2.0 THERMAL CONSTANTS

The thermal constants used in the analysis are grouped Into specific thermal properties of cask materials and effective thermal properties which were derived from correlations. The latter pertain to natural convection as well as to effective radiation In confined spaces.

2.1 Svecific Thermal Properties

Table VIII-2 lists the thermal properties of the cask materials. Specific emissivities are listed in Table VIII-3.

2.2 Effective Thermal Properties

Natural convection coefficients are expressed in terms of the equation

C C (T1 - T0) (1)

whereas radiation coefficients are defined by

hR = 0.174 x 10 8 F (T 12 + T02) (T1 + TO0 (2)

T and T1 are sink and source temperatures, respectively. For the analyses of this report temperatures are in *F or *R and the vaWles for C, b, and F are listed in Table VIII-4. The coefficients are in the units of Btu/hr-ft2 - F.

3.0 ANALYTICAL METHOD

The thermal analyses of the cask are based primarily on digital computer program results with hand calculations providing a significant amount of input to the computer programs used.

VIII- 9

TABLE VI1I-2

MATERIAL PROPMM

Latent Heats/\., Den Spe He em tu Conducvteaersture Trans. Temjierstuw.

Stainless Steel (304) 494.0 0.120 32.0 7:74 32.0

Rd. 1 0.139 732.0 9.43 212.0 12.6 923.0 13.0 1292.0

Lead Rat. 1 703.6 0.0305 32.0 20.0 68.0

0.0315 211.0 19.6 208.0 10.26 521.5 0.0338 621.3 18.3 400.1 0.0340 621.7 16.9 498.9 0.0328 1832.0 14.5 581.0

12.1 630.0 9.68 717.1 9.02 799.3 3.71 980.1 1.65 1276.0

Uranium 1132.0 0.0275 32.0 15.0 32.0 Ref. 1 0.0350 662.0 20.1 732.0

0.0480 1225.4 27.6 1472.0 0.3300 1234.4 0.0450 1243.4 0.0450 1418.0

Dummy. Cavity Fluid* 0.011 1.24 - 0.137 -

Air 0.08053 0.14 0.014 32.0Rd. 1 0.018

0.029212.0 7532.0

Helium Rd. 1& 2 0.011 1.24 0.079 0.0

0.0100 200.0 0.119 400.0 0.137 600.0 0.132 300.0 0.167 1000.0 0.197 1400.0

NeUtr Shield fled (wee Raf. 1 60.0 1.0 - - -.

Fuel " 1.0 36.5 - -Ref. 1& 3

6061-oT Aluminum 119.0 0.20 32.0 96.8 - 167.0 1080.0 Raf. 1 & 3 0.23 392.0

Neunm= Shield Fluid 0.08033 0.24 - - -(AT same as air /•f. 1

* This dummy fluid is used to calculate the fuel element temperature ri0e in a helium environment using a single point model (Section 4.1.S)

tE fifective volumetric heat capacity for typcal UO and Zr. Based only an active fuel length. Conductuvity of fuel is not used. 2

(Continued on page VIII-11.) vm-s0

**Continuation of footnote The product of these two dazdW variables results in 36.5 BTU/ft 3/OF which is actually the volumetric specific beat used in the analysis. This heat capacity was established by volumetric weighting of the Zr clad and U02 fuel heat capacities of a PWR fuel pin assuming a conservatively low teiperatwe of 2120F. Following equation provides the effective beat capacity.

SCeff • ccrVc + /f-fVf = CC, ro2_r_2, +o/fCf ri 2

Vc +Vf ro 2

VYhere:

Ceff = effective volumeztric heat capacity, 36.5 BIU/ft 3/°F

/'1c = clad density, 408.9 #/ft3 (Ref. 3)

Cc - clad specified heat .0732 MTU/lb/IF (Ref. 3) ",f = fuel density - 92M T.D., 608.8 #/ft3 (Ref. 1) Cf h fuel specific beat, .063 BTU/lb/OF (Ref. 1)

ri = fuel pin inner radius .1867 in. (PWR)

ro - fuel pin outer radius, .211 in. (PWR)

Vc - volume of clad per unit length of pin

Vf = volume of fuel per unit length of pin

TABLE VIII-3

EMISSMITIES

Material

Fuel pin

Surfaces exposed to fuel (aluminum, can and can head)

Neutron shield interior Head cavity

Cask surface before and after fire Cask surface during fire

All dry gaps except those with. aluminum surfaces

Aluminum surfaces

Stainless steel surfaces opposite aluminum

E

0.4

0.2

0.8 0.8

0.5 0.8 0.5

0.2

0.5

References

3

6

5

5

5, 6

5 (10 CaR 71) 5, 6

6

5, 6

VIII-If

TABLE VIII-4

CONVECTION AND RADIATION FACTORS

Application C b F

ID single point fuel temperature - -- 0.1514

2D single point

fuel temperature .... 0.1087

1D aluminum to can gap .... 0.1668

2D aluminum to can gap .... 0.167

ID can to inner shell gap - ._ 0.3351

Other gaps not covered above .... 0.333

Wet neutron shield- 45.0 0.3333

Dry neutron shield 0.09 0.3333 0.7

Head cavity -- 0.6207

Outer surface before and after fire 0.18 0.3333 0.5

Outer surface during fire 0.18 0.3333 0.7347

VIII-12

3.1 Hand Calculations

Hand calculations wereperformed to compute:

a. The natural convection coefficients applicable to the fluid

in the neutron shield.

b. The effective radiation form factors internal to the cask

including a single point fuel element temperature rise model.

c. The effective conduction length applicable to the single

point fuel element temperature, rise model.

The calculations are based mainly on the recommendations In the "Cask

Designers Guide" (S) and on references 6 and 7.

3.2 Computer Programs

Detailed steady-state and transient thermal analyses of the cask were (9)

performed with the TRUMP digital computer program supplemented by

the mesh generator program FED (10). Explicit temperature distributions

within the fuel element under dry conditions were calculated using a (11)

modification of a program developed by 1. S. Watson The modified

program is identified as FETA (Fuel Element Thermal Analysis).

3.2.1 TRUMP-FED

TRUMP solves a general nonlinear parabolic partial differential equation

describing flow in various kinds of potential fields, such as fields of

temperature, pressure, and electricity and magnetism.

Thermal problems may include heat transport by conduction, free and

forced convection, radiation, and mass flow.. Heat may be produced or

absorbed by internal heat sources and sinks and phase changes. Boundary

VIM-13

conditions may include insulation, specified surface temperatures or

heat fluxes, or heat transfer by radiation, free convection, or forced

convection. Thermal properties, heat generation rates, mass flow rates,

and surface heat transfer coefficients may be tabulated functions of time

or temperature. Surface or external temperatures may be tabulated

functions of time. Temperature-dependent variables at one spatial

position may be functions of temperatures at other spatial locations.

Initial conditions may vatry with spatial position.

The FED computer program reduces the effort required to obtain the

necessary geometric input for problem solutions using the heat transfer

code TRUMP. FED can properly zone any body of revolution in one, two,

or three dimensions. Rectangular bodies can be only approximated by

using a very large radius of revolution corppared to the total radial thickness

and by considering only a small angular segment in the circumferential

direction.

In FED the regions of a common material are divided Into four-sided

areas. The boundaries of these areas are the required FED input. Each

area is subdivided into volume nodes and the geometrical properties

are calculated. Finally, FED connects adjacent nodes to one another,

using the proper surface area, interface distance, and, if specified,

radiation form factors and interface conductances.

3.2.1.1 Geometric Model

The cask, as evaluated by the ?RUMP program, has been represented by

two models. The first considers one-dimensional radial heat transfer at

the location of maximum power density. Figure VIII-3 shows the nodal

network (29 nodes) used in this model. It is noted that the fuel has

been represented thermally as an annulus which preserves the volume of

VIII-1 4

FIGURE VIII-3 ONE DIMENSIONAL NODAL ýtTWORK"

NODE NO. - Iomwmm W

AMBIENT TEMPERATURE

NEUTRON SHIELD JACKET ( 28. 29)

NEUTRON SHIELD (26)

SS OUTER SHELL ( (, 24 .() )

LEAD (19,20,21.22)

URANIUM ( , 16, 17, 18)

GAP

SS INNER SHELL ( &.12. (@) _- GAP

~ZZ7-SSCAN (.8® ,,. GAP

ALUMINUM (0J. 4, ®) HELIUM CAVITY (2)

"FUEL (SIMULATED) (1)

NOTE: CIRCLED NODES ARE ZERO VOLUME NODES

SCASK CENTERLINE

VIII-1 s

fuel and cladding. The outer radius of this annulus encloses an area 2

equal to the fuel element cross section (8. 375 x 8.375 in • The fuel

element structure has been assumed conservatively to be non-existent.

The- second model is shown in Figure VIII-4. Explicit identification of

the nodes is provided in Figure VIII-S. It represents the upper half of the

cask as an axisymmetric body of revolution starting at the center of

the active fuel. The model consists of 324 nodes. The fuel and cavity interior

have been approximated as in the one-dimensional model. Supporting

structure has been omitted for conservatism. Gaps within the cask

(e. g., the closure head to cask gaps) have been represented explicitly.

Perfect contact has been assumed at the seating surfaces between the cask,

the can and the heads. The outer closure bolt contact surface with the

head has been modeled by a continuous perfect contact annulus.

3.2.1.2 Internal and Boundary Conditions

The entire decay heat load has been concentrated in the fuel for simplicity

and conservatism. In the one-dimensional model the heat load corresponds

to the axial maximum. In the two-dimensional model it is the uniform

design average heat load. A solar heat load has been represented in

all cases except during the fire by a heat generation rate in the very

thin surface nodes of the cask. The balsa crash barriers have been

simulated by insulating areas on the surface of the cask.

The heat transfer within the cavity from fuel to cask is modeled using a

combination of radiation and conduction heat transfer. For the one-dim

ensional TRUMP analysis a single point radiation, conduction model

with a heat transfer area of the fuel element envelope (one side 8.375

inches long) has been generated on the basis of detailed FETA analyses.

In the two-dimensional model the heat transfer coefficients have been

normalized to the cask internal surface rather than the outer surface

of the fuel element. This in effect distributes-the decay heat load

evenly over the cask internal surface which is conservative with respect

VIII-16

C C

-- - - -- --

-- ----- rrrrnn----- - I A A A �------- -

--

- - -

MATERIAL LEGENDS

*WATER

STAINLESS

LEAD

DEPLETED

ALUMINUM

SIMULATED

STEEL NOTEI SEE FIGURE 3EM-5 FOR DETAILED NODE IDENTIFICATION

URANIUM

FUEL

FIGURE Mn-4 TWO-DIMENSIONAL NODAL NETWORK OF CASK END

C

I--

I

431

VIII-18

FIGURE =l-5 SHEET!

20 NODE IDENTIFICATION

to.

400

0~ro

a N C

N C

C

I"

ci

-4 -U ii -4 0 z

a"

C

m 'C

0

urn

m -4 C'

(

Be

It;k-

Li N a

HMt

(I: Id

I

to cask end temperatures. Maximum fuel element temperatures are

established by the one-dimensional model and are thus not affected

by this assumption.

In the neutron shield, natural convection has been assumed to be effective

over, only half the circumference of the cask. Turbulent convection has

been Judged to exist because the effective convection length exceeds

9 inches (7). No convection heat transfer has been assumed in the head cavity because of an optimistic radiation heat transfer coefficient and

the piping, valves and gages in this area.

Radiation in the neutron shield has been taken to be one-dimensional.

The reference heat transfer area is that of the cask outer shell. not the

shield Jacket. In the head cavity radiation is assumed to occur only

in an axial direction. A configuration factor accounting for the connecting

walls between the parallel disk radiating surfaces has been considered.

Radiation and natural convection heat transfer on the surface of the cask

.have been modeled as recommended in the "Cask Designers Guide" (5). All the insulating crash barrier is assumed removed at the inception

of the fire. Temperatures in the head area are thus maximized. As

already noted, the balsa crash barriers have not been explicitly modeled

for the thermal analysis because they act essentially as insulation.

They have been considered in the steady state analys.s by providing an

insulating boundary condition at all cask surfaces covered by balsa wood.

The fire analysis is performed with the neutron shield Jacket intact. This

is the design basis and is also more-likely than complete stripping of

the structure.

VIII-21

Thermal radiation augments the conduction-heat transfar across all gaps

within the cask.

3.2.2 FETA

The FETA (Fuel Element Thermal Analysis) program provides an approximate

two-dimensional steady-state temperature distribution applicable to

the pins of a square array dry fuel element. Heat transfer is by radiation

and conduction. Uniform or variable wall temperature of the surrounding

enclosure may be analyzed. Each pin Is assumed to have one temperature.

All active fuel pins are assumed to generate identical power, however,

any number of pin locations may have zero power. The emissivity within

the array is constant but may be different from that of the enclosure.

FETA is essentially an extension of a fuel elembnt program developed by (11)

Watson

3.2.2.1 Analytical Considerations

The heat transfer in a dry fuel element consists of radiation, conduction

and convection. FETA analyses have been restricted to radiation and

conduction which are readily modeled in two dimensions. Natural convection

may be significant but has conservatively not been incorporated in the

program because of the much greater complexity of an appropriate mathematical

model, the large number of additional uncertainties introduced and what

is currently judged to be a small return on the added effort required

to evaluate natural convectiorf explicitly. Some approximation of natural

convection may be achieved in FETA by adjusting conductivity accordingly.

VIII-22N-

The computation of temperature in FETA is based on an energy balance of a

single fuel pin considering the heat transfer to surrounding pins as far

as two rows removed (Figure VIII-6). Radiation heat transfer occurs

between the reference pin and the four closest (primary) pins, four diagonal

pins, and eight (secondary) pins located two rows from the reference pin.

Conduction heat transfer occurs between the reference pin and the adjacent

four primary pins as well as four diagonal pins. With the exception of the

comer pins, heat transfer at the boundary of the element occurs by

means of one conduction link and one combined radiation link between the

enclosing wall and each pin of the outside row of the fuel element.

One combined radiation link and three conduction links are utilized at

each comer of the fuel element.

The energy equation for each pin of unit length is:

rA 4 N F (T 4 -Tj) (3)

- J=l T where

qi heat generated in pin i, BTU/hr-ft

Ai = surface area of pin i, (v D), ft2/ft

a = Stefan-Boltzmann constant, 0.1713 x 10-8 BTU/hr-ft 2- R

Ti surface temperature of pini, 0 R

T = surface temperature of pin j, 0R

F = general radiation exchange factor

k = gas conductivity, BTU/hr-ft-aF

A = area for conduction heat transfer between pins I and J, ft2 /ft

S = conduction length between pins I and J, ft

VII-23

Figure VI1I-6

HEAT TRANSFER LINKS IN FUEL ELEMENT

-= Conduction Links

--- .-dlaaton Links

R = Reference Pin

P = Primary Pin

D = Diagonal Pin

S = Secondary Pin

VIII-24

N = number of radiation links applicable to pin 1, (N = 16 for internal pins, N = 6 for outside corner pins, N = LO for outside row pins, excluding corners).

N' = number of conduction links applicable to pin 1, (N = 8 for internal pins, N = 6 for all outside row pins).

The general radiation exchange factor is evaluated by

F 1 42) (4)

where

El, E2= surface emissivities, E =E2 for heat transfer between pins; E is pin emmissivity and E2 is wall emissivity for pfn to wall heat transfer.

F =- configuration factor. (See Reference 1 for Justification of using F).

The configuration factors, F are those described in Reference 11

Conduction heat transfer for pins within the fuel element array is approximated by four links to primary pins and four links to diagonal pins. The primary effective area is the projected area of a pin. For the diagonal links the effective area is the difference between the surface

area of a rod and the projected area used for the primary links. The corresponding effective conduction lengths are the moan distances

between pin surfaces. In effect:

Primary A = D (5)

Diagonal A . D = 0.S71 D (6)

Primary 4 = P 4 D (7) a 4

Diagonal = 1. 414 p -- Z D (8) -J 4

VMII-25

Conduction to the enclosing wall is simulated by two primary links and

one diagonal link for comer pins. For the remaining pins of the outer

row-only one primary conduction link is utilized.

The above procedure is considered sufficiently conservative because

convection is not considered. It is judged that the above conservatism

combined with pessimistic emissivities compensates for the possible

unconservatism mentioned in Reference 5.

The wall temperature used in the FETA analysis is the maximum established

in the TRUMP analyses.

4.0 CALCULATIONS AND RESULTS

"This section provides the details of hand calculations and computer

calculations applicable to the thermal analysis of the cask.

4.1 Hand Calculations

4.1.1 Surface Conditions

Under normal steady-state conditions the combined convection-radiation

heat transfer coefficient of the outer surface is established by equations

(1) and (2) and the factors in Table VIII-4. No special hand calculations

are required except that during the fire the emissivity of the cask surface

is given by:

F +- = 0-1. =0.7347 (9) +_ + 1 -1

1 (2.9

The solar heat load is based on double the daily average normal heat flux

2 (5) of 144 Btu/hr-ft applied to the projected area of the cask.

VIII-26

The completely distributed surface heajt flux in the computer calculations

is then 92 Btu/hr-ft 2. With a surface node thickness of 0.0002 ft. 5 3 this results in a surface node heat generation rate of 4.58 x 10 Btu/hr-ft

4.1.2 Wet. Neutron Shield Conditions

Natural convection in the water filled neutron shield is based on a

conventional turbitlent convection coefficient applicable to a vertical

wall (6, 7) which yields

hN 90 (AT) 1/3 (10)

for an average temperature of 325 F. It is noted that the convection

coefficient actually increased with higher fluid temperature. To account

for an effective convection surface area which is half of the computer

model area, the coefficient has been reduced to 45 in the computer

problem. The turbulent correlation has been used since the active

convection height exceeds 9 inches.

4.1.3 Dry Neutron Shield Conditions

With a dry neutron shield, the natural convection coefficient is assumed

to be the same as on the outer surface (5)

h0C = 0.18 (AT) 1 13 (11)

However, to account for the difference of effective heat transfer area and

heat transfer area in the computer problem, the effective natural convection

coefficient Is 0.09.

Radiation heat transfer in the neutron shield is assumed to act only radially.

For surface emissivities of 0.8, the radiation exchange factor based on the

inner surface of the neutron shield is 0. 7.

VIII.-2 7

4.1.4 Gan Conditions

All gaps within the cask are evaluated with simple conduction and one

dimensional radiation. With a surface emissivity of 0.5, the radiation

exchange factor is 0. 333 for most gaps. The ID can to inner shell gap

and the aluminum to can gap are exceptions. With the ID can to inner

shell gapthe variation of surface area has been considered in establishing

a radiation exchange factor of 0.3351. With the aluminum to can gap;

the differing emissivities result in a radiation exchange factor of 0.1668

for the more exact 1D analysis and 0.167 for the 2D analyses. The

smaller surface establishes the reference area for all radiation heat

transfer calculations.

4.1.5 Cask Cavity Conditions

A single point fuel element model considering radiation and conduction

has been uped for all TRUMP calculations. The radiation exchange

factor for this model is based on wall and peak fuel pin temperatures

obtained from a FETA analysis considering only radiation heat transfer.

F 10.63(3413) (1.2) 4 (8.375)

= .1514

[-1 W -l _1081')] (12) .173x10 1586a -10

The effective conduction length is obtained by using the above radiation

exchange factor with wall and peak fuel pin temperatures from a second

FETA analysis considering radiation and heat conduction through helium

with a conductivity of 0.137 Btu/hr-ft-°F.

k (T1 - T0 ) (13)

p•. F•(Tr1 - T0)

.137 (980 - 621) 09256 ft.

1299.6 - .1514 (.173 x 10-8) (14404 - 10814)

VIII-28

In the TRUMP analysis, conduction from the wall to the fuel occurs

through two series connections (Figure VIII-3). The conduction length

for use in TRUMP is thus .04628 ft.

The above calculations apply to the ID analysis. The reference area is

established by the fuel element periphery, 4 x 8.375 inches. A constant

conservatively low helium conductivity applicable to a helium temperature

of 6000F has been assumed. The reference internal heat load is 1.2 x 10.63 kw.

The fuel element volume consists of 204 12 ft long 0.422 in. diameter rods.

The maximum heat generation rate for 97% of this volume which accounts

for fuel dishing and gaps is 18860 Btu/hr-ft3 . To establish the single

point model parameters any reasonable surface temperature may be used

in the FEMA analysis because the model parameters are practically

independent of temperature. This has been confirmed by FETA check

problems, after obtaining TRUMP results.

For the 2D TRUMP analysis the single point fuel element model is also

used. However, the energy from the fuel element is distributed over

the entire cavity surface instead of Just the periphery of the element.

This procedure is a conservative approximation of the very complex

radiation, convection and conduction heat transfer occurring in the cavity.

The reference heat transfer surface is now the cavity surface with each

section receiving a portion of the. total heat load. To maintain the

appropriate temperature differentials, the increase of the reference

heat transfer surface must be offset by a reduction of heat transfer

coefficients. Tle adjustment factor, 0.7182, is the ratio of half the

fuel element surface (2HL = 16.750 ft 2 ) to the cavity surface (23.322 ft)

in the 2D TRUMP.model. Application of the adjustment factor to the

effective conduction length results in t/2 = 0.06444 ft. Similarly the

radiation exchange factor becomes F= 0.1087 for the 2D analysis.

VIII-29

/

In the 1D analysis two series connected conduction lengths act in

parallel with one radiation link. In the 2D analysis one conduction

length Is in series with many parallel connected conduction lengths

and the combination of all conduction lengths act in parallel with many

radiation links. The reference area for any connection to a cavity

surface node is the cavity area of that node. The reference area for the

single connection between the fuel node and the cavity fluid node is

the entire cavity surface area.

4.1.6 Head Cavity Conditions

For the head cavity radiation is assumed to occur only in the axial

direction with a reflective configuration factor, % of 0.9 (6) based on a

parallel disc diameter to length ratio of 4.77.

With a surface emissivity of 0.8 and an area ratio of approximately 1.0,

the radiation exchange factor applicable to all head cavity connections

is:

F 0. 2 - = 0.6207 (14)

Conduction through air contributes negligibly to the total heat transfer

in the head cavity.

4.2 Computer Program Input

4.2.1 FED Input

The FED input consisted of 97 parts defining the two-dimensional TRUMP

problem geometry. FED was used primarily to calculate each of the 324

nodal volumes and most of the nodal connections and provide a deck of

cards forming part of block 4 and block 5 input to TRUMP. Convection and

radiation factors for block 5 were added by hand. In some cases, where

the volume boundaries did not provide the desired effective node connection

VIU-30

areas, .the FED output was modified by hand.

4.2.2 TRUMP Input

Both the one-dimensional and two-dimensional TRUMP problems consisted of steady-state problems and five step continuation problems to simulate the fire transient. The latter consisted of a pre-fire steady-state, a 0 - 0.5 hr transient, a 0.5-3 hr transient, a 3-10 hr transient and a post-fire steady-state. problem. The use of contintiation problems permits varying temperature dependent material properties with time. Modes of heat transfer during a transient as well as edit times may also be changed. Thisfeature was used to simulate rupture of the neutron shield with attendent introduction of radiation heat transfer.

The one-dimensional nodal volumes and nodal connections were computed by hand. With the exception of node 1 which has its nodal point on the outer surface, each nodal point was located at the arithmetic mean radius of the volume. Except for nodes 1 and 29 all surfaces were represented by zero volume nodes.

The two-dimensional problem nodes were generated by FED. All nodal points were located at the geometric center of the nodal volume as defined by FED. The cask surface nodes which provide the solar heat load simulate surfaces with high accuracy. Convection and radiation heat transfer across othcr surfaces within the cask were calculated using temperatures which occur at some distance from the surface inside the bounding riode. This approximation is necessary because of the large size of the problem but is also considered acceptable because maximum temperatures in the cask are established by the one-dimensional

problem.

VIII-31

4.2.3 FETA Innut

The reference design fuel element used a pin emissivity of 0.4 and a wall emissivity of 0.2 With a pin diameter of 0.422 in. (15 x 15 array, 204 rods

active) and a square pitch of 0.563 in, S as defined in Reference 11 is 2.67.

The resulting configuration factors as defined in Reference 11 are F1 2 = 0.1263, F13 = 0.08536, and F14 = 0.01917. A nominal pin to wall clearance

of 0.223 in. was used to calculate the conductance from the outer row

of rods to the aluminum wall.

4.3 Normal Operation Results

The temperature distribution within the cask under normal operating conditions is relatively flat in all material regions with the most significant

temperature rises occurring on the surface of the cask, at the gaps and

through the fuel element.

A plot of the radial temperature distribution at. the axial location with

the highest power density is shown in Figure VIII-7. Table VIII-5

provides a complete listing of the temperatures at each node for the

two-dimensional analysis of the upper end of the cask at design average

power under normal steady state conditions. The location of each node is

established in Figure VII-5. Special attention is directed towards

nodes 241, 290, 296, and 298 which represent respectively the can

seal node and the three nodes in the vicinity of the outer closure seal.

4.4 Hypothetical Fire Results

A summary of the cask thermal response during the fire and after the fire

has already been presented in Figure V111-1. Figure VIII-8 shows the

radial temperature distribution at the axial location of maximum power

density just after the fire and under post-fire steady-state conditions.

VIII-32

FIGURE VIII-7 ONE DIMENSIONAL STEADY-STATE TEMPERATURE DISTRIBUTION

(TRUMP MODEL)

LI III I0

0.6 0.8 1.0

RADIAL DISTANCE FROM CASK CENTERLINE. FT.

4.,

.4 WZ Zii, TEMP. p II1.2 1.4

K

C

1200

1000

eoo

o - - - - - - fl - - -

I-

Li.

'Ii

o'"

I,-

600

400

200

0

'U

IL

a SM

z :3

n,�1.6 U.It

J_ JU

z

IAMBIENT TEMP.

I

i.!II

1 .wNU l . .

a 41 U~s " .1 ob E 00.0 0 -4 g a. *UJ A4.6* #4I's " 190 a d.4 am A h~.*b .0 a ~ . 'dham.d h AOfta a)mb.a w SO W&.4.MN *0da. goS 5 5 C . b . "e CFO Ce C O 5aO * * * e * * * * e * .

a,3 a n~. 6 -d "I's- sNe466 goft %1, mean *wam..u .4.3~ .01 -j -40a I-166 of-i *."a moa L01 "6.02

COt.1.3T 0 , P .1'A -01 see p...... 0CC COO OO .0S a C S O @ fm63066U4 436-4 1 h4 a vo. z.15 10 1 .1 .4 MW 00. is A r j1 Co3 .3 .36 fe 1..0 A- .1 r. %$A to 61 t.3 .11 -4 a.ms dI'I

66~~~ -4 4ý' as94 us0 "4 0. 0*1 CDI. %014. 14 410 Od -A 04 3 AU~ 049 WNW d A. Pyf .8.. ra~- . mi W I4 G % A

If %*m .a La Cs iatsa 4 3as rum oQLl- ra 124 42 es ab a a de o Ca .. a.a af% mmmc a a V4 0- ?a rs 2 .2as tom mC3 a a c M 1 :p -. a a " s-a 6%~a on .2 Ci es

1. 4 tdao W a~elS 1 q~ sAo- 4 NNN L L &6'" a& 18.1..4.1&A-1.1. w A " so Is .0.1 e ,aaaa - 6V AAa P -a o0 Ua- bAS P0 f

d ft1.4 4 -4.1 4 IVq -4W.-4 -4 -4 4 to-4 wd I -4 4 N-4 4N 4 1

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Cr. C C. FIG•JRE VIII-8

ONE DIMENSIONAL TRANSIENT TEMPERATURE DISTRIBUTION

(TRUMP MODEL) 1200

10O00 = . . . . .

END OF FIRE

800 L BA.

'-

V600 c n _u

400 POST FIRE SS

2 gal

SAMCIENT z TEMP.

0 ,1 . II1 ,I I I I , 0.2 0.4- 0.6 0.8 1.0 1.2 1.4

RADIAL DISTANCE FROM CASK CENTERLINE. FT.

As in the case of normal operation, the gaps influence the thermal gradient

considerably. Examination of the maximumn fuel temperature transient

indicates that a steady-state evaluation of fuel element temperature

distribution is acceptable. In addition maximum transient fuel temperatures

do not appear to be significantly higher than post-fire steady-state

temperatures. A substantial amount of lead, almost 100%, appears

to have melted at the axial location of maximum power density 1.25

hours after the start of the accident. All lead has solidified 1.75 hours

after the start of the accident.

Table VIII-6 provides supplementary transient results obtained from the

two-dimensional analysis of the upper end of the cask. In contrast to

the extensive lead melting noted in the ID analysis, there appears to be

no molten lead at the head end of the cask (See node 124). Some lead

melting may occur at the other end of the- cask where there is less structural

steel. A maximum external closure seal temperature of 648°F is reached

0.75 hours after the start of the accident. The can seal reaches a

maximum temperature of 5640F 2.5 hours after the start of the accident.

On the basis of the entire transient calculations it is concluded that

steady-state conditions are essentially reached within 48 hours after

the start of the accident assuming no special external factors.

4.5 Cold Operation

An explicit evaluation of the cask under cold conditions in still air has

not been considered necessary because the cask cavity fluid is helium.

A uniform temperature of -40 F, the minimum specified by regulations,

is considered acceptable as long as the neutron shield contains appro

priate anti-freeze.

VIII-36

(Cr,

Location Fuel Al

NAods 1 4

850 S08

850 508

850 S08

850 509

851 512

851 517

856 540

865 570

891 617

914 643

941 665

949 666

929 634

000 565

Inner Inner Outer Outer Outer Can Pb Al Pb Can Beal Head Beal Bolt Bolt 10 24 88 124 41 241 269 290 306 308

497

497

498

Soo

504

510

538

569

612

637

657

657

624

555

201

316

396

472

541

606

006

588

562

542

512

492

443

374

511

511

511

512

514

518

537

565

608

633

654

655

623

553

285

301

34S

397

450

501

534

536

532

523

503

484

426

338

460

460

460

461

465

472

507

553

610

630

636

626

574

490

355

356

358

364

375

393

449

493

542

560

561

545

478

379

369

369

374

385

403

425

466

485

526

554

567

556

492

390

286

301

350

417

490

564

648

642

595

554

499

464

390

300

278 269

349 480

448 632

543 753

631 850

709 • 929

696 750

655 666

589 576

543 525

486 466

451 431

379 364

293 284

TABLE VITI-6

TRANSIENT TWO-DIMENSIONAL TEMPERATURES AT SELECTED NODES

Time (bra)

(,4 -,4

0 (at)

.I

.2

.3

.4

.5

.75

1.0

1.5

2.0

3.0

4.0

0.0

Poet-fire as

"I

p.

Maximum Fuel Pin Temperatures

Maximum fuel pin temperatures are provided primarily for information. The values given in Figure VIII-2 are conservatively high. No time limit is associated with the transient temperatures because they are close to the post-fire steady state temperatures. A change of wall temperature of 1°F has been found to result in approximately 1/2°F change of hottest

pin temperature.

4.7 Thermal Expansion and Contraction

Thermal expansion of the neutron shield fluid is accommodated by the provision of an expansion tank on the truck. In the cask and head cavities, the thermal expansion of helium and air results in only a

minor increase of pressure.

The effects of differential expansion on gap sizes have been evaluated explicitly on all gaps except in the head area and between the uranium shield and the stainless steel inner shell wheie nominal cold dimensions

were used. The nominal cold gap dimension between the uraniu.t and stain

less steel shell is 0.0265 inches and is the result of the rninal machine

dimensions of 14.400 diameter of the stainless steel shell, and 14.453 diam

eter of the uranium shell. he diamnters have tolerances of t .010 and ! .015 respectively. Dur.ing manufacture dimnsi oil inspections are made

to assure that all dinensions are within the specified tolerances. The

uranium shall is assembled over the stainless steel shell and the gap is

equalized by temporary sbius placed in the gap. The shins remain in place

until the lead pouring operation is complete and the cask body has return

ed to room teeperature. At this point the concentric relationship is main

tained by the surrounding lead shield. The bottotm uranium cap and the outer

bottom ecd forging are then welded into place. 7he bottom end forging is

machined to provide radial restraint of the uranium cap thereby maintaining

VIII-38

4.6

the concentric relationship. Under. steady-state conditions when the

internal temperatures are high and beat transfer is outward the gaps will

actually be reduced thus providing lower internal cask temperature than

predicted. During the fire the gaps will increase thus impeding the flow

of energy to the cavity to a greater degree than in the analysis. The

use of miform n4,a1 gaps axound the lid with a conservatively high

beat transfer through the tie-down bolts is considered appropre. Var

iations in these gaps do not alter the temperature distribution siniifi

cantly at the top of the cask.

4.8 Uncertainty Factors and Design .Magins

The entire thermal analysis of the cask has been directed at achieving

conservative results by the use of ua -mzn heat loads and low cavity

beat capacities. Although nocmirl dimensions have been used through

out the analyses, physical pbencna such as blaokwns and rutures

have been simulated in the most severe manner possible. No part of

the analysis depends upon artificial cooling during a transient. Inde

pendent 1D and 2D therml analysis of the cask assure further conserva

tisn of the temperatures predicted. It is therefore anticipated that

the actual temperatures in the cask will be somewhat lower than those

predicted by ths analysis.

4.9 General Conclusions

The analyses performed indicate that no unusual thermal response

characteristics occur. Under normal operation the cavity may locally

reach a temperature of 1013 0 F which corresponds to the maximum fuel

surface temperature (ID analysis). The actual maximum temperature will

more likely be 860°F (2D analysis). Under the same conditions the.

maximum local neutron shield water temperaturd may reach 3520 F

although the bulk fluid temperature will more likely not exceed 296 F.

The inner gasket temperature under normal conditions may reach 375°F

VIII-39

whereas the outer seal may reach 309*F. maximum local temperature on the sur

face of the cask may be as high as 3400F.

As a result of the hypothetical fire transient, the maximum fuel pin sur

face temperature is not expected to exceed 11020 F. Gasket temperatures are

not expected to exceed 564OF and 6480F at the internal and external seals

respectively. A negligible amount of lead melting may occur at the end of the

cask. In the central portion of the cask, however, almost all of the lead may

melt in regions, of high power density. Post fire steady-state temperatures

may eventually be reached if all neutron shield water is lost in an accident

less severe than the hypothetical fire transient.

Under extreme cold conditions in still air, a uniform postulated tempera

ture of -40°F should not adversely affect the cask provided sufficient anti

freeze is added to the neutron shield.

4.10 Effect of Personnel Barrier on Thermal Performance

The personnel barrier description is provided in NLI Drawing No. 70514F.

The solid roof portion of the barrier extends over a 96 inch arc with a 48

inch radius. Approximately 128* of the cask surface is included in the sector

formed by this arc.

A conservative simplified analysis of the barrier effect on the cask

temperature under normal conditions has been performed as follows. It was

assumed that the solid portion of the personnel barrier roof absorbs incident

solar heat but does not permit any decay heat transfer from the cask surface

through its included angle. A temperature 350 higher than the maximum neutron

shield jacket temperature of 340*F given the SAR was thus computed. If it is

assumed that the barrier eliminates radiation heat transfer from the effected

portion of the cask surface but does not reduce natural circulation, this

results in a 2*F lowering of the jacket temperature provided in the SAR.

VIII-40

On the basis of these calculations it is concluded that increasing the

shield jacket temperature by 35*F accounts conservatively for any adverse

effect on the personnel barrier on the cask thermal performance. It is to be noted that this increase in temperature is attenuated towards the center of

the cask because of temperature dependent heat transfer. The effect of a 350F

increase in cask peak surface temperature is insignificant.

4.11 Effect of Consolidated PWR Fuel Rods

The temperature of the hottest fuel rod during normal operation and

during a hypothetical fire accident must be determined for the consolidated

fuel. The cask surface temperature is also of Interest, even though it is

independent of the fuel form. Methods for calculating the temperature in a

triangular array of rods, as in a consolidated fuel canister, are not as well

developed to date as methods used for square arrays so an additional margin of conservatism has been added by using 600 watts as the fuel heat load instead

of the value of 564 watts, calculated. The tables of temperature versus cool time has also been provided and compared to the design basis fuel analysis of

the SAR. These comparisons show that the temperatures developed in a cask containing consolidated W14 x 14 fuel cooled 12 years will be significantly

lower than temperatures for the SAR design basis intact fuel. The fuel

parameters of the W14 x 14 fuel are given in Table VIII-7.

TABLE VIII-7 CONSOLIDATED PWR FUEL ANALYSIS PARAMETERS

INTACT CONSOLIDATED PWR W14 x 14

Number of Rods 204 358

Burnup (MWD/MTU) 40,000 25,000

Cool Time 150 days 12 years

Heat Rate (kw) 10.6 0.600 (564 calculated by ORIGEN)

Ambient Temperature 130OF 130OF

Cask Cavity Gas Helium Helium

Rod Array Square Triangular, Modeled as square

VIil-41

TABLE VIII-8 "SCOPE ANALYSIS"

MAXIMUM STEADY STATE TEMPERATURES

TYPE

W15 x 15 Intact

W14 x 14 Consolidated

W14 x 14 Consolidated

W14 x 14 Consolidated

W14 x 14 Consolidated

W14 x 14 Consolidated,

In air

PWR Intact

COOL TIME (yrs.)

0.4

5

8

10

12

12

HEAT (kw)

10.630

1.300

0.780

0.650

0.600

0.600

SAR ANALYSIS FOR DESIGN BASIS PWR 150 days 10.6

The maximum steady state normal operation rod temperatures for

consolidated and intact fuel forms are given In Table VIII-8. Inspection of

this table shows that the normal operation hottest pin will be at 353°F,

significantly lower than the 1013 0 F calculated in the SAR.

The maximum fire temperatures for consolidated and intact fuel forms are

given in Table VIII-9. The results in this table show that the maximum rod

temperature due to a fire accident (536*F) for the consolidated fuel is less

than the temperature of normal operation for the SAR design basis PWR fuel.

Tables VIII-8 and VIII-9 also include the temperatures for W14 x 14 fuel

cooled 12 years if shipped in air instead of Helium. The temperature in air

is included for comparison purposes since most of the published experiments to

date have been for air filled cavities.

The cask surface temperature during normal operation with consolidated

W14 x 14 fuel is 216°F. This is much less than the design basis PWR fuel

surface temperature of 340°F.

The SCOPE computer printouts for the W14 x 14 consolidated and W15 x 15

intact fuel are presented in Appendix C of to this Section.

VIII-42

TEMP (F)

1022.

478.

387.

362.

352.

426.

1013.

11-1/

TABLE VIII-9 "SCOPEu ANALYSIS

MAXIMUM PIN TEMPERATURES (FIRE ACCIDENT)

TYPE COOL TIME(yrs) HEAT (kw) TEMP(*F) W15 x 15 Intact 10.4 10.630 1203. W14 x 14 Consolidated 5 1.300 680. W14 x 14 Consolidated 8 0.780 576. W14 x 14 Consolidated 10 0.650 547. W14 x 14 Consolidated 12 0.600 536. W14 x 14 Consolidated,

In Air 12 0.600 556.

SAR ANALYSIS FOR DESIGN BASIS PWR PWR Intact 150 days 10.6 1102

The temperature calculations performed in this anlysis were made by the SCOPE code (ORNL/CSD/TM-149, TTC-0316, by J.A. Buckholz). This code generates temperature distributions comparable with those produced by the HEATING-5 code commonly used for cask safety analysis, but in the radial dimension only. The maximum temperatures occur at the fuel midplane, and this is the location at which the radial temperature profile was calculatd. The code version used by Nuclear Assurance Corporation has been benchmarked against th HEATING-5 code. The calculations of rod temperatures are made in SCOPE via the Wooten-Epstein correlation method, which treats square arrays of rods. It is to be.expected that this method will cause the central rod temperature to be overpredicted for consolidated fuel since the number of layers (conccentric squares) of rods required for a square lattice is larger than for a triangular lattice, and-the temperature if dependent upon the number of layers.

Shipment of consolidated fuel in the NLI-1/2 cask results in lower temperatures, under all conditions, than temperatures caused by the design basis intact PWR fuel. Thus shipment of the consolidated fuel is safe from a thermal standpoint.

VIII-43

4.12 Effect of Metallic Fuel Rods

The temperature'of the hottest fuel rod during normal operation and

hypothetical fire accident were determined using the SCOPE program. The

fuel is placed in an aluminum basket containing three cylindrical holes,

7 sound rods or less per location. Each group of 7 sound rods is

contained in a transfer basket which is modeled in SCOPE by using the

canister option of the program. For this analysis, the cask is assumed

to be inside of the shipping container. The intact, design basis PWR

parameters are compared to the metallic fuel parameters in Table VIII-lO.

Table VIII-10

Metallic Fuel Analysis Parameters

Number of Rods

Burnup (MWD/MTU)

Cool Time

Heat Rate (kW)

Ambient Temperature

Cask Cavity Gas

Rod Array

204

40,000

150 days

10.6

130OF

Helium

Square

Metallic Fuel

21

1,600

365 days

0.75

130"F

Air

Triangular, Modeled as Square

The analysis results are summarized in Table VIII-11 and are compared to

the design basis PWR assembly, demonstrating that the metallic fuel

developed significantly lower temperatures than the 150 day cooled PWR

assembly. The SCOPE computer printouts for the metallic fuel are

presented in Appendix D of this Section.

Revised Oct. 1986 Feb. 1987 Oct. 1990 VIII-43a(l)

Table VIl-l .

Metallic Fuel Rod MaxIumTemperatures

Cool Time Normal Operation Hypothetical

(LDays) Heat (kW) Temp (°F) Tem (°F)

Metallic Fuel 365 0.75 376" 603*

Intact PWR 150 10.60 1013* 1102"

The comparison of normal operation maximum rod temperatures shows that

rod temperatures developed in normal operation and fire accident for the

metallic fuel are much less than the temperature for the design basis PWR

fuel.

The cask surface temperature during normal operation with metallic fuel

with a decay heat of .750 kW is 210*F. This is much less than the design

basis PWR fuel surface temperature of 340'F. Thus shipment of the

metallic fuel is safe from a thermal standpoint.

Up to six individually encapsulated failed metallic fuel rods may be

shipped in a specially designed six hole failed fuel aluminum basket.

This aluminum basket will be limited to a maximum heat load of 30 watts

(5 watts per rod). A SCOPE analysis has been performed on the six hole

failed fuel basket to determine the maximum basket temperature. The

SCOPE computer printouts for the failed fuel basket are presented in

Appendix C.

Using a different basket comprised of three tubes, up to three,

individually encapsulated, failed fuel rods may be transported in the

same aluminum basket used for transporting the 21 sound fuel rods. This

basket can be constructed using aluminum or stainless steel Type 304.

For either case, the heat load per rod is 5 watts which brings the total

heat load for the 3 element basket to 15 watts. This is enveloped by the

six hole failed fuel aluminum basket, which corresponds to a maximum heat

load of 30 watts. Substituting the 1/8 inch stainless steel liners for

the 1/8 inch aluminum liner for the three hole basket, the change in the

maximum temperature for the 15 watt heat load is negligible.

Page Added VIII-43a(2) Oct. 1990

4.13 Effect of PWR or BWR Rods

The maximum heat load in the cask with up to 25 PWR rods is 1.65 kW. The maximum heat load in the

cask with 25 BWR rods is 4.0 kW. The temperature of the hottest rod in normal operation and

hypothetical fire accident conditions were calculated using the SCOPE program. The fuel rods are placed

in a rod holder that is inserted into the cask to support the fuel. No credit for heat conduction was taken

for this holder;, the fuel rods were modeled free in helium. The parameters used in this analysis are

compared to the design basis PWR assembly in Table VIII-12.

Table VIJ-12

PFW ROD ANALYSIS PARAMETERS

Number of Rods

Burnup (MWD/MTU)

Cool Time (days)

Heat Rate (kW)

Ambient Temperature

Cask Cavity Gas

Rod Array

intacR

204

40,000

150

10.6

130°F

Helium

Square

The thermal analysis results are summarized in Table VIII-13 and are compared to the design basis PWR

assembly, demonstrating that 25 PWR or 25 BWR rods developed significantly lower temperatures than

the design basis PWR assembly. For the content condition of 18 PWR rods with specific power of 60

kW/kgU and a cooling time of 300 days, the decay heat load is 0.9 kW, which is less than and enveloped

by, the 1.65 kW analyzed for the 25 PWR rod content condition.

Page Added May 1987

Revised Jan. 1990

Oct. 1991 Apr. 1992

VIII-43b Feb. 1996

P-WR Rods

25

60,000

150

1.65

130OF

Helium

Square

13WR Rods

25

75,000

150

4.0

130°F

Helium

Square

TABLE VIII-13

PW AND BWR ROD MAXITHU TENPERATURE

Normal Operation Temperature OF)

Hypothetical Accident

Temperature (F)

Intact PWR PWR Rods (25)

BWR Rods (25)

10.6 1.65 4.0

1013 393

617

1102 556 697

The comparisons of maximum rod temperatures for up to 25 PWR or 25 BWR rods show that

temperatures developed in normal operation and hypothetical fire accident conditions

are much less than the maximum rod temperatures for the design basis PWR fuel. The

actual heat source from 25 PWR rods is less than 1.65 W, but 1.65 kW is used as a

bounding case.

The cask surface temperature during normal operation with PWR rods is 229°F and 236*F

with BWR rods. These are much lower than the design basis PWR fuel cask surface

K.- I temperature of 340*F. Thus, shipment of up to 25 PWR rods or 25 BWR rods is safe

from a thermal standpoint.

Page Added May 1987 Revised Jan. 1990 Oct. 1991 VI11-43c

I

4.14 Effect of Mark 42 Fuel Assemblies The maximum temperature of a Hark 42 fuel assembly during normal opera.

tion and during the hypothetical fire accident were determined using the SCOPE program. The intact, design basis NR parameters are compared to the Mark 42 fuel assembly parameters in Table VIII-14.

Table VIII-14 Mark 42 Fuel Assembly Analysis Parameters

Intact ML1 Mark 42 Fuel Assembly Cool Time 150 days 1245 days Heat Rate (kW) 10.6 .0.45 Ambient Temperature 130oF 1300F Cask Cavity Gas Helium Air

The thermal analysis results are summarized in Table VIII-15 and are compared to the design basis NR fuel assembly, demonstrating that the Hark 42 fuel assembly developed significantly lower temperatures than the 150 day cooled PWR assembly. The SCOPE computer printouts for the Hark 42 fuel assembly are presented in Appendix E of this Section.

Table VIII-15 Mark 42 Fuel Assembly Maximum Temneratures

Cool Time Normal Operation Hypothetical Das Heat M) Te L (oF)

Mark 42 Fuel 1245 0.45 253 410 Intact PIR 150 10.6 1013 1102

The comparison of normal operation maximum fuel temperatures shows that fuel temperatures developed in normal operation and during the fire accident for the Hark 42 fuel assembly are much less than the temperatures for the design basis P1R fuel.

The cask surface temperature during normal operation with the Hark 42 fuel assembly is 1410F. This is much less than the design basis PWR fuel surface temperature of 340'F. Thus, shipment of the Mark 42 fuel assembly in the NLI-1/2 cask is safe from a thermal standpoint.

VIII-43d Page added August 1988

Effect of Mark 22 Fuel Assemblies

The maximum temperature of a Mark 22 fuel assembly during normal operation and during the hypothetical fire accident were determined using the SCOPE program. The intact, design basis PWR parameters are compared to the Mark 22 fuel assembly patameters in Table VIII-16.

Table VIII-16 Mark 22 Fuel Assembly Analysis Parameters

Intact PJR Two Mark 22 Fuel Assemblies Cool Time 150 days 150 days Heat Rate (kW) 10.6 3.451 Ambient Temperature 1300F 1300F Cask Cavity Gas Helium Air

The thermal analysis results are summarized in Table VIII-17 and are compared to the design basis PWR fuel assembly, demonstrating that the Mark 22 fuel assembly developed significantly lower temperatures than the 150 day cooled PWR assembly. The SCOPE computer printouts for the Mark 22 fuel assembly are presented in Appendix F of this Section.

Table VIII-17 Mark 22 Fuel Assembly Maximum Temperatures

Cool Time Normal Operation Hypothetical (Days) Heat (W) Temu (*F) Temp (*F)

Mark 22 Fuel 150 3.451 773 805 Intact PWR 150 10.6 1013 1102

The comparison of normal operation maximum fuel temperatures shows that fuel temperatures developed in normal operation and during the fire accident for the Mark 22 fuel assembly are much less than the temperatures for the design basis PWR fuel.

The cask surface temperature during normal operation with the Mark 22 fuel assembly is 243°F. This is much less than the design basis PWR fuel surface temperature of 340°F. Thus, shipment of the Mark 22 fuel assembly in the NLI-1/2 cask is safe from a thermal standpoint.

Page Added VIII-43e February 1990

4.15

5.0 REFERENCES

1'. Edwards, A. L., "A Compilation of Thermal Property Data for Computer Heat Conduction Calculations", UCRL-50589, Feb. 24, 1969.

2. Transactions ASME, Vol. 76 pp. 967-985, 1954

3. Scott, D.B., "Physical and Mechanical Properties of Zircaloy 2 and 4", WCAP-3269-41, May 1965.

4. Kreith, F., "Principles of Heat Transfer", Second Edition, International Textbook Co., 1965.

5. Shappert, L. B., "Cask Designers Guide", ORNL-NSIC-68, Feb. 1970.

6. McAdams, W.H., "Heat Transmission*, Third Edition, McGraw-Hill Book Co., 1954.

7. Lauer, B. E., "How to Evaluate Film Coefficients for HeatTransfer Calculations", Technical Manual reprinted from THE OIL AND GAS JOURNAL, 1953.

8. The Aluminum Association "Aluminum .Standards & Data", 1972-1973.

9. Edwards, A. L. ,"TRUMP: A computer Program for Transient and Steady-State Temperature Distributions in Multidimensional Systems", USAEC Report UCRL-14754, Rev. II, July 1, 1969 (Errataissued Feb. 4, 1971 and Aug. 27, 1971).

10. Schauer, D.A., "FED: AComputer Program to Generate Geometric Input for the Heat Transfer Code TRUMP", USAEC Report UCRL-50816, Feb. 10, 1970.

11. Watson, J. S., "Heat Transfer from Spent Reactor Fuels During Shipping: A Proposed Method for Predicting Temperature Distribution in Fuel Bundles and Comparison with Experimental Data", ORNL-3439, May, 1963.

VIII-44

APPENDIX A

-' Section VIII

THERMAL ANALYSIS

The modification to the truck cask consists of removing the inner container.

The aluminum basket will then be made to a slightly larger diameter to fill

the cask cavity. From a qualitative analysis, the modified design should be

thermally less limiting than the design with the Inner container. The reason

for this is that one helium gap and the stainless steel inner container will

be replaced with an equal thickness of aluminum and the aluminum has a

higher thermal conductivity than helium and stainless steel. To prove this

the thermal resistance of the added basket thickness will be compared to the

resistance of the basket-Inner container design.

The basket-inner container resistance to be evaluated consist of the gap from

the basket to the Inner container, the thickness of the Inner container and the gap

from the inner container to the cask cavity. The gaps Include both radiation

and conduction resistances.

K..

VIMI-AI

/

I.. Summary Of Temperatures

Material

Aluminum

Helium Gap

S. S. Container

Air Gap

Inner Shell

Average Temperature

522.2

517.5

512,.3

439.6

365.1

The above temperatures from the 2 dimensional TRUMP analysis at the

cask centerline will be used to calculate the thermal resistance for

both the modified design and the original design. Nominal cold gap

thicknesses will be used.

K..

VIII-A2

Node

6

8

10

11,12

14

I. Cask With Inner Container

A. Resistance of helium gap from basket to inner

1. * Conduction across the helium gap

gap thickness = X =12.625 - 12,55, 2

container

.0375 in. - .003125 ft.

helium conductivity k - .119 + 5.17.S - 400 (.137 - .119) 600 - 400

= .13 BTU/hr. ft.°F

Conduction resistance =R = X = .003125 .02404 hr.ft.* k .13 OF/BTU

2. Radiation across helium gap

hr = .174x 10-8 F ( T1 2 + T02 ) (TI + TO)

Ti 522.2 +460 - 982.2 0 R

To= 512.3 + 460 = 972.3 0 R

F = .167

hr .174 x 10-8 (.167) (982.22 + 972.32) (982.2 + 972.3)

= 1.083 BTU/hr. ft.2 OF

Radiation resistance Rlr . j --. 9234 ir hr 1.083

The radiation and conduction resistance act Is parallel. There

fore the combined, resistance is calculated as follows:

VfII-A3

RIT = Rr Ric Rl r +R1 c

= .9234 (.02404) .9234 + .02404

=" .0222 .9474

RT = .0234 RiT

B. Resistance of SS Can

Thickness of Can = X .275" .0229'

SS conductivity @ 312.3

k= 9.43 + Si2.3- - 212 (12.6- 9.43) 923 - 212

= 10.77 BTU/hr.ft.OF

R = .0229 - .00212 k 10.77

C. Air gap from inner container to cask cavity

1. Conduction through air gap

X= 13.375 - 13,175 = .1" - .00833' 2

@ T 439.6 0 F

k= .018+ 439.6 -212 (.029 - .018) 752 - 212

= .0226 BTU/hr.ft.°F

R3c = .00833 .3686 .0226

VIII-A4

I

2. Radiation across air gap

hr = .174 x 10-8 F (T1 2 + T02 ) (T1 + To )

T1 ='512.3 +460 -972.3°R

To = 365.1 +460,= 825.1 0 R

hr = .174 x 10-8 (.3351) (972.32 + 825.12 ) (972.3 + 825.1)

= 1.704

R 1 .5869 R3r = 1.704

the radiation and conducttion resistance act in parallel. There

fore the combined resistance is calculated as follows:

R3T R3r R3c R3r + R3c

= .5869 (.3686) .5869 + .3686

- .2163

.9555

R3T = .2264

D. Total Resistance

The total resistance of the original basket and inner container is the

sum of the individual resistances.

RT = RlT+R2 + R3T

= .0234 + .00212 + .2264

RT r .252

VIII- AS

II. Cask without Inner Container

The total resistance of the new basket arrangement will now be calculated.

The resistance includes the increased aluminum thickness and the

helium gap from the basket to the cask cavity.

A. Resistance of Increased Aluminum Thickness

Increased aluminum thickness 13.28 - 12.55 2

X= .365" = .0304'

Aluminum conductivity k = 96.8 BTU/hr.ft.OF

R;= -2L = .0304 - .000314 k 96.8

B. Helium gap from basket to inner container

1. Radiation across helium gap

Radiation coefficient = hr

hr = .174 x 10-8 F (T12 + T0

2 ) (Ti + To )

F= .167

T= 512 +460 =9720 R

T - 365 + 460 = 825oR 0

hr - .174 x 0o- (.167) (9722 + 8252 ) (972 + 825)

hr = .85 BTU/hr. ft.2 oF

Radiation resistance = R2r 1 -1.176

VIII-A6

2. Conduction across helium

x- 13.375 - 13.28 2

• .095" = .00792'

Conductivity @ 440OF

k - .119 + 440 -400 (.137 - .119) 600 - 400

= .1226

Conduction resistance = R~c = X = .00792 .0646 k .1226

The radiation and conduction resistance act In parallel. There

fore the combined resistance is calculated as follows:

RiT -R;r R'c

= 1.176 (.0646) 1.176 + .0646

= .0759 1.24

R2ýT -. 0612

C. Total Resistance

The total resistance of the new basket is the sum of the Individual

resistances.

RT = RI + R2T

= .000314 + .0612

= .0615

VIU-A7

2 o Therefore, the resistance of the modified basket (.0615 hr. ft. F/BTU)

is much less than the original basket (.252 hr. ft2 0F/BTU) and becaus•->

of this, the temperature of the fuel and basket will be less than the

-original basket and fuel.

VII-A8

APPENDIX B

Section VIII

Thermal Analysis

During shipment the cask cavity of the 1/2 LWT Cask is normally dry. The cavity is dried by forcing water out the cavity drain line using compressed air. To remove the residual moisture the cask cavity is evacuated by using

a vacuum pump.

At low decay heat loads the dryout procedure is not necessary providing the maximum fuel temperature is below the saturation temperature at 365 psig (maximum allowable operating pressure) of 4400 F. Mafimum fuel temperature is limited to 410OF allowing for a 30 0 F fuel temperature increase at the end of the fire accident transient. Limiting the fuel temperature to 410OF results in the maximum decay heat load of 2.02 KW for which cask cavity dryout procedures are not

necessary.

Hand calculations were used to calculate the temperature distribuiion in the cask for the post fire steady state condition. The calculations were performed for the alternate

cask configuration (configuration B).

The cask surface temperatures were chosen (235 0 F, 225 0 F, 220 0F), which preliminary calculations indicated would bracket the required surface temperature. A heat balance was written

VIII-Bl

for the cask surface and the cask decay heat for each of

these surface temperatures was calculated. A distributed

solar heat rate of 92 BTU/hr. ft 2 is included in the heat

balance. The decay heat was then used to calculate the

temperature distribution in the cask starting at the surface

and working inward to the fuel. One dimensional heat transfer

was assumed for the analysis. The cask geometry is shown

schematically in figure 1.

Material thermal properties, that were used, are listed

in table VIII-2 of the LWT 1/2 Safety Analysis Report.

Emissivities and convection and radiation factors are listed

in tables VIII-3 and VIII-4. The effective fuel conduction

length and radiation exchange factor were calculated in

section 4.1.5, pg. VIII-28 of the Safety Analysis Report and were

used in this analysis. The ambient temperature was 130 0 F.

After the fuel temperatures were calculated, they were

plotted against the rated decay heat. The rated decay heat

is the decay heat value, used in the calculations, divided by

1.2. This was done to include the axial peaking factor. This

plot is included as figure 2. The decay heat corresponding to

a maximum fuel of 410OF is then read from the graph as 2.02 KW.

A suirmary of surface temperatures, fuel temperatures and decay

heat is shown in table 1.

VIII-B2 y..

- /

Calculations show that leaving residual moisture in

the fuel cavity of the NL 1/2 LWT cask will have no effect

on cask safety if the decay heat load is below 2.02 KW.

Even with the 300F increase during the fire accident transient,

the maximum fuel temperature will be no higher than the satura

tion temperature at 365 psig which is the maximum allowable

operating pressure.

VIII-B3

BASKET

FUEL

HELIUM GAP

INNER SHELL

URANIUM

LEAD

FIGURE 1 - CASK

TABLE i

GEOMETRY

SURFACE TEMOF

220

225

235

FUEL TEMP°F

376.4

405

460

DECAY HEAT-KW

2.05

2.42

3.18

RATED DECAY HEAT-K4

1.71

2.02

2.65

VIII-B4

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APPENDIX C

SECTION VIII

SCOPE Computer Printouts

VIll-Cl

TITL[: 0' CA1:.I 'EF (REj FUELL, 12 YEARS COOLFDt Hil-1l2 Crl¶.K )

DES(C1DIIC'. Of bASTE PATERIAL

(PUR)

YEA RS AAT ISCUFT

W.TTSIASSY NISECIASSY PISECIASSY

NELEM---TYPE OF WASTE MATENIAL (SHOWN Ig RNACKETS) BU ------ AVERAGE SURNUP (EDIT FOR OOOKEEPING PURPOSES ONLY; TIME----COOLING TIFf (AGE Of FUEL SINCE DISCOIADGE) WHEAT---DECAY HEAT GIVEN OFF nY THE WASTE MATERIALS NOV USED DHEAT--DECAT HEAT SIVEN OfF PT EACH ASSEMBLY (ON CANISIER); SRC .... NEUTRON SOURCE (EDIT FOR BOOKEEPING PURPOSES ONLY; SRCG--- PHOTON SOURCE (EDIT FOR BOOKEEPING PURPOSES ONLY;

NO LONGER USED)

If iERO NOT USED IF ZERO 40 LONGER USED) NO LONGER USED)

DESCDIPTIOI Gf WASTE CONTAINER

1 35s

13.0C• * GC

( SS ) I.CHES INCHES

lEiT FEET FEET

DESCRIPTION OF INSERT

! ( AL )

.22C

.13C IhCHES

.•CC INCHES 644c MiPCHES

13 ( HE ) *!SC IkCHES

13 ( HE ) 2.190 INCHES

1 13.3? INCHES

ITYPE---FLAG: I FOR SQUARE CA4ISTERS5 2 FOR SQUARE ASSEPOLIES (WITHOUT CAN) NPINS---NUMPER OF FUEL PINS PER ASSEMBLY (IF "ELI" DENOTES PUS OR BUR) MCAN----ITYPE OF MATERIAL USED FOR CANISTERS (NO CAN USED If MCA49O) ODCAN---OUTSIVE DIN OF CANISTER (MCAN3O,) OP WIDTH 9F FUEL ASSEMBLY 1NCAhMO) IKCAN---WALL THICKNESS OF CANISTER (IF PCINGToOI HTCAN---LINGTH OF CANISTER (OR FUEL ASSEMBLY) HTVOID--PORTION OF CANIST1R NOT OCCUPIED VV WASTE MATERIAL (IF WCAN.GT,.) HTFUEL--PORTION OF FUEL ASSEMBLY CONTAINING U02 (If MfLEM DENOTES PUR OR BUR)

IINSRT--TYPE Of MATERIAL USED FOR INSERT (SHOWN IN BRACKETS) EMINST-SURFACE EPISSIVITY Of THE INSERT MATERIAL (DIIENSIONLESS) TffINST--THICKNESS OF INSERT B!TWEEN ASSEMBLIES (INCLUDES TPOISN) TPOISN--THICKNESS OF NEUTRON POISON IMBEDDED IN INSERT WATL BETWEEN ASSEMBLIES TKCGAP--THICKNESS OF GAP BETWEEN CANISTER AND INSERT MCGAP---TYPE PF GAP ATMOSPHERE (SHOWN IN BRACKETS) TKIGAP--TWICKNESS OF GAP BETWEEN INSERT AND THE INNER SHELL 0IGAP---TYPE OF GAP ATMOSPHERE (SHOWN IN URACKETS) WTHICK--THICKNESS OF THE INSERT BETWEEN CANISTER ANO INNEP SHELL NfLEM---NNUMBER OF ASSEMBLIES COR CANISTERS) PER CASKS If RERO. PERFORN SEARCH CASKDb--INSIDE DIAMETER OF THE CASK (CALCULATED BY CODE If USER ENTERS 0.0)

DESCRIPTION OF INNER 9 OUTER SNELL AND THE OUTSIDE LINER

6 6 6

,500 .!7S .250

( SS ) ( SS ) ( SS )

INCHES INCHES IhCHES

MISHLL--TYPE Of MATERIAL USED FOR THE INNER SHELL (SHOWN IN BRACKETS) MOSHLT--TYPE OF MATERIAL USED FOR THE OUTER SHELL (SHOWN IN BRACKETS) MOLIN---TYPE Of MATERIAL USFD FOR OUTSIDE LINER AND FINS (If REQUIRED) TKISHL-:IHICKNESS OF INNER SHELL TKOSHL--THICKNESS OF OUTER SHELL IKOLIN--IHICKNESS OF OUTSIDE LINER

DESCPIPTIOI. OF NEUTRON AND G60NA SHIELDS

M45PLD 15 (HCO2) NHSHLD--TYPE OF MATERIAL USED FOR NEUTRON SHIELD (SHOWN I'd BRACKETS) P.(SHLD 1 ( Pb ) NCSHLD--TYPE Of MATERIAL USED FOR GAMPA SHIELD (SHOWN IN BRACKETS)

DESCRIPTtOl? OF HEAT TRAINSFER PARAMETERS FOR FINS (9 CASK$

t ( SS ) 4.ron ItiCHES

.537

NFIN---TYPE OF PATERIAL USED FOR FINS (IF RQouIEklD SPF.IN---SPACi1G BETWE!N FINS ( EPISF---SURFACE EMISSIVITY Of THEi fl( DIMFNSI(INlESS)

HEL[n jigs

1 1I4E wi4E A I 10HE aTI SR C

SREG

21 22C�C.

1�.C.O

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'TYPE NPItlS MCAN OtC hA I i CA t NT C Ai HTVnN b HilFUEL

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MISHL MOSHL MOLIh TKISHL TKOSHL IKOLIIJ

rflf: $1If11./ C I

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x

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DIV~tiSIONIk~iCNES) a 2 5o

5.3"') oc75

2.130 .soc .550 * coo 6.640

8.000 7.820

14.0000 .0000

13. 370

THERMAL PARAMETERS A'OIEFIT TEMP = 1!0.G00 (DEG.F) S0LAA INIJSALNCE a 1 (t-YES,?-No) TOTAL DECAY hEAT 2 .6C') (KxdS

"~ANI"U' STEADY STATE TE"PER4DTURES

DEGHEE S-F

o -L I NE R I

216.33

N-SHIELDI- I 0-T

.05 217.S3*. 1.2G

0-SHELL P9-S

217.19 .26 211!15

HIELI I-SHELL l-T D -T

.36 219.36 021

"HE-GAP T D-T

?3!031 10.95

IIISER T T D-T

24.*31 .000

HE-GAP T D-T

261.66 234.35

CANI STER I D-T

266.39 4.7!

FUEL PIN I D-T DEG-C

--- - - - - - - - - -352.0? R5.69 177.82

C

SURFACE T

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I L-C

(PUP ) MUDI"I

YEARS WATIS/CUFT WATTSIASS NISECIASSY PISECIASSY

"IELEM---TYPE OF WASTE MATFRIAL (SHOUN IN BRACKETS) OU ----- -AVERAGE BURNUP (EDIT FOR GOOKEEPING PURPOSES ONLY; TIME .--- COOLING TIMF (AGE OF FUEL SINCE DISCHARGE) WHEAT*-DECAY HEAT GIVEN OFF BY THE HASTE MATERIAL; NOT USED DHEAT-;-DECAT HEAT GIVEN OFF BY EACH ASSEPBLY (0 CANISTER); SRCN-•--NEUTRON SOUPCE (EDIT FOR UOOKEEPING PURPOSES ONLY; SRCG-;--PNOTON SOURCE (EDIT OR BOOKEEPING PURPOSES ONLY;

NO LONGER USED)

If ZERO NOT USED IF ZERO NO LONGER USED) NO LONGER USED)

DESCRIPTIO4 OF WASTE CONTAINER

ITYPE kPI4S MCAV OCCAM TKCICA NTCA% 4TVY'lil NTIJ[EL

2 •04

.•3

1•.k

I H INCHES INCHES fcET

FEET FEET

DESCRIhIO'4 OF INSERT

E'II?. I

TKII!ST TPOISN IKCrAP MCGAP TI IGAP 'qIGAP WTHICK hELEN CASKID

5

.C30

13

13

:,19C 1

( AL )

INCHES I4CHES INCHES ( HE ) I4CHES ( HE ) IVICHES

13.37 INCHES

ITYPE--LFLAG: I FOR SQUARE CANISTERS; 2 FOR SQUARE ASSEMIBLIES (WITHOUT CAN) NPINS-:ZNUMBER OF FUEL PINS PER ASSEMBLY (If PELEM DENOTES PUR OR OUR) "MCAN*--TYPE OF MATERIAL USED FOR CANISTERS INO CAN USED IF MCANuO) ODCAH---OUTSIDE DIM OF CANISTER ("CAN1O), O WIDTH OF FUEL ASSEMBLY INCAMa0) TKCAN---UALL THICKNESS OF CANISTER (IF MCAN.GT.O) HTCAN-;-LENGTH OF CANISTER (O FUEL ASSEMBLY) HTVnID--PORTION OF CANISTER NOT OCCUPIED BY HASTE MATERIAL (If OCAN*GTO) HTFUEL-;-PORTION OF FUEL ASSEMBLY CONTAINING UO2 (IF NELEM DENOTES PUN OR BUR)

MINSRT--TYPE OF MATERIAL USED FOR INSERT (SHOWN IN BRACKETS) ErINSt--SURFACE EMISSIVITY OF THE INSERT MATERIAL (DIuENSIOkLESS) TKINST--IHICKNESS OF INSERT BETWEEN ASSEMPLIES (INCLUDES TPOISN) TPOISN--THICKNESS OF NEUTRON POISON IMB0ODED IN INSERT PAIL BETWEEN ASSEMBLIES TKCGAP--THICKNESS OF GAP BETWEEN CANISTER AND INSERT MCGAP--:TYPE OF GAP ATMOSPHERE (SHOWN IN BRACKETS) TWICAP--THICKNESS Of GAP PETUEEN INSERT AND THE INNER SHELL MIGAP-YTYPE OF GAP ATNOSPIERE (SHOWN IN BRACKETS) WTHICK--THICKNESS OF THE INSERT BETWEEN CANISTE% AND INNER SHELL NELEM--;-NUMOER Of ASSEMBLIES (OR CANISTERS) PER CASKI IF ZERO, PERFORM SEARCH CASKiD-;INSIDE DIAMETER OF THE CASK (CALCULATED BY CODE If USER ENTERS 0.0)

DESCRIPTIOfl Of INNER I OUTER SHELL AND THE OUTSIDE LINER

6 6

4750

C SS ) ( SS ) C SS ) INCHES ICHES I•CHES

MISHL---TYPE Of MATERIAL USED FOR THE INNER SHELL (SHOWN IN BRACKETS) MOSHL---TYPE OF MATERIAL USED FOR TNH OUTER SHELL (SNOWNIN BRACKETS) MnLIN---TYPE OF MATERIAL USED FOR OUTSIDE LINER AND FINS (if REQUIRED) TKISHL--THICKNESS OF INNER SHELL TvOt HL--HINCKNESS OF OUTER SHELL TKOL•U-;THICK4ESS OF OUTSIDE LINER

DtfCRIrTIO'I Of %ELTaOFl AND GA.'rA SHIELtS

MNSI*L P 15 (H20 ) MISbLO--TYPE OF MATERIAL USED FOR NFUTROM SHIELD (SHOWN IN PRACKETS) "NSS4L1 I ( Pol ) MCSNLD--TYPE OF MATERIAL USED FOR GANMA SHIElD (SNUWN IN BRACKETS)

DESCQI1I|IIZ Of HEAT TlaASFER PARAMETERS FOR Files (2 CASK)

£ SS ) MFI4 ---- TYPE !F 4ATERIAL USED FOR FINS (IF REQUIRED) SPFFto---bPACItIG OETWEEN FINS E|ISf---SliRFACE EMISSIVIIY PF IHE( S (DI'lkSIloIl[SS)

F L! 14 Du

WHEAT SNE AG

SRCG

21 40oCC.

.50 pn

HISHL POSHL MCLI'.

1KOhL TI OL I*.'

~P4F It

(

TITLE: ONfE ASSEIPLY (WISXlS FUFL, .5 YEARS COOLED. NLI-1/2 CASK )

DESCR|P|IIt; OF WASTE MATERIAL

I

i

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PAN IPUp STEADY STATE TEMPERATUIRES

DEGREES-U

SUN FACE O-LINER1 1 D-T

N-SHIELD I D-T

3303 313082 .69 342.52 8 .70

0-SHELL

346.01 3.49

P8-SHIELD

351.27 5.26

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55.303? 2e60

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INSERT T D-I

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543.o?6 .flO

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1021.73

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(Q (

--------------

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This page Intentionally left blank.

.APPENDIX D

SECTION VIII

Metallic Fuel Computer Printout

Revised Oct. 1986 Feb. 1987

VIII-DI

I

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C

(PUN )

YEARS WATTSICUFT WATTSIASSY NISECIASSY PISECIASSY

MELER---TYPE OF WASTE MATERIAL (SHOWN IN BRACKETS) lU .----- AVERAGE BURNUP (EDIT FOR BOOKEEPING PURPOSES ONLY; TIME ---- COOLING TIEf (AGE OF FUEL SINCE DISCHARGE) WHEAT---DECAT HEAT GIVEN OFF BY THE WASTE MATERIAL; NOT USED DHEAT---DECAY HEAT GIVEN OFF Of EACH ASSEMBLY (OR CANISTER); SRCH----NFUTRON SOUICE (EDIT FOR BOOKEEPING PURPOSES ONLY; SRCG ---- PHOTON SOURCE (EDIT FOR OOKEEPING PURPOSES ONLY;

NO LONGER USED)

IF ZERO NOT USED If ZERO NO LONGER USED) NO LONGER USED)

DESCRIPTION OF WASTE CONTAINER

I

S '.53

.125 12.00

.00 17.00

C AL ) INCHES INCHES FEET FEET FEET

DESCRIPTION OF INSERT

S ( AL ) .,zf, ,-63 INCHES .103 INCHES .440 INCHES

14 (AIR ) .!sr) INCHES

14 (AIR ) .125 INCHES

13.3? INCHES

ITYPE---FLAG: 1 FOR SQUARE CANISTERS; 2 FOR SQUARE ASSEMBLIES (WITHOUT CAN) NPINS---NUnMER OF FUCL PINS PER ASSEMBLY (IF HELEM DENOTES PUN OR BUR) "NCAN----TYPE OF MATERIAL USED FOR CANISTERS (HO CAN USED IF MCANsO) ODCAN---OUTSIDE DIN OF CANISTER (MCAN1O), OR WIDTH OF FUEL ASSEMBLY (MCAN-O) TKCAN---WALL THICKNESS OF CANISTER (IF MCAN.GT.O) "HTCAN---LENGTH OF CANISTER (OR FUEL ASSEMBLY) "HTVOID--PORTION OF CANISTER NOT OCCUPIED BY WASTE MATERIAL (IF HCAN.GT.O) "HTFUEL--PORTION OF FUEL ASSEMBLY CONTAINING U02 (IF MELEm DENOTES PUN OR BUR)

MINSRT--TYPE OF MATERIAL USED FOR INSERT (SHOWN IN BRACKETS) EMINST--SURFACE EqMISSIVITY OF THE INSERT MATERIAL (DIMENSIONLESS) TKINST--THICKNESS OF INSERT BETWEEN ASSEMBLIES (INCLUDES TPOISN) TPOISN--THICKNESS OF NEUTRON POISON IMOEDDED IN INSERT MATL BETWEEN ASSEMBLIES TKCGAP--THICKNESS OF CAP BETWEEN CANISTER AND INSERT MCGAP---TTPE OF GAP ATMOSPHERE (SHOWN IN BRACKETS) TKIGAP--THICKNESS Or GAP BETWEEN INSERT AND THE INNER SHELL HIGAP---TYPE OF GAP ATMOSPHERE (SHOWN IN BRACKETS) UTHICK--THICKNESS or THE INSERT BETWEEN CANISTER AND INNER SHELL NFLEM---NU99ER OF ASSEMBLIES (OR CANISTERS) PER CASK; IF ZERO, PERFORM SEARCH CASKID--INSIDE DIAMETER OF THE CASK (CALCULATED BY CODE IF USER ENTERS 0.0)

DESCRIPTION OF INNER I OUTER SHELL AND THE OUTSIDE LINER

4IS1L 6 MOSHL 6 MOLIN 6 TKISHL .500 TKOSHL 0Ts TKOLIS) .?t

( SS)

(SS)

INCHES INCHES INCHES

MTSHL---TYPE Of MATERIAL USED FOR THE INNER SHELL (SHOWN IN BRACKETS) HOSHL---TYPE OF MATERIAL USED FOR THE OUTER SHELL (SHOWN IN BRACKETS) MOLIN---TYPE Of MATERIAL USED FOR OUTSIDE LINER AND FINS (IF REQUIRED) TKISHL--THICKNESS OF INNER SHELL TKOSHL--THICKNESS OF OUTER SHELL TKOLIN--THICKNESS OF OUTSIDE LINER

DE0CqRITION OF NEUTRON AND GAMMA SHIELDS

MNSHLD 15 (HO ) MNSHLD--TYPE OF HATERIAL USED FOR NEUTRON SHIELD (SHOWn IN BRACKETS) MGSHLb I ( PD R MGSHLb--TYPE OF MATERIAL USED FOR GAMMA SHIELD (SHOWN IN BRACKETS)

DESCRIPTION OF HEAT TRANSFER PARAMETERS FOR FINS (S CASK)

6 ( SS ) e.'in INCHES lt97.

mFIN ---- TyPE OF MATERIAL USED FOR FINS (IF REQUIRED) SPFIN---SPACIN6 BETWEEN FINS EMISF---SURFACE EMISSIVITY OF THE FINS (DIMENSIONLESS)

""E|CR IPT "0 .O.fI dIe' . de •1 ,, "A.KtI/.AiL WAAS, TJEi ,. M ATAE l9JI)

DESCRI"•TIOWI OF J/ASTC M.ATERIAL

MELEM Ott

WHEAT bHE AT

SRCIs

21 16"0. 2 * 30

*1

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C

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APPENDIX E

SECTION VIII

Mark 42 Fuel Assembly Computer Printout

VIII-El Page added August 1988

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DESCRIPTION OF WASTE CONTAINE

NELE--'.-TYPE OF WASTE MATERIAL (SHOWN IN BRACKETS) lU•":';-AV[RAGE BURNUP (EDIT FOR BOOKEEfPING PURPOSES ONLY; TIHE*';::-COOLING gTIrE (AGE OF FUEL SINCE DISCHARGE)

lEA1-•DECAY HEAT GIVEN OFF fiBY THE WASTE MATERIAL; NOT USED D91EAT'-':DECAY HEAT GIVEN OFF BY EACH ASSE4flLt (OR CANISTER); SRCtI--=NEUTfRON SOURCE (EDIT FOR BOOKEEPING PURPOSES ONLY; SRCG-•-;P OTON SOURCE (EDIT FOR BOOEEPING PURPOSES ,ONLY; Ar

"NOLONGER USED)

IF ZERO NOT USED If ZERO NO LONGER USED) NO LONGER USED)

ITYPE---FLAG: I FOR SQUARE CANISTERS; 2 FOR SQUARE ASSEMBLIES (WITHOUT CAN) NPINS:-:NUPflER Of FUEL PINS PER ASSEMBlLY (IF NELEN DENOTES PWR AO OUR) M"PTYS' tTPE OF MATERIAL USED FOR CANISTERS (NO CAN USED IU MCAH.O) ODCA:---OUTSIDE DIM OF CANISTER tMCAN),O) ON WIDTH Of FUEL ASSENULV IMCANSO) TKCAH-:l--ALL THICKNESS OF CANISTEO (IF MCAN.GT.O) "NTCA"-.-LENGTN Of CANISTER (OR FUEL ASSEMOLY) "HTVOID--PORTION OF. CANISIER NOT OCCUPIED BY WASTE MATERIAL (IF MCA9.GT.o) "HTFUEL:-PORTION OF FUEL ASSEMBLY CONTAINING U02 (If RELEN DENOBES PUR ON OUR)

NINSRT.--TVPE OF MATERIAL USED FOR INSERT (SHOWN IN BRACKETS) S91NS 7 -- SURFACE EMISSIVITY Of THE INSFAT MATERIAL (DIMENSIONLESS) IKISTttTNICKNESS Of INSERT BETWEEN ASSfEOLIES (INCLUDES TrOISm) IPOISNH:THICKNESS Of NElUTRON POISON INIEDDED IN INSERT RAIL BETWEEN ASSEMBLIES IKCGAP--TsICKNESS OF GAP BETWEEN CANISTER AND INSERT MCGAP-Z-TYPE nF GAP.ATMOSPHERE (SHOWN IN BRACKETS) TKIGAP--TNICKNESs OF GAP BETWEEN INSERT AND THE INNER SHELL MIGAP-L-IYPE Of GAP ATMOSPHERE (SHOWN IN BRACKETS) UTHICK-'-IslI(KNItS OF THE INSERT BETWEEN CANISTER AND INNER SHELL NFL1M-:FN'-ilPIR 01 ASSEIMLIES (OR CANISTERS) PER CASK; If ZERO, PERFORM SEARCH CASKI|D'INSIDE DIAMETER OF THE CASK (CALCULATED BY CODE IF USER ENTERS 0.0)

ITYPE NPIllS MCA4 ODCAII IKCA" MICA" "HTwnlb "HIFUEL

I

6 4.Z3

14.75 .00

1 2.67

( SS I INCHES lIe c ES FEET FEET FEET

DESCRIPTIO*1 qF INSERT

MINSRT EMINST TKINST I PC)ISN

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HELEN CASrin

6 .•53

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14 d14

I

13.3? INCHES

( SS )

INCHES INCHES INCHES (AIR ) INCHES (AIR ) INCHES

au LnI1u.I Or JOINER a OUTER SHELL AND THE OUTSIDE LINER frt & ................

6 C SS ) 6 ( SS ) 6 ( Ss )

• Oo INCHES .I7S INCHES • ?S0 INCHES

MIlSOIL-;ýTVPE OF MATERIAL USED FOR THE iNNER SHELL (SNOWN IN BRACKETS) M:OSNL---TYPE OF MATERIAL USED FOR THE OUTER SHELL (SHOWN IN BRACKETS) RMOLlN-:;TVPE OF MATERIAL USED FOR OUTSIDE LINEN AND FINS (IF REQUIRED) TKISHL•THICKNESS Of INNER SNELL TKOSHL:"-TICKNESS OF OUTER SHELL TKOLI":-THICKNESS OF OUTSIDE LINER

DESCRIPTION OF NEUTRON AND GANHA SHIELDS

is (H02) MINSHLB-ýTYPE Of MATERIAL USED FOR NEUTRON SHIfLD (SHOWN IN BRACKEIS) 1 ( Pa ) NGS"Lb-.TYPE OF MATERIAL USED FOR GAMMA SHIELD (SHOWN IN BRACKETS)DEISCRPTION OF HEAT TRANSFER PARAMETERS FOR FINS (3 CASK)

6 C SS ) 1.101 IUACOES

.SUqnrFIN---=lYPE Of M4ITRIAL USED FOR FINS (IF RIQUlRDel SPIH.-:-SPACING OFIilEN FINS EMISF-•SSUFACE ENISSIVITY • FINS (DIi•ncSiOntc.s

TITLE: NL1.1i2 WITH A MAqK42 ASSEMBLY: INSOLANCE: 130 DEG ANBIENT

DESCRIPTION OF WASTE MATERIAL

C

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Auus 198

APPENDIX F

SECTION VIII

Mark 22 Fuel Assenbly Computer Printout

VIII- FPage Added February 1990

S C o P 1 INPUT (iN CARD-IMAGE FORMAT) FOLLOW:

SCOPE INPUT FOR MLI-1/2 CONTAINING TWO MARK 22 ASSEMBLIES * tmSiBBm B UUUUUZRUUUUUUUZUZZZS ZZZZSUEZZZUUZUZZUUUUUEUUEUZ3UUUUN

MELEN DURNUP TIME WHEAT DHEAT SRCN SRCG ITYPE KPINIS 21 253000 150 0 1726 7.87.5 9.75.15 1 2

MCAN ODCAN TKCAN HTCAN HTVOID NTFUEL 5 4.000 0.25 14.67 1.67 12.5

MINSRT ENINSAT TKINSRT TPOISM TKCGAP TIIGAP WIHTICK 6 0.58 0.625 0.125 1.5243 0.215 0

NELEM CASKID 2 13.37

* MISHL HOSHL MOLIN MFIN TKISHL TXOSHL TMOLlN MGSHLD 6 6 6 6 0.500 0.875 0.25 3

* SPFIN EMISF EMISC TFNMAX WGHTMX TAMB NSOLAIU 4 0.59 0.59 400 500 130 1

* KTRANS RHONS TCNS1 TCNS2 EPNS2 (FOLLOWING DATA IS FOR NS4FR) 5 0 0 0 0

GENERAL FORMAT FOR SHIELDING DATA: LTYPE, NUMPTS* (NASSYS(I,LTYPE),TKG(I,LTYPE),TKNCI,LTYPE). I=I,NUKPTS)

SHIELDING DATA FOR NLI-1/2 WITH TWO MARK 22 ASSEMBLIES:

3 1 2 4.875 5.0

* TERMINATION FLAG (FOR THIS PARTICULAR CASE): 0

Page Added VIII-F2 February 1990

PROPERTIES OF MATERIALS CURRENTLY IN THE SCOPE DATA LIBRARY

MATERIAL DENSITY CONDUCTIVITY BEAT CAPACITY TEMPERATURE LIMIT (L8/CUFT) (BTU/HR/FT/F) (8TU/LO/F) (DEGREES F)

1 2 3 4 5 6 7 8 9

10 11 12 13 14 15 16 17 18 19 20 21 22 23

p8 FE U CU AL SS NA LI

PB-L comC ALSI DOwA

HE AIR H20 HUL3 SHOT NULl VLVG NUL2 PWR BUR HLWC

708.56 4M.26

1189.25 559.35 168.49 494.43 S45.00 30.00

1"6.60 707.60 170.00 62.00

.01

.08 62.43 81.20

370.00 284.00 212.00 203.00 220.90 199.30 113.00

19.3000 26.0000 15.0000

210.0000 129.0000

9.5200 38.0000 20.0000

.4400 18.0000 80.0000

.0760

.1200 .0360 .3920 .3000 .3500

1.2000 .7000 .6000

1.0000 1.0000 .2500

.0320

.1200

.0280

.0950

.2280

.1200

.3000 1.0000

.1560

.0320 .2000 .5260

1.2400 .2600

1.0000 .0660 .0950 .0660 .1600 .0660 .1000 .1000 .2200

618 1950 1450 1730 1050 1800 1400 1400 1200

618 1065 600

1400 1400

250 10o 1400 1000 1290 1000 1650 1650 932

NOTE: THIS EDIT WILL BE PRINTED ONLY ONCE, EVEN TNCOUGH THE USER MAY HAVE MULTIPLE SETS OF INPUT DATA.

Page Added February 1990

CAPITAL COST (SLI)

7.500 5.000

22.500 1.00 .700

10.000 .500

27.500 4.700

.400

.880

2.500 .000 .000 .000 .000 .000 .000 .000 .000 .000 .000 .000

VIII-F3

TITLE: SCOPE INPUT FOR NLI-1/2 CONTAINING TWO MARK 22 ASSEMBLIES

DESCRIPTION OF WASTE MATERIAL

HELEN 21 (PWR ) NELEM-..TYPE OF WASTE MATERIAL (SHOWN IN BRACKETS) DU 253000. MWINT BU ...... AVERAGE URNUIP (EDIT FOR BOOKEEPING PURPOSES ONLY; NO LONGER USED) TIME 150.00 YEARS TIME ---- COOLING TIME (AGE OF FUEL SINCE DISCHARGE) WHEAT 0.01100 WATTS/CUFT WHEAT---DECAY HEAT GIVEN OFF BY THE WASTE MATERIAL; NOT USED IF ZERO DHEAT 1.7E103 UATTS/ASSY DHEAT---DECAY HEAT GIVEN OFF BY EACN ASSEMBLY (OR CANISTER); NOT USED IF ZERO SRCN 7.91105 N/SEC/ASSY SRCN ---- NEUTRON SOURCE (EDIT FOR BOOKEEPING PURPOSES ONLY; NO LONGER USED) SRCG 9.8E+15 P/SEC/ASSY SRCG ---- PHOTON SOURCE (EDIT FOR BOOKEEPING PURPOSES ONLY; NO LONGER USED)

DESCRIPTION OF WASTE CONTAINER

ITYPE NPINS NCAN OOCAN TKCAN HTCAN NTSABS HTFUEL

DESCRIPTION

MINSRT EMINST TKINST TPOISN TKCGAP TKIGAP WTHICK NELEN CASKID

DESCRIPTION

MISHL MOSHL MOLIN TKISHL TKOSHL TKOLIN

DESCRIPTION

MNSHLD MGSHLD

DESCRIPTION

MFIN SPFIN EMISF EMISC

1 ITYPE---IuCIRC CANISTERS, 2xSQUARE ASSYS (NO CANS), 32SOR ASSYS WITH SQR CAMS 2 NPINS---N.UMBER OF FUEL PINS PER ASSEMBLY (IF MELEN DENOTES PWR OR B3R FUEL) 5 C AL ) MCAN ---- TYPE OF MATERIAL USED FOR CANISTERS (NO CAN USED IF MCANzO)

4.00 INCHES ODCAN---OUTSIDE DIAN OR WIDTH OF CAN (MCAN.GT.O), OR WIDTH OF ASSEMBLY (MCAN=O) .250 INCHES TKCAN---WALL THICKNESS OF CANISTER (IF MCAN.GT.0) 14.67 FEET HTCAN---LENGTH OF CANISTER (ITYPEul OR 3), Ol LENGTH OF FUEL ASSEMBLY (ITYPE=2) 1.67 FEET HTSABS--LENGTN OF INTERNAL SHOCK ABSORBERS HOLDING ASSYS OR CANISTERS IN CASK

12.50 FEET HTFUEL--ACTIVE LENGTN OF U02 FUEL, OR NT OF-SOLID WASTE MATL IN CAN (FULL IF 0)

OF INSERT

6 c SS ) NINSRT--TYPE OF MATERIAL USED FOR INSERT CSMMOIJ I1 BRACKETS) .580 ENINST--SURFACE EMISSIVITY OF THE INSERT MATERIAL (DIMENSIONLESS) .625 INCHES TKINST--THICKNESS OF INSERT BETWEEN ASSEMBLIES (INCLUDES TPOISN) .125 INCHES TPOISN--THICKNESS OF NEUTRON POISON IMBEDDED IN INSERT MATL BETWEEN ASSEMBLIES

1.524 INCHES TKCGAP--THICKNESS OF GAP BETWEEN CANISTER AND INSERT (AIR-FILLED IF * 0) .215 INCHES TKIGAP--THICKNESS OF GAP BETWEEN INSERT AND THE INNER SHELL (AIR-FILLED IF 0 0) .000 INCHES WTNICK--THICKNESS OF THE INSERT BETWEEN CANISTER AND INNER SHELL

2 NELEM---NUMBER OF ASSEMBLIES (OR CANISTERS) PER CASK; IF ZERO, PERFORM SEARCH 13.37 INCHES CASKID--INSIDE DIAMETER OF THE CASK (CALCULATED BY CODE IF USER ENTERS 0.0)

OF INNER A OUTER SHELL AND THE OUTSIDE LINER

6 ( SS ) MISNL---TYPE OF MATERIAL USED FOR THE INNER SHELL (SHOWN IN BRACKETS) 6 ( SS ) MOSHL---TYPE OF MATERIAL USED FOR THE OUTER SHELL (SHOWN IN BRACKETS) 6 ( SS ) MOLIN---TYPE OF MATERIAL USED FOR OUTSIDE LINER AND FINS (IF REQUIRED)

.5D0 INCHES TKISHL--TNICKNESS OF INNER SHELL

.875 INCHES TKOSHL--TNICKNESS OF OUTER SHELL

.250 INCHES TKOLIN--TNICKNESS OF OUTSIDE LINER

OF NEUTRON AND GAMMA SHIELDS

15 (H20 ) MNSHLD--TYPE OF MATERIAL USED FOR NEUTRON SHIELD (SHOWN IN BRACKETS) 3 ( U ) MGSHLD--TYPE OF MATERIAL USED FOR GAMMA SHIELD (SHOW IN BRACKETS)

OF HEAT TRANSFE& PARAMETERS FOR FINS (1 CASK)

6 C SS ) NFIN ---- TYPE OF MATERIAL USED FOR FINS (IF REQUIRED) 4.000 INCHES SPFIN---SPACING BETWEEN FINS

.590 ENISF---SURFACE ENISSIVITY OF THE FINS (DIMENSIONLESS)

.590 ENISC---SURFACE ENISSIVITY OF THE CASK (DIMENSIONLESS)

Page Added February 1990

VIII-F4

CASK DESIGN PARAMETERS

TFNMAX £00.0 DEG.F TFNMAX--NAXII4UJ ALLOWABLE SURFACE TEMPERATURE 'GHTNX 500.0 KILO.LIS iGHTMX--NAXIIUM ALLWABLE WEIGHT OF LOADED CASK

TANS 130.0 DEG.F TAM .---- JTSIDE AMBIENT TEMPERATURE NSOLAR I INCLUDE NSOLAR--SOLAR HEATING IN NORMAL STEADY-STAT9 CALCULATION (INCLUDE/IGNOREaI/0)

KIND OF TRANSIENT THERMAL ANALYSIS TO BE PERFORMED

ITRANS S OPTION SELECTED BY USER; OTHER POSSISLE OPTIONS INCLUDE:

KTRANSU4 ... ASSUME LIQUID WATER NEUTRON SHIELD IS LOST MANY HOURS BEFORE START OF FIRE (ORIGINAL SCOPE DEFAULT) KTRANSxS ... ASSUME LIQUID NEUTRON SHIELD IS LOST INSTANTLY AT THE START OF THE FIRE (STANDARD NRC SCENARIO) KTRANS96 ... ASSUME A SOLID NEUTRON SHIELD, USE DATA BELOW, AND SWITCH TO NEW THERMAL CONDUCTIVITY DURING FIRE

THERMAL CHARACTERISTICS OF SOLID NEUTRON SHIELD (THE FOLLOWING DATA IS NOT USED UNLESS KTRASs6)

&HONS 62.43 LBM/CUFT RHONS--NONINAL DENSITY OF THE SOLID NEUTRON SHIELD TCNSI .3920 BTU/HR/FT/F TCNSI--NONINAL THERMAL CONDUCTIVITY OF THE SOLID NEUTRON SHIELD TCNS2 .3920 BTU/HR/FT/F TCNS2-,-THERMAL CONDUCTIVITY OF TNE SOLID NEUTRON SHIELD DURING AND AFTER THE FIRE CPNS2 1.0000 STU/LD/DG.F CPNS2--NEAT CAPACITY OF THE SOLID NEUTRON SHIELD DURING AND AFTER THE FIRE

VIII-F5 Page Added February 1990

S C 0 P I -- THE SNIPPING CASK OPTIMIZATION AND PARAMETRIC EVALUATION CCOE, VERSION 1.2 (BY J. A. BUCHOLZ, ORML)

TITLE: SCOPE INPUT FOR NLI-1/2 CONTAINING TWO MARK 22 ASSEMBLIES

COMPONENT DIMENSIONS (INCHES):

O-LINER a N-SHIELD a O-SHELL a G-SHIELD s I-SHELL a I-CASK a I-GAP a U-POISM a Y-INSRT x C/F-GAP a O-CANSTR a I -CANSTR a Q-LENGTH a

.250 5.000

.875 4.875

.500 13.370

.215

.125

.625 1.524 4.000 3.500

150.000

THERMAL PARAMETERS:

SOLAR INSOLANCE * I (INCLUDED) AMBIENT TEMP a 130.000 (DEG.F) TOTAL DECAY HEAT a 3.452 (KV)

NOMINAL STEADY STATE TEMPERATURES (DEGREES-F)

SURFACE T

242.72

O-LINER OT TMKAX)

........ o4.94. .. 2?. 242.94

N-SHIELD DT T(MAX)

3.65 246.59

O-SHELL DT T(MAX)

1.10 247.69

PB-SHIELD OT T(MAX)

5.13 252.87

GAP INSERT DT T(NIN)

119.73 373.75

INSERT OT T(MAX)

17.97 391.72

UAP CAN DT TCNIN)

191.27 583.00

CANISTER OT TCMAX)

.08 583.08

FUEL PIN DT T(MAX) DEG.C

190.08 773.15 411.73

CM) OUTSIDE DIAMETER OF THE CASK BODY a 36.37 INCHES; NO EXTERNAL COOLING FINS WERE REQUIRED. TOTAL WGNT OF CASK LOADED VITH 2 FUEL ASSEMBLIES ON CANISTERS IS APPROXIMATELY 57310 LBS.

Page Added February 1990

VIII-F6

I-SHELL OT T(KAX)

1.13 254.00

PREFIRE TEMPERATURES (DEGREES-F)

TIME... (.N.M.S)

RADIUS (FT)uu,

.00 0. 0. 0

.00 0. 0. 0

IMIT TEMP RISE

TIME.. .(.N.S)

999.0 999.0.0

MAX TEMP RISE: AT TIME (MRS):

MAX FIRE TEMP:

•....... INlSERT .... *--GAP--* ---- I .SKELL ---- * ... G.SHIELD .... 0 .... O.SHELL ---- * .... K.SH iELD ---- * .... O. L IER ...... *

.4153 .4813 .5392 .5571 .5783 .5987 .8272 1.0050 1.0421 1.0779 1.3030 1.4946 1.5050 1.5154

REFERENCE STEADY-STATE TEMPERATURES USED IN CALCULATING TEMP RISE (ASSUMES N-SHLC PRESENT & NO SOLAR HEATING): 358.95 348.81 340.98 209.23 206.64 204.24 201.08 199.14 196.64 194.30 192.28 190.25 189.77 189.30

INITIAL STEADY-STATE TEMPERATURES AT START OF 30-N]ITi FIRE (SEE ASSUMPTIONS :;R ITRANSzS): 358.95 348.81 340.98 209.23 206.64 204.24 201.08 199.14 196.64 194.30 192.28 190.25 189.77 189.30

.00 .00 .00 .00 .00 .00 .00 .00 .00 .00 .00 .00 .00 .00

ULTIMATE POSTFIRE TEMPERATURES (DEGREES-F)

....... INSERT .... *-'GAP--.---. I.SNELL ----.... G.SHIELD .. '..*...O.SHELL .... *.'.N.SHIELD .... * ---- O.LINER ......

ULTIMATE STEADY-STATE TEMPERATURES, LONG AFTER THE 30-NIMUTE FIRE IS OVER: 438.12 427.98 420.14 315.02 312.43 310.03 306.87 304.94 302.43 300.09 245.17 190.25 189.77 189.30

TRANSIENT TEMPERATURES (DEGREES-F)

....... INSERT ---- *--GAP-- ---- I.SHELL .... *---G.SHIELD ----. .... O.SNELL ---- * ---- N.SHIELD .... • .... O.LINER .----

108.08 108.09 108.10 171.22 173.85 176.35 180.27 192.69 260.88 347.43 710.32 1074.52 1094.58 1114.",

7.33 7.33 7.17 1.20 1.18 1.14 1.00 .62 .52 .50 .50 .50 .50 .50

467.04 456.90 4"9.08 380." 380.49 380.59 381.34 391.83 457.52 541.73 902.60 1264.77 1284.35 1303.73

DURING AND AFTER THE 30-MINUTE FIRE:

MAX GAMMA SHIELD TEMP .......... 392 DEG.F a 200 DEG.C MAX BASKET INSERT TEMP ......... 67 OEG.F a 242 DEG.C MAX FUEL PIN CLAD TEMP ......... 805 DEG.F a 429 DEG.C

AS PER 1OCFR73, SECT 71.73. THE EFFECTS OF SOLAR HEATING WERE NEGLECTED PRIOR TO. DURING, AND AFTER THE 30-NINUTE FIRE.

VIII-F7 Page Added February 1990

This page intentionally left blank.

APPENDIX G

SECTION VIII

Failed Metallic Fuel Basket Computer Printout

Page Added Oct. 1990VIII-GI

S C 0 P B INPUT (II CARD-INAGE FORMAT) FOLLOWS:

Evaluation of ILI-1/2 Containing 6 Failed Fuel Rods (30 WATTS TOTAL BEAT LOAD)

N MELEN BURNUP TINE WHEAT DHEIT SICI SICG ITYPE IPINS 21 1600 2 0.0 5.0 6.31+4 1.12+14 1 1

' ECAN ODCAV TICAN M CAN HTVOID HTFIJL 5 3.0 0.125 12.00 0 12.00

M MIS[? EMINSRT TKINSRT TPOISM 5 0.22 0.01 0.0

NELEN CASKID 6 13.37

M NISEL NOSEL MOLII NFIN 6 6 6 6

TKCGAP TIIGAP ITHICK -0.01 3.75 0.0

TKISBL TKOSEL TIOLII 0.5 0.8175 0.25

NGSILD 1

SPFIN EXISF EXISC TFIMAX WGITNI 1 0.5817 0.50 750 200

KTRANS REGIS TCIS1 TCYS2 CP1S2 6 0.371 0.0154 0.0154 0.24

GENERAL FORMAT FOR SHIELDING DATA: LTYPE, NUMPTS, INASSYSIILTYPE)oTKGIILTYPE).TKN(I,LTYPE), :lolNUXPTS)

SHIELDING DATA FOR 6.5-INCH PB-METAL CASKS (ASSUMES A 5-INCH NEUTRON SHIELD)

11 6 4.875 5

* TERMINATION FLAG (?OR THIS PARTICULAR CASE): 0

Page Added' Oct. 1990 VIII-G2

PROPERTIES OF KATERIALS CURREITLY IN THE SCOPE DATA LIDIARY

KATERIAL DENSITY CONDUCTIVITY HEAT CAPACITY TEMPERATURE LIKIT WLBICUFT) ISTUIIIIRFTI) (ITUILII)f (DEGREES F)

7

2 3 4 5 6 7 8 9

10 I1 12 13 14 15 16 17 18 19 20 21 22 23

?I FE

U CU

AL SS IA LI

Pl-L COIC ALSI DONA HE

AIR 120 HUL3 SHOT HULl

HLVG IUL2 PVR IeV HLYC

701.56 488.26

1189.25 559.35 168.49 494.43 45.00 30.00

146.60 707.60 170.00 62.00

.00

.01 62.43 11.20

370.00 284.00 212.00 203.00 220.90 199.30 113.00

19.3000 26.0000 15.0000

210.0000 80.0000 9.6900

33.0000 20.0000.

.4400 11.0000 80.0000

.0760 .1200 .0360 .3920 .3000 .3500

1.2000 .7000 .6000

1.0000 1.0000

.2500

.0320

.1200

.0280

.0950

.2280

.1200

.3000 1.0000

.1560

.0320 .2000 .5260

1.2400 .2600

1.0000 .0660 .0950 .0660 .1600 .0660 .1000 .1000 .2200

618 1950 1450 1730 1050 1800 1400 1400 1200

618 1065 600

1400 1400

250 1000 1400 1000 1290 1000 1650 1650

932

NOTE: ?HIS EDIT VILL IE PRINTED ONLY ONCE. EVEN THOUGH THE USER KAY HAVE MULTIPLE SETS OF INPUT DATA.

Page Added Oct. 1990

CAPITAL COST ISL)

7.500 5.000

22.500 1.400

.700 10.000

.500 27.500 4.700

.400 .880

2.500 .000 .000 .000 .000 .000 .000 .000 .000 .000 .000 .000

VIII-G3

TITLE: Evaluatiot of 1LI-1/2 Coutaihial 6 failed hel tods 130 IITTS TOTAL BEAT LOD)

DESCIIPTIOI OF WASTE KATEIAIL

MILEN 21 (FIR ) NLEI--ffl OF VAST! MATERIAL 1SHO1N IN BRACKETS) 11 1600. KIDII? 1O ------. AVERAGE BUR1V I (EDIT FOR BOOKEEPING PURPOSES ONLY: N0 LONGER USED) TINE 2.00 TEARS TIHI---COOLING TIN! (AGE OF FUEL $IKE DISC1ARGE) VIE!T 0.O+E00 VITTSICUT VWEITT---DECAY HEAT GIVEN OfF BY THE WASTE NATERIAL: IOT USED IF ZERO 0111T 5.01+O0 VATTSiASST DIEI?---DECAY 51AT GIVEN OFF IT EACH ASSEMBLY 1OR CANISTER): NOT USED IF ZERO SICI 6.3.+04 NISEC/SSY SRC ---- NEUTRCN SOURCE IEDIT FOR EOOIEEPING PURPOSES OILY: NO LONGER USED) SROG 1.1E+14 PISECIASSY SRCG ---- PHOTON SOURCE (EDIT FOR 3OOIEEPIIG PURPOSES ONLY: N0 LONGER USED)

DESCIIPTION OF WASTE CONTAINER

ITYPE 1 ITYPE---I;CIRC CANISTERS. 2'SQUARE ISSYS (NO CARS), 3:St ASSYS VITH SQ1 CANS IPIIS. I NPINS---NUNDER OF FUEL PINS PEI ASSEMBLY (IF HELEN DEVOTES PVW OR BVI FUEL) NCAN 54 AL I 1CA1 ---- TYPE Of MITEIIAL USED FOR CANISTERS (NO CAN USED If lCAN:O) ODCAI 3.00 INCHES ODCAI---OUTSIDE DIIN OR WIDTH OF CAN INCAIN.GT.0). O WIDTH OF ASSEMBLY (KC1A=O) TICAN .125 INCHES TKCAN---AILL THICKNESS OF CANISTER IIF XCAX.GT.0 ITCAN 12.00 FEET HTCAN---LENGTH OF CAIUSTER 1ITYPE;I 01 3), 0 LENGTH OF FUEL ASSEMBLY (ITT'!Ez) HTSAIS .00 FEET 8TSAUS--LEIGT! OF INTERNAL SHOCK ABSORBERS HOLDING ASSYS OR CANISTERS IN CASK ITFPUEL 12.00 FEET BTPEL--ACTIVE LENGTH OF U02 FUEL, OR IT OF SOLID WASTE NATI IN CAN (FULL IF 0)

DESCRIPTION OF INSERT

HINSRT 5 ( AL I N!NSRT--TYPE OF MATERIAL USED FOR INSERT (SHOVW IN BCRIKETS EIaIST .220 ENINST--SURFACE UISS:J:TY -2 72-; :533RT MATERIAL (DINENSIONLESS) TINS? .010 INCHES TKINST--TIICKUESS OF INSERT BETWEEN ASSEMBLIES (INCLUDES TPOISN) TPOISN .000 INCHES TPOISN--TIICIIESS OF NEUTRON POISON IMBEDDED IN INSERT KNT! BETWEEN ASSENBLIES TKCGIP -.010 NCIES TKCGAP--TIICINESS OF GAP BETWVEE CANISTER AND IISERIT (H-FILLED IF 0 )1 TIIGAP 3.750 INCHES TIIGAP--THICKNESS OF OAP BETWlEE INSERT AID THE INNER SNELL (AII-FILLED IF ) 0) 1T7I1C .000 INCHES VIICK--TIICIiESS OF THE INSERT ETWVEEI CAIISTER AND IiNER SHELL NELEN 6 NELE---NUNBER OF ASSEMBLIES [Ot CANISTERS) PER CASK: If ERMO. PERFORM SEICH CASIID 13.37 INCHES CASKID--INSIDE DIAMETER Of THE 6ASK !CALCULATED BY CODE IF USER ENTERS 0.0)

DESCRIPTION Of INNER & OUTE S.ELL AND TIE OUTSIDE LINER

MISIL 6 SS I MISL---TYPE OF MATEIIAL USED FOR THE INNER SHELL (SHOWN I1 BRACKETS) MOSIL 6 SS I MOSHL---TYPE OF MATERIAL USED POR THE OUTER SHELL (SHOVN IN BRACKETS)

OLIN 6 SS I MOLII---?TYPE OF MATERIAL USED FOR OUTSIDE LINER ALD FINS 1IF REQUI1ED1 TKISHL .500 INCHES TKISNL--TICKNIBSS OF 1INER SNELL ?KOSIL .875 INCHES TKOSIL--TNICKIESS Of OUTER SNELL TIOLIN .250 IICHES TIOLIN--THICKIESS OF OUTSIDE LINrS

DESCRIPTION OF NEUTRON AID GAMMA SHIELDS

KISELD 0 ISOLID I ISHfL.--TvPE Of ATER':AL USE POR NKEUTRION SHIELD (DESCRIBED BELOW IF SOLID) NGSILD 1 I PB ) NGSILD--TYPE OF NATER:AL USED FOR GAMMA SHIELD (SHOVW !I BRACKETS)

DESCRIPTION OF HEAT TIAISFEI PARANETEIS FOR FINS 1 CASK!

NKIX 6( I S M ) .FIN.---- YPE OF NATERIAL USED FOR FINS (IF REQUIRED) .SPFII 1.000 INCHES SPFII---SPACIIG IBETEEI FINS

ENISF .587 ENISP---SURFACE ENISSIVITY OF THE FINS IDIMENSIONLESS) ENISC .500 ENISC---SURFACE NISSIVITY oR ':. ; ASK (DIMEISIONLESS)

Page Added Oct. 1990 VIII-G4

CASK DESIGN PARAMETERS

750.0 DEG.? 200.0 KILO.LBS 100.0 DEG.F

I INCLUDE

TFNMAI--MAIIMUK ALLOWABLE SURFACE TEMPERATURE IGHTKI--KAIIKUK ALLOWABLE WEIGHT OF LOADED CASK TAI B---- OUTSIDE AMBIENT TEMPERATURE ISOLAR--SOLAR HEATING I1 NORMAL STUDY-STATE CALCULATION (IICLUDE/IGNORE:I/0)

KIND OF TRANSIENT THERMAL ANALYSIS TO BE PERFORMED

KTRANS 6 OPTION SELECTED BY USER: OTHER POSSIBLE OPTIONS INCLUDE:

KTRANS=4 ... ASSUME LIQUID WATER NEUTRON SHIELD IS LOST MANY HOURS BEFORE START OF FIRE (ORIGINAL SCOPE DEFAULTJ KTRANS:5 ... ASSUME LIQUID NEUTRON SHIELD IS LOST INSTANTLY AT THE START OF TE FIRE (STANDARD IRC SCEIARIO) KTRAKS=6 ... ASSUME A SOLID NEUTRON SHIELD, USE DATA BELOW, AID SWITCH TO NEW THERMAL CONDUCTIVITY DURING FIRE

THERMAL CHARACTERISTICS OF SOLID NEUTRON SHIELD (THE FOLLOWING DATA IS NOT USED UNLESS KTRIAS=6)

LBMICUFT BTUIHRIFTIF BTU/HI/FTIF BTU/LB/DG.F

IRONS--NOIKNAL DENSITY OF THE SOLID NEUTRON SHIELD TCNSI--NOKINAL THERMAL CONDUCTIVITY OF THE SOLID NEUTRON SHIELD TCKS2--THERKAL CONDUCTIVITY OF TiE SOLID NEUTRON SHIELD DURING AID AFTER THE FIRE CPIS2--HEAT CAPACITY OF THE SOLID NEUTRON SHIELD DURIIG AND AFTER THE FIRE

Page Added VIII-5 Oct. 1990

TFIKAI

TANB KSOLAR

IRONS TCNS1 TCNS2 CPIS2

.07 .0154 .0154 .2400

S C 0 P I -- T11 SNIPPING CASI OPTIIIUTIOl ANND PARIITlC EVALUATION COol, VESImON 1.2 IT J. A. IUCIOL, OuILI

TITLE: balmatiom of ILI-1I/ Contajisig 6 failed Fuol Rods 130 VITTS TOTAL BEAT LOAD)

COCNP01EN DIKEISIONS UIICID):

O-5I]1t "

O-SHELL G-SIIELD I-SHELL I-CASK I-GAP V'POISI 3-INSRT elF-GAP 2

0-CIISTI '-CANSTR Q- LEIGTR

.250 5.000

.075 4.075

.500 16.580 3.750 .000 .010 .010

3.000 2.750

143.750

THEIL PAIRNETERS:

SOLAR IISOUNCE I (INCLUDED) AXBIEIT TEIP 100.000 (DEG.F1 TOTAL DECAY HEUT .030 (M[I

INOINAL STEADY STATE TERPEIATUIES tOEIBEES-F)

0-LINER OT TINAi.

.00 l11.16

INSERT OT T(Nil.I

.02 215.59

I-SHIELD DT TAl INi)

A6.11 207.26

O-SHELL OT TIKAI)

.00 207.2'

GAP CAN DT ThAIN)

.09 2!5.61

PB-SHIELD DT TMill

.03 :07.30

CANISTER OT T(INil)

.............. .30 215.60

I-SHELL D? ?0(711)

M• 207.)1

FUEL PIN OT ?INAII 130.P

2.:2 2!1.10 103.22

('1 OUTSIDE DIAINETER Of THE CASI BODY : 39.51 INCHES: 1O EXTERNAL COOLING FINS VERE REQUIRED. TOTAL VGI? OF CASS LOADED NITA 6 FUEL ASSENBLIES OR CANISTERS IS APPROIIITSLY 30441 LIBS.

Page Added Oct. 1990 VIII-G6

SURFACE

181.16

GAP INSERT DT Till)

1.27 215.58

PRIME11 TEPER1AUES IDEGIES-?)

.3--- ...-- .. "33 . ... 7.1. . 7.. I.957 -.1 -'" .175 .... 1... 52 1 . 1--O.33ln 1.----

• 3012 .3451 .3713 .690t .7120 .7325 .9574 1 .1118 1 .1751 : .2 "'? 1 .4353 1 .3283 1.6311 1.1492

REFERENCE STEADY-STATE TENPERATURES USED 11 CALCULATIEG TEP RISE (1553S I-SEL PRESENT 1 10 Soul IEATII;:: .0• 0. 0. 0 138.63 131.67 138.67 127.00 127.00 127.00 126.91 126.97 126.96 126.96 112.09 100.98 100.97 100.9?

NITIAL STEADY-STITE TENPERATURES AT START OF 30-NIIUTE PilE (SEE ASSUpTIONS pot M113.6): 133.68 138.67 1383.67 127.00 127.00 "127.00 126.98 126.97 126.96 126.96 112.09 100.91 100.97 :00.r

.00 .00 .00 .00 .00 .00 .00 .00 .00 .00 .00 .00 .00 .30

ULTINATE POSTFIRE TIEKPERATRES (DEGIEES-F)

A ------- 1139IT ---- I- P---... I.SRELL ---- A --- .SKIIIL ---- I---- O.flBIL ---- 2 .... E.SlIELD ---- I ---- O.l ...... ---

ULTIKATE STEADY-STATE TENI(PRATUIES. LONG AFTER Tie 30-KINUTE FIRE IS OVER: !38.61 131.67 138.67 127.00 127.00 1.27.00 126.91 126.97 126.96 126.96 112.09 100.93 100.97 100;97

TRAISIENT TEMPERATURiS (DEGIEES-F)

---- ..... SEIT ...---- A " ---- I.SIELL ---- I"---G.SBIELD ---- 2 ---O.SIELL ...--- .. 1.31IL ---- a --- M.INER ......-

2.86 2.36 2.16 3.40 3.41 3.41 3.41 3.41 3.41 3.41 757.39 1322.11 1322.16 1322.90 21.67 23.67 23.67 4.17 4.17 4.17 4.08 4.00 4.00 3.92 .50 .50 .50 .50

141.54 141.53 141.52 130.41 130.41 130.40 130.39 130.38 130.37 130.37 169.91 1423.79 1423.13 1423.11

DURING AND AFTEi THE 30-KIIUTE FIRE:

KAI GAKXA SMIELD TEN? ........ 130 DEC.! z 54 DEG.C NAi BASKET INSEIT ?E ......... 142 DE0.! z 61 DEG.C MAX FUEL Pig CLAD TEK? ......... 145 DEG.1 a 63 DEG.C

AS PER 10CF173. SECT 71.73. TIE EFFECTS OF SOLR WEATING 1ER iELECTED PRIOR TO. oullIs. AgD AFTER TuE 3i-MINU Fill.

VIII-G7Page Added

Oct. 1990

?IKE... MI.

RADIUS (FT):z)

.00 0. 0. 0

lIlT TEN? RISE

TIME... M.I.S)

999.0 999.0.0

MAI ?EN? RISE: AT TIME IB1SI:

NAi FIRE TIM:

This page intentionally left blank.

Section MX

SECTION DC

SHIELDING ANALYSIS

K..

This page intentionally left blank.

SECTION IX SHIELDING

Table of Contents

INTRODUCTION

1.0 SUMMARY

2.0 DESIGN FUEL SOURCE TERMS

3.0 METHODS OF ANALYSIS

3.1 Shield - Dose Point Description

3.2 Primary Gamma Dose Rates

3.3 Neutron Dose Rates

3.4 Secondary Gamma Dose Rates

3.5 Sound Scatter

3.6 Hypothetical Accident Conditions

4.0 RESULTS

4.1 Normal Conditions

4.2 Hypothetical Accident Conditions

4.3 Shielding Analysis for Consolidated PWR

4.4 Shielding Analysis for Metallic Fuel

4.5 Shielding Analysis for PWR or BWR Rods

4.6 Shielding Analysis for the Mark 42 Fuel

4.7 Shielding Analysis for the Mark 22 Fuel

5.0 REFERENCES

APPENDIX A

APPENDIX B

APPENDIX.C

Type Fuel Rods

Assembly

Assembly

Cask Configuration "B"

Supplemental Analysis for High Burnup PWR Fuel

Supplemental Analysis for Fermi-I Fuel

and EBR-II Blanket Fuel

Revised Feb. 1987 May 1987 Dec. 1988 Feb. 1990 Oct. 1990 Feb. 1991i

IPag e

iv

IX- 1

IX-4

IX-4

IX-4

IX-l1

IX-17

IX-21

IX-21

IX-22

IX-25

IX-25

IX-31

IX-34

IX-37

IX-37d

IX-37h

IX-37k

IX-38

IX-Al IX-Bi

IX-C1

Table of Contents (cont'd)

Pagie

4.2.3 FETA Input .......................................... VIII-32

4.3 Normal Operations Results ........................... VIII-32

4.4 Hypothetical Fire Results ........................... VIII-32

4.5 Cold Operations .................................... VIII-36

4.6 Maximum Fuel Pin Temperatures ....................... VIII-38

4.7 Thermal Expansion and Contraction .................. VIII-38

4.8 Uncertainty Factors and Design Margins .............. VIII-38

4.9 General Conclusions ................................. VIII-39

4.10 Effect of Personnel Barrier on Thermal Performance.. VIII-40

4.11 Effect of Consolidated PWR Fuel Rods ................ VIII-41

4.12 Effect of Metallic Fuel Rods ....................... VIII-43a

4.13 Effect of PWR and BWR Rods ........................... VIII-43b

4.14 Effect of Mark 42 Fuel Assemblies ................... VIII-43d

4.15 Effect of Mark 22 Fuel Assemblies ................... VIII-43e

5.0 REFERENCES

APPENDIX A Cask Configuration "B" ........................... VIII-Al

APPENDIX B 2 KW, Residual Moisture .......................... VIII-B1

APPENDIX C SCOPE Computer Printout .......................... ViII-C1

APPENDIX D Metallic Fuel Computer Printout .................. VIII-DI

APPENDIX E Mark 42 Fuel Assembly Computer Printout .......... VIII-El

APPENDIX F Mark 22 Fuel Assembly Computer Printout .......... VIII-FI

APPENDIX G Failed Metallic Fuel Basket Computer Printout .... VIII-GI

Revised Feb. 1987 May 1987 Dec. 1988 Feb. 1990 Oct. 1990 ii Feb. 1991

List of Tables (cont'd)

Top and Bottom - Metallic Fuel Dose Rates

(Hypothetical Accident)

Metallic Fuel Dose Rates - Total

PWR Assembly and PWR or BWR Source Strengths

25 PWR Rod Dose Rates

25 BWR Rod Dose Rates

PWR Fuel and Mark 42 Fuel Source Strengths

Mark 42 Fuel Assembly Dose Rates

Mark 22 and Mark 42 Fuel Source Strengths

Mark 22 Fuel Assembly Dose Rates

IX-19

IX-20

IX-21

IX-22

IX-23

IX-24

IX-25

IX-26

IX-27

List of Figures

NLI 1/2 Reference Design Shield Model - Normal

Design Dose Point Locations for NLI 1/2 Reference Cask

Top and Top - Side Shield - Dose Point Relationship

Page

IX-7

IX-12

IX-16

Page Added May 1987

Revised Dec. 1988 Feb. 1990 Oct. 1990 Apr. 1992iii

IX-37b IX-37c

IX-37e

IX-37f

IX-37g

IX-37h

IX-37J

IX-37k

IX-371

I

No.

IX-1

IX-2

IX-3

INTRODUCTION

The shielding model used to calculate the radiation levels around the top corner

of the cask is not an exact representation of the actual top corner arrangement.

Minor design modifications were made in this area subsequent to the shielding

analysis. The following sketch shows the differences between the two

configurations. It can be seen that configuration changes contribute to the

shielding adequacy and, therefore,. a revision to the corner analysis is not

necessary.

111-ý~~

Revised Feb. 1991 iv

"*Final design configuration is addressed in Appendix A.

riL .....----. -- " . / 1

r -~m am -- m-m-

V

*0

a' 0

V5 c.

Revised Feb. 1991

i

This page Intentionally left blank.

Section IX

SHIELDING

1.0 SUMMARY

The purpose of this section is to define the total gamma and neutron dose rates to be expected outside the NLI 1/2 spent fuel shipping cask. The fuel to be shipped and the corresponding fission product source inventories are discussed in Section III. Detailed dose rate calculations have been performed for the fuel parameters identified in Section In for a 40,000 MWD/MTU PWR fuel assembly with an average operating specific power of 40 kw/kgU,

and an initial enrichment of 3.35 w/o U-235 in 0.454 MT of uranium.

The NLI spent fuel shipping cask shield has been designed to insure that the. radiation emanating from the cask is effectively reduced to levels equal to or less ,than those levels currently specified by the DOT or AEC hazardous materials shipping regulations. The pertinent radiation standards controlling

the shield design are:

From Section 173.393 of Reference 1 for normal conditions of transport

(1) 1000 millirem per hour at 3 feet from the external surface of the package(closed transport vehicle only);

(2) 200 millirem per hour at any point on the external surface of the car or vehicle (closed transport vehicle only);

(3) 10 millirem per hour at 6 feet from the external surface of the car or vehicle; and

(4) 2 millirem per hour in any normally occupied position in the car or vehicle, except that this provision does not apply to private motor carriers."

"IX-1

From Section 71.36 of Reference 2 which specifies the standards for hypothetical accident conditions for a single package

"(a)•.

(1) The reduction of shielding would not be sufficient to increase the external radiation dose rate to more than 1000 millirems per hour at 3 feet from the external surface of the package.

The shield materials .used in the cask design have been selected and arranged to minimize the cask weight while maintaining overall shield effectiveness. Lead and depleted uranium were chosen as effective gamma radiation shields, and a water Jacket on the outside of the cask was provided to efficiently moderate the neutron radiation.

The total neutron and gamma dose rates calculated for the design fuel loading for design-shipping conditions are shown in Table IX-1. It can be seen that the maximum dose rate is located on the fuel axial midplane, and does not exceed the limits specified above. The total neutron and gamma dose rates under hypothetical accident conditions arb also given in Table DC-1. Again, it can be seen that the maximum expected dose rate is located on the fuel axial midplane and is within the specified limits.

The dose rates given in Table DC-1 include those from neutrons and gammas originating in the fuel, from neutrons and gammas sca--ered from the ground, from secondary gammas resulting from neutron capture in the shield, and from neutrons originating from fissions in the depleted uranium shield. All other. sources of dose are insignificant. The details of the calculations and results are described in the following sections.

IX-2

C

* Dose Point Behind Vehicle Cab.

'1

0�

I-a

I-'

TABLE IX-I

NLI 1/2 SHIPPING CASK SUMMARY OF MAXIMUM DOSE RATES

mrem/hr

C

FUEL NIDPLANE TOP BOTTOM

66 from 6' from . 6' from Cask 3' from Personnel Cask 3' from Personnel Vehicle Cask 3' from Personnel

Surface Cask Shield Surface Cask Shield Cab Surface Cask Shield

Normal Conditions

Gamma 27.8 3.58 12.9 1.94 0.56 13.2 1.44

Neutron 25.0 5.01 29.5 5.71 1.41 37.5 5.29

Total 52.8 8.59 42.4 7.65 1.97 50.7 6.73

49 CFR 173 Limit 10 10 2 10

Hypothetical

Accident Conditions

Gamma 1438 400 63 14.5 72 13.1

Neutron 991 387 154 53.0 224 62.0

Total 2429 787 217 67.5 296 75.1

10 CFR 71 Limit 1000 1000 1000

2.0 DESIGN. FUEL SOURCE TERMS

Section III of this report defines the PWR and SWR fuel design and operating conditions on which the cask design is based and develops the corresponding neutron and gamma source strengths. The results, summarized in Table II-3,1indicate that both source terms are larger for the 1 PWR assembly cask loading then for the 2 BWR assembly cask loading. Detailed gamma and neutron dose rate calculations were therefore performed only for the limiting 40,000 MWD/MTU PWR assembly.

Table ID-2 gives the PWR design information used as a basis for the cask shield design. The energy distributions of the primary gamma and neutron sources are given in Table IX-3. The basis for these distributions is given in Section I1. In addition to the primary neutron and gamma sources originating in the spent fuel being shipped, secondary neutron and gamma sources occur because of interactions of the primary radiation with the cask materials. These sources and the resulting doses are discussed later in this section.

3.0 METHODS OF ANALYSIS

Presented below is a discussion of the techniques used to perform gamma and neutron dose rate calculations leading to the design of the NLI spent fuel shipping cask. Included is a definition of the various computer codes used in the cask design , a description of the configuration and composition of the source and shield materials used, and the location of the various dose points considered.

3.1 Shield - Dose Point Description

The shield overlay shown in Figure TX-i and described in Table IX-4 was used to represent the NLI 1/2 cask for normal shipping conditions. Table DC-5 defines the material overlay densities and elemental composition. All

IX-4

Table IX-2

PWR REFERENCE DESIGN INFORMATION

Loading *454 kgU/assembly

Iritial Enrichment

Average Specific Power

Average Burnup

Cooling Time

Total Gamma Energy

Neutron Source

*3.35 w/o

40 kw/kgU

40,000 MWD/MTU

150 days

3.074 x 1016 MeV/sec

7.55 x Joe n/sec

*UraniLum loading and initial enrivtnent values were subsequently increased to 475 kgU/assembly and 3.7 w/o respectively. Increasing the uranium loading frum 454 Kg to 475 Kg could result in a combined gamma and neutron source strength increase of a similar magnitude (about 4.6%). The only shielding problem this could cause is in the doe at the truck cab (see table IX-l). This small Increase, however, is more than offset by the fact that a uniform axial source distribution was used for the direct gamma dose calculations. If the correct axial distribution were used, it would decrease the direct ga, contribution by a factor of 2 (see page IX-27). This would result in a decrease in total dose at the truck cab of 9.Wv which mre'•tan compensates for a 4.6% increase in source strength.

DC-S

Table DC-3

DESIGN BASIS PWR ASSEMBLY GAMMA AND

NEUTRON ENERGY DISTRIBUTIONS

Gamma

Energy Group Mev

0.1 - 0.4

0.4 -0.9

0.9 - 1.35

1.35 - 1.8

1.8 -2.2

>2.2

Relative Distribution

150 Days After Shutdown

1.2 x 10-2

0.97526

1.8 x 10"3

9.4 x 10-4

1.0 x 10-2

<2 x 10-4

Assumed Gamma Energy MeV/7

0.30

0.80

1.25

1.50

2.00

Gamma Decay Source MeV/sec

3.6888 x 1014

2.9979 x 1016

5.5332 x 1013

2.889 x 1013

14 3.074 x 10

Total 3.074 x 10 1 6

Neutron

Energy Group MeV

<0.1

0.1 -0.4

0.4 - 0.9

0.9 - 1.4

1.4 - 3.0

>3.0

Relative Distribution

Combined Spectrum

0.031

0.163

0.150

0.383

0.273

Total

rX-6

Neutron Source n/sec.

0

0.24 x 108

1.23 x 108

1.13 x 108

2.89 x 108

2.06 x 108

7.55 x 108

14,

16

19

21

22

23

24

2 27,

28

SEE FIGURE IX -3 FOR DETAIL ARRANGEMENT

29 waou

FIGURE IX - 1 NLI 1/2 REFERENCE DESIGN SHIELD MODEL

NORMAL TRANSPORTATION MODE

IX-7

TABLE DC-4

SHIELD REGION AND IDENTIFICATION

REGION

1

2

3

4

5

6

7

a

9

10

11

12

13

14

15

16

17

18

19

20

21

22

23

24

25

26

27

28

29

Dc-8

MATERTAL

Dry Fuel Region

Void

Aluminum

Steel

Void

Steel

Void

Uranium

Lead

Steel

Cold Water

Steel

Balsa Wood

Steel

Void

Steel

Uranium

Steel

Void

Void

Dry Fuel

Void

Void

Steel

Void

Steel

Uranium

Steel

Balsa Wood

THICKNESS cm

12.05

4.7625

3.5075

0.635

0.3175

1.27

0.0668

6.985

5.3975

2.2225

12.7

0.635

40.64

2.8575

9.8425

3.175

7.62

1.905

40.1075

7.2

365.76

8.89

30.1625

1.27

0.3175

3.4925

7.9375

3.175

40.64

TABLE DC-5

NLI 1/2 REFERENCE CASK SHIELD DENSITIES USED IN THE SHIELD DESIGN *

Shield Material

Fuel Region.

Aluminum Region

Stainless Steel Regions

Depleted' Uranium

Lead Shield Region

Water in Shield Tank*'

Balsa Wood

* See Figure lX-1 for shield ** Taken at,-. 360 0F

Number Densit, 4 (atoms/cc) x 10Element

Oxygen

Chromium

Iron

Zirconium

Nickel

U-235

U-238

Pu-239

Pu-240

Aluminum

Chromium

Iron

Nickel

U-235

U-238

Lead

Hydrogen Oxygen

Hydrogen

Carbon

Oxygen

arrangement

0.012975

0.0000085

0.0000307

0.00405

0.0001566

0.0000395

0.00605

0.0000303

0.0000083

0,059006

0.01674

0.0606

0.00988

0.0001053

0.04785

0.033

0.05919 0.02960

0.0040938

0.003024

0.0017628

IX-9

materials thus defined conform with standard material compositions

specified in References 3, 4, 5 and 6 for the materials of construction K.

specified in Section VI. For example, the U-235 content of the depleted

uranium shieldtaken as 0.22 w/o U-235, is in accordance with depleted

uranium material composition specifications given in Reference 3. The

water density wa4 based on calculated water tank temiperatures of

approximately 360°F. Of course, these temperatures are not expected to

be attained and maintained, and thus these densities will assist in

insuring a conservative shield design. The maximum expected temperature

is 3S2°F for the shield tank. The density of balsa wood used was

conservatively assumed to be 0.12 g/cc (See Reference 6) composed of

60% cellulose (C 6 HI 005 ) and 35% Lignin (C2 0 H 2 0 06 ).

It will be noted that all shield regions were assumed to maintain cylind

rical geometries. All regions except the fuel region are cylindrical. The

fuel region cross-sectional area was calculated based on dimensi~ns

specified in Table 111-2 (PWR-Westinghouse) and the equivalent circular

cross section was used in the dose rate calculations. This cylindrical

configuration will tend to underestimate th; gamma and neutron dose

rates adjacent to the assembly corner and will tend to overestimate the

dose rates adjacent to the assembly side. Calculations, described later,

were performed to determine the magnitude of this effect.

Dose points for the normal shield design were chosen and placed in accordance

with conditions specified in 49 CFR 173M discussed in Section 1.0 above.

Thus, dose points were placed at the fuel midplane on the surface of the

shield Jacket and 10 feet off the cask centerline (i.e., 6 feet off the external sur

face of the vehicle) and on the cask centerline on the surface.of the balsa wood

IX-10

I--.

IX-11

placed on the cask top and bottom, 6 feet off the personnel barrier at

the cask top and bottom, and 197 inches off the cask top surface

(to obtain dose rates to be expected behind the vehicle cab). These dose

points are shown in Figure IX-2.

3.2 Primary Gamma Dose Rates

The QAD-PS computer code was used as the principal tool for calculating

the primary gamma dose rates. The QAD-P5 computer code, a version of

the QAD-IV computer code (7) calculates both uncollided and collided gamma (and neutron) dose rates, energy deposition and fluxes for a volumetric

source represented by a number of point isotropic sources in a user

specifi~d shield configuration. For each dose point the straight distance

and attenuation in each shield 'material is calculated for each source point.

The total dose is obtained by summing the contributions from each source

point.

The code QAD has had extensive use in industry for dose rate calculations

and has been shown to yield satisfactory results. W, 9) Calculations

performed on similar shield configurations with hand methods and other

computer codes have shown agreement with results generated by the QAD

code. The gamma cross sections used in the QAD code were obtained from

Reference 10, whereas the neutron removal cross sections were obtained from

Reference 11. It should be noted, as will be discussed later, that the

neutron portion of this code was used only for obtaining relative neutron

dose rates off the cask surface.

The buildup factors used to calculate the ratios of uncollided-to-collided dose

rates were obtained from Reference 12. The general equation used for the

buildup factor calculation is

B(;or,E) = [tsr]i [E].J + or°E) 1=O J=O + e rE

for the ringes 0'or £15 and 0.5 MeV -E £10 MeV.

C

4.5

'7

173"

FIGURE IX - 2 DESIGN DOSE POINT LOCATIONS FOR NLI 1/2 REFERENCE CASK

K

E is the energy and gor is the optical thickness of the shield material traversed by the gamma photon. The coefficients C and correction

terms cJ were obtained from Reference 12.

The above equation defines buildup factors effectively for only one element.

Uranium was chosen for the basic analyses since (1) it is the most effective

primary gamma attenuating material in the cask, and (2) the QAD code can

handle buildup for one shield material per computer calculation.

To define the actual expected buildup effect for the NLI cask shield complex,

additional hand calculations were performed to determine the actual buildup

that might be experienced on the cask side, top, and bottom. A weighted

mean buildup factor was calculated for a 2 MeV gamma (this is the major

dose contributor) using Taylor's buildup equation. (13) The side buildup factor was calculated for the gamma traversing the region directly adjacent

to the fuel midplane whereas the top and bottom buildup factors were

calculated for the gamma transversing the region directly adjacent to

the fuel axial centerline. The resulting bi•Idup factor was compared

to the uranium equivalent buildup factor calculated for the same cross

sections using the equation given above. This comparison shows that the

actual buildup factor may increase the side dose rates by 57% and the top

and bottom dose rates by 45%. This increase can be attributed to the presence of a substantial quantity of light material (i.e., steel, shield

water). The gamma flux to dose rate conversion factors given in Table IX-6

were used in the QAD program. These values are consistent with the

information in References 14 and 15.

The source and shield configuration described in Figure DC-1 and Tables

DC-4 and DC-S was input into the QAD code and dose rates calculated at

DC-1 3

Table IX-6

GAMMA FLUX-TO-DOSE RATE CONVERSION FACTORS

Gamma Enerqgy MeV

0.3

0.8

1.25

1.50

2.00

Flux-to-Dose Conversion Factor

mr/hr/photon/cm 2 - s ec

5.7669 x 10-4

1.4878 x 10-3

2.1524 x 10-3

2.447 x 10-3

2.9995 x 10-3

DC-1 4

I i "4

the dose points specified in Figure IX-2. The source given in Table IX-3 was considered to be uniformly distributed over the active fuel volume.

The source &'oss section was circularized to have the same area as that of the actual fuel assembly cross section. This assumption would be expected to overpredict the cask surface doses off the sides of the fuel assembly and underpredict the surface doses off the corners of the

assembly. To evaluate this effect additional detailed dose rate calculations were performed with QAD. In these calculations the source region was arranged to mock up the actual fuel cross-sectional configuration with the dose points off the comer of the assembly. The results on the cask surface and 10 feet from the centerline were compared to those for the

cylindrical source configuration. The surface gamma dose rates were found to be 11% higher than for the circularized source while the 10 foot

dose rates were 6% larger.

To confirm the adequacy of the shielding at other locations off the end and side of the cask, additional OAD calculations were performed for dose points placed along the surface of the cask as shown in Figure IX-3.

The results of this analysis, given in detail in Section 4.0, show that the dose rates off the top of the cask do not anywhere exceed that at the centerline (dose point 3, Figure DC-2) and that the dose rates off the side do not anywhere exceed that at the fuel midplane.

All of the above calculations were performed assuming the source to be distributed uniformly throughout the fuel region. In addition, dose rates were determined for an axial gamma source distribution taken to be the same as the burnup distribution given in Tigure I13-1 for 300 days operation.

This is a conservative representation of specific power and therefore of gamma source strength distribution for the reference fuel conditions. The results show that the midplane surface and 10 feet dose rates will

DC-IS

- -- - - a - - - - -

ACTUAL CONFIGURATION

CALCULATION MODEL

Ss

55

S4 . BALSA WOOD

S3I s3 I

-,, -- .

Is Si AT TRANSVERSE I CENTERLIN.E

FIGURE IX - 3 TOP AND TOP - SIDE SHIELD - DOSE POINT RELATIONSHIP

Ix-t-

increase by about 10% and 5% respectively for the 18% increase in peak source. These increases were included in the final result. The topside and top dose rates would decrease approximately 55%; however, no credit was taken for this decrease.

3.3 Neutron Dose Rates

The principal analytical tool for the determination of neutron dose rates from the shipping cask is the DTF-IV(1 6) computer code. DTF-IV is a onedimensional multigroup code which solves the Boltzman transport equation, including anisotropic scattering, by the discrete ordinate method. The DTF-IV code was developed by the Los Alamos Scientific Laboratory for use in calculating neutron flux through shield media. The code uses the same theory and equation for neutron flux calculation as the ANISN code (17)

which has been checked with several simple experiments described in

References 18 and 19.

For the present analysis the P1 -S 4 approximation to the transport equations was used with a 16 group neutron energy spectrum. The spectrum, taken fromtha of ansn an Roch,(20) from that of Hansen and Ro (20) is given in Table DC-7. The appropriate material cross sections were obtained from the Los Alambo Scientific Laboratory and are based on a U-235 fission spectrum.

The cask surface neutron fluxes determined by DTF-IV were converted to dose rates using the conversion factors in Table DC-7. (15, 21) The dose rates off the cask surface were obtained by using relative dose rate fall-off factors obtained from the neutron removal option of the QAD-PS program described previously. This option uses neutron removal cross sections to calculate neutron flux distributions.

The DTF-IV radial dose rates were checked with. several computer analyses performed with the ANISN code. The energy group specification, for cross

DC-1 7

Table EC-7

STANDARD GROUP STRUCTURE USED IN DTF-IV FLUX CALCULATIONS

Group

1

2

3

4

S

6

7

8

9

10

11

12

13

14

15

16

Minimum -Energy

3.0 MeV

1.4 MeV

0.9 MeV

0.4 MeV

0.1 MeV

17.0 keV

3.0 keV

555.0 eV

100.0 eV

30.0 eV

10.0 eV

3.0 eV

1.0 eV

0.4 eV

0.1 eV

0.001 eV

IX-1 8

Neutron Flux to Dose Rate Conversion Factors

mrem/hr/n/cm2-sec

0.13

0.13

0.127

0.071

0.030

0.008

0.00431

0.00455

0.0050

0.00555

0.00521

0.00492

0.00476

0.00463

0.00439

0.00374

sections and flux calculations used in the ANISN code is similar to the

DTF-IV distribution defined in Table lC-7. However, the cross section

data was generated, assuming a Cf-252 spontaneous fission spectrum,

using the CSCROS (23) computer code, which is a program containing

the GAM U1( 2 4 ) and THERMOS (25) codes. GAM II calculates the fast

neutron spectrum up to the P-3 or B-3 approximation In 99 fine groups

from 14.9 MeV to 0.414 eV. THERMOS solves the one-dimensional integral

transport equation in 30 groups with an upper energy cutoff at 1. 85 eV.

The results of the ANISN calculations were in satisfactory agreement with

the DTF-IV results.

In addition to calculation of the flux distribution resulting from the spontaneous fission and (a, n) reaction neutron source in the fuel, the DTF-IV

code was used to calculate the additional flux due to subcritical multi

plication (i.e., fission) in the fuel and the similar effect due to fissioning

in the uranium shield material. The surface fluxes and doses include all

of these effects.

DTF-IV problems were set up to calculate the flux distributions radially at

the fuel midplane and axially at the top and bottom of the cask using the

shield description of Figure IX-1 and Tables IX-4 and IX-S. The source

was considered to be uniformly distributed throughout the cylindrical fuel

region. The radial DTF-IV problems used a buckling height of 12 feet.

The x-y extrapolated buckling dimensions used for the axial calculations

were chosen such that the fuel assembly centerline fluxes were consistent

with the same fluxes obtained with the radial DTF-IV problems. This

resulted in x-y buckling dimensions of 43 x 43 cm.

The effect of circularization of the fuel region was discussed for gamma

dose rates in Section 3.2. A similar analysis using the neutron removal

option of QAD-P$ yields a maximum dose increase 10 feet off the cask

IX-19

centerline of 4%.

Based on the end of life axial burnup distribution given in Figure III-1

and the specific source strength of Figure Irr"3 the axial neutron source

distribution has a peak to average ratio of 1.2S. While earlier in life burnup distributions would have greater peak to average ratios, the average

burnup, and therefore total neutron source, would be considerably

less than at end of life. The neutron source axial peaking factor

of 1.25 was conservatively applied directly to the off the side neutron

dose rates.

The ýalculations described in Section 3.2 of gamma dose variations

around the top and top-side of the cask also yielded relative neutron

doses. These results show that the maximum neutron doses off the side and off the ends occur either on the cask axial centerline or at the fuel

midplane.

In addition to the detail computer calculations performed around the side and top-side of the cask, hand calculations were performed to

define the neutron dose rate contribution off the ends from neutrons

streaming through the cask cavity. The calculations were performed for

the dry fuel assembly for (1) direct radiation emanating from fuel pins

through the assembly interior to a worst point as compared to the

computerized homogeneous system, (2) radiation scattering through the

fuel assembly channels interior to the assembly, and (3) radiation scattering

through the assembly-aluminum gap. The calculations were performed

using equations defined in References 14 and 26. It was calculated that

the direct neutron dose rates off the ends should be increased by

approximately 100% to account for neutron streaming. These results

are included in the reported top and bottom neutron dose rates in Section 4.0.

IX- 20

3.4 Secondary Gamma Dose Rates

The secondary capture gamma sources and subsequent dose rates on the side and ends of the NUr spent fuel shipping cask were calculated using the ANISN computer code along with TAPEMAKER, (27)a routine to prepare a group independent cross section tape for ANISN. The ANISN code was used to calculate fast and thermal neutron fluxes in the NLI cask, and these results were then used by TAPEMAKER to generate capture gamma cross sections for each region. Then these results were put back into ANISN to calculate capture gamma source strengths and resulting capture gamma dose rates. The =AD-PS code was used to predict the capture gamma fall off from the cask side and end surfaces. For these calculations the source was taken to be located at the cask outer shell since this was found to be a major source of secondary gammas. These results are reported in Section 4.0.

3.5 Ground Scatter

Scattering of neutrons and gamma photons from the ground could contribute to an increase in the total dose rate off the cask surface. The amount of this contribution has been calculated for both gamma photons and neutrons using the NUSALE code. This code is a portion of the SOSC code developed for NASA-Goddard Space Flight Center. The NUSAL code integrates radiation scattered to a receptor by a differential area of the scattering material. Angular-differential scattering is determined according to the albedo formulae. The code results are in good agreement with graphic data published by A.B. Chilton. (2 9 )

The cask was taken to be horizontal 5.25 feet above the ground, which was taken to be concrete. The receptor was 3 feet above the ground, 10

IX-21

feet from the cask centerline. The neutron (and photon) calculation

was carried out over a ground projected area defined by a distance 1 10

meters from both source and receptor.

The cask was defined as four point sources spaced at 3.0 feet along the

cask axis. Each of these four sources were defined as having 25% of

the total source strength. The energy spectral distributions (normalized

to 1 mremihr) used in the ground scatter calculations are given in Tables

IX-8 and IX-9. These distributions were obtained from the primary gamma

and neutron radial shielding analysis of the cask with water inside the

cavity rather than the can and aluminum blocks. These distributions

differ from those for the cask with the dry can by less than 6%.

The neutron differential scattering was determined according to the

empirical relationship of Y. T. Song (30) which was derived from an

analysis of the Monte Carlo data of F. J. Allen. (3 1 ) Foi thermal

neutrons the relationship of R . L. French and M. B. Wells (32) was

used. The angular differential dose albedo for -gamma photons was *

obtained from the work of A . B. Chilton. (29)

3. 6 HYPothetical Accident Conditions

In order to show compliance with the requirements of 10 CPR 71,

calculations similar to those described in sections 3.2 and 3.3 were

performed for the expected post-hypothetical accident conditions of the

cask. These conditions are loss of shield jacket water, loss of balsa wood

and melting with complete loss of all lead shielding.

The calculations were actually performed for a cask design which

omitted the aluminum blocks inside the can and the I inch can itself. The

results are conservatively applicable to the final design with the aluminum

IX-22

Table IX-8

NORMALIZED NEUTRON DOSE DISTRIBUTION

E min.

3.0 MeV 1.4 MeV

0.9 MeV

0.4 MeV

0. 1 MeV

17.0 KeV

3.0 KeV

555.0 eV

100.0 eV

30.0 eV

10.0 eV

3.0 eV

1.0 eV

0.4 eV

0.1 eV

"0.001 eV

Group

-I1 *0.63-.1 is 0.63 x 10

IX-23

Normalized Group Dose (mrem)

0.184

0.340

0.128

0.151

0.63 -1w

1.26 -2

3.84 -3

5.13 -3

6.10 -3

5.22 -3

3.53 -3

4.33 -3

1.73 -3

2.56 -3

2.34 -3

8.69 -2

Total - 1.0 mrem

1

2

3

4

5

6

7

8

9

1-0

11

12

13

14

15

16

Table IX-9

NORMALIZED GAMMA PHOTON DOSE

Photon Energy (Mev)

0.8

1.25

1.50

2.0

Group Normalized Dose (mrem)

0.0031

0.0068

0.0176

0.9725

Total = 1.0 (mrem)

IX-24

Group

1

2

3

4

and the can in place. f

Primary gamma and neutron doses at the cask side and end surfaces and

3 feet off these surfaces were determined using the techniques previously

described. The shield configuration was as described in Figure DC-1

and Table DC-4 except that the aluminum, can wall, lead and shield

Jacket water were omitted for the radial calculations while the balsa

wood was omitted for the axial calculations. The end dose points were placed on the outer surfaces of the top and bottom heads. For the purpose

of defining the cask side surface (and the 3 foot off of the surface point)

it was assumed that the shield jacket was collapsed onto the cask outer

shell.

4.0 RESULTS

4.1 Normal Conditions of Transport

The calculated dose rates off of the sides and ends of the cask for normal

conditions are given in Table IX-10 and are discussed below.

The primary gamma dose rates on the fuel midplane as calculated by QAD

are 11 mrem/hr on the cask surface and 1.6 mrem/hr ten feet from the

centerline (i.e., 6 feet from the nearest accessible surface). To conservatively

obtain the maximum dose rate, these must be increased by the appropriate

amounts to account for the higher buildup factor due to the presence of

water, the comer peaking of the actual fuel cross section and the axial

peaking of the actual source distribution. These factors were discussed

previously and are 57% for the buildup factor, 11% and 6% for the comer

effect at the surface and 10 foot point respectively and 10% and 5% for

the axial effect at the surface and at 10 feet respectively. The results are

21 mrem/hr on the surface and 2.8 mrem/hr 10 feet from the centerline.

DC-25

STABLE IX-1O

NLI 1/2 CASK MAXIMUM DOSE RATES (Normal Conditions of Transport)

mremAr

Fuel Midplane Off End on Axial Centerline

Source TOP Bottom Category 6' from 6' from 61 from

Cask Personnel Cask Personnel Vehicle * Cask Personnel Surface Shield Surface Shield Cab Surface Shield

Primary Gamma 21 2.8 12.5 1.79 0.56 12.9 1.29

Neutron 25 3.34. 29.5 4.42 1.41 37.5 4.00

Secondary Gamma 6.8 0.61 . 0.37 • 0.033 0.37 0.033

Ground Scatter Gamma 0.17 0.12 0.12

Ground Scatter Neutron 1.67 1.29 1.29

Total Gamma 27.8 3.58 12.9 1.94 0.56 13.3 1.44

Total Neutron 25 5.01 29.5 5.71 1.41 37.5 5.29

Total 52.8 8.59 42.4 7.65 1.97 50.8 6.73

D bPoint Behind VehicleCab (See Figure IX-2)

(. (.

0)

!

.

The neutron dose rates from DTF-IV are 18.1 and 2.57 mrem/hr on the surface

and 10 feet from the centerline respectively both at the fuel midplane.

These must be corrected by 9% and 4% respectively for the comer effect and

25% (conservatively applied to both surface and 10 foot doses) for axial

peaking. The results are neutron doses of 25 and 3.34 mrem/hr at the

surface and 10 feet from the centerline respectively.

The results of the secondary gamma dose rate calculations yield midplane dose rates of 6.8 and 0.61 mrem/hr at the surface and 10 feet from the cask

centerline respectively.

The ground scatter calculations described in Section 3.S yield backscatter

dose factors of 0.05 for gamma photons and 0.5 for neutrons. These values are

the effective fractional increase in the 10 foot direct gamma and neutron dose due to ground scatter. Conservatively applying these to the. total maximum dose

derived above yields 10 foot ground Scatter dose'rates of 0. 17 mrem/hr and 1.67

mrem/hr for gammas and neutrons respectively.

Adding all the components gives combined dose rates on the midplane of

the fuel of 52.8 mrem/hr on the cask surface and 8.59 mrem/hr 10 feet

from the cask centerline. Table IX-1l summarizes the detailed derivation

of the 10 foot dose rate.

The direct gamma dose rates off the top of the cask on the axial centerline

were calculated to be 8.62, 1.23, and 0.38 mrem/hr at the balsa wood

surface, 6 feet off the personnel barrier and behind the truck cab respectively.

Off the bottom the dose rates are 8.85 and 0.89 mrem/hr on the balsa wood

surface and 6 feet off of the personnel barrier respectively. The buildup

correction factor applied to the gamma dose rates for the axial cases was

previously defined as 1.45 and should be applied to the above numbers.

The above numbers are based on a uniform axial source distribution. If

the correct shape was utilized the dose off of both ends would decrease by

IX-27

TABLE IX- 11

DETAILED DERIVATION OF DOSE RATE FUEL MIDPIANE 10 FEET FROM AXIS

(Normal Conditions of Transport)

Base Value or Percent Dose Rate

Correction mrem/hr

Primary Gamma

Base Calculated Value 1.6 Buildup Factor Correction 57%

Comer Correction 6% Axial Peak Correction 5%

Total 75%* 2.80

Primary Neutron Base Calculated Value 2.57 Comer Correction 4% Axial Peak Correction 25%

Total 30%* 3.34

Secondary Gamma 0.61

Ground Scatter Gamma

Direct Gamma Dose 3.41 Backscatter 5% 0.17

Ground Scatter Neutron

Direct Neutron Dose 3.34 Backscatter 50% 1.67

Total 8.59

*Total correction is product of individual factors

IX-28

about a factor of 2. Also note that no credit is taken for any attenuation

by the truck cab itself.

Including a 100% increase for streaming, the dose rates off the top would be

29.5, 4.42 and 1.41 mrem/hr on the balsa wood surface, 6 feet off the

barrier, and behind the vehicle cab respectively. Off of the bottom the

dose rates are 37.5 and 4.0 mremi/hr on the balsa wood surface and 6

feet off the barrier respectively.

The results of the secondary gamma dose rate calculations yield end dose

rates of 0.37 and 0.033 mremAhr on the balsa wood surface and 6 feet

off the barrier respectively. These dose rates apply to both the top and

bottom of the cask, since both have similar shield configurations.

Ground scattering dose rates were not calculated off of the ends. However

an upper limit estimate made considering the results of the side

calculations yields the results given in Table DC-10.

The total dose rate six feet off the personnel barrier is seen to bd 7.65

and 6.73 mrem/hr off the top and bottom respectively.

The gamma and neutron dose rates on the cask surface at the dose points defined in Figure DC-3 are given in Table IX-12. The neutron dose rates

were normalized to the radial and axial DTF-IV result.. It should be

noted that these results include all applicable axial gamma and neutron correction factors for buildup, streaming, and corner peaking. Secondary

gamma dose rates have not been included since they are small and would

decrease approximately as the neutron dose decreases. It should further

be noted that these dose rates do not account for any shielding offered

by the balsa wood. 'The results show that none of the dose rates exceed

the assembly axial or transverse centerline dose rates. !.

IK-29

TABLE IX-12

END AND SIDE DOSE RATES NORMAL CONDITIONS OF TRANSPORT

Gamma Dose Rate** Neutron Dose Rate*** Dose Point* mrem/hr mrem/hr

T-1 34.0 154.0

T-2 33.2 152.2

T-3 31.1 147.9

T-4 27.9 141.2

T-S 21.0 115.9

T-6 13.8 80.0

T-7 12.4 37.0

T-8 .3.7 5.9

T-9 1.4 1.2

T-10 2.0 1.2

T-11 0.3 0.3

T-12 0.1 0.1

S-1 19.3 19.7

S-2 1.7 S.0

3-3 0.S 8.3

S-4 0.3 4.7

S-5 0.7 2.3

3-6 1.0 2.6

S-7 2.0 4.0

S-8 0.6 1.6

* See Figure DC-3 ** Gamma dose rate include appropriate material buildup correction

factors (1.57 for side and 1. 45 for top) and side comer peaking correction factor of 1. 12.

** Neutron side dose rate includes an approgriate corner peaking

correction factor of 1.09 and a top streaming correction factor of 2.0. IX-30

4.2 Hypothetical Accident Conditions

The results of the shielding analysis for post-hypothetical accident conditions are summarized in Table IX-13. The primary gamma dose rates

-off the side of the cask are 985 mrem/hr on the surface and 261 mrem/hr three feet off of the surface. These values should be increased by 10% to account for axial peaking. The buildup correction factor applicable to the case of a uranium and steel composite compared to the equivalent uranium value is 1.30. The corner correction factor for the accident case has been calculated to be approximately 1.01. The net effect is maximum primary gamma dose rates of 1423 and 377 mremAr on the surface and three feet from the surface respectively.

The neutron dose rates increase significantly because of the assumed loss of all water. The DTF-IV results are 785 and 204 mrem/hr at the surface and three feet from the surface respectively. Each of these are conservatively increased by 25% due to axial peaking of the neutron source, and 1% due to assembly corner peaking.

Based on a comparison of the surface neutron fluxes calculated for the accident case with those calculated for the normal case the secondary gamma dose rate is estimated as 15 mrem/hr on the surface and 4.1 mrem/hr at three feet. The ground scatter dose rates are take: to be the same relative fraction of the direct doses as was found for the normal case.

Adding all the components yield a maximum dose rate 3 feet off of the surface of 787 mrem/hr. The detailed derivation of this number is summarized in Table IX-1 4.

K-J

IX-31

TABLE DC-I 3

NLI 1/2 CASK MAXIMUM DOSE RATES (Hypothetical Accident Conditions)

_ _ _ _ __,_ _ mrem/hr Fuel M1dplane Off End on Axi 1 Centerline

Source I To_ Bottom

Category 3' from 3' from 3' from

Surface Cask Surface Cask Surface Cask

Primary Gamma 1423 377 34 7.6 40 7.0

Neutron 991 258 154 36 224 41

Secondary Gamma 15 4.1 29 6.4 32 5.6

Ground Scatter Gamma --- 19 0.5 --- 0.5

Ground Scatter Neutron --- 129 17 -- 21

Total Gamma 1438 400 63 14.5 72 13.1

Total Neutron 991 387 154 53 224 62

Total 2429 787 217 67.5 296 75.1

(.

S

0

TABLE IX-14 I

DETAILED DERIVATION* OF DOSE RATE FUEL MIDPLANE 3 FEET FROM SURFACE

(Hypothetical Accident Conditions)

Base Value or Percent Dose Rate

Correction mrem/hr

Primary Gamma

Base Calculated Value 261

Buildup Factor Correction 30%

Axial Peaking Correction 10%

Corner Peaking Correction 1%

Total 44% 377

Primary Neutron

Base Calculated Value 204

Axial Peaking Correction 25%

Corner Peaking Correction 1%

0 Total 26% 258

Secondary Gamma 4.1

Ground Scatter Gamma

Direct Gamma Dose 381 Backscatter 5% 19

Ground Scatter Neutron

Direct Neutron Dose 2S8 Backscatter 50% 129

Total 787

*Total correction is product of individual factors.

IX-33

The primary gamma and neutron doses off of the ends are given in Table IX-13

and are very similar to the normal shipping conditions dose rates, except that

the accident dose rates were obtained at three feet off the surface with the

balsa wood removed from the cask.

The maximum dose rate three feet from the cask surface for the design basis

PWR assembly is seen to be at the midplace of the fuel. This result is less

than the 10 CFR 71 limit of 1000 mrem/hr.

4.3 Shielding Analysis for Consolidated PWR Type Fuel Rods

Analyses of the effects of the shipment of consolidated fuel in the NLI 1/2

cask have been performed for criticality, shielding, structural, and thermal

effects. The consoldated fuel modeled in the criticality, shielding, and

structural analyses is 1415 x 15 fuel cooled for two years, with an initial

enrichment of 3.7 w/o U-235 and a burnup of 40,000 MWD/MTU. These values are

considered to be representative of consolidated PWR spent-fuel shipments. The

thermal analysis has been performed (see Section VIII) for W14 x 14 fuel

cooled 12 years to add an additional margin of conservatism because the

thermal behavior of consolidated spent fuel is currently being investigated.

The gamma and neutron sources generated by a PWR assembly (150 days cooled)

and a conslidated fuel canister (730 days cooled) are given in Table IX-15.

This Table compares SAR values obtained for the design basis PWR assembly from

Section III and values calculated from the LOR-2 version of ORIGEN-2 used at

Babcock and Wilcox in Lynchburg, Virginia. The gamma source for the canister

is 69 percent of the source for an assembly, and the neutron source is 140

percent of the source for an assembly. Note that the consolidated fuel

canister, because of its longer cool time, contains a smaller quantity of

radionuclides than the design basis PWR assembly (cooled 150 days). The

resulting dose rates may be obtained for canisters by multiplying the assembly

dose rates contained in the NLI 1/2 SAR by the appropriate percentages. The

results of these calculations are shown in Table IX-16.

IX-34

C

TABLE IX-15

PWR ASSEMBLY AND CONSOLIDATED FUEL SOURCE STRENGTHS

Decay Time, days

Thermal Output, Kw

Gamma, MeV/sec

Neutrons, n/sec

Fission Products, Ci Total

0-1

La) U'

SAR Design Basis PWR Assembly#

150

10.63

3.074 x 1016

7.55 x 108

PWR Source Strengths from LOR-2 Design Basis Design Basis Canister of Rods

PWR Assembly* PWR Assembly* of 2 PWR Assemblies

150 730 730

10.2 3.3 6.6

2.96 x 1016 1.03 x 1016 2.05 x 1016

6.23 x 108 4.37 x 108 8.74 x 108

2.17 x 106 6.63'x 105 1.33 x 106

65

69

140

61

*W15 x 15 assembly; 475 KgU; 3.7 w/o; 40,000 MWD/MTU

% Source of Canister at 730 days Source of Assembly at 150 days

(C

TABLE IX-16

CONSOLIDATED FUEL BOSE RATES

Normal Operation: (6 feet from Personnel Shield)* As-sembly dose rates in parentheses)

Source

Primary Gamma Neutron Secondary Gamma Ground Scatter Gamma Ground Scatter Neutron Total Gamma Total Neutron

Total

Fuel Mldplane

1.93 (2.8) 4.68 (3.34)

0.42 (0.61) 0.12 (0.17)

2.34 (1.67)

2.47 (3.58)

7.01 (5.01)

9.48 (8.59)

Top

1.24 (1.79) 6.19 (4.42)

0.02 (0.03)

0.08 (0.12)

1.81 (1.29)

1.34 (1.94)

7.99 (5.71)

9.33 (7.65)

Bottom

0.89 (1.29) 5.60 (4.00)

0.02 (0.03)

0.08 (0.12)

1.81 (1.29) 0.99 (1.44)

7.41 (5.29)

8.40 (6.73)

Hypothetical Accident: (3 feet from Cask Surface)*

Source

Primary Gamma Neutron Secondary Gamma Ground Scatter Gamma Ground Scatter Neutron Total Gamma Total Neution

Total

Fuel Mtdplane

260 (377)

361 (258) 2.8 (4.1) 13 (19) 181 (129)

276 (400)

542 (387)

818 (787)

Top

5.2 (7.6)

50.4 (36) 4.4 (6.4)

0.3 (0.5)

24 (17)

10.0 (14.5) 74 (53)

84 (67.5)

Bottom

4.8 (7.0)

57.4 (41)

3.9 (5.6)

0.3 (0.5)

29.4 (21)

9.0 (13.1)

87 (62)

96 (75.1)

*Based on Table IX-lO, Page IX-26

I X-36

The maximum calculated Normal Operation dose rate of 9.5 trem/hour is less than the limit of 10 mrem/hour at six feet, and the maximum Hypothetical Accident dose rate of 818 mrem/hour at three feet is less than the limit of 1000 inrem/hour. Thus, the gauma and neutron sources for consolidated fuel are acceptable.

4.4 Shielding Analysis for Metallic Fuel Analysis of the neutron and gamnm radiation dose rates is performed for metallic fuel by calculating the ratio of source strength for metallic fuel to source strength for intact PWR fuel and applying this ratio to the dose rates from PWR fuel. The neutron shield tank is drained for these shipments so the dose rates are calculated from the hypothetical accident -- loss of neutron shielding -- analysis presented in section 4.2. The dose rate ratios are computed as follows:

PWR Gamma Source: 2.25 x 1016 HEV/sec Metallic Fuel Gamma Source: 1.253 x 1015 4EV/sec

Gamma Ratio 0.056

PWR Neutron Source: 4.609 x 108 n/sac Metallic Fuel Neutron Source: 2.289 x 105 n/sec

Neutron Source a 4.97 x 10-4

The neutron ratio is also used for secondary gammas because secondary gammas are produced by interactions of neutrons with shielding material.

4.4.1 Fuel Midplane Dose Rates: The PWR loss of neutron shield analysis calculates dose rates 3

feet from the cask surface. Normal operation dose rates were calculated at 6 feet from the cask package by applying a l/r dose rate falloff relationship. The cask has a 47.125 inch outside diameter so the distance from the centerline of the cask to the 3 foot dose point is 59.6 inches. The cask may be located inside the personnel barrier on a truck trailer or inside a seagoing container. For conservatism, these packages are ignored and the cask surface is used as the package boundary for these calculations, so the 6 foot dose point is 95.60 from the cask centerline. The geometric falloff from the 3 foot distance dose point for hypothetical accident to the 6 foot normal operation location is thus 59.6/95.6 - 0.623.

Revised IX-37 Oct. 1986

Feb. 1987

The source ratios for gammas and neutrons and the geometric falloff (for hypothetical accident to normal operation) are applied to the fuel midplane dose rates for intact PWR fuel shown in Table IX-13, page IX-32. The primary gamma dose rates used in the normal operation calculation is not the value given in Table IX-13 because this assumes that the lead gamma shield is melted, which is not the case in normal operation. Instead, the primary gamma dose rate at the cask surface, 21 mrem/hr., is adjusted to the desired 6 foot location by applying the 1/r relationship from 23.6 inches to 95.6 inches, a factor of 0.25. Thus, the dose rate for normal operation with intact PWR fuel at the desired dose point is 21 mrem/hr x 0.25 = 5.25 mrem/hr. This is larger than the normal operation dose rate from primary gammas listed in Table IX-1O (2.8 mrem/hr) because of the difference in definitions of the package boundary. The metallic fuel primary gamma dose rate for normal operation is thus 0.056 x 5.25 - 0.294 mrem/hr. The groundscatter gamma is 5% of the primary gamma dose rate as specified on page IX-27, or 0.05 x (0.294 + 0.001) = 0.015.

Table IX-17 Fuel Midplane

Metallic Fuel Dose Rates (mrem/hr)

Hypothetical Accident Normal Operation Intact PWR Metallic Metallic

Hypothetical Acc. Fuel Fuel (3' from Cask) Ratio (3' from Cask) (6' from Package)

Primary Gamma 377 0.056 21.11 0.294 Neutron 258 4.97x10"4 0.128 0.080 Secondary Gamma 4.1 4.97x10-4 0.002 0.001 Ground Scatter

Gamma 19. 0.056 1.064 0.015 Ground Scatter

Neutron 129. 4.97x10" 4 0.064 0.040 Total Gamma 400. 22.18 0.310 Total Neutron 387 0.192 0.120

TOTAL 787. 22.372 0.430

4.4.2 Top and Bottom Dose Rates The normal operation dose rates are calculated by applying the

metallic fuel/PWR fuel source ratios directly to the dose rates given in Table IX-1O (Normal Operation) and Table IX-13 (Hypothetical Accident). The results

Page Added Oct. 1986 IX-37a Revised Feb. 1987

are listed in Table IX-18 (Normal Operation) and IX-19 (Hypothetical Accident).

4.4.3 The results of these calculations are summarized in Table IX-20. Inspection of these results shows that dose rates from metallic fuel cooled two years is at most 2-1/2 percent of the dose rate from one PWR assembly cooled 150 days. This shows that transportation of metallic fuel is adequately shielded under all conditions.

Table IX-18 Top and Bottom

Metallic Fuel Dose Rates (mrem/hr) Normal Operation

Primary Gamma Neutron Secondary Gamma Ground Scatter Gamma Ground Scatter Neutron

Total Gamma Total Neutron

TOTAL

TOP PWR Meta11ic

1.79 0.100 4.42 0.002 0.033 0.000 0.12 0.007 1.29 0.001 1.94 0.107 5.71 0.003

7.65 0.110

TOP PWR Metal lic

1.29 0.072 4.00 0.002 0.033 0.000 0.12 0.007 1.29 0.001 1.44 0.079 5.29 0.003

6.73 0.082

Table IX-19 Top and Bottom

Metallic Fuel Dose Rates (mrem/hr) Hypothetical Accident

Primary Gamma Neutron Secondary Gamma Ground Scatter Gamma Ground Scatter Neutron

Total Gamma Total Neutron

TOTAL

TOP PWR Metal lic

7.6 0.426 36. 0.018 6.4 0.003 0.5 0.028 17. 0.008 14.5 0.457 53. 0.026

67.5 0.483

TOP PWR Meta51 ic

7.0 0.392 41. 0.020 5.6 0.003 0.5 0.028 21. 0.010 13.1 0.423 62. 0.030

75.1 0.453

I

Page Added Oct. 1986 Revised Feb. 1987

I I II

4

IX-37b

Table IX-20

Metallic Fuel Dose Rates Totals - mrem/hr

Fuel Midplane

Top

Bottom

Normal Operation

0.430 (5.0%)

0.110 (1.4%)

0.082 (1.2%)

Hypothetical Accident

22.372 (2.81)

0.483 (0.7%)

0.453 (0.6%)

Percents of intact PWR dose rates are shown in parentheses.

Page Added Oct. 1986 Revised IX-37c Feb. 1987

4.5 Shieldiny Analysis for PWR or BWR Rods

The gamma and neutron sources generated by a PWR assembly and by 25 PWR rods or 25 BWR

rods are given in Table IX-21. The gamma source for PWR rods is 40 percent of the high burnup

PWR assembly gamma source. The neutron source is 60 percent of the high burnup PWR

assembly neutron source. The dose rates for the rod shipments may be obtained by multiplying

the high burnup assembly dose rates by the appropriate percentages. The results of these

calculations are shown Table IX-22. Similar calculations comparing the BWR rods to the design

basis BWR assembly are shown in Table IX-23. In the case of the content condition of 18 PWR

fuel rods with specific power of 60 kW/kgU and a cooling time of 300 days, the dose rates will

be significantly below the dose rates calculated for the 25 PWR fuel rod content condition based

on the significantly smaller source terms as listed in Table IX-21.

The results given in Table IX-22 show that the dose rates from 25 PWR rods are much less than

the dose rates from the high burnup PWR assembly. Similarly, the dose rates given in Table

IX -23 for 25 BWR rods are much less than the design basis assembly. The shipment of up to 25

PWR or 25 BWR rods at 150 days cool time is, therefore, safe even though the burnups are

greater than the design basis PWR and BWR assembly burnups.

Page Added May 1987

Revised Jan. 1990 Feb. 1991 Oct. 1991 Apr. 1992

IX-37d Feb. 1996

-N 6 &- - -

Table IX-21

PWR ASSEMBLY AND PWR OR BWR ROD SOURCE STRENGTHS

Decay Time, days

SAR High Burnup 15RA0m* 150

Thermal Output, kW 10.63

Gamma, MeV/sec 1.00 x 101,

Neutrons, n/sec 1.86 x 10' Fission Products, 2.31 x I o" Ci Total

150 *

1.65 (16%)

4.04 x 1015 (40.4%)

1.12 x 109(60.3%)

2.84 x 105(12.3%)(5)

18 WR od(2)

300 '

0.9(8.5%) 1.94 x 101S (19.40/)

7.56 x 107(4%)

1.69 x 105 (7%)

150

4.0(38%)

7.54 x 10" (75%)

3.5 x 108 (19%)

6.03 xl0' (26%)P')

SAR Design Basis

150

10.63

3.074 x 1016

7.55 x 10'

2.17 x 106(6)

Source of Rods at 150 days Source of Assembly at 150 days

Notes: ( Calculated with the LOR2 version of ORIGEN2 which yields conservative results. Values in parentheses represent percentage as

compared to SAR High Burnup PWR Assembly. Twenty three rods at 60,000 MWD/MTU and two rods at 65,000 MWD/MTU were considered.

(2) Calculated with SAS2(H).Values in parentheses represent percentage as compared to SAR High Burnup PWR Assembly. (3) Calculated with the SAS2 sequence of the SCALE4 code package (NUREG/CR-200). Values in parentheses represent percentage

as compared to SAR High Burnup PWR Assembly. Twenty-five rods at 75,000 MWD/MTU were considered. C Calculated with the LOR2 version of ORIGEN2.

(5) Total Fission Product Curies for rods were not listed in the Final Report data tables, but were obtained directly from the computer results.

(6) From Table IX-15.

Q C C

TABLE IX-22

25 PWR ROD DOSE RATES

Normal Operation: (6 feet from Personnel Shield)*

(High Burnup PWR Assembly dose rates in parentheses)

Source Fuel Midplane ToP -- Bottom

Primary Gamma 0.36 (0.91) 0.23 (0.58) 0.17 (0.42)

Neutron 2.09 (3.47) 3.36 (5.57) 3.06 (5.07)

Secondary Gamma 0.12 (0.20) 0.01 (0.01) 0.01 (0.01)

Ground Scatter Gamma 0.02 (0.06)' 0.02 (0.04) 0.02 (0.04)

Ground Scatter Neutron 1.05 (1.73) 0.99 (1.63) 0.99 (1.63)

Total Gamma 0.50 (1.17) 0.26 (0.63) 0.20 (0.47)

Total Neutron 3.14 (5,20) 4.35 (7.20) 105 (6.70)L

TOTAL 3.64 (6.37) 4.61 (7.83) 4.25 (7.17)

Hypothetical Accident: (3 feet from cask surface)*

Source Fuel Midplane Top Bottom

Primary Gamma 49.7 (123. ) 1.0 ( 2.5) 0.93 ( 2.3)

Neutron 129.0 (213. ) 26.8 (44.4) 30.4 (50.5)

Secondary Gamma 0.8 ( 1.3) 1.2 (2.1) 1.0 (1.8)

Ground Scatter Gamma 2.5 (6.2) 0.04 (0.1) 0.08 (0.2)

Ground Scatter Neutron 64.5 (107. ) 12.5 (20.9) 15.6 (25.9)

Total Gamma 53.0 (131. ) 2.24 ( 4.7) 2.0 ( 4.3)

Total Neutron 193.5 (320. 39.3 (65.3) 46.0 (76.4)

TOTAL 246.5 (451.) 41.5 (70.0) 48.0 (80.7)

Based on Table IX-10, Page IX-26 Page Added

May 1987 Revised Jan. 1990

IX-37f Feb. 1991

TABLE IX-23

25 BWR ROD DOSE RATES 150 Days Decay Time

Normal Operation: (6 feet from Personnel Shield)*

(High Burnup PWR Assembly dose rates in parentheses)

Source Fuel Midplane TOp Bottom

Primary Gamma 0.68 (0.91) 0.44 (0.58) 0.32 (0.42)

Neutron 0.66 (3.47) 1.06 (5.57) 0.96 (5.07)

Secondary Gamma 0.04 (0.20) 0.00 (0.01) 0.00 (0.01)

Ground Scatter Gamma 0.05 (0.06) 0.03 (0.04) 0.03 (0.04)

Ground Scatter Neutron 0.33 (1.73) 0.31 (1.63) 0.31 (1.63)

Total Gamma 0.77 (1.17) 0.47 (0.63) 0.35 (0.47)

Total Neutron 0.99 (5.20) 1.37 (7.20) 1.27 (6.70)

TOTAL 1.76 (6.37) 1.84 (7.83) 1.62 (7.17)

Hypothetical Accident: (3 feet from cask surface)*

Source Fuel Midplane Too Bottom

Primary Gamma 92.25 (123. ) 1.88 ( 2.5) 1.73 ( 2.3)

Neutron 40.47 (213. ) 8.44 (44.4) 9.60 (50.5)

Secondary Gamma 0.25 ( 1.3) 0.40 (2.1) 0.34 (1.8)

Ground Scatter Gamma 4.65 ( 6.2) 0.08 (0.1) 0.15 (0.2) Ground Scatter Neutron 20.33 (107.) 3.97 (20.9) 4.92 (25.9)

Total Gamma 97.15 (131.) 2.36 ( 4.7) 2.22 ( 4.3)

Total Neutron 60.80 (320.) 12.41 (65.3) 14.52 .4)

TOTAL 157.95 (451.) 14.77 (70.0) 16.74 (80.7)

Based on Table IX-lO, Page IX-26

Page Added May 1987 .

Revised Oct. 1991

IX-37g Apr. 1992v

4.6 Shielding Analysis for the MARK 41 Fuel Assembly This analysis is performed to allow the shipment of a MARK 42 fuel

assembly. The analysis is based upon gamma sources from ORIGEN-2 with the CANDU library and neutron sources based upon calculations made by Savannah River Plant personnel. A decay time of 1245 days is used in the analysis, although the actual decay time will be somewhat longer. The gamma and neutron sources, generated by the design basis PWR assembly, the high burnup assembly in Section IX Appendix B, and the MARK 42 assembly, are given in Table IX-24.

Table IX-24 PWR FUEL AND MARK 42 FUEL SOURCE sTR 1s

SAR Design Basis High Burnup MARK 42

Decay Time, days 150 450 1245 Gamma, MeV/sec 2.96 x 1016 1.0 x 1016 5.8 x 1014 Neutrons, n/sec 6.23 x 108 1.86 x 109 1.2 x 109

As can be seen from Table IX-24, the gamma source from the MARK 42 assembly is 2.O of the design basis assembly gamma source and the neutron source is 193% of the design basis PWR assembly neutron source. The MARK 42 neutron source is, however, only 65% of the neutron source of the high burnup assembly. The dose rates at two meters resulting from the NLI-1/2 containing one MARK 42 assembly are 3.11 mrem/hr at the cask midplane, 5.58 mrem/hr at the top and 5.16 mrem/hr at the bottom. These dose rates are below the requirements of 10 CFR 71.

4.6.1 Gamma Dose Calculation The gamma dose rates at the fuel radial midplane were obtained using

the XSDRNPM computer code. The XSDRNPM computer code solves the Boltzman transport equation for one-dimensional geometries. The XSDRNPM radial gamma results have been corrected for axial peaking by applying a 1.10 axial peaking factor. A backscatter factor of 1.05 for the gammas as described on page IX-27 was also app ied. The gamma dose rates at the top and bottom of the cask were obtained by applying the 2.0% source ratio to the design basis PWR assembly doses. The resulting dose rates can be found in Table IX-25.

IX- 37h Page added August 1988

4.6.2 Neutron Dose Calculation The neutron dose rates at the fuel radial midplane were obtained using

the XSDRNPH computer code that is accurate for neutrons which comprises the primary radiation source for the MARK 42 assembly. The XSDRNPM radial neutron results have been corrected for axial peaking by applying a 1.10 axial peaking factor. A backscatter factor of 1.5 for the neutrons as described on page IX-27 was also applied. The neutron dose rates at the top and bottom of the cask were obtained using the same method as described in Section IX Appendix B. This method corrects for the axial peaking at the ends of the cask. A 501 decrease of the design basis NI assembly dose rates was used at the cask ends in this analysis. The dose was then adjusted for the MARK 42 assembly by a factor of 1.93 (the ratio of the MARK 42 neutron source term to the design basis PN assembly neutron source term). The resulting dose rates can be found in Table IX-25.

Page added IX-37i August 1988

Table IX-25

HARK 42 FUEL ASSEMBLY DOSE RATES. (mjm/hr)

Normal Oneration: (2 meters from Personnel Shield)*(PWR assembly dose rates in parentheses)

Primary Gamma 0.02 (2.80) Neutron 1.75 (3.34) Secondary Gamma 0.46 (0.61) Ground Scatter Gamma 0.00 (0.17) Ground Scatter Neutron 0.88 (1.67) Total Gamma 0.48 (3.58) Total Neutron L . TOTAL 3.11 (8.59).

Hypothetical Accident: (1 meter from Cask (PWR assembly dose rates in parentheses)

SourceFuel Midip1ane Primary Gamma 0.01 (377.) Neutron 176. (258.) Secondary Gamma 0.22 (4.1) Ground Scatter Gamma 0.00 (19.) Ground Scatter Neutron 87.8 (129.) Total Gamma 0.23 (400.) Total Neutron W.., (387.) TOTAL 264. (787.)

0.04

4.27

0.03

0.00

1.24

0.07

5.58

(1.79)

(4.42)

(0.03)

(0.12)

(1.29)

(1.94)

(7.65)

0.03 (1.29) 3.86 (4.00)

0.03 (0.03)

0.00 (0.12) 1.24 (1.29)

0.06 (1.44)

5.16 (6.73)

Surface)*

Bottom0.15 34.7

6.18

0.01

16.4

6.34

57.4

( 7.6)

( 36.)

( 6.4)

( 0.5)

( 17.)

(14.5)

(67.5)

0.14

39.6

5.40

0.01

20.3

5.54

65.4

( 7.)

(41.) ( 5.6)

(0.5) (21.)

(13.1)

(75.1)

*Based on Table IX-l0, Page IX-26

Page added August 1988

IX-37J

Shielding Analysis for the MARK 22 Fuel Assembly

This analysis is performed to allov the shipment of two Mark 22 fuel assemblies. The analysis is based upon gamma sources from the LOR-2 version of ORIGEN-2 available from Babcock and Wilcox. A decay time of 150 days is used in the analysis, although the actual decay time will be somewhat longer. The gamma and neutron sources, generated by the Mark 42 assembly and the Hark 22 assemblies, are-given in Table IX-26.

Table IX-26 HARK 22 AND KARK 42 FUEL SOURCE S'hIMGTHS

Two Hark 22 MARK 42

Assemblie Fuel Assembly Decay Time, days 150 1245 Gamma, MeV/sec 9.753 x 1015 5.8 x 1014

Neutrons, n/sec 7.872 x 105 1.2 x 109

As can be seen from Table IX-26, the gamma source from the Hark 22 assembly is 1682 percent of the Hark 42 assembly gamma source and the neutron source is 0.065 percent of the-Hark 42 assembly neutron source. The dose rates at 2. meters resulting from the NLI-l/2 cask containing two Mark 22 assemblies are 0.18 mrem/hour at the cask midplane, 0.70,mrem/hour at the top and 0.53 mrem/hour at the bottom. These dose rates are below the requirements of 10 CFR 71.

All dose rates were obtained by applying the previously mentioned source ratios to the Mark 42 dose rates. A factor of 16.82 was applied to the primary gamma dose rate and a factor of 6.5 x 10'4 was applied to the neutron, secondary gamma and ground scatter neutron dose rates. The ground scatter gamma dose rate was taken to be 5 percent of the primary gamma dose rate. The resulting doses can be found in Table IX-27.

Page Added February 1990 IX-37k

4.7

Table IX-27 ffRK 22 FUEL ASSEMBLY DOSE RATES (ures/hr)

Normal Ooeration: (2 (Mark 42 assembly dose

Source

Primary Gamma

Neutron

Secondary Gamma

Ground Scatter Gamma

Ground Scatter Neutron Total Gamma

Total Neutron

TOTAL

Hvyothetical Accident:

meters from Personnel Shield)* rates in parentheses)

0.17 (0.01)

0.00 (1.75)

0.00 (0.46)

0.01 (0.00) 0.00 (0.88)

0.18 (0.48)

9oo t2.3 0.18 (3.11)

0.67

0.00

0.00

0.03

0.00

0.70

0.70 0.70

(0.04)

(4.27)

(0.03)

(0.00)

(1.24) (0.07)

(5.58)

(1 meter from Cask Surface)*

(Mark 42 assembly dose rates in parentheses)

Source Primary Gamma

Neutron

Secondary Gamma

Ground Scatter Gamma Ground Scatter Neutron Total Gamma

Total Neutron

TOTAL

Fuel Mid~~lane

0.34 (0.02)

0.11 (176.)

0.00 (0.22)

0.02 (0.00)

0.06 (87.8)

0.36 (0.23)

2,U1 -26. 0.53 (264.)

*Based on Table IX-25, Page IX-3j

IX-371Page added February 1990

0.50

0.00

0.00

0.03

0.00

0.53

0.53

(0.03)

(3.86)

(0.03)

(0.00)

(1.24)

(0.06)

(5.16)

2.52

0.02

0.00

0.13

0.01

2.65

9..U 2.68

(0.15)

(34.7)

(6.18)

(0.01)

(16.4)

(6.34)

(57.4)

2.35

0.03

0.00

0.12

0.01

2.47

D.0a 2.51

(0.14)

(39.6)

(5.40)

(0.01)

(20.3)

(5.54)

(65.4)

5.0 REFERENCES

1. Federal Register, Department of Transportation, Hazardous Materials Regulations Board, Radioactive Materials and other Miscellaneous Amendments, Volume 33, Number 194, Washington, D. C., Part II, Title 49, Transportation, Friday, October 4, 1968.

2. United States Atomic Energy Commission, Rules and Regulations Title 10 - Atomic Energy, Part 71, "Packaging of Radioactive Material for Transport", December 31, 1968.

3. Book of Standards, Part 7, ASTM, 1970.

4. Metals Handbook, 8th Editon, American Society of Metals.

5. Alcoa Aluminum Handbook Aluminum Company of America, Pittsburgh, Pa., 1959.

6. Merritt, F. S.. Standard Handbook for Civil Engineers, McGraw-Hill Book Company, New York, 1968.

7. Malenfant, R., "QAD - A Series of Point-Kernel General Purpose Shielding Programs", LA-3573, April 1967.

8. "Design and Analysis Report", IF 300 Shipping Cask, General Electric.

9. Shappert, L. B., "Cask Designer's Manual", ORNL-RSIC-68, February 1970.

10. Hubbell, I., "Photon Cross Sections, Attenuation Coefficients, and Energy Absorption Coefficients from 10 KeV to 100 Gev", NSRDS-NBS29, U.S. Department of Commerce, August 1969.

11. Chapman, G. T.,Storrs, C. L., "Effective Neutron Removal

Cross Sections for Shielding", AECD 3978, September 1955.

12. Capo,'M.A., APEX-510, 1958.

13. Jaeger, R. G., et al., "Engineering Compendium on Radiation Shielding, Shielding Fundamentals and Methods", Volume I, Springer-Verlag, Inc., New York, 1968.

(

"IX-38

14. Rockwell, Theodore (III), Editor, "Reactor Shielding Design Manual", D. Van Nostrand Company, Inc., New York, 1956.

15. USAEC, Rules and Regulations, Title 10 - Atomic Energy, Part 20, "Standards for Protection Against Radiation", December 31, 1970.

16. Lathrop, K. D., "DTF-IV - A Fortran Program for Solving the Multigroup Transport Equation with Anisotropic Scatted ng-, LA-3373, 1965.

17. Engle, W. W. (14, Mynott, F. R., "ANISN, A One-Dimensional Discrete Ordinates Transport Code with Anisotropic Scattering, (to be published).

18. 'Shielding Benchmark Problems", ORNL-RSIC-2S, A. E. Profio, Editor, August 1970.

19. Trubey, D. K., "Kernal Methods for Radiation Shielding Calculations, Nuclear Engineering and Design, Radiation Shielding (2) ", Vol. 13, (1970) Vol. 3.

20. NAA-SR-11980, Vol. IM. 0

21. "Protection Against Neutron Radiation Up to 30 Million Electron Volts", Handbook 63, U. S. Department of Commercb, NBS, 1957.

22. Proceedings, Third International Symposium, "Packaging and Transportation of Radioactive Materials", CONF-710801, Richmond, Washington, August 16-20, 1971.

23. CSCROS Users Manual, Draft, Computer Sciences Corp. N.Y.

24. Joanou, G. D., Dudek, J. S., "GAM-II, A B3 Code for the Calculation of Fast Neutron S;-.ctra and Associated Multigroup Constants", GA-4265, September 1963.

25. Honeck, H., "THERMOS, A Thermalization Transport Theory Code for Reactor Lattice Calculations", BNL-5826, 1961.

26. Fisher, E., "The Streaming of Neutrons in Shields, Nuclear Science and Engineering, 1, p.222-238 (1956).

27. "ANISN, A One-Dimensional Discrete Ordinates Transport Code, RSIC Computer Code Collection", ORNL-CCC-82.

IX-39

This page intentionally left blank.

SECTION IX

APPENDIX A

An analysis has been made to determine the change in the neutron and gamma

flux resulting from removing the Inner container, closure head redesign, and

repositioning the fuel in the cask. The analysis was based on the same neutron

and gamma source strengths that were used in the original SAR except that only

two energy groups ( 1.0 and 2.0 mev) were used for the gamma calculations

and three energy groups ( 1.15, 2.2 and 4 mev) for the neutron calculations.

Appropriate correction factors were incorporated in order to bring the method

of calculation for the alternate configuration in agreement with the original

computer analysis.

IX-A1

APPENDIX A

The following changes have been made in shielding thicknesses and fuel

position for the alternate shipping configuration.

Table 1

I. Shield Thickness

1) Closure Heads a) Total thickness of

Stainless Steel b) Thickness of Uranium

2) Bottom of Cask a) Total thickness of

Stainless Steel b) Thickness of Uranium

MI. Fuel Position

1) Distance from active fuel to top surface of the cask

2) Distance from active fuel to bottom surface of the cask

These changes will produce the following

longitudinal centerline of the cask.

Original Configuration

3.15 in.

3.00 in.

3.13 in.

3.13 in.

20 in.

28.4 in.

Alternate Configuration

3.5 in.

4.0 in.

3.22 In.

3.15 in.

28 in.

21.4 in.

surface radiation values at the

ID-A2

Table 2

Surface Dose Rates

Neutron Total (mrem/hr) (mrem/r) (mrem/hr)

Original Head 34 154 188

Alternate Head 0.74 85.8 86.54

Original Bottom 13.0 104.6 117.6

Alternate Bottom 20.0 147.0 167.0

The radiation level through the bottom of the alternate configuration, Lncluding

the 1/2" stainless steel Fuel Support Stand, would be 167.0.

There will also be a slight Increase In the side dose rate due to the removal

of the •" thick Inner container. The maximum side dose rate for the original

configuration was only 39 mrem/hr. therefore, the removal of the inner con

tainer which provided I" thick stainless steel shielding in this area will not

result in excessive dose rates.

IX-A3

The dose level shown in Table-2 were calculated as follows

(except for the original head values which were calculated by computer for the K

original configuration SAR):

The square fuel cross section was converted to an equivalent circular

cross section and the flux from each Incremental fuel segment to a point detector

at the surface of the cask was calculated. The flux from all these fuel segments

was then summed up to determine the total flux at the surface of the cask for

both neutron and gamma sources as follows

S• detector

a) Gamma

D =Sv (E) CR (E)

.K=l,(

s0

Z ! BI(E) e CF u XX 2 R1I T-1

dR, ) dZK

IX-A4

uX = (ult.1 + U2 t2 + Uf tf) R/z

u = linear absorption coefficient

Uf = linear absorption coefficient of the fuel

t = shield thickness

tf = shield thickness of fuel for source segment considered

R = distance from source segment to detector point on the cask

surface

CF correction factor, adjusts attenuation to agree with computer

code results used in original analysis

Sv = volumetric source strength

CR = gamma flux-to-dose-rate conversion factbr

*rn'v r

tf

"200 50

Z 2Sv (E) CR (E) '7( CF e- ~Xp dRI T 2 R11K

X = ( 1 tl +Z 2 t 2 + 2f tf) R/Z

S= macroscopic cross section

f = macroscopic cross section of the fuel

t = shield thickness

shield thickness provided by the fuel for the source

segment considered

R = distance from the source segment to the detector point on

the cask surface

ZK

DC-AS

b) Nei

CF = correction factor, adJusts calculations to agree

with computer code results obtained from original

analysis

Sv = volumetric source strength

CR neutron flux-to-dose-rate conversion factor

In summary, it can be seen that the dose rate at the surface of the outer

closure head of the alternate configuration will be less than for the original

configuration. The dose rate will Increase on the bottom of the alternate

configuration, including the 1/2" stainless steel Fuel Support Stand, to a

level of 167 MREM/HR.

IX-A6

REFERENCES:

"Principles of Nuclear Reactor Engineering" - Glisstone June 1957.

IX-A7

This page intentionally left blank.

SECTION IX - APPENDIX B SUPPLEMENTAL ANALYSIS FOR HIGH BURN-UP PWR FUEL

This appendix presents supplemental analyses of gamma and neutron dose rates for high burn-up PWR fuel assemblies that meet the specific conditions of Appendix A, Section III.

1.0 Summary

This analysis was performed In order to allow shipment of a specific high burn-up PWR assembly. The analysis is based on the results of an ORIGEN calculation that describes an assembly with burn-up of 58,600 MWD/MTU with a cool time of 450 days. The results of the analysis are generalized to allow shipment of other assemblies whose parameters are within bounds of the specific assembly as given in Appendix A, Section 111. This resulted in a fuel assembly gamma source one-third (1/3) the cask design gamma source strength and a corresponding neutron source of approximately two and one-half (2-.5k times the cask design neutron source strength. For normal and accident conditions of transport, gamma dose rates were established by taking one-third of the gamma dose rates established in paragraph 3.2 of Section IX. The neutron dose rate was established s~eparately for the top and bottom of the cask and for radial mid-plane locations, a radial ESDRNPM calculation was performed to establish acceptable normal dose rates with the 1.0 weight percent boron in the neutron shield with the (new) neutron dose rate in the cavity. The new neutron dose rate of 26,1 mrem/hr, results in a total dose rate .(gamma and neutron) of 35.2 mrem/hr at the cask surface, or 6.4 mrem/hr at six feet from the personnel barrier. For the post accident condition, similar calculations show a total dose rate (gamma and neutron) of 450 mrem/hr at three feet from the cask surface. Both total dose rates are below the requirements of 49 CFR 173 and 10 CFR 71 for both normal transport and accident conditions.

IX-B1

2.0 Ganma Dose Calculation

The gamma source for the high burn-up assembly is 1.0 x 1016 MeV/sec. comparison of dose rates with the high burn-up source and the SAR gamma source of 3.074 x 1016 MeV/sec (Table IX-2) results In dose rates of (total gamma):

Normal Operations

SAR Source

High Burn-up Sodrce

*from personnel barrer

Hypocnectcal Accider

SAR Source High Burn-up Source

Fuel Mtdolane

Cask 6 feet* Surface

27.8 3.58 9.04 1.17

Fuel Midolane it Cask 3 feet

Surface

1438 400 468 130.

Too End

Cask 6 feet* Surface

12.9 1.94

4.20 0.63

Too End Cask 3 feet

Surface 63 14.5 "20.5 4.7

Bottom End Cask 6 feet*

Surface

13.3 1.44 4.33 0.47

Bottom End

Cask 3 fee: Surface

72 13.1

23 .4 4.3 \-,

3.0 Neutron Dose Calculation

The neutron source for the high burn-up source is 1.86 x 109 neutrons/ sec. This is significantly higher than the design basis source of 7.55 x 108 n/sec (Table IX-2). Consequently, a I weight percent boron solution is added to the neutron shield tank to maintain dose levels ALARA. Axial and radial neutrod shielding calculations were performed using the XSDRNPM code and the 123-group GAM-THERMOS cross-section library collapsed to 50 energy groups. SAR calculations use the 16-group Hansen-Road cross-section library. Flux to dose conversation factors were obtained from ANSI/ANS 6.1.1-1977 (NGGG).

Axial calculations were performed to verify that a decrease in dose rate at the top and bottom of the cask results when axial peaking of the

IX-82

/

neutron source is considered. A factor of 2 decr'ease is quoted on page IX-29, but this factor was conservatively not applied in the original analysis. Calculations performed using XSORNPM and a 20 percent peaking factor show a*55 percent decrease at the cask ends. (Although a peaking factor of 25 percent is applied to the fuel mldplane (radial) dose rate claculations, the 20 percent peaking factor was used in the axial calculations because it results in a smaller decrease at the cask ends.) These calculations confirm the factor of 2 decrease. The cask end doses listed below are based on the SAR (Table IX-1O) with the end dose rate decrease, conservatively set at 50 percent, applied (total neutron):

Nomal Operation

Top End Bottom End CasK Surface 6 feet* Cask Surface 6 feet-*

SAR Source 14.8 2.9 18.81 2.7 High Burn-Up Source 36.5 7.2 46.3 6.7

*from personnel barrier

Hypotheti cal Acci dent

Top End Bottom End Cask Surface 3 feet Cask Surface 3 feet

SAR Source 77 26.5 112 31 High Burn-Up Source 190 65.3 276 76.4.

Radial calculations were performed using a *circularized" fuel region as was done in Paragraph 3.3 of this section and the Kcornern corrections described (4%) were applied to the results. A 1.0 weight percent boron addition to the neutron shield tank water is included to maintain dose levels ALARA. All water is assumed to have been lost in the hypothetical accident scenario. Subcritical multiplication is included and results in an overall 70 percent increase in neutron population in the fuel region

IX-B3

and depleted uranium shield. Results were adjusted for ground scatter, and a neutron source axial peaking factor of 25 percent was conserva

tively applied.

The density of water at 3600F was used in the shield tank even though the temperatures for an actual shipment will probably be significantly less. The density of. water at 32*F was used in the fuel region to conserva

tively maximize the moderation of neutrons which in turn maximizes subcritical multiplication. A one percent increase was applied to dose rates to account for the higher 11ear power rating of the fuel, assuming

a fuel pellet stack height of 142.8 inches. Therefore, dose rates are

conservative for standard length (144 inches) fuel assemblies.

RADIAL NEUTRON DOSE RATE AT: Cask Surface .6 feet* 3 feet Normal Operation 26.1 5.20 Hypothettcal Accident 818.0 - 320.00

*from personnel barrier

Adding the gamma and neutron contributions for the high burn-up shipment results in total dose rates for normal and hypothetical accident

scenarios:

Normal Operation

Fuel Midplane Top End Bottom End

Cask Surface 6 ft. Cask Surface 6 ft. Cask Surface 6 ft.

Total Gamma 9.04 1.17 4.20 0.63 4.33 0.47

Total Neutron 26.1 5.20 36.5 7.2 46.3 6.7

Total 35.2 6.4 40.7 7.8 50.6 7.2

All dose rates at six feet from the personnel barrier are less than the 10 mrem/hr limit by more than 20 percent.

IX-84

Hypothetical Accident

Fuel Midplane

Cask Surface 6 ft.

468

818

1286

130

320

450

Top End

Cask Surface

20.5

190

211

6 ft.

4.7

65.3

70.0

Bottom End

Cask Surface

23.4

276

300

All dose rates at three feet are less than the 1000 mrem/hour limit by more

than 50 percent.

Based on the above analysis, the shipment of high burnup fuel is permitted,

provided that the assembly exhibits parameters of gamma and neutron source

terms equal to or less than those described above (and also in Section

111-6.0), as verified by ORIGEN analysis of the prospective assembly.

4.0 Detailed Neutron Shielding Analysis

Results:

NORMAL OPERATIONmrem/hour

Source Neutrons/sec

Normal with/w/o Boron 3.298-6 10.5 2.0

HYPOTHETICAL ACCIDENT

("Essentially Intact" Shield Tank, wall, page viii-2 of SAR) 1.235-4 328.8 129

*Neutron source of 7.55 x 108 n/sec or 2.064 x 106 n/sec/cm of active fuel length.

IX-B5

Total G amma

Total Neutron

Total

6 ft.

4.3

76.4

80.7

Surface* Dose Rate

Off-Cask Dose Rate

9

Surface Dose Rate - F x source per cm active fuel x

where F = Flux Peaking Factor (25%)

+ Corner Correction (4%) + 100%

- 129%

HA Normal with 1 w/o Boron

mrem/hr source n/sec

Ratio - 31.3 with low density shield water

Off the cask surface (3 ft. HA and 6 ft. normal), a 50 percent ground scatter is added.

increase due to

From the NLI-1/2 SAR:

Normal: Dose6 ft - 0.134 x Dose surface x 1.50 - 0.200 x dose surface

Hypothetical Accident: Dose3 ft - 0.260 x Dose surface x 1.50

surfacea 0.391 x dose

(From ratios of doses given in Tables IX-10 and IX-13 for fuel midplane)

1) Primary neutron dose rates are increased by 4 percent for "corner"

correction (related to the circularization approximation for the fuel region). An axial peaking factor (20 percent in VEPCO high burnup

analyses, 25 percent in SAR) increase is also applied.

2) A secondary neutron contribution of 50 percent was conservatively applied

in the SAR.

Radial Neutron Shielding Analysis:

1. Model the cask in the radial direction as a set of concentric circular

regions. The inner, square, fuel region can be modeled as a circular area whose area is equal to the edge length (8.88") squared, for a radius of

5.01 inches (12.725 cm).

I X-B6

2. Material Boundaries:

Radius (cm)

12.725 16.847

16.986

18.256

18.324

25.309

30.706

32.929

45.629

46.264

Interval

1-7 8,9 10 11,12 13 14-19

20-24 25 26-29 30

F

7

2

1

2

1

6

5

1

4

1

3. Interval 30, the outermost, is constructed with a thin annulus of steel on its outer layer. This annulus is used for neutron dose computation, which is conservative because the flux here is always greater than the surface flux. The average flux in the outer annulus is calculated by XSDRNPM.

4. Material constituents are:

Nuclide

-922351

922381

555081

400000

555011

130000

260001

240003

280003

Density (atoms/barn-cm)

2.700-5

5.200-3

3.373-3

3.744-3

4.165-2

6.026-2

6.042-2

1.673-2

8.367-3

IX-B7

Region

1

2

2

3

3

4

5

6

7

8

Material

Fuel Al uminum Void

Steel Void Uranium Lead Steel Water Steel

Mixture

1

(Fuel)

2

3 (Steel)

Material

U-235 U-238 Oxygen

Zi rconi um

Hydrogen

Aluminum

Iron Chromium Nickel

Density (atoms/barn-cm)Mixture

4

(Depleted uranium)

5

6

(Water)

Material

U-238

U-235

Lead

Oxygen Hydrogen Boron

Nuclide

555381

-555351

820000

80003

10001

50000

For Hypothetical Accident case, the density of the shield tank water is set to zero to represent its loss due to a tank puncture.

*When water is at 360 0F. "**From NAC-1 SAR.

IX-B8

4.773-2

1.052-4

3.300-2

2.960-2*

5.919-2*

None (unborated tank)

5.577-4** (Borated tank 1 w/o boron)

SECTION IX

APPENDIX C

SUPPLEMENTAL ANALYSIS FOR FERMI-i FUEL AND EBR-II BLANKET FUEL

The measured dose rates from Fermi-I and EBR-II fuel may be used to estimate the normal operation (no) and hypothetical accident (ha) dose rates for the cask. The first step is to calculate the dose rate from a typical PWR fuel assembly with 40,000 14WD/MTU burnup and 150 day cool time. This was done using the ISOSHLD shielding code, which uses point-kernel methods to calculate dose rates from various geometries. The actual radionuclide mixture present in PWR fuel was represented as an equivalent source of Cs-137, whose 662 keV gamma is representative of the gamma spectrum of the fuel. The fuel was represented as a right circular cylinder whose cross sectional area was 8.5 inches squared. The fuel was treated as a uniform medium consisting of Uranium at a density of 2.65 g/cm3 , which equates to 453 kgU and a 144 inch active fuel length. TMe Cs-137 source was calculated from the design basis gamma source of ;.074 x 101 MeV/sec or 4.64 x 1016 Cs-137 gammas/sec, which requires 1.41 x 10 Curies. The dose rates at six inches and three feet calculated are 213000 R/hr and 38,9000 R/hr respectively.

Table IX-l1 lists the maximum normal operation dose rate at six feet as 8.59 mrem/hour, including neutrons and gammas. The low burnups of the Fermi-i and EBR-II fuel do not produce significant quantities of neutron emitters so use of the total dose rate is conservative. A dose rate of 787 mrem/hour is listed in Table IX-14 for the maximum hypothetical accident dose rate. These values are scaled to find equivalent dose rates for Fermi-1 fuel by applying the equation:

Drom.- 12 R/hour x 16 assemblies x Dpwr 38900 R/hour per cask

using the dose rate measured at three feet. The EBR-II dose rates are similarly calculated from the dose rate measured at six inches and a cask load of four cans. The resulting dose rates are

Normal Op. Hypothetical Accident

Fermi-1 0.042 mrem/hour 3.9 mrem/hour

EBR-II 0.001 mrem/hour 0.074 mrem/hour

The estimated dose rates are much less than the allowed limits, thus the cask shielding is adequate for these fuel types.

Page Added IX-Ci Oct. 1990

This page Intentionally left blank.

Section X

SECTION X

CRITICALITY ANALYSIS

This page intentionally left blank.

SECTION X

CRITICALITY ANALYSIS

Table of Contents

A

B

C

D

E

F

APPENDIX G

INTRODUCTION AND SUMMARY

THE FRESH FUEL ASSUMPTION

THE MOST REACTIVE WATER/FUEL RATIO

GENERATION OF CROSS SECTION FOR NEUTRON TRANSPORT

CALCULATIONS

CRITICALITY OF PWR LOADING

TEMPERATURE EFFECT ON REACTIVITY

THE DRY FUEL

AN INFINITE ARRAY OF CASKS

BWR LOADING

CONSOLIDATED PWR FUEL

CRITICALITY EVALUATION FOR METALLIC FUEL

CRITICALITY EVALUATION FOR PWR OR BWR RODS

CRITICALITY EVALUATION FOR MARK 42 FUEL

CRITICALITY EVALUATION FOR MARK 22 FUEL

CRITICALITY EVALUATION FOR FERMI-I AND EBR-II FUELS

REFERENCES

Additional Criticality Analysis Mark 42 Fuel Assembly KENO Input and Output

Mark 42 Fuel Assembly NITAWL Output

Mark 22 Fuel Assembly KENO Input and Output

Mark 22 Fuel Assembly NITAWL Output

Computer Imput and Output for Fermi-1 Fuel and

EBR-II Blanket Fuel

Criticality Analysis for Connecticut Yankee

Stainless Clad Assemblies

a

i

I

1.0 2.0

3.0

4.0

5.0

6.0

7.0

8.0

9.0

10.0

10.4

10.5

10.6

10.7

10.8

11.0

APPENDIX

APPENDIX

APPENDIX

APPENDIX

APPENDIX

APPENDIX

Paqe X-1

X-1

X-3

X-5

X-8

X-11

X-11

X- 13 X-15

X-18

X-22

X-22

X-22d

X-22m

X-23

X-36

X-A-i

X-B1

X-C]

X-D1

X-El

X-F1

X-G1

Revised Feb. 1987 May 1987 Dec. 1988 Feb. 1990 Oct. 1990

List of Tables

Fuel Assembly Description

The Host Reactive Water/Fuel Ratio

The 16 Energy Groups Used in Transport Calculation

Temperature Effect on Reactivity

Effective Multiplication Factor of Basket Concept

1c vs Pitch (XSDRNPM, Sn)

Summary of Criticality Evaluation, for Metallic Fuels Mark 42 Fuel Criticality Results Summary

Summary of Criticality Evaluation for Mark 42 Fuel

Assemblies

Mark 22 Fuel Densities

Mark 22 Fuel Criticality Results Summary

Summary of Criticality Evaluation for Mark 22 Fuel

Assemblies Fissile Class I

List of Figures

ANISN Layout for NLI PWR Cask

An Infinite Array of PWR Casks

ANISN Layout With Basket NLI-1/2 Cask

k vs Number of Rods

X-5 keff vs Number of Rods

X-6 Nl-1/2 Cask and Metallic Fuel Contents Normal Fuel

Configuration

X-7 NL-1/2 Cask and Metallic Fuel Contents Failed Fuel Configuration

X-8 Mark 42 Assembly Cross Section

X-9 Mark 42 Assembly Longitudinal View

X-lO Mark 42 - Actual Loaded Cask Geometry

X-ll Mark42 4 KENO Model

Revised Feb. 1987 Feb. 1990 ii

X-1

X-4

X-3

X-4

X-5

X-6

X-7

X-9 X-9

X-1O

X-ll

X- 12

X-2

X-4

X-6

X-12

X- 16

X-21

X-22a

X-22c

X-22f

X-22m

X-22n

X-22o

X-1 X-2

X-3

X-4

X- 9

X- 14

X- 17

X- 19

X- 19

X-22b

X-22c

X-22g

X-22h

X-221

X-22J

List of Figures (cont.)

Mark 42 - KENO Model, Hypothetical Accident

Mark 42 - KENO Model, Hypothetical Accident

Mark 22 Assembly Cross Section

Hark 22 Actual Loaded Cask Geometry

Mark 22 KENO Model

Mark 22 KENO Model, Hypothetical Accident

NLI 1/2 KENO Model (Fermi-I Fuel Shown)

KENO Geometry

Fermi-I Normal Operation

Fermi-I Hypothetical Accident

EBR-II Normal Operation

EBR-II Hypothetical Accident

Page Added Feb. 1990

Revised Oct. 1990iii

X-12

X-13

X-14

X-15

X-16

X-17

X- 18

X-19

X-20

X-21

X-22

X-23

X-22k

X-221

X-22p

X-22q

X-22r

X-22s

X-28

X-29

X-30

X-31

X-34

X-35

This page Intentionally left blank.

Section X

CRITICALITY ANALYSIS

1.0 INTRODUCTION AND SUZ.'1,4ARY

This section describes the criticality safety analysis performed for the

NLI 1/2 spent fuel shipping cask. Although the cask is designed so that

under hypothetical accident conditions water will not enter into the can

containing the fuel, it was conservatively assumed for this analysis

that the can was filled with water. Consideration was also given to

possible changes in fuel rod configuratIon within the cask and to the

interactions of an array of casks surrounded by water. All of the

requirements of 10 CER 71.33 thru 71.38 with respect to criticality

safety are shown to have been satisfied.

Cask loadings of either 1 FWR assembly or 2 BWR assemblies were

examined. The. fuel assemblies considered have the specifications

indicated in Table X-1 and are conservatively consistent with the design

basis assemblies described in Section III, Table MII-1. The results

of this analysis yield a maximum keff of 0.94 for 1 FIWR assembly and a

much lower keff for 2 BWR assemblies in the isolated cask. If an array

of casks is considered with each crushed up against the others (so that

there is effectively zero neutron leakage at the outer boundary between

casks) and with water within each cask as well as between casks the

resulting kIs 0.94. (In the implausible situation of an infinite array of

water filled crushed casks ard no surrounding water the k iS 0. 98).

2.0 THE FRESH FUEL ASSUMPTION

Fuel'assemblies with zero burnup were used in this analysis. Aside from a

fuel assembly containing burnable poison elements, a regular fuel assembly

X-1

TABLE X-1

FUEL ASSEMBLY DESCRIPTION

PWR ypeBW~R Type

Initial enrichment, w/o *3.35 2 * 65

Active length, in 144 144

Rod array 15 X IS 7x7

Number of rods 204 49

Rod pitch, in 0.563 0.738

Rod diameter, in 0.422 0.563

Clad thickness, in 0.0243 0.032

Clad material Zircaloy Zircaloy

Pellet diameter, in 0.3649 0.487

"Fraction of uO 2 0.94 1.0

theoretical density

Dished end pellets No NO

Total uranium content, MTU * 0.457 0.208

*Uranium content and initial enrichment values ware subsequently increased to 0.475 NTU and 3.7 w/o respectively. These increased values result in a Keff of 0.95.

A PWR fuel assembly configuration containing additional irradiated fuel rods inserted and secured in the guide thimbles is permissible provided the initial uranium content of the assembly does not exceed 495 Kg and the maximum average initial U-235 enrichment does not exceed 3.35 w/o. There will be no increase in reactivity since the increase in uranium content coupled with the decrease in enrichment results in a mass of U-235 less than the design basis fuel assembly.

X-2

p-

always decreases in reactivity with burnup. Therefore the most reactive

condition of the.fuel which a cask is expected to accommodate is fresh or

near zero burnup fuel.

Although the reactivity of a fuel assembly with burnable poison pins

increases with burnup until most of the burnable poison is depleted , its

peak reactivity never exceeds, in practical power reactor applications,

the initial reactivity of a fuel assembly without poison pins. Fuel

assemblies with poison pins were therefore not included in this analysis.

3.0 THE MOST REACTIVE WATER/FUEL RATIO

The fuel assembly design parameters that have been used by severil reactor

manufacturers in this country are given in Table M-2. Some'variation can be

seen in fuel dimensions and configurations. To account for this variation, a

most reactive water/fuel ratio was established for each fuel pin cell.

Introduct/on of this most reactive ratio also protects the cask against

any possible rise of reactivity due to the temporary displacement of fuel

pins within the fuel can under hypothetical accident conditions.

As described in Section VIII, the aluminum blocks placed inside the fuel

can to serve as a heat transfer medium will not melt under hypothetical

accident conditions.* The most reactive water/fuel ratio is limited by

the amount of water available in the fuel can when the cask is flooded

with water. This ratio, in turn, is restricted by the across flats dimension

of the aluminum blocks. This dimension is 8 3/4 inches.

The search for the most reactive water/fuel ratio was dccomplished by performing

a series of fuel pin cell calculations with the LEOPARD (1) code. LEOPARD

X-3

db

TABLE X-2

THE MOST REACTIVE WATER/FUEL RATIO

PWR Fuel Pin Cell at 320F

Pitch Water/Fuel in Ratio

0.5630 (normal) 1.69 1.4063

0.5949* 2.05 1.4212

0.6368 2.54 1.4260

0.6437 2.62 1.4255

0.6506 2.71 1.4248

BWR Fuel Pin Cell at 320F

Pitch Water/Fuel in Ratio

0. 7380 (normal) 1.59 1.3536

0.7914 2.03 1.3661

0.8084 2.17 1.3665

0.8251 2.32 1.3654

*Maximum pitch allowed by constraint of aluminum blocks.

X-4

analyzes a repetitive lattice of fuel pin cells each consisting of concentric cylindrical regions of fuel, gap, clad, and moderator. Similar to MUMT and iRTE, LEOPARD performs a unit cell calculation and calculates an infinite medium spectrum for a given set of homogenized constituents. It then uses the spectrum to reduce 54 energy group fast cross sections to few-group fast cross sections. In a similar manner, it weighs each of 172 thermal groups to form one energy group thermal cross section. This set of macroscopic cross sections is then used in calculating the multiplication factor of the system.

In this calculation, the fuel pin size was maintained constant. By varying the amount of water associated with a fuel pin, the fuel pin cell reactivity varies. Results are presented in Table X-2. The asterisk indicates the largest uniform pitch which the fuel pins can possibly attain within the barrier of the aluminum blocks. For the BWR type fuel pin cell, the most reactive water/fuel ratio is 2.17. In the case of PWR type fuel pin cell, the most reactive ratio is limited by the 'largest possible pitch and, therefore, is 2.05. The largest uniform pitch was then entered into the cask calculation. This analysis assumes that the most reactive condition of the cask corresponds to the most reactive 6ondition of the fuel pin cells.

Although the above calculations were performed for Westinghouse and General Electric type fuels, the resulting most reactive water/fuel ratios also represent the most reactive conditions for the fuel designs described in Section III.

4.0 GENERATION OF CROSS SECTIONS FOR NEUTRON TRANSPORT CALCULATIONS

In view of the small size of the cask, neutron transport theory was applied in determining the reactivity of the cask. Sixteen groups of cross sections

X-5

Table X-3

THE 16 ENERGY GROUPS USED IN TRANSPORT CALCULATION

Group Number

1

2

3

4

5

6

7

8

9

10

11

12

13

14

15

16

Group Lower Energy Limit, eV

3.0 x 106

1.4 x 10 6

9.1 x 105

4.1 x 105

1.1 x-105

1. x 10 4

3.4 x 103

5.8 x 102

1.0 x 102

2.9 x 101

1.1 x 101

3.06

1.86

1.13

0.41

0

X-6

(15 fast and 1 thermal) for each material region of the cask were generated

using GAA-II( 2 ) and TIiEROS( codes. The 16-group neutron energy

structure is given in Table X-3.

GAM-fI is a code for calculation of fast neutron spectra and the associated

multi-group constants. The code contains 99 energy groups, ranging from

0.414 eV to 14.9 MeV. The neutron scattering kernel may be calculated

by the code up to B3 or P3 approximations. Each of the scattering kernels

is correct up to the sixth moment for anisotropic scattering in the center

of-mass system. Thus, the energy angle correlation is preserved for

slowing down in all nuclides in the medium. The inelastic scattering

and (n, 2n) processes are also included in the calculation. The resonance

absorption is treated quantitatively with the integral method developed

by L. W. Nordheim. (4)

THERMOS code solves the one-dimensional integral transport equation in

30 energy groups. The code computes the scalar thermal neutron spectrum

as a function of position in a lattice by solving numerically the multi

thermal-group integral transport equation with isotropic scattering. All

cross sections are weighted by the spectrum and edited into one single

group from zero to 0.4 eV.

Both fast and thermal cross sections for the homogenized UO 2 ' clad

and moderator were averaged over a neutron spectrum of the flooded fuel

assembly with the most reactive water/fuel ratio. The fast cross sections

for the nuclides present in the shielding materials were weighted by the

infinite medium spectrum of each material region. The thermal cross section

of these nuclides were weighted by the spectrum in the flooded fuel

assembly. These cross sections were then ready for use in spatial

calculations.

X-7

5.0 CRITICALITY OF PWR LOADING

Tho oJe ANISN( )was employed for spatial calculation of the cask and

for d:..'*:i.mininq its effective mutiplication factor. ANISN is a one-dimen

sior.-!, multi-group transport program which solves the Boltzmann trans

port equation by the method of discrete ordinates. The code allows

anisotropic scattering which may be approximated in any chosen order

of Legendre expansion.

S8 quadrature coefficients which were derived for use in the Carlson

Sn method(7) were input to ANISN. These coefficients are distinctive

in that the direction angles are restricted to symmetrical orientations

about each space point. Although other sets are available, the Carlson

Sn quadratures are the most commonly used.

Figure X-1 shows a layout of the PWR cask. There are nine regions

representing, explicitly-, the following materials:

Re'yion Description

1 Homogenized fuel with the most reactive

water/fuel ratio.

2. Aluminum

3 1/4" thick stainless steel can

4 1/8" gap filled with water

5 $1/2" thick stainless steel wall of the inner

cylinder

6 2 3/4" thick depleted uranium

7 2 1/8" thick lead

X-8

FIGURE X-1 ANISN LAYOUT FOR NLI PWR CASK

X-9

Region Description

8 7/8" thick stainless steel wall of the

outer cylinder

9 5" water shield

The fuel region was treated as a homogeneous mixture of UO 2 , clad,

moderator and structure materials all within a fuel assembly and has an

equivalent diameter of 10.02 inches. This value was obtained by

transforming the square fuel assembly to a right cylinder by conserving

the fuel volume. The geometric transformation will conservatively under

estimate the neutron leakage by approximately 10% . The fuel assembly

was assumed to be of infinite length and, therefore, the axial neutron

leakage was neglected. The U-235 content of the depleted uranium which

serves as a shielding material was taken as 0.22%.

The basic calculation assumed that the cask was filled with 1000F

water and that the fuel was also at 1000 F. The effective multiplication

factor as calculated by P1, S8 ANISN is 0.944 for the PWR cask. It

is estimated that the BWR cask will have an effective multiplication

factor of 0.85 (Section X-9.0).

Previous experience in the transport calculation on a geometry and

material compositions similar to those shown in Figure X-1 indicates

that P1 and P2 approximations yield results of k effwithin 0.1%. As the

anisotropic scattering Is calculated with high and higher order approximations,

the result converges to the true value of the multiplication factor. For

X-10

this analysis, the P1 approximation appears to give a converged keff. As to the order of angular quadrature, S2 and S8 produce little different in reactivity , ("-0.002).

6.0 TEMPERATURE EFFECT ON REACTIVITY

The multiplication factor of the cask was "evaluated at 100°F as mentioned in Section X-5 .0. At temperatures other than 100 F, the change of reactivity was estimated by running several LEOPARD fuel assembly aell problems. In addition to the capability described in Section X-3. 0, LEOPARD has an option to include the non-fueled cells in a fuel assembly. Outside the basic fuel pin cell, there is provision for describing an additional (nonlattice) region which is allowed to affect the spectrum but does not affect any of the calculations of spatial effects. The thermal flux In this region relative to the thermal flux in the cell moderator is an input item and, in this case, the ratio was estimated from experience with similar PNR fuel assemblies. Inclusion of the non-lattice region and, consequently,. the presence of more water Will make the temperature effect more

pronounced.

Table X-4 lists the fuel assembly reactivity at 32, 100, and 270°F. The reactivity change due to the change of water temperature Is negligibly small in the temperature range of our Interest. From 100 to 320, Ap is +0.04% and +0.02% for PWR and BWR type fuel respectively.

7.0 THE DRYFUfEL

If there is no water in the inner cavity of the cask as in the case of normal transport conditions, the cask reactivity would decrease substantially. According to Reference 8, UO2 containing less than 5 w/o U-235 cannot become critical if no moderating material is present. Transport of the dry

X-11

TABLE X-4

TEMPERATURE EFFECT ON REACTIVITY

Temperature

32°F 270°F

270°F

PWR Type Fuel Assy 3.5% Enriched

k

1.4356

1.4348

1.4280

BWR Type Fuel Assy 2.65% Enriched

k 1n56

1.3665

1.3661 1.3618

X-12

fuel will, therefore, pose no criticality problem.

8.0 AN INFINITE ARRy OF CwSKS

The effective multiplication factor presented in Section X-5.0 is the

result of the ANISN problem which has a zero flux condition at the outer

boundary of the cask.

To conservatively simulate the condition of an infinite array of casks,

additional ANISN problems were run with a reflection or zero current outer

boundary. It has been found (See Section X-9. 0) that the 5 inch water

shield is thick enough to serve as an infinite reflector and there is no

practical difference between keff and k. values of the cask. The value

given in Section 5.0 therefore is applicable to an infinite array of undamaged

casks.

If the casks are so damaged that the water jackets are collapsed or

crushed, the casks may be placed somewhat closer to one another.

The cask has shoulders (33 inches in diameter) at its top and bottom

With the water jackets crushed and neglecting the impact absorbers on

each end, the distance between two adjacent casks is maintained by the

shoulders and the closest center-to-center distance is, therefore,

33 inches (Figure X-2). Two damaged conditions were examined. The

first case assumes that the space between casks is filled with water,

while the second assumes the space between casks cohtains no

water. The fuel'conditions are the same as prescribed in the preceding

sections for a cask loading of 1 FWR fuel assembly. The resulting

infinite multiplication factor of the cask conservatively represents the

multiplication factor of an infinite array of damaged casks. The kvalues

are as follows:

X-13

FIGURE X-2 AN INFINITE ARRAY OF PWR CASKS

X-14

Cask k

With water between -

casks (3.5 inch thick) 0.944

Without water between casks 0.979

The higher k in the second case is due to the fast neutron interactions

between adjacent casks. Any water between the casks will thermalize

the fast neutrons and since thermal neutrons are unable to penetrate deep

into a neighboring cask, the k is lower in the first case.

Note must be taken that a flooded fuel assembly was assumed in all

calculations presented in this chapter. As pointed out in Section X-l. 0,

the fuel can in which the fuel assembly is loaded is designed to be

water-tight so that this, in itself, is a conservative assumption.

That a fuel can can be flooded while the space between casks is dry is

unlikely. The k of 0.979 is therefore a very conservative upper bound

of the cask reactivity.

9.0 BWR LOADING

CriUcality" analysis of the 2 BWR loading (and the 1 PWR loading) was

performed for the cask with a previous internals concept which utilized

a stainless steel fuel basket in place of the can and aluminum blocks

presently used. The methods of analysis were identical to those described

in the preceding sections. The ANISN layout for these analyses is

shown in Figure X-3. Note that the basic cask dimensions are Identical

to those given in Figure X-1.

The results of these analyses are given in Table X-5. Since the difference

in reactivity for the PWR loading resulting from the change in internals

-- X-1 S

- . - -a

TABLE X-5I--

EFFECTIVE MULTIPLICATION FACTOR OF BASKET CONCEPT

Number of Fuel Assys in Cask

1lPWR

SPWR

2 3WR

Fuel Enrichment w/o

3.35

2.50

2.65

Fraction of

3L.2 Densit¥

0.94

1.00

1.00

/ e

X-16

Cask k

0.92S

Cask

k elf

0.925

0.865

0.821

4.52" 0.07

FIGURE X-3 ANISN LAYOUT WITH BASKET NLI 1/2 CASK

X-17

S. ... . . ----. . --.. . ... *°....-.. ....- -..-.-....-...--.... .

design is small compared to the margin for the BWR loading, it was concluded

that reanalysis of the BWR loading was not necessary. If the same increase in

reactivity of the PWR loading applies to the BWR loading, the effective

multiplication factor of the present BWR cask will be no higher than 0.85.

10.0 CONSOLIDATED PWR FUEL

Consolidation of fuel consists of packing the fuel rods from up to two PWR

assemblies in a stainless steel canister. The non-fuel bearing components of

the assemblies are packaged seperately.

Analyses of the effects of the shipment of consolidated fuel in the NLI-1/2

cask have been performed for criticality, shielding, structural and thermal

effects. The consolidated fuel modeled in the criticality, shielding, and

structural analyses ifs W15 x 15 fuel cooled for two years, with an initial

enrichment of 3.7 w/o U-235 and a burnup of 40,000 MWD/MTU. These values are

considered to be representative of consolidated PWR spent fuel shipments. The

thermal analysis has been performed for W14 x 14 fuel cooled 12 years to add

an additional margin of conservatism because the thermal behavior of

consolidated spent fuel is currently being investigated.

10.1 Analysis Method

The process of rod consolidation makes it difficult to guarantee a specific

number of rods in a canister, because some rods may be bowed, dented, or

otherwise at variance with their original manufactured dimensions. The effec

tive multiplication factor, k-effective, depends on the number of rods per

canister and the resulting rod pitch. The k-effective for design basis fuel

canisters is therefore variable, but k-effective may be bounded by calculating

it for a standard assembly (W15x15, 204 rods) and for a canister consolidated

2 to 1 (408 rods). It is also necessary to establish the trend of k-effective

between these bounds. To establish this trend, k-infinity was calculated at

various rod pitches (with an equivalent number of rods) and plotted-in Figure

X-4. This figure shows that the standard assembly is undermoderated, and that

k-infinity decreases as the number of rods is increased from 204 (standard

X-18

r

I

200 300 400Equivalent Number of Rods / 8.5" ID Canister

500

Figure X-5 - Keff vs. Number of Rods

Hypothetical Accident

(N Standard "-'i15x15

(204 Rods)

I

I I

I

,Estimated

F I0 200 300 40C

-- 2to 1 Consolidation

I I I

I I

I. I

0 500

Equivalent Number of Rods / 8.5" ID Canister

X-19

1.40

1.20 a

P

-55-

SFigure X-4 - Kwys. Number of Rods *a-- Hypothetical Accident Scenario one ormal Operation

Square Lattice

WI I"'W15x15 2.o \Cnsliato m (204 Rods)

T g L

I I I (Triangular Lattice , I I!

,I .. . . ,• I

k.

1.00

0. s

4 a

U

I

I

S

I

I

0.90

0.80

keff

0.70

0.60

a

4

4l

I

assembly) to 408 (2 to 1 consolidation). A calculation of k-effective

for a standard assembly in the cask was performed and yielded k-eff =

0.91564 t 0.00371 for Normal Operation conditions. (A similar

calculation for Transportation Hypothetical Accident conditions yielded

K-effective - 0.93092 t 0.00260.) Rather than calculate k-effective at

each possible number of rods between 204 and 408, an estimated k-eff was

calculated by multiplying k-eff for a standard assembly by the ratio of

k-infinity for the desired number of rods, N, to the k-infinity for the

standard assembly:

kest(N) a keff(Std. Asbl.) x ka(N) Koo(Std. Asb7.)

The calculated k-effectives and k-estimateds are plotted in Figure X-5

and the specific results are shown in Table X-6. A verification check

was made by calculating k-effective at 373 rods/canister, which is the

most loosely-packed canister expected in actual consolidation operations.

10.2 Results and Discussion

Inspection of Figure X-5 and Table X-6 shows that:

(1) k-effective decreases as the number of rods increases, i.e., k-effective is less for any consolidated fuel canister containing more rods than a standard asembly.

(2) k-effective for Normal Operations for consolidated canisters is less than k-effective in the Hypothetical Accident scenario, because the Hypothetical Accident scenario employs a larger rod pitch with a larger k-infinity.

(3) k-effective for the worst case, the standard assembly Hypothetical Accident, is 0.93092 t 0.00260. A calculation of ks, which includes corrections for bias and uncertainties give ks - 0.950 for the Hypothetical Accident scenario and ks = 0.936 for the worse case Normal Operation conditions (Std. Asbl.).

Thus, the consolidated fuel canister is shown to be less reactive than

the standard assembly, under all conditions, and is subcritical for all

conditions.

X-20

TABLE X-6

km vs. Pitch (XSDRNPM, Sn)

keff (Keno)

0.93092 * 0.00260

0.91564 1 0.00371

0.907 est.

0.868 est.

0.802 est.

0.752 est.

0.71336 1 0.00255

0.635 est.

0.520 est.

0.8484

0.8068

0.7899

0.7334

0.6670

0.6488

0.6301

0.6024

0.5681

kc

1.4403650

1.4213351

1.4073366

1.3476443

1.2444021

1.1683589

1.1627481

0.98522996

0.80736114

Equivalent No. of Rods (8.508 ID can)

204, H.A. (8.870 car.)

204, Normal Operations

237

275

323

352

373, loosest expected packing (1.85 to 1)

2 to 1

459, maximum packing

Pitch (CM)

1.504 square

1.430 square

1.400 square

1.300 square

1.200

1.150

1.200

square

square

triangular

1.147

1.082 triangular rod dla. 1.072

# rods - Can area (8.50")2 Cell area

Can ID O 8.50" 2 x can wall a 0.18"

T.3-

8.870 cavity gives 0.1" clearance

X-21

10.3 Comouter Codes

The calculations for k-infinity were performed using the XSDRNPH module of SCALE-2'J

Calculations for k-effective were performed using HITAWL, XSDRNPN, and KENO-IV

modules of SCALE-2. Dancoff correction factors and effective moderator cross

sections for NITAWL were calculated by the NULIF code provided by Babcock and

Wilcox.

10.4 Criticality Evaluation for Metallic Fuel

The metallic fuel has only natural enrichment, 0.711 w/o uranium-235. Metallic

natural uranium fuel cannot achieve criticality in any geometric configuration with

(light) water as a moderator. Neither is an array of packages containing natural

uranium fuel critical, with or without light water moderation. Table X-7

demonstrates that the loaded cask meets the criteria established for Fissile Class

I packages, as defined by 10 CFR 71.38. A description of the fuel and its

configuration in the cask are provided in the following section.

10.5 Criticality Evaluation for PWR or BWR Rods

The enrichment limit for a BWR rod shipment is 5.0 w/o uranium-235. Th]i-'

enrichment is higher than that of the design basis BWR fuel assembly. Therefore,

an analysis must be performed to demonstrate that a shipment of these rods is

critically safe.

A series of runs was performed with the CSAS25/KENO V.a sequence, modeling an

infinite array of NLI-1/2 casks containing 25 BWR rods, enriched to 5.0 w/o U-235,

in the PWR basket. The pitch between rods in the cask was varied, and the keff

values recorded.

The peak keff was determined to occur at a pitch of 3.0 centimeters. This

configuration represents the optimum pitch for the BWR rods in the NLI-1/2 cask.

This configuration resulted in a k*f value of 0.6790 ± 0.0027, which gives a keff

+ 2o value of 0.6844. This value is well below the accepted limit of 0.95;

therefore, the NLI-1/2 can safely transport 25 BWR rods with an enrichment of 5.0

w/o uranium-235.

Revised Oct. 1986 May 1987 Jan. 1990 Feb. 1991 Oct. 1991 X-21a

The enrichment limit for-a PWR rod shipment is 4.9 w/o uranium-235. This

enrichment is highe* than the design basis enilthment limit. During the shipment

of PWR rods, it is hypothetically possible for the rods to attain their optimum or

most reactive pitch. Therefore, it is necessary to calculate the optimum pitch for

a PVR rod and determine the reactivity of an array of 25 rods.

The XSDRNPH discrete ordinates code was used to determine k, values for various rod

pitches. A pitch of 1.53 centimeters was used as a starting point. The pitch was

increased by a small distance and k. was calculated for. that case. The pitch was

steadily increased and k. calculated until a maximum value of k, was determined.

The values for each pitch analyzed are listed below:

1.53 1.4983526

1.56 1.5044017

1.58 1.5076288

1.60 1.5103494

1.62 1.5125901

1.64 1.5143920

1.66 1.5156379

1.68 1.5165032

1.70 1.5169899

1.72 1.5170627

1.74 1.5168190

1.76 1.5162339

1.78 1.5153357 1.80 1.5141438

1.82 1.5126569

1.84 1.5109039

A peak in k. was seen at a pitch of 1.72 centimeters. This is the optimum pitch

for the PWR rods. KEND-IV was used to model a square 5 x 5 array containing 25

rods at optimum pitch inside the NLI-1/2 cask. This case gave a kff of 0.52478

: 0.00254, which corresponds to a k. of 0.545. tA homogenized, cylindrical array

of 25 rods yielded a k6 of 0.540, indicating that the array geometry does not have

a strong effect on reactivity. Therefore, the KLI-1/2 cask can safely transport

25 PWR rods with an enrichment of 4.9 w/o uranium-235.

Page Added Jan. 1990

Revised X-22 Oct. 1991

Table X-7 SUMMARY OF CRITICALITY EVALUATION FOR METALLIC FUELS

FISSILE CLASS I

NORMAL CONDITIONS

Number of undamaged packages calculated to be subcritical (Fissile Class I must be infinite; Fissile Class II must be at least 25; and Fissile Class III must be at least identical shipment.)

Optimum interspersed hydrogenous moderation (required for Fissile Class 1)

Closely reflected by water (required for Fissile Class II and III)

Package size, cm3

Infinite*

yes

yes

60,240

ACCIDENT CONDITIONS

Number of damaged packages calculated to be subcritical (Fissile Class I must be at least 250; Fissile Class II must be at least 10; and Fissile Class III must be at least 1.

Optimum interspersed hydrogenous moderation,

full water reflection

Package size, cm3

Other Transport Index (must not exceed 10 for Fissile Class II)

250"

yes

60,240

Not Applicable

"Natural Enrichment Uranium Fuel cannot go critical in any configuration, with or without (light) water moderation. No combination of damaged or undamaged packages can result in criticality.

Page Added Oct. 1986 Revised Feb. 1987 Jan. 1990 X-22a(])

Figure X-6 NL-1/2 CASK. & METALLIC FUEL CONTENTS NORMAL FUEL CONFIGURATION Page Added

Oct. 1986

X-22b

4

Fi gure X-7 NL- 1/2 CASK AND METALLIC FUEL CONTENTS

FAILED FUEL CONFIGURATION

X-22c

Page Added Oct. 1986

10.6 CrieicalIMy £vuluatfon for HAJC ,2 Fuel

A MARK 42 fuel assembly consists of three concentric fuel tubes wi:h PuO2 -Al powder mecallurgy cores containing a total of 3.3SO kg of plutonium as shown in Figure X-8 and Figure X-9. The cores are clad with type 6063 aluminum and previously surrounded an inner target which has been removed. The plutonium was initially enriched to contain 78.28 w/o Fu 18.58 w/o pu2 4 0 . 2.27 v/o1au 0.71 v/o P2U2 and 0.15 v/o Pu

Cross-section sets were prepared by NITAWL using infinite dilution. Dancoff factors are not applicable for this assembly because of the annular cylinder geometry and resonance calculations were conserve. tively ignored in these analyses. The cross-section sat used for these analyses is the 27 group version of the ENDF-B/IV library. These cross-sections were not collapsed or weighted, and were passed directly from NITAVL to KENO.tV. The output for NITA1WL is given in Appendix A. Oak Ridge National Laboratory has validated the codes and cross-section set used in the analyses.as shown in Reference 9.. that no bias is required when analyzing plutoniLu systems.

The Mark 42 will eventually be sectioned to several twenty inch long segments. 'Jhen :he Mark 42 fuel is :ransported in the NLI-1/2 cask following sectioning, two quadrants of the four quadrant "Rockwell basket* will contain the sectioned assembly. The criticality analyses were conservatively performed for two adjacent fuel assemblies as shown in Figure X-10. The cask geometry that was modeled in KENO-IV is shown in Figure X-l. Several analyses were performed to determine the effect of different parameters on the reactivity of the fuel. The normal operation analysis utilized the initial concentrations of plutonium isotopes including neutron poisons such as P-s240 (18.58 w/o) and Pu242 (0.71 w/o). Aluminum was also included in the fuel modeling.

Page Added Aug. 1988 Page Revised Dec. 1988* X-22d Oct. 1990

This analysis yielded a keff of 0.556. In the off-normal analysis. th* neutron poisons, such as pu24 and pu -- ..ere cOns.r.atively

removed to account for possible inaccuracies in the pluconiu Isotope measurements for the fresh fuel. Aluminum in the fuel was also removed. which does not have much effect since aluminum is almost c€ansparenct-o neutrons. This analysis resulted in a kef Of 0.612. The input and output for this limiting case are given in Appendix 3.

Alternate configurations of the fuel were also modeled to simulate a hypothetical accident scenario. The fuel was evaluated as two cubes either centered in opposing quadrants or closely packed in adjacent quadrants as seen in Figures X-12 and X-13 respectively. All of the accident configurations yielded a keff lover than 0.612. These analyses show that the initial annular configuration of the Mark 42 fuel assembly is the most reactive one. The criticality results for the different cases are presented in Table X-S. Table X-9 demonstrates that the loaded cask meets the criteria established for Fissile Class I packages, as determined by 10 CFiR 71.57. The results of the criticality evaluation for the NLI-1/2 cask containing Mark 42

fuel show that kef. does not exceed 0.612.

Table X-8 Mark 42 Nal Criticality Results Summry

•t~q keff

Normal Operation 0.556 Off-normal Operation 0.612 Centered - Opposing quadrants 0.321 Closely packed - Adjacent quadrants 0.285

Page Added Aug. 1988 Page Revised Dec. 1988 X-22e Oct. 1990

Table X-9 SUMHAl• OF CRITICAI.TY. EVAIUATION FOR MARK 42 FUEL ASSEMBLIES

FISsnz CLss i

NORMAL CONDITIONS

Number of undamaged packages calculated to be subcritical (Fissile Class I must be infinite)

Optimum interspersed hydrogenous moderation

(required for Fissile Class I)

Closely reflected by water (required for Fissile

Class II and III)

Package size, cm3

ACCIDENT CONDITIONS

Number of damaged packages calculated to be

subcritical

Optimum interspersed hydrogenous moderation,

full water reflection

Package size, cm3

Other Transpbrt Index

Page Added August 1988

Infinite

yes

yes

60,240

Infinite

yes

60,240

Not applicable

X-22f

ova

a06

n a

0 n 60

is i

0 -. 11 n x

Q

2 Pu :ui

Pg

20 0 an'4

(AI

CA(4-4 r-�Ip

z 0 C U)

C,

C

vi

0�J

Ii

0 C '4 2 C,

S 'I

gq I-

a, I-

C

I

x

I

I.'U.-

Figure X-10

&ARK 42 - ACTUAL LOADED CASK mEom=T

Stainless rSteel

Depleted Uranium

Lead

Water

Page Added December 1988

X-221

Figure X-11

MARK 42 - KENO MDDEL

Specular

Page Added December 1988

X-22J

Y

X

Figure X-12

MARK 42 - KEmO MODEL, Hypothetical Accident

Specular Reflector

X-22k Page Added December 1988

Figure X-13

MAM 42 - KMO MODEL

Hypothetical Accident

Specular Reflector

Page Added December 1988

X-221

Fuel