7 ksi.

Rev. 0 4/992.10.3-2

Cask in the Horizontal Orientation

* Stresses due to 25 psi Pressure

The stress due to 25 psi internal pressure is the same in both horizontal and vertical orientations (4.5 ksi).

* Stress due to 25 psi pressure, 2G vertical, 5G lateral accelerations, and lOG longitudinal

forward acceleration (shown in Figure 2.10.3-8)

Again, the stress due to 25 psi internal pressure is 4.5 ksi.

The vertical and lateral accelerations are combined such that, g = (2.02 + 5.02)1/2 = 5.4G. When calculating the stress due to a 5.4G inertia load, it is conservatively assumed that the weight of the outer shell, resin, and aluminum containers is uniformly distributed over the 160 in. length and over a 600 arc only. Therefore, the equivalent pressure applied to the outer shell is:

Weight of outer shell: 11.2 kips Weight of resin: 13.9 kips Weight of alum. Containers: 2.5 kips

Ptrsvcrsc = (11.2 + 13.9 + 2.5)(5.4)(1000)(360)/(7rt)(96.5)(160)(60) = 18.4 psi

For the loading due to 1OG longitudinal forward acceleration, the weight of the resin and "aluminum boxes is modeled as an additional pressure on the forward (lid side) inner surface of the outer shell.

Weight of resin: 13.9 kips Weight of alum. Containers: 2.5 kips Plongitudinal = (13.9 + 2.5)(10)(1000)/(7t)(48.25' - 42.25 2) = 97 psi

The effect of the outer shell dead weight is accounted for by using a lOG gravitational load in the longitudinal direction. The maximum stress intensity for this load case is 28.3 ksi.

* Stress due to 25 psi pressure, 2G vertical, 5G lateral accelerations, and lOG longitudinal backward acceleration (shown in Figure 2.10.3-9)

The loading due to 25 psi internal pressure, 2G vertical, and 5G lateral is the same as it is in the previous case.

For the loading due to lOG longitudinal backward acceleration, the weight of the resin and aluminum boxes is modeled as an additional pressure on the forward (rear side) inner surface of the outer shell.

Rev. 0 4/992.10.3-3

Weight of resin: 13.9 kips Weight of alum. Containers: 2.5 kips Pequipment = (1:3.9 + 2.5)(10)(1000)/(it)(48.252 - 42.252) = 97 psi

The effect of the outer shell dead weight is accounted for by using a lOG gravitational load in the longitudinal direction. The maximum stress intensity for this load case is 24.4 ksi.

* Stress due to 25 psi and 1 foot side drop (35G)

The loading scheme applied to the model to simulate a 1 foot side drop is depicted in figure 2.10.3.10.

The stress due to 25 psi internal pressure is the same in both horizontal and vertical orientations (4.5 ksi). When calculating the stress due to a 35G inertia load, it is conservatively assumed that the weight of the outer shell, resin, and aluminum containers is uniformly distributed over the 160 in. length and over a 60' arc only. Therefore, the equivalent pressure applied to the outer shell is:

Weight of outer shell: 11.2 kips Weight of resin: 13.9 kips Weight of alum. Containers: 2.5 kips

Pequipment = (11.2 + 13.9 + 2.5)(35)(1000)(360)/(7t)(96.5)(160)(60) 119.5 psi

An internal pressure of 25 psi and additional 119.5 psi pressure acting in the inner surface of the outer shell is applied in this load. The stress results from this additional 119.5 psi load is approximated by an internal pressure and applied on the full 360' inner surface of the outer shell. Therefore, the stress due to the this 35G inertia load can be ratioed from the 25 psi internal pressure case and is:

(T = 4.5 (119.5)/25 = 21.51 ksi

The total stress intensity is (4.5 ksi + 21.51 ksi) 26.01 ksi.

Rev. 0 4/992.10.3-4

Summary of the Maximum Stress Intensities

Based on the above calculations the stress intensities are summarized in the following table:

2.10.3.6 Analysis Results

All the above calculated maximum stress intensities are less than the allowable stress of 33.75 ksi (1.5 Smn, SA-516, Gr. 70 at 250 0F).

Rev. 0 4/99

Loading Stress Intensities

25 psi Internal Pressure 4.5 ksi 25 psi + 3G Down 9.1 ksi

(Cask in Vertical Orientation) 25 psi + 15G Down 27.7 ksi (Vertical End Drop)

25 psi + 2G vertical and 5G Lateral + lOG Longitudinal (Forward) 28.3 ksi

(Cask in Horizontal Orientation) 25 psi + 2G vertical and 5G Lateral +

lOG Longitudinal (Backward) 24.4 ksi (Cask in Horizontal Orientation)

25 psi + 35G Down 26.0 ksi (Side Drop)

2.10.3-5

2.10.3.7 References

1. ASME Code, Section II, Part D, 1995 including 1996 addenda.

2. ANSYS Engineering Analysis System, Users Manual for ANSYS Revision 5.2, 1995.

Rev. 0 4/992.10.3-6

REV. 0 .4/99

1_'I X,See Figure 2.10.3-4 for details

See Figure 2.10.3-3 for details

TN-68 OUTER SHELL, FINITE ELEMENT MODEL

FIGURE 2.10.3-2 FINITE ELEMENT MODEL

OUTER SHELL REV. 0 4/99

REV. 0 4/99

-4 "1 11 11 cc

r"

rM

I-.€

"I. I

1 1

Simulation of weld connecting outer shell and closure.plate by coupling nodes at this locations

FIGURE 2.10.:3-4 FINITE ELEMENT MODEL

TOP CORNER

REV. 0 4/99

REV. 0 4/99

FRONT/ TOP END

p - 25 psi

p - 2 5 + 2 9 - 5 4 p s i -\

ACCELERATION

REAR/ BOTTOM END

p - (2.5 + 13.9) x 3 x 1000/ (48.252- 42.252)w;1w 29 psi

FIGURE 2.10.3-6 25 psi & 39 DOWN (CASK VERTICAL) i

REV. 0 4/99

FIGURE 2.10.3-7 25 psi & 159 DOWN (CASK VERTICAL)

REV. 0 4/99

REV. 0 4/99

FIGURE 2.10.3-9 25 psi & lOg

BACKWARD INERTIA LOAD REV. 0 4/99

FIGURE 2.10.3-10 25 psi & 359 DOWN

INERTIA LOAD REV. 0 4/99


APPENDIX 2.10.4

TABLE OF CONTENTS

Page

FRACTURE TOUGHNESS EVALUATION OF TN-68 CASK

Introduction ...................................................................................... 2.10.4-1

Fracture Toughness Evaluation of Containment Boundary ............. 2.10.4-1 Fracture Toughness Evaluation of the Gamma Shield ..................... 2.10.4-1

R eferences ...................................................................................... 2.10.4-11

LIST OF FIGURES

Locations of Fracture Toughness Evaluations (TN-68 Gamma Shield) Charpy V-Notch Test Results for SA-266 Class 2

Rev. 0 4/99

2.10.4

2.10.4.1 2.10.4.2 2.10.4.3 2.10.4.4

2.10.4-1 2.10.4-2

2.10.4-i


Rev. 0 4/99

APPENDIX 2.10.4

"FRACTURE TOUGHNESS EVALUATION OF TN-68 CASK


This appendix documents the fracture toughness requirements of the TN-68 containment

boundary and also calculates the allowable flaw sizes of the gamma shield and welds. The

results of the flaw sizes can be used to develop an appropriate inspection program and to select

an appropriate technique to properly inspect the cask. It can also be used as an initial screening

criteria to disposition any indications that are detected during inspection.

2.10.4.2 Fracture Toughness Evaluation of Containment Boundary

The TN-68 containment boundary material is a ferritic steel and is therefore subject to fracture

toughness requirements in order to assure ductile behavior at the lowest service temperature (LST) of -20'F. The containment boundary materials (including lid bolts) are selected to meet

the fracture toughness criteria of ASME Code Section Ill, Division 3(1), Subsection WB. The

cask shell and bottom plate are 1.5 inches thick, the flange is 7.5 inches thick, and the lid plate is

5 inches thick. Therefore, by interpolating between values provided in Table WB-2331.2-1 of

the Section III, Subsection WB (Para. WB-2300), the nil ductility transition temperatures (TNDT)

of the containment boundary materials are:

Shell and bottom plates: -80'F

Flange: -133 0F

Lid plate: -126 0F

In addition to determining the nil ductility transition temperature, charpy v-notch testing is

performed at a temperature no greater than 60'F above the TNDT. The acceptance criteria is that

the material exhibit at least 35 mils lateral expansion and not less than 50 ft-lbs absorbed energy.

This testing is sufficient to ensure that the containment boundary materials will not be

susceptible to brittle fracture at -20°F.

The fracture toughness requirements of the lid bolts meet the criteria of ASME Code, Section III,

Division 3, Subsection WB (Para. WB-2333). Charpy v-notch testing is performed at -20'F.

The acceptance criteria is that the material exhibit at least 25 mils lateral expansion (Table W\VB2333-1).

2.10.4.3 Fracture Toughness Evaluation of the Gamma Shield

The gamma shield shell is forged from SA-266 Class 2 material. The bottom shield plate is

constructed from either SA-105 forging or SA-516 Grade 70 plate material, and the top shield

plate (plate welded to the bottom of the lid) is constructed from either SA-266 Class 2 material, or SA-516 Grade 70 material. The main function of the gamma shield is to provide shielding. It

is not part of the containment, and its shielding properties are not temperature dependent.

Rev. 0 4/992.10.4-1

Preliminary charpy test data of the samematerial (SA-266) from a similarly sized shield shell has been provided by one of the material manufacturers for the shield shell, and the results are tabulated below.

Charov V-Notch Test - Results for SA-266 Class 2

Temperature Specimen No. Absorbed Energy (ft-lbs) Avg. of 3 specimens

00C (32-F) V1 63 V2 60

-10°C (14°F) V3 56 V4 50

-20°C (-4°F) V5 45 V6 40

-30 0C (-22-F) V7 18 V8 20

-400 C (-40°F) V9 17 V10 10

The TN-68 package is designed for an ambient temperature of -20'F. As can be seen from the materials testing, even at temperatures as low as -20'F the gamma shielding has relatively good charpy impact properties. It is unlikely that the gamma shield would reach -20'F, since the heat load of the fuel would keep the cask temperatures elevated.

The shielding material is not part of the containment boundary and it is unlikely that the gamma shield would reach -20'F. A fracture of the gamma shield will have no safety implications. However, Transnuclear has performed a fracture mechanics evaluation of the TN-68 package gamma shield based on a service temperature of -20'F. The work includes the following:

* Methodology * Loadings * Material fracture toughness * Fracture toughness criteria * Primary stress criteria * Allowable flaw size calculations * Conclusions * NDE Inspection Plan

Rev. 0 4/992.10.4-2

Methodology

The allowable flaw sizes were determined using linear elastic fracture mechanics (LEFM)

methodology from Section XI of ASME(2) Code (1989). Flaws in the welds, if they occur, are

welding defects, rather than initiated cracks. There is not an active mechanism for crack

initiation and growth at any of the weld locations. Thus, the calculated allowable flaw sizes can

be used for screening the gamma shield during fabrication.

Loadings

The following table lists the maximum membrane and bending stresses at the gamma shield

under normal and accident conditions. Figure 2.10.4-1 shows the selected locations on the

gamma shield numbered 1 through 6 for fracture toughness analysis. These locations were

selected to be representative of the stress distribution in the gamma shield with special attention

given to areas subject to high stresses and weld locations. The maximum stress may occur at a

different location for different load combinations (bolt preload, pressure, temperature, lifting

load, fabrication, end drop, CG over comer drops, side drop and slap down drop).

Rev. 0 4/992.10.4-3

Summary of Stress Components (TN-68 Gamma Shield)

Normal Condition Accident Condition Axial Stress Hoop Stress Axial Stress Hoop Stress Residual

(ksi) (ksi) (ksi) (ksi) Stress Location am Ub am 'b am ab am ab (ksi)

(Fig.2.10.4-1) 1 -0.27 1.82 6.68 0.80 -4.96 0.72 -25.79 0.19 8.0 2 0.15 0.15 -0.91 0.18 2.36 2.13 -9.05 0.83 36.0 3 7.14 5.71 7.02 5.76 14.42 7.65 9.55 8.27 0 4 -4.43 1.62 -6.21 1.24 -15.71 9.37 -14.30 2.32 0 5 0.76 1.65 1.26 1.28 4.61 10.2 7.68 4.09 0 6 0.45 1.54 -0.43 0.53 2.31 8.86 0.88 3.13 8.0

The gamma shield welds at locations 1 and 6 are partially stress relieved. However, the lower gamma shield weld to the bottom shield plate (location 2) does not undergo stress relief. Weld residual stress is included in the calculations for all weld locations. The weld residual stress is reduced due to the stress relief at all weld locations except the weld at location 2.

Weld residual stresses are steady state secondary stresses. The ASME Code does not prescribe limits for weld residual stresses. These stresses are displacement (or strain) controlled, and are self equilibrating through the weld thickness. For the purpose of this calculation, residual stresses are conservatively assumed to be a constant tensile magnitude of 36 ksi at location 2. This value corresponds to the minimum specified yield stress of the base material (SA-266, Class 2). For other welds, 'which have been stress relieved, it is conservatively assumed that not all of the weld residual stress is relieved during the stress relief process. A stress value of 8 ksi has been included for welds at locations 1 & 6 for fracture toughness evaluations. This is similar to the procedure used in evaluation of reactor pressure vessels to account for the potential for residual stress after post weld heat treatment.

The K factor due to residual stresses is applied with a safety factor of 1, as recommended in ASME, Section XI, Appendix H, Paragraph H-7300. The total K1 (applied) is determined from membrane, bending, and residual stresses.

Rev. 0 4/992.10.4-4

Material Fracture Toughness

"The gamma shield shell is a forged cylinder, nominally 6 inches thick by 180.15 inches long,

made from SA-266, Class 2 material. The welding at the top flange and bottom plate may be

performed using SAW, FCAW, or SMAW processes.

Figure 2.10.4-2 shows a summary of the Charpy impact testing data tabulated above. The actual

data points are shown along with a smoothed line that connects the average values at each test

temperature. This data demonstrates that a lower bound Charpy impact value of 18 ft-lbs is

appropriate for an exposure temperature of -20'F.

The Charpy impact measurement may be transformed into a fracture toughness value by using an

empirical relation given below (Ref.3):

Kid = [5E(Cv)]" 2 = 51,960 psi-(in)112

Where

Kid = Dynamic Fracture Toughness (based on crack arrest), psi -(in)12

E = Modulus of Elasticity, 30 x 106 psi Cv = Charpy Impact Measurement, 18 ft-lbs

For conservatism, the above calculated Kid was reduced by another 10% to 47 ksi-(in)1 2

(corresponding to 15 ft-lbs charpy values at -20'F) for fracture toughness evaluations.

Both the FCAW and SMAW electrodes used in the gamma shield weldments are alloyed with

manganese, nickel, chromium, and vanadium. They are essentially matching filler metals for alloys such as ASME SA-533 Gr. B, the most commonly used reactor pressure vessel steel. The higher alloy content of the FCAW and SMAW electrodes and their typical usage in applications where good toughness is required indicate that the expected fracture toughness values for the

FCAW and SMAW weld fillers are as good as or better than that of the SA-266 material. Use of the fracture properties from the wrought material for locations at or near the weld joints is conservative.

Fracture Toughness Criteria

Using the rule of Section XI, IWB-3613, the limiting fracture toughness values are reduced by a

factor of 4110 for the normal condition and 42 for the accident condition, to define the limiting

allowable Kallowable. That is,

Kallowable < Kia/(4 10) = 14.86 ksi-qin for normal conditions

Kanlowable < K ic /(2) = 33.2 ksi-4in for accident conditions

Rev. 0 4/992.10.4-5

Where:

Kia - the available fracture toughness based on crack arrest

K ic = the available fracture toughness based on crack initiation

The Kid value (47 ksi-in" 2) calculated above is the available fracture toughness based on crack arrest for the corresponding crack tip temperature and is conservatively used for the following normal and accident condition fracture toughness evaluations.

Primary Stress Criteria

ASME Section XI, IWB-3610 requires that any flaw evaluation include verification that the primary stress limits of ASME Code Section III continue to be met for the flawed component. The following formula is used which conservatively assumes that the available cross section is equal to the original thickness minus the allowable flaw depth.

a., = t(l - S/San)

Where:

awI = allowable flaw depth based on ASME Code Section III limits

t = original local thickness S = maximum calculated local stress intensity S~al = allowable stress intensity per ASME Section III.

All stresses are considered to be pure tensile membrane stresses. The stresses increase linearly with decreasing wall thickness.

Allowable Flaw Size Calculation

Using the above load definitions and fracture toughness values, a series of allowable flaw size calculations was performed using the method from ASME Section XI, Article A-3000, "Method for K, Determination". The total applied stress intensity Ki (applied) is determined from the membrane, bending, and residual stresses as:

K1 (applied - normal) = K, (membrane) + K, (bending) + KI/410 (residual)

K1 (applied - accident) = K, (membrane) + K, (bending) + KI/42 (residual)

The K1 due to residual stresses is applied with a safety factor of 1, as recommended in ASME Section XI, Appendix H, Paragraph H-7300.

Rev. 0 4/992.10.4-6

- Surface Flaws

For purpose of analysis, the postulated surface flaws are oriented in both the axial and

circumferential directions. The results of the surface flaw size calculations are shown in the

following table.

- Subsurface Flaws

The above discussion addressed the determination of allowable flaw sizes for flaws that are

connected to the surface of the shield shell. The shell or weld could also contain subsurface

defects.

An evaluation of allowable subsurface defects was performed using the same linear elastic

fracture mechanics (LEFM) techniques as described above for surface defects. For this case, a

center cracked panel (CCP) model was used to evaluate an assumed flaw length. The flaw must

be sufficiently embedded such that treatment as a subsurface flaw is justified. In general, if a

flaw is closer to the surface than 0.4 of its half-depth, it must be considered a surface flaw.

The results of the subsurface flaw size calculations are shown in the following table.

Rev. 0 4/992.10.4-7

Allowable Surface Flaws Denth (inches)(TN-68 Gamma Shield)

Location Normal Condition Accident Condition I to Axial -L to Hoop ±I to Axial I- to Hoop

Stress Stress Stress Stress. 1 -- 0.25 ....

(0.33) (0.33) (0.35) (0.35) 2 0..30 0.41 0.23 0.56

(1.16) (1.16) (1.13) (1.13)

3 0.40 0.41 0.63 1.05 (3.60) (3.60) (4.20) (4.20)

4...... -

(5.18) (5.18) (3.56) (3.56) 5 -- 2.25 1.63 1.64

(3.85) (3.85) (2.79) (2.79) 6 0.37 -- 0.37 -

(0.58) (0.58) (0.31) (0.31)

Allowable Sub-Surface Flaws (Embedded) Depth (inches) (TN-68 Gamma Shield)

Location Normal Condition Accident Condition 2- to Axial I- to Hoop _1 to Axial IL to Hoop

Stress Stress Stress Stress 1 -- 0.40 ....

(0.33) (0.33) (0.35) (0.35) 2 0.56 0.64 0.48 0.92

(1.16) (1.16) (1.13) (1.13) 3 1.38 1.40 1.72 2.58

(3.60) (3.60) (4.20) (4.20) 4...... -

(5.18) (5.18) (3.56) (3.56) 5 4.00 3.94 2.26 3.00

(3.85) (3.85) (2.79) (2.79) 6 0.68 -- 0.60 -

(0.58) (0.58) (0.31) (0.31) Note: 1. "--" Indicates that the allowable flaw depth is not limited by fracture mechanics calculation. 2. "( )" Indicates the allowable flaw depth determined by primary stress criteria 3. The actual maximum allowable flaw size is limited by the smaller of the allowable flaw size

limited by fracture mechanics or primary stress criteria.

Rev. 0 4/992.10.4-8

Specific conservatisms included in the above analysis are listed below:

"* All factors of safety on applied stress required by ASME Section XI are included in the evaluation.

"* Weld residual stresses are treated as constant tensile stresses normal to the flaw orientation. Flaws are assumed to be long (infinitely long or full circumference)

"* Lower bound material properties are used.

Conclusions

The gamma shield is not part of the containment boundary. Cracks postulated in the gamma shield will not propagate into the containment boundary due to the geometry of the cask. If the gamma shield were to fracture along the length or around the circumference or around the weld between the gamma shield and top flange, there is no credible mechanism that would result in the gamma shielding separating from the containment vessel. The top shield plate is welded to the lid and is captured by the containment vessel. Therefore, if the weld were to completely fail, the shield plate would still remain inside the confinement boundary and would not lose its shielding capability. Therefore, even if a fracture were to occur in the gamma shield shell or the weld between the gamma shield and top flange or top shield plate or weld between top shield plate and lid, there would be no safety significance, since containment would be maintained, and shielding would remain in place. The one exception is in the region of the weld of the gamma shield shell to the bottom plate. In this region, if the weld were to completely fail, the bottom plate could become detached and have an impact on the shielding capability of the cask.

NDE Inspection Plan

The results of the fracture toughness analysis show that the flaws in the gamma shield shell and top and bottom shield plates which would result in unstable crack growth or brittle fracture are larger than those generally observed in forged steel and plate components. No special examination requirements on the gamma shield shell, top and bottom shield plates are required.

The flaw sizes in the welds that could result in brittle fracture at -20'F will be detected by NDE methods and repaired. If the welds at locations 1 and 6 were to completely fail, there would be no compromise to safety. Therefore, only PT or MT of the final surface is specified.

If the bottom plate weld fails, the bottom plate could become detached and have an impact on the shielding capability of the cask. The minimum allowable flaw sizes for surface and subsurface are 0.23 in. and 0.48 in., respectively. Therefore, the following NDE will be used to ensure that defects of the minimum flaw sizes calculated are detected and repaired prior to use.

Rev. 0 4/992.10.4-9

* PT or MT at weld preparation surfaces (base metal) * PT or MT at root pass 0 PT or MT for each 0.375 inches of weld a PT or MT at final surface

The liquid penetrant or magnetic particle method will be in accordance with Section V, Article 6 of ASME(4 Code.

Rev. 0 4/992.10.4-10

2.10.4.4 References

1. ASME Section III, Division 3, Containment Systems and Transport Packagings for Spent Nuclear Fuel and High Level Radioactive Waste, 1998.

2. ASME Code Section XI, Welding and Brazing Qualifications, 1989.

3. NUREG/CR-1815, Recommendations for Protecting Against Failure by Brittle Fracture in

Ferritic Steel Shipping Containers up to Four Inches Thick.

4. ASME Code Section V, Nondestructive Examination, 1995 with 1996 addenda.

Rev. 0 4/992.10.4-11


Rev. 0 4/99

FIGURE 2.10.4-I LOCATIONS OF FRACTURE TOUGHNESS EVALUATIONS (TN-68 GAMMA SHIELD)

REV. 0 4/99

Figure 2.10.4-2 Charpy V-Notch Test Results for SA-266 Class 2

70

60

40

w 30

S-n.0..•'-.

-40 -30 -20 -10 0 10 20 30

Test Temperature, F

Rev. 0 4/99

TN 68 TRANSPORT PACKAGING APPENDIX 2.10.1 TABLE OF …Shock Loads (Cask Horizontal) 2.10.1-60 Local...

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