ML19341C535

Chapter 8 to Univ of Tx Triga Mark 1 RSAR, Safety Analysis.

ML19341C535
Person / Time
Site:	05000192
Issue date:	01/31/1981
From:	TEXAS, UNIV. OF, AUSTIN, TX
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NUDOCS 8103030704
Download: ML19341C535 (27)

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I .L l

8. SAFETY ANALYSIS In this section an analysis of abnormal operating conditions will be made with conclusions concerning the effects on safety to the reactor, the public, and the operations personnel, as a consequence of any abnormal opecations.

The abnormal conditions that will be analyzed are:

1. Clad rupture

2. loss or reactor coolant

3. Reactivity accident 8.1. FISSION PRODUCT RELEASE In the analysis of fission product releases under accident conditicns, it is assumed that a fuel element in the region of highest power density fails.

8.1.1. Fiesion Product Inventory Tabic 8-1 gives the inventory of radioactive nobic gases and halogens in the TRIGA Mark I after continuous operation at 250 kW for four years (i.e., 1 MW-yr).

8.1.2. Fission Product Release Fractions The release of fission products from U-ZrH fuel has been studied at some length. A summary report of these studies (Ref, i ) indicates that the release from the U-Zril l.6 fu 1 meat at the steady-state operating l

l a-1 l

810bh0 3 0 kCidLl

temperatures is principally through recoil into the fuel-clad gap. At high temperatures (above 400*C or 500*C), the release mechanism is through a dif-fusion process and is temperature-dependent, unlike recoil.

TABLE 8-1 NOBLE CAS AND IIALOGENS IN THE REACTOR Ir.ct oi t

_ _ . . _ _ . _ . . _ ,f ._ . . . - Q t n t *. r/ (C1) hr-83 I a.'20 i

.x-83m 1 .20 31-84 .' , )nn Br-85 [ 2,150 Kr-85m 2,150 1

Kr-85 113 Kr-87 5,400 Kr-88 7,700

, Kr-89 9,750 Kr-90 10,850 Kr-91 7,350 I-131 5,950 Xe-131m 48 I-132 8,850 I-133 14,350 Xc-133m 350 Xe-133 14,350 I-134 16,100 1-135 13,400 Xe-135m 4,050 Xe-135 13,850 1-136 12,950 Xe-137 12.550

. Xe-138 11,700 Xe-139 11,800 Xe-140 8,100 8-2

For the accident considered here, it is assumed a fuel element in the region of highest power density fails in water and that the peak fuel temperature in the element is less than 300*C. At this temperature, the

~

-5 For the purpose long-term release fraction would be less than 1.5 x 10 .

of this analysis it is also assumed that 100% of the noble gases and 50% of the halogens are released from the highest power density fuel element in which 2.6% of the total power is generated.

It is important to note that the release fraction in accident conditions is characteristic of the normal operating temperature and not the temperature during the accident conditions. This is because the fission products re-leased as a result of a fuel clad failure are those that have collected in the fuel-clad gap during normal operation.

Other assumptions concerning the transport from the fuel to the exit of

, the stack are:

1. 100% of the noble gases released from the fuel are transported to the building exhaust stack.

2. 10% of the halogens released from the fuel are in the form of organic compounds and all of these halogens escape from the tank water.

3. Only 1% of the balance of the halogens escapes from the tank water.

4. There is no plate-out of any of the fission products.

5. The stack radiation monitor fails to place the ventilation system

. in the emergency mode (and that the reactor operators also fail to do so).

6. The effective building ventilation rate is 15 air changes /hr.

(This is greater than the actual release rate but results in larger dose rates.

8-3

The net effect of these assumptions is that for the accident condition, the fraction of the noble gases released from the building is:

-5 -2 ~7 f,;g = 1.5 x 10 x 1.0 x 2.6 x 10 = 3.9 x 10 ,

and of the halogens:

-5 f = 1. 5 x 10 H x 0.5 x (0.1 x 1.0 + 0.9 x 0.01) x 2.6 x 10-2

~

= 2.1 x 10 .

8.1.3. Downwind Dose Calculations The minimum roof 1cvel dilution factor was calculated, in Sec-tion 5.4 3 to be 4.2 x 10- sec/m . This is based on mixing in the lee of the building when the wind velocity is ) m/sec.

The calculation of whole body gamma doses and thyroid doses downwind from the' point of release was accomplished through the use of the computer code C DOSE (Ref. 2 ). In this code the set of differential equations describing the rate of production of an isotope through the decay of its precursors and the rete of removal through radioactive decay and removal by the ventilation system is integrated for each ecmber of the chain. The release rate qf to the environment for the ich isotope at time tg, in hours is:

q (t) = gg Q1 (t) (1/V)/3600 ,

f where Q (t) = the concentration of the ich isotope in Ci/m ,

3 1/V = the building leakage rate tn (m /hr)/m ,

g = 1-c, g c = the filter efficiency for the ith isotope.

l 8-4

I l . - - .

l i

The quantity Q g (t) is the concentration of the ith isotope in the discharged air at the time, t. This concentration is given by Qg (t) = f Q ff W e where Q g (0) = the concentration of the ith isotope as found in Table 8-1, A = the decay constant for the ith isotope, and 1

f = the release fraction to the reactor hall.

f The concentration downwind at a distance x for the ith isotope is calculated from Q '(t,x) = q (t-T) * $(x)e ,

f 1 where T = the transit time from the release point to the dose point, hr,

$ = the dilution factor at the distance x, sec/m .

The whole body gamma ray dose rate for the ith isotope, Dy , at the distance x and time t is caleclated, assuming a semi-infinite cloud, through the expression:

D (t,Xf) = 900 Ef1Q '(t,x) ,

where Eg = the average gamma ray energy per disintegration, MeV, and the constant includes the attenuation coefficient for air as well as the conver--

i 1

sion factors required.

O 8- 5

Internal dose rates, in this case the dose rate to the thyroid, are

. calculated by:

. D dy (t,x) = 3600 B-Q '(t,x)K g ,

where B = the breathing rate, m 3/sec, and Kg = the internal dose effectivity of the ith isotope, rem /Ci.

The values for the breathing rate are given in Table 8-2 and are taken from USAEC Regulatory Guide 4.

The average gamma ray energy per disintegration and the internal dose ef fectivity for each isotope considered are given in Table 8-3.

The decay products of these isotopes are also included in the calcula-tion; however, their contribution to the dose rates are small and therefore the data for these isotopes were not included in the table.

8.1.4. Downwind Doses The whole body gamma dose and thyroid dose in the lee of the building are shown in Tabic 8-4. These doses are trivial in nature.

TABLE 8-2 ASSUMED BREATIIING RATES '

Time (hr) Breathing Rate (m3/sec) 0 to 8 3.47 x 10-4 8 to 24 1.75 x 10-4 Over 24 2.32 x 10-4 4

8-6

TABLE 8-3

" AVERAGE CAMMA RAY ENERGY AND I?RERNAL DOSE

'EFFECTIVITY FOR EACH FISSION PRODUCT ISOTOPE Isc> tope Eg(MeV) Kg (rem /Ci)

Bit-83 -2 0.92 x 10 Pr-84 1.87

'l-131 0.40 1.486 x 10 6 I-132 1.96 5.288 x 10 1-133 0.56 3.951 x 10 I-134 3.02 2.538 x 10 4 I-135 1.77 1.231 x 10 1-136 2.91 Kr-83m -3 0.8 x 10 Kr-85m 0.16 Kr-85 0.4 x 10~

Kr-87 1.07 Kr-88 2.05

, Kr-89 2.40 Xe-131m 0.82 x 10 -2 Xc-133m 0.37 x 10' Xe-133 0.29 x 10-Xe-135m 0.46 Xe-135 0 25 Xe-137 1.22 Xe-138 1.57 8-7

. TABLE 8-4 DOWNWIND DOSES FROM FISSION PRODUCT RELEASE

, Distance (m) Whole Body Gamma (mrad) Thyroid (mrem)

Accident condition 0 (Release of fission products from one .24 2.66 fuel element)

\

ti. 2 LOSS OF REACTOR C00!>.NT 8.2.1 Summary The reactor will operate at a calculated maximum power density of 6 kW/ element when the reactor power is 250 kW and there are 63 elements in the core, all of which are standard TRICA fuel. If the coolant is lost immediately af ter reactor shutdown, the fuel temperature (see Fig. 8-1) will rise to a maximum value of s275'C. The stress imposed on the fuel element clad by the internal gas pressure (see Fig. 8-2) is about 1200 psi when the l fuel and clad temperature is 275*C and the yield stress for the clad is about 37,000 psi. Therefore, it can be concluded that the postulated loss-of-coolant accident will not result in any damage to the fuel, will not result in release of fission products to the environment, and will not requi.e emergency cooling.

8-8 i

L 2000 -

l 1800 -

1 COOLING TIME (SEC r

. 1400 - 0 3

! 10 1200 -

4 i P l  !

O 5 ,

Q 1000 -

5 I

i i . t" j 800 -

1 j 5 1

i 600 -

~

400 -

200 -

'.s /

! o

+ i i . . . , , ,

0 5 to 15 20 25 30 35 40 45 OPERATING POWER DENSITY-KV/ ELEMENT I

EL-1872 Fig. 8-1. Maximum fuel temperature versus power density after loss of coolant for various cooling times between reactor shutdown and coolant loss l 8-9

= - . - , - - , - - - - - - - - - - .

5 10 _

ULTIMATE STRENGTH

_,~% s

's N N

N YlELD STRE."GTH

\\

\

4 \

10 -

g N

m \

\

e, \

m m

\

N e N O N N

STRESS IMPOSED ON CLAD 10 -

2 10 , , , , , , ,

400 600 800 1000 1200 TEMPERATURE (*C) EL-1873 Fig. 8-2. Strength and applied stress as a function of temperature, U-ZrlII .65 fue l. , fuel and clad at same temperature 8-10 l

./

j 105 . Tm = 500 600 70' 800 90 100 0

T S

5 T>

ca 5

5 10 -

8 s

e U

x 5

5 w

d E

taJ 10 -

5.

3 c>

b T 6-MA/. FUEL TEMPERATURE 10 2 AFTER WATER LOSS ('C)

. 0 10 20 :0 40 S0 60 70 IWG ' ;

FlX iTf-KW/EL: MENT EL-0706A Fig. 8-3. Cooling times after reactor shutdown l

necessary to limit maximum fuel tempera-ture versus power density l 8-11 I

l

If 'the reactor tank is drained of water, the fission product decay heat wiil be removed through the natural convective flow of air up thrcagh the reactor core. If the decay-heat production is sufficiently low because of a low fission product inventory or a long interval between reactor shutdown and coolant loss, the flow of air will be enough t i maintain the fuel at a temperature at which the fuel elements are undamaged. The following analysis shows that:

1. The maximum temperature to which the fuel can increase is 900*C without substantial yiciding of the clad or subsequent release of fission products.

2. This temperature will never be exceeded under any conditions of coolant loss if the maximum operating power density is 22 kW/ element or less.

3. For maximum operating power densities greater than 22 kW/ element, emergency cooling can be provided to ensure that the fuel-element temperature does not exceed 900*C. The required emergency cooling time as a function of maximum operating power density is shown in Fig. 8-3.

8.2.2. Fuel Temperature and Clad Integrity The strength of the fuel element clad is a function of its temperature.

The stress imposed on the clad is a function of the fuel temperature as well as the hydrogen-to-zirconium ratio, the fuel burnup, and the free gas volume within the element. In the analysis of the stresses imposed on the clad and strength of the clad the following assumptions will be made:

, 1. The fuel and clad are at the same temperature.

2. The hydrogen-to-zirconium ratio is 1.65.

8-12

3. 'The free volume within the element is represented by a space 1/8-in. high within the clad.

4. The reactor contains fuel that has experienced burnup equivalent to only about 4.5 MW-days.

The fuci element internal pressure p is given by p"ph+Efp P air '

where p = hydrogen pressure, h

pgp = pressure exerted by volatile fission products, and pg = pressure exerted by trapped air.

For hydrogen-to-zirconium ratios greater than about 1.58 the equilibrium hydrogen pressure can be approximated by p =

h xp [1.767 + 10.3014x - 19740.37/(T g)] (atmosphere's) , (2) where x = ratio of hydrogen atoms to zirconi.im atoms, and Tg= fuel temperature (*K).

This expression was derived from least-square fits to the data of Dee and Simnad (Ref. 3 ). For Zril l.65 the hydrogen pressure becomes ph= 1.410 x 10 exp [-19740.37/(T g)] (atmospheres) .

The pressure exerted by the fission product gases is given by e

, RTg P fp " f y E ,

(3) 8-13

where f = fission product release fraction.

n/C = number of moles of gas evolved per unit of energy produced, moles /W-day,

-2 R = gas constant, 8.206 x 10 liters-atmospheres / mole 'K, V = free volume occupied by the gases, liters, and E = total energy produced in the element, W-day.

The fission product release fraction (Ref. 4 ) is given by f= 1.5 x 10-5 + 3.6 x 103 exp -1.34 x 104 /(Ty) du , (4) where T = fuel temperature in the differential volume of the element during normal operation, *K, and v = fuel volume normalized to 1.

The fission product gas production rate n/E is not independent of power density (neutron flux) but varies slightly with the power density.

. The value n/E = 1.19 x 10 -3 moles /W-day is accurate to within a few per-cent over the range from a few kilowatts per element towellover40kN/

element. The free volume occupied by the gases is assumed to be a space 1/8-in. (0.3175-cm) high at the top of the fuel so that 2

V = 0.3175 x r f

, (5) where gr = inside radius of the clad (1.822 cm).

For utandard TRICA fuel the maximum burnup is about 4.5 W-days /

element.

, Finally, the air trapped within the fuel element clad would exert a pressure ph a = RTg/22A (6) 8-14

where it'is assumed that the initial specific volume of the air (22.4 liters /

moles) is present at the time of the loss of coolant. Actually, the air forms oxides and nitrides with the zirconium so that after relatively short operation the air is no longer present in the free volume inside the fuel element clad.

For Zril fuel burned up to 4.5 W-days / element, with a maximum l.65 operating temperature of 600'C, the internal pressure as a function of maximum fuel temperature gT is 0

p = 1.410 x 10 exp (-19740.37/T ) + 3.66 x 10 g Tg (amospheres) 9 or p = 2.073 x 10 exp (-19740.37/T ) + 5.38 x 10 -2 . (7) g Tg (psi)

The stress imposed on the clad by the gases within the free volume inside the clad is

~

S= (p) ,

(8) where r = clad outsido radius (= 1.873 cn'),

t = clad thickness (= 0.051 cm) .

If Eqs. 1 and 8 are combined, the stress can be rewritten as S = 36.7 p 10

= 7.61 x 10 exp (-19740.37/T g ) + 1.97 Tg (psi) . (9)

In Fig. 8-2 this imposed stress is plotted as a function of maximum fuel temperatures. Also plotted are the yield and ultimate strength of the type 304 stainless steci clad. The ultimate strength of the clad is not

, execcced if the maximum fuel temperature is maintained below about 950*C l 8-15 l

and the yield strength is not exceeded for any fuel temperatures below

. about 920*C. The Ilmit is set at 900*C, slightly below the yield point and well below the rupture point.

8. 2. 3. After-Heat Removal Following Coolant Loss It is assumed that the reactor operates continuously at a constant power density IcVel P so that the maximum inventory of fission products is available to produce heat after the reactor is shut down. The power density after reactor shutdown P is given by P = 0.1 P, [(t + 10) * - 0.87 (t + 2 x 107)-0.2j x {I .3 cos [2.45 (0.0261 - 0.5)] } ,

(10) where P = operating power density, W/cm ,

t = time af ter reactor shutdown, sec, L = distance from the bottom of the fuel region, cm.

At the time that the coolant is lost from the core the fuel and its surroundings are assumed to be at a temperature of 27'C. This is not necessarily true, for an accident can be postulated in which the coolant loss in the mechanism by which the reactor is shut down. (For the standard non-gapped fuel elemant, under normal conditions, the time to cool down from operating temperatures is a matter of one to two minutes.) Although such an accident does not appear to be conceivable, calculations indicate that: if it is assumed that the average fuel temperature at the time of coolant loss is equivalent to the operating average fuel temperature, the maximum temperature after the coolant loss is not appreciably different (2% - 4% higher) from that calculated assuming 27'C fuel initially.

The af ter-heat removal will be accomplished by the flow of air through the core. To determine the flow through the core the buoyant forces were 8-16 l

equated to the friction, end, and acceleration losses in the channel as shown in the expression Apb = Ap7 + Apc + Api + Ap a *

(II)

The buoyant forces are given by Apb"P o - 9dt = pt-p L -S g

-pVj (12) o oo ,

~

where p g, ,p 3

= the entrance, mean, and exit fluid densities, respec-tively, L = the effective length of the channel (= L +Lf + L') ,

Lg , L g, L' = the length of the channel adjacent to the bottom end reflector, fuel, and top end reflector plus ten channel

, hydraulic diameters, respectively.

The friction losses in the flow channel are given by L 2 Ap g = f * (

F D 2 i e 2ge A f f where the summation is over the lower unheated length, the heated length, and the upper unheated length, f

p

= the friction factor (= 23.46/R )(Ref. 5 ),

D = the hydraulic diameter (= 0.0601 ft),

A = the flow are.i through the core per element (= 0.0058 ft ),

- g = 4. i 7 x 10 ft/hr .

8-17

The sum of the exit and inlet losses, using appropriate expansion and

, contraction coefficients, is given by

~ 2 ope +OE i

" ( K) 2 2go A n

2' vith IK = k = 1.57 .

where k is appropriate expansion or contraction coefficient from regions of area A .

The acceleration losses are given by

• AP .

(14) a 2 1 0 gA c

By substituting the appropriate expression in Eq. 11, using the de'fi-nition of the Reynolds nuet+r, and L = 2.40 ft, L, = 0.29 ft, fL = 1.25 ft, and L' = 0.87 ft, one obtains

\

10 w + 0.665 E + 0.153 (0.700 _ 0.149)

P l P 0 + (0.153 P O N 0 1/

(15)

-2 x 10 w+ 1.25 p + 0.889 py - 2.139 9 =0 ,

0 with the flow w in units of Ib/hr and p the viscosity in units of lb/hr-ft, l

i l

l 8-18

~

The properties of air for use in Eq. 15 are expressed as pf = 40/Tg (Ib/ft3)

. and ug= 5.739 x 10-3 + 7.601 x 10-5 T1 - 1.278 x 10" T g -(16)

(Ib/hr-ft) , ,

where T is the appropriate teraperature in 1

'R.

The heat transfer coef ficient was calculated through the relationship N = 6.3 {R 51000I u \a j

= 0.806 R a

• R >1000\ ,

\a j e

where N = the Nusselt number = hD /k, e

4 p2gSATc /pkL, R, = the Rayleigh number = D e p b = the heat transfer coefficient, Btu /hr-ft *F, k = the thermal conductivity of the laminar film, Btu /hr-ft *F, S = the volumetric expansion coefficient, *F~ ,

AT = the temperature rise over the channel length, L('F),

c = the specific heat of air (Btu /lb *F).

p The expression for the Nusselt number was derived from the work of Sparrow, Loeffler, and llabbard (Ref. 6 ) for laminar flow between triangular arrays of heated cylinders.

8-19 l

The enermal conductivity and specific heat are given by k - 2.377 x 10' + 2.995 x 10 -5 T - 4.738 x 10 ~9 T 2

(Btu /hr-ft *F)

~

• (17)

-8 and c p

= 2.'e13 x 10 ' - 1.780 x 10-6 T + 1.018 x 10 T (Stu/lb *F), '

where T is the appropriate temperature in *R.

These two expressions, as well as that given for the dynamic viscosity of air in Eq. 16, are least-square fits to the data presented by Etherington (Ref. 7 ).

TAC 2D (Ref. 8 ), a two-dimensional transient-heat transport computer code developed by Gulf Energy & E'nvironmental Systems, was used for calcu-lating the system temperatures after the loss of tank water. The parameters derived above were programmed into the calculations.

The maximum temperatures reached by the fuel are plotted as a function of operating power density in Fig. 8-1 for several cooling or delay times between reactor shutdown and loss of coolant from the core. For reactor operation with maxtmum power density of less than 22 kW/ element, loss of coolant water immediately upon reactor shutdown would not cause the maximum fuel temperature to exceed 900*C. Operation at maximum power densities greater than 22 kW/ element will not result in fuel temperatures above 900*C, if the coolant loss occurs sometime after shutdown, or if emergency cooling I

is provided. (The time required between shutdown and the beginning of air cooling depends on power density.)

In Fig. 8-3, the data presented in Fig. 8-1 were replotted to show the time required for natural convective water cooling or emer;2ncy cooling, after reactor shutdown, to produce temperatures no greater than a given value. Thus, for example, for a reactor in which the maximum operating power 8-20

4 density is 10 kW/ element, there must be an interval of at Icast 1.3 x 10 sec (or 3.6 hr) between reactor shutdown and either the loss of tank water from the core or the cessation of emergency cooling.

8.2.4. Radiation Levels Even though the possibility of the loss of shielding water is believed to be exceedingly remote, a calculation has been performed to evaluate the radiological hazard associated with this type of accident (see Table 8-6).

Assuming that the reactor has been operating for a long period of time at 250 kW prior to losing all of the shielding water, the radiation dose lates at two different locations are listed below. The first location (direct radiation) is 18 f t above the unshicided reactor core, at the top of the reactor tank. The second is at the top of the reactor shield; this loca-tion is shielded from direct radiation but is subject to scattered radia-

. tion from a thick concrete ceiling 9 ft above the top of the reactor shicid.

The assumption that there is a thick concrete ceiling maximizes the reflected radiation dose. Normal roof structures would give considerably less back-scattering. Time is measured from the conclusion of a 250 kW operation.

Dose rates assume no water in the tank.

The above data show that if an individual does not expose himself directly to the core he could work for approximately 16 hours1.851852e-4 days <br />0.00444 hours <br />2.645503e-5 weeks <br />6.088e-6 months <br /> at the top of the shield tank I day after shutdown without receiving a dose in excess of that permitted by AEC regulations for a calendar quarter.

TABLE 8-5 CALCUlm ED RADIATION DOSE RATES l FOR LOSS OF REACTOR POOL WATER

[

Direct Scattered Radiation Radiation

. Time (r/hr) (r/hr) 3 10 sec 2.5 x 10 0.65 l 9 l . 1 day 3.0 x 10" 0.075 1 week 1.3 x 10 0.035 3

1 month 3.5 x 10 0.01 8-21

For persons outside the building, the radiation from the unshielded o core would be collimated upward by the shield structure and, therefore, 4

would not give rise to a public hazard.

O 8.3. -REACTIVITY ACCIDENT The rapid insertion of the total excess reactivity in the reactor system is postulated. The' method of inserting this reactivity is through the rapid removal of a control rod or experiment. This reactivity insertion is the most serious that could occur. It is also the normal pulsing condi-tion and the analysis is presented here as a point of information since it is not actually an accident condition.

The sequence of events leading to the postulated reactivity accident is:

1. The reactor is just critical at a zero power level, t

2. Upward force is applied to a high worth control rod or experi-ment causing it to be ejected from the core and to introduce the total excess reactivity of the core; i.e., $3.00.

The consequences of the above sequence of events are:

1. An increase in reactor power to a maximum power of approximately

1200 MW.

2. A maximum energy release of approximately 16 MW-see when the maximum fuel temperature of 539"C is reached.

h

3. Stresses in the stainless steel cladding of approximately 1650 psi.

l These pressures are caused by expansion of the air and fission 8-22 l

l I , - _ - , . ,. - , . ..

_. .. = . . _ _ . -. _ _ - - - __ . _

T' 4

, product gases and the hydrogen release from the fuel material.

Neither of the preceding stress values will-cause cladding rupture.

l The analysis of this accident is conservative in a number of ways, some of which have been indicated in the reactor design bases (Section 3).

For example, the equilibrium pressure of hydrogen over the fuel is not achieved during a pulse or step insertion of reactivity.

Analysis i It was assumed that the reactor is just critical at zero power level with a fuel and coolant temperature of 20*C. Additional input parameters are summarized in Table'8.5.

Calculations of reactor transient conditions were performed with the '

PULSE computer code using the preceding initial conditions and input para-meters. PULSE is a reactor kinetics code based on the Fuchs-Nordheim-Scalletar model and developed by General Atomic.

TABLE 8-6 REACTIVITY TRANSIENT INPUT PARAMETERS Reactivity insertion, $ 3.0 Prompt fuel temperature coefficient, f -1.1 x 10~

8, % 0.70 t,psec 43

• '~""

Cp (fuel)""lement e 817 + 1.6 T fuel

'~"*"

, Cp (water)""lement c 879 1

i Thermal resistances, "C, MW:

5.29 x 10 0 I

, Fuel to cooling channel Coolant to pool 1.42 x 10 3 8-23

-- m . ~ . . ~ .m.--.. - ,

t The PULSE calculations indicate that the average fuel temperatures in the U-ErH1 .65 e re w uld he 297'C. This temperature would occur at )

l approximately 1.2 seconds af ter initiation af the traasient. The peak-to-average power ratio used in these calculations was 2.21. Using this peak.

to average power ratio and considering the energy release during the tran-sient coupled with the volumetric heat content of the fuel, the maximum fuel temperature was obtained on the average temperature computed by PULSE.

~

This maximum temperature was 539*C. The reactor power level af ter the tran-j .sient with no control rod insertion was calculated to be less than 1 MW.

I It has been shown in the reactor design bases that this power level poses no a safety problems to the core. Of course, reactor shutdown would be initiated immediately by both power level and period trips or by manual scram and would be achieved even with the most reactive rod stuck out of the core.

During the time of peak fuel temperature the stress on the clad from the pressure prcduced by the expansion of air and fission product gases and the hydrogen released from the fuel is less than the strength of the clad material and therefore there is no loss of clad integrity.

4 ..

Calc'ulation of the fission product gases in a fuel element of the highest power density gives a total of 3.1 x 1021 atoms of stable and radio-active gases produced for continuous operation at 250 kW for four years. If the release fraction is taken as 1.5 x 10 -5 as discussed in Section 8.1, then I i

' 21 3.1 x 10 N

gp

=

23 (1.5 x 10-5) = 7.7 x 10-8 moles.

6.02 x 10 l

i The partial pressure exerted by fission product gases is i

i P

fp

= 1. , bI = 7.7 x 10~ EI 1p V V i

i e

8-24

, l

! i I

E _- _ __ _ . _. . _ _ ._ _ . --. _ _- a

where initially the volume V is taken as a 1/8-in. space between the fuel

. and reflector end piece. This is conservative since the graphite reflector pieces have a porosity of 207 The volume then is 3

V = wr h = w(1.80) 0.317 cm = 3.23 cm ,

From this, one obtains

-8 P = 2.40 x 10 RT .

f The partial pressure of the air in the element is P

air 2 4 x 0 = 4.46 x 10 RT .

The total pressure exerted by the air and fission products is e

-5 P; = (4.46 + 0.002) x 10 RT = 4.46 x 10 -5 RT = P,g .

Also we have P = 14.7 psi .

As an upper limit, assuming an air temperature of 539'c or 812*K (equal to the peak fuel temperature), one obtains I

Pj = (14. 7) = 44 psi .

8-25

. . .. - = - . .. . . .. . . . - -

i

! I i

The equilibrium hydrogen pressure over ZrH at 39'C is negligible.

1.65

. The total internal pressure then is-P = 44 psi-t" h+ .

1 Assuming' no expansica of the clad, the stress produced in the clad by this pressure is <

S= P P = 36.75 P = (36.75) (44) = 1620 psi

=f*02 t t

For a reactivity insertion of $3.00, the clad surface temperature would

~

be approximately equal to the saturation temperature of the water which is'

{'

113*C at a pressure of 23.4 psia. ,At this temperature, the ultimate tensile strength for type 304 stainless steel is.approximately 70,000 psi. Compar-

. . ing this~ strength with the stress applied to the cladding during the

. reactivity insertion, it is seen that the strength of the material far

) exceeds the stress which would be produced. Therefore there would be no loss of clad integrity or damage to the fuel as a result of the reactivity.

1 i accident. - -- ,

4 J

2 1

l

+

5-26 I

I

Chapter 8 References

1. Foushee, F.C. , and R.H. Peters, " Summary of TRIGA Fuel Fission Product Release Experiments," Gulf Energy & Environmental Systems Report Gulf-

. EES-A10801, 1971 p. 3.

2. Lee, E., R.J. Mack, and D.B. Sedgeley, "CADOSE and DOSET - Programs to Calculate Environmental Consequences of Radioactivity Release,"

Gulf General Atomic Report GA-6511 (Rev.),1969.

3. Simnad, M.T. , and J.B. Dee, " Equilibrium Dissociation Pressures and Performance of Pulsed U-ZrH l'ucls at Elevated Temperatures," Gulf General Atomic Report GA-8129,1967.

4. Foushee, op. cit.

5. Sparrow, E.M., and A.L. Loeffler, " Longitudinal Laminar Flow Between Cylinders Arranged in a Regular Array, AICLEJ. 5, No. 3, 323 (1959).

6. Sparrow, E.M. , A.L. Loef fler, Jr. , and H. A. Hubbard, " Heat Transfer to Longitudinal Laminar Flow Between Cylinders," Trans. ASME J. of Heat Transfer, Nov. 1961, p. 415.

. 7. Etherfngton, H. (ed.), Nuclear Engineering Handbook, 1st ed., McGraw-11111 Book Co. , New York 1958, p. 9-1.

8. Peterson, J. F., " TAC 2D, A General Purpose 3 Two-Dimensional Heat-Transfer Computer Code - User's Manual," Gulf General Atomic Report GA-8869, 1969.

8-27 f

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