ML20199B117

Ro:On 860501,improper Handling of Experiment Discovered Which Could Have Resulted in Serious Operational Event. Caused by Insertion of $1.09 Reactivity.Postulated Accidents & Results & Related Info Encl

ML20199B117
Person / Time
Site:	05000128
Issue date:	06/04/1986
From:	Feltz D TEXAS A&M UNIV., COLLEGE STATION, TX
To:	Johnson E NRC OFFICE OF INSPECTION & ENFORCEMENT (IE REGION IV)
References
NUDOCS 8606170035
Download: ML20199B117 (41)

Text

.

\

(

8 TEXAS ENGINEERING EXPERIMENT STATION THE TEXAS A&M UNIVERSITY SYSTEM COLLEGE STATION, TEXAS 77843-3575 #

1 WesL J 4 June 1986 NUCLEAR SCIENCE CENTER 409/845-7551 Mr. Eric Johnson, Director g)g v

@ 3 DW/2 Division of Reactor Safety and Projects Region IV

( 3 gg U.S. Nuclear Regulatory Commission 8 611 Ryan Plaza Drive, Suite 1000 Arlington, Texas 76012 RE: NSCR Reportable Occurrence of 1 May 1986 and USNRC Region IV Enforcement Conference, 21 May 1986

Dear Mr. Johnson:

Information is provided in regard to concerns that the incident involving the insertion of $1.09 reactivity due to improper handling of an experiment at the NSCR could have resulted in a serious operational event had the sample been of larger reactivity. This was obviously of concern to several USNRC represen-tatives participating in the enforcement conference of 21 May 1986. To aid in evaluations of the severity of the insertion of large reactivity at both low -

power and high power operation, the results of postulated accidents taken from the Safety Analysis Reports for the Torrey Pines TRIGA Mark III Reactor, the Puerto Rico Nuclear Center TRIGA-FLIP Reactor (PRNC), and the Texas A&M Nuclear Science Center Reactor (NSRC) are provided. These studies demonstrate that the TRIGA fuel design has self limitina safety features for incidents involving large reactivity insertions. In all cases the safety lioit of TR:GA FLIP fuel was not in danger of being exceeded. The studies of the Torrey Pines and the PRNC reactors are conservative when compared to the NSCR as these are 2 megawatt reactcrs as the NSCR is licensed for 1 megawatt operation. The 2.2 meg wett power trips assumed in the studies would have a de of 1.15 megawatts for the NSCR, thus, producing an earlier scram signal. The sample remosal time for the experiment involved in the ircidents at the NSCR was measured and fcund to be approximately 1 second. Thus, the 0.3 second removal time cf the Torrey Pines study is very conservative. The NSCR limiting reactivity insertions to prcduce maximum core temperatures of 950 C is also a conservative approach. The safety limit temperature of 1150 C for TRIGA FLIP fuel provides a safety margin of 200 C in the NSCR study.

8606170035 860604 PDR ADOCK 05000128 S PDR L

RESEARCH AND DEVELOPMENT FOR MANKIND l

e ,

f Mr. Eric Johnson Page 2 The following is a brief description of the postulated accidents and the results of each study:

PRNC TRIGA-FLIP Reactor (2 MW)

Assumes a transient rod ejection accident of $1.93 worth from 2.2 N operation followed by a fuel temperature scram. (Note: High power trips have been dis-abled by operator error in switching to the pulse mode), thus, a much worse case than the NSCR incident at 550 Kw).

Results: Maximum Power 950 N Energy Release 14.4 N-sec (within 0.15 sec.)

Maximum Fuel Temp. 860'C Maximum Clad Temp. 850'C No fuel damage Torrey Pines TRIGA Mark III Reactor (2 MW)

Assumes operator error in the removal of a $1.71 worth sample at low power (1 watt). Removal time was 0.3 seconds and the reactor scrams at 2.2 N. A scram signal delay of 5 msec occurs before rods are inserted of $13 reactivity worth.

Time to insert the rods is assumed to be 1 second.

Results: Maximum Power 2,000 N Energy Release 25.2 N-sec Maximum Fuel Temp. 1005 C Maximum Clad Temp. 67*C No fuel damage Note: (This is a conservative analysis of the NSCR low power incident).

NSCR (1 MW)

Assumes accidental pulsing from 300 watts and 1 Ms resulting in maximum fuel temperature of 950 C. Beginning of fuel life and end of fuel life cases are examined. Other assumptions; pulse insertion time of 15 msec, scram delay time of 15 msec and rod drop time of 0.985 sec.

g - -

3 I

Mr. Eric Johnson Page 3 NSCR (1 MW) cont'd Results: (Beginning of life which applies to the NSCR incidents)

Insertion from 300w to produce 950 C fuel temp. $2.36 Insertion from 1 Mw to produce 950*C fuel temp. 53.65 Note: 950 C is the rnaximum core fuel temperature protected by the limiting safety system temperature scram. The maximum fuel temperature safety limit is 1150 C for FLIP fuel. Thus, a 200*C safety margin is assumed in the NSCR accident analysis.

Na fuel damage.

The NSC staff is presently making additional evaluations of the self limiting safety features of the NSCR. One evaluation underway is to determine the value of reactivity insertion required to produce the safety limit temperature of 1150 C. It has been shown that to produce maximum ~ fuel temperatures of 950 C requires insertions of $3.35 from power and $2.35 frora just critical. Both values are in excess of the maximum allowed single experiment reactivity of

$2.00. Thus, these are worse cases than the operational errors reported. In all cases these evaluations involved steady state operation of the reactor.

The results of additional evaluations will be provided as soon as they are com-pleted. I hope these accident evaluttions will be of help in your review. If you have questions please contact me.

Sincerely, LJ Donald E. Feltz Director DEF/ym Enclosures cc: R. Martin, USNRC, Region IV J. Gagliardo, USNRC, Region IV L. Constable, USNRC, Region IV D. Hunter, USNRC, Region IV D. Powers, USNRC, Region IV D. Tondi, USNRC, Washington H. Richardson, Texas A&M University C. Erdman, Texas A&M University K. Peddicord, fexas A&M University F. Jennings, Texas A&M University

rl PUERTO RICO NUCLEAR CENTER TRIGA-FLIP REACTOR

.6 SAFETY ANALYSIS REPORT

. 4, 6 ' REACTIV!TY ACCIDENT -

Several reactivity accidents have been analyzed. They include:

1. The sudden addition of a fuel element bundic to the critical core

2. The run-out of all the control rods

3. The ejection of the transient rod when the reactor is operating -

in the steady-r. tate mode.

Of the three classes of accident, the ejection of the transient rod during

, operation at power would have the greatest consequences and will there-fore be discussed first.

4, 6.1 TRANSIENT ROD EJECTION ACCIDENT The transient rod ejection accident considered here is the most se rious that could occur with any reasonable degree of probability.

The sequence of events and condition of the reactor leading to this acci-dont are:

1. The reactor is in the steady-state mode improperly operating at 110% of full powe , 2.2 MW, juet below the point at which

') the two powe r level scrams would bc initiated.

2. .The transient rod is fully inserted, but the cylinder-dash pot

~

is in its fully out position.

3. The operator turns the mode switch to " Pulse".

l 4. The interlock that prevents energizing the pulse circuit when the reactor powe r is above 1 kW fails.

5, 7..m ope rator fire s the pulse rod.

6. Wl,en the fuel temperature at the thermocouple locations reaches 800"C the reactor is scrammed.

I The accident, then, presupposes two independent f;.ilures in the system: first, the failure of the lkW interlock, and, second, the f ailure of the operator in pulsing the reactor when it ic operating in the steady-state mode. The operator als'o was in error in operating at 2. 2 MW as well' as using poor technique in not using the transient rod as a safety rod in the s teady-sta.te mode. This last point, however, is not a necessary requirement to operating in the sicEdy-state mode.

4-6-1

n. ,- , , - , - - n ~~- - ,,,n - , - - . - - -

7 ,

g l

.r The consequences of the above described sequence of events are:

1. ' An increase in reactor power from 2. 2 MW before pulse initia-

~

tion to a maximum power of slightly over 950 MW (see Fig.

4, 6.1. ) .

. 2. An energy release within 0.15 seconds from pulse initiation of 14. 4 MW-sec at which time the maximum fuel temperature occurring at the periphery of an element next to a water filled control rod channel is 675"C.

. 3. A further energy release at a rate decreasing from 9. 5 MW to

3. 3 MW at 10. 5 seconds after pulse initiation at which time the temperature at the position of the thermocouple in the in-strumented element will have reached 800 C, initiating a scram.

4. A stress of 4300 psi in the clad caqsed by the expansion of the air and fission product gases and the release of hydrogen from the fuel material.

5. As the strength of the clad at the temperature of the clad (850 C) is greater than 10,000 psi, the integrity of the fuel will be maintained.

.

• There were several conservative factors built into this analysis,

-[] the most important of which are:

, 1. The one-dimensional reactor kinetics calculations assumed that the heat removal continued at about the same rate after the pulse as before. As a large fraction o'f the core would be i

in the film boiling regime after the transient, the average fuel temperature would increase faster than predicted in this cal-culation and therefore the scram would occur sooner.

2. It was assumed that the thermocouple element was in the posi-l tion next to the water channel left by a control rod. Although this is the position where the highest power densities occur in a pulse, the highest fuel temperature at the the rmocouple

, position would occur in the element in which the average power density is highest, and this may be a different element. If this is the case, the temperature scram signal would occur earlier than 10. 5 seconds after the pulse.

The analysis was based on beginning-of-life conditions and on 800 C fuel temperature scram set-point. The conclusion is that the clad integrity is maintained during the postulated accident. At the end-of-life, however, the negative temperature coefficient is smaller and the fuel and clad

. temperatures rc sulting from the postulated accidern would be more severe.

w.

Changed 11/25/69 4-6-2 -

i i

e -

s T he refo r e, the set point for the fuel temperature scram should' be

()#

reduced to 600 C to allow for these lifetime changes while remaining within the accident boundaries already investigated.

1. 6. 2 ANALYSIS The steady state conditions at 2. 2 MW were calculated using the computer code STAT developed by Gulf General Atomic and described in Section 2. 2. 3 Fuel and water temperature changes with power density 7 ..

9 j 4_ (,- Za I I/25/6<J

i l'

., ii '

O .

1000 P

'O v

~

,. p 100 -

, I -

v .."

,,,,- . ,s e.

W O

=

~

_n. - .

e .,

o .

o

<r '

w e 10 - )

~

f5kb _

+

t 1!

! s i

i i i iI i j_ i i i l i i i i l

' 0.01 0.1 3 1.0 10 TIME AFTER' PULSE INITIAT10N (SEC)  ?

p

/* *%',

t

's.

. n f

I s __

6 t'

i l

. Fig. 4. 6.1--Power afte r a maxinium pulse f rom 2. 2 MW

u. Y t

4-6-3 v

s ,

/ ,

r

-,a. _t

..t ,-._._,.___,-.__.._.___;x...,, .- _ ,

were investigated and the results are plotted in Figs 4. 6. 2 and 4. 6. 3.

Fig. 4. 6. 2 shows the radially averaged and peripheral fuel temperature

@ as'a function of local heat flux to the coolant. The fuel temperature at the periphe ry decreases with time as the gap between fuel and clad closes.

Fig. 4. 6,3 shows the average bulk water temperature in a cooling channel as a function of the average power density in the fuel adjacent to the cool-ing cliannel. At 2. 2 MW the average fuel temperature in the core is found in Fig 4. 6. 2 to be 314 C as the aveiage heat flux is 2. 22 x 105 btu /hr ft2 ,

From Fig 4. 6. 3 we find the average bulk water temperature to be 66 C as the average core power density is 23. 2 kW/ element.

These steady-state temperatures and power density values were then used as initial conditions for a combined heat transfer reactor kinetics calculation using a computer code BLOOST developed by Gulf Gene ral Atomic. Additional input parameters were:

1. The reactivity insertion in the pulse .

6k/k = 1. 35% in 0. I second

2. The prompt fuel temperature coefficient

-5 -

a = -(2. 65 x 10 + 2.10 x 10-7 T fuci,) (oC) 1 3),n

'. 3. The delayed neutron fraction p = 0. 70%

f

4. The prompt neutron lifetime 4 = 16 u sec

5. The specific heat of the fuel and water F -

c p = (772. 5 + 1. 516 Tfuel) (watt-sec/ element) i c

W = 762. 9 (watt-sec/ clement)

P

,i The transfer of heat from the fuel to the coolant channel and from the coolant channel to the pool water was approximated by the use of the data calculated to e stimate the the rmal resistances required as input to i

B LOOST. These thermal resistances were:

Changed 7/28/69 4-6-4

, - _ - ,, --- ~

- 4:

g. -g ~ "

• s da

's. 600 c_

500 -

w ~

^ 400 - --

v ,f W

s ' ' ,

RADIAL AVERAGE FUEL TEMP h R 300 L__ ,

, -c w -

c_

t.

z W

200 -__/ "

X-

. FUEL TEMP AT EDGE 7

. 100 9

~

y

( -

R -' . t n t U ,. ,,,,

0 I I

0 1 2 3 4 5 q/A X.'10 -0 BTU /HR/FT2 ,f.y 3

, , ,; ,,gg e> . , ,

>. c 5'- t .

Fig. 4. 6. 2--Fuel element tempe ratu e vs local heat flux, PRNC 4-rod, 9 5 elements ' ^

~

i (e2Q .

h .

.80 70 -

60 - PRNC 95 ELEMENTS: 4-ROD.

U .TkN = 38 C L 50 -

w A "

S 40 e- g a.

5 r

30 -

20 -

10 -

0 ' ' '

I, O 10 20 -30 40 50 l-P/N AVERAGE PCWER. DENSITY IN ELEMENT (KW/ELEr)

Fig. 4. 6.3- Average water channel temperature vs average powe r density in element

r _

From the fuel to the cooling channel 1. 0 x 10 C/MW O end frem the coolant ie the geel 1. 4z x 10 3

b/Mw The fuel thermal resistance as calculated is predicated on sub-cooled boiling in the channel. As will be shown subsequently, however, shortly after the conclusion of the pulse, film boiling begins, and this resistance increases drastically (by an order of magnitude). As this inability to remove the after pulse heat efficiently is not accounted for in the BLOOST calculation, the tempe ratures calculated are lower and the reactor powe r level decline is slower.

In Fig 4. 6.1 is shown the reactor powe r as a function of time af ter the initiation of the pulse. Within 0. I seconds the reactor power reaches a peak value of 950 MW and by O.15 seconds the prompt release of 14. 4 MW-sec had concluded. The power level from that point in time decreases for the next 3. 5 seconds from 9. 5 MW to about 3. 3 MW at which level it remains until 15 seconds after the pulse. The transient rod would be automatically dropped no later than 15 seconds after the pulse began if nothing else caused the rods to drop. As we will show, a tempe rature scram will occur before this timed rod drop so that 15 seconds is the latest time of interest.

The reactor energy release in the pulse was used to dete rmine

~

the temperature distribution at the conclusion of the prompt burst in an element next to a water filled control rad position. Because the powe r peaking in the fuct cell next to the water is 30% greater than in a cell surrounded by fuel, the peak temperatures occur in such an element.

From two-dimensional diffusion theory calculations it was found that the average power density in an element next to a control rod ' wate r channel is 1.4 times g reate r than tho average in the core. There is an axial peaking factor of 1. 3 in the element (it is assumed that the power is distributed as a chopped cosine) and there is a radial peaking factor of 2. 5. These factors, coupled with the tempe ra-ture distribution in an element during steady state operation and the energy release in the transient, made possible the calculation of the temperature di.stribution in the element next to the water channel at O.15 seconds after pulse initiation. In Fig 4. 6. 4 the tempe rature is plotted at the axial mi:1 plane of this element at t = 0 and t = 0.15 seconds.

This temperature distribution was used as input to the compute r code TAC, a two-dimensional transient heat transport code developed by Gulf General Atomic. The afte r pulse power density, Q, was explicitly described by l  !

4-6-7

• g:

= 'T;,_-

ge . I,qf s

.1<

h!

-800

.l.

--600

- AFTER PULSE ,

u _ N '

e r

L_  ?,fw w

w .

I .O 3 400 -

g,<, .

< N,

$. BEFORE PULSE

a. -

,c. r w

.'e;y - -

-K 200 _

i

i. u>)

3 l

I '

0 8 - 1 0 0.5- 1.0 1.5 '2.0 --,

FUEL RADIAL DIST. (CM)-

Fig. 4. 6.4- Fuc1 tempe rature in maximum power density element before and after a 1.287o pulse from 2. 2 MW 4-6-s j

T n

? 0.26c 4. 75(r/r o)

(d Q = 2. 04 x 10 7 "(t + 0.15) 0. 3034- cos (0. 78mc/L) 0.<t_'_S a - - - - -

0. 26 e 4.75 (r/ro )2 -

7

= 1. 25 x 10 cos(0, 78 nx/L)  ; 5 <t< 15.

Where Q is the local power density in Blu/hr ft t is the time after the end of the pronipt burst (sec) x is the axial distance from the center of the fuel element L ir< the length of the fuel meat r is the radial distance from the center of the fuel r is the radius of the meat o

The volumetric heat capacity and the the rmal conductivity of the fuel material we re i

3o (cpp)f = ( 3 7. 58 + 0. 24 5T) Etu/ft - F kg = (10. 4 2 - 6. 42 x 10~ T) Blu/hr ft- F .

The clad was assumed to be in contact with the fuel with no thermal re sistance between the two. The thermal properties of the stainless steel clad were taken to be 3o (c p) = 59. 4 Btu /ft - F p c o

k c

= 10 Hu/hr ft- F .

The properties of the coolant strearn were obtained from the work desc ribed in Sec . 2.2.3. As during the operation jor,t prior to the tran-s ie nt , the cooli'y took place through sub-cooled boiling, the heat flux

/

4-6-9

0-can be described as a function of the temperature difference between the

-h fuel element surface and the saturation temperature. Above the initial heat flux the film boiling heat transfer was taken from the work of Ellion.

The heat flux as a function of the wall-saturation temperature difference used to derive an explicit heat transfe r coefficient is shown in Fin 4. 6. 5 (the dashed curve was used to approximate Ellion's values). The mass flow rate was also derived from the steady state calculations and is given by.

W = 7. 9 ( eT - T.) /h r 3

where Te is the exit water temperature ( F) and Ti is the inlet water temperature (100 F).

The results of this calculation are plotted in Fig 4. 6. 6 and show the maximum fuel temperatu re, the maximum clad temperature and the temperature at the location of the ther mocouple in an instrumented cle-ment.

The time history in this original calculation extended to 15 seconds -

after the transient at which time there would have occurred a scram of the transient rod from the automatic timer.

()

• i The dashed portions of the curves in Fig 4. 6. 6 show the tempera-tures of interest at 15 seconds after the pulse: the fuel temperature at the hottest point is 955 C and the clad is not over 94S C. O From the curve showing the temperature at the location of the thermocouple it can be seen that the scram initiated by 800 C measured fuel ternperature will occur at 10. 5 ccconds after the transient. At this time the maximum fuel temperature would bc 860 C and the stainless steel cladding would be 8500 C, .The fuel temperature is well below the maximum temperature to which TRIGA fuel has been subjected in thou-sands of transients and is less than the peak fuel temperature in a trans-1 ient from low power (because of the temperature dependence of the prompt tempe rature coef ficient). It is only necessary to show that the stress on l the clad produced by the hot air, fission product gaser,, aild hydrogen released from the fuel is less than the strength of the clad.

i Calculation of the fission product inventory in the element with the highe st power dennity, i. c. , 1. 5 x 2000 kW/95 elements = 31. 6 kW/

clement, indicate a total of 6. 7 x 10-2 moles of stable gases produced

• Ellion, M . E. , "A Study of the Mechanirm of Poiling Heat Transfo r,"

JPL Memorandum No. 20-88, March 1, 1954.

I Changed 11/11/69 4-6-10 r w ~ w,-------m-em ---,,.,-v , , .~v - - - - -e . - - - - ---,,-e- ,.-,we sc,-

. ,m .

u. - g; .

6 10 FROM ELLION N (JPL MEMO #20-88, 3/1/54)

N

\

APPROX TO

^ \

u r 10 5 -

N ELLION /

w _ \ f N /

s 5 _ V m

v s 5 _ FROM STEADY b d STATE ANALYSIS L

~

~

I 10 4 -

a 10 3 10 100 "1000 10,000 Ty-TSAT ( F)

Fig. 4. 6. 5- Heat flux vs wall superheat used in transient heat removal calculation

t'

'J.. - 'h ,

1000 -

ll

//

!/ /

F

~

800 -

/ FUEL w

3 T/C t

5 m

k. I CLAD MAX FUEL 600

- TMERM0 COUPLE LOCATION MAX CLAD 4

400 ' ' ' '  ! ' ' ' '  ! ' ' ' '

O.1 1.0 10 100 TIME AFTER END OF PROMPT EXCURSION (SEC)

Fig. 4. 6. 6- Minimum fuel tempe rature, maximum 'clad tempe rature and temperature at thermocouple location vs time after reactivity accident excursion

r

, o in 8 MW-years of operation. The fission product gases that appear in

() the gap between the fuel and clad comprise the total number of moles

. produced times the release fraction ( see Se ction 4. 4. d) or 2 -4 -5 N = 6. 7 x 10 x 7 x 10 = 4. 7 x 10 mole s .

fP The partial pressure of these gases ic P = 4. 7 x 10~ RT fP V where R is the universal gas constant, and the free volume in the element is assumed to be only that equiva-lent to a 1/8-in. longitudinal space inside the element, that is 3

V :. tr(1. 73)2 2. 854 = 3. O cm .

._ ) From this we find Pg = 1. 6 x 10~ RT .

The pressure exerted by the air trapped inside the cladding is RT -5 P = = 4. 4 6 x 10 RT.

air 22. 4 x 103

/

The total pressure exerted by the air and fission products is

-5 P = (4. 46 + 1. 6 ) x 10 RT 1

= 6, 0 6 x 10~ RT = 1.36 P ,

air.

Also, we have T

P = 14.7 psi .

air 273 Chany,cd i 1/11/M 4-6.13

O O and for a temperature of 860 C, or 1133 g C/4 1133 P = 1. 36 x 14. 7 x 273 = 83 psi.

1 The equilibrium hydrogen pressure P yg oo over ZrlI I. 65, at 860 'C is 51 psi. The total internal pressure is then, P 'l, = PII + P1 = 51 + 8 3 = 134 psi, assuming no expansion of the clad to accommodate the entrapped gases.

The stress produced in the clad by this pressure is S=P

[rh l= 134i I 0.707\ '= 4750 psi.

Tl t (0.020j The yield strength of 304 stainicss steel at the clad temperature 850 C is over 10,000 psi. The re fo re, one can conclude that the accidental reactivity excursion postulated here would not result in damage to the fuel or loss of integrity of the fuel element cladding.

If one equates the yield stress of the stainless steel clad with the stress imposed by t,he internal gases one finds that the fuel-clad syt, tem must be at about 9 50 C before the clad would yield. In Fig 4.6. 6 it is seen that such tempe ratures are not reached untilabout 15 seconds after the exeyrsion at which time the tempe rature at the thermocouple position is almost 900 C, that is,,100 C over the scram set point. Also, we have assumed that the thermocouple was in the element next to the rod position where the local power density is highest. Iloweve r, as the thermocouple would actually be located in the position where the steady state temperature would be highe s t, that is, where the average power density in the element would be highe st, the temperature at the thermocouple location would be higher than if the thermocouple were located next to a water channel. The tempe rature scram, with the thermocouple element in the cente r position, would be expected to occur even earlier than 10.5 seconds after the pulse.

• Simnad, M. T. , and J. B. Dee, " Equilibrium Dissociation Pressures and Performance of Pulsed U-ZrII Fuels at Elevated Temperaturcs,"

General Dynamics, General Atomic Division Report GA-817.9, August 1967. I 7 p.

Changed 11/11/69 4-6-14

If the temperature scram set-point is reduced to 600 C, the scram is initiated approximately 2 sec after the prompt excursion compared with

10. 5 sec for the 800 C set-point. The o

earlier scram reduces the clad o

temperature about 250 C (from 850 C) for beginning-of-life conditions.

Analysis has shown that the maximum after-pulse fuel temperature increases approximately 150 C (to 825 C vs 675 C in Fig. 4-6-4) when the end-of-life temperature coefficient is used in the transient analysis.

Thus for end-of-life conditions, with a 600 C fuel temperature scram o o set-point, the peak clad tempez ature should be about 750 C (100 C le::s than the 850 C derived above for beginning-of-life conditions with an 800 C sc ram set-point. )

4 e

11/25/69 4-6-11a

r- -

.,* * :J

TORREY PINES TRIGA MARK III REACTOR qg SAFETY ANALYSIS REPORT ,

8. 2. REACTIVITY ACCIDENT 8.2.1. Summary The rapid insertion of a large amount of positive reactivity is postulated to occur through the rapid withdrawal of a maximum worth ~

experiment from the water flux trap nea rest the core center of flux. It is assumed that the force applied to withdraw the experiment is constant and the withdrawal is completed in O. 3 seconds. The reactor power before the accident is ve ry low (1 W) and the total reactivity insertion in the accident is 1. 2 % 6k/k. When the power level passes 2. 2 MW a scram signal is generated which results in the insertion (after 5 msec delay) of a total of

' 9.1% Sk/k in I sec. In summary, the sequence of events and the condition of the reactor before and during the accident are:

1. The reactor is initially at a power of 1 W and the end of core life temperature coefficient applies.

2. An experiment worth 1. 2% Sk/k is withdrawn from a flux trap in 0. 3 sec. (This is the largest experiment that may be placed in the core without being locked in place. The withdrawal time of 0. 3 see is physically realizeable through rnanual removal yet it is short enough that this is effectively a step insertion.)

3. When the reactor power reaches 2.2 MW a signal initiates a sc ram. (This has a much smaller effect on the termination of the transient than does the shut-down through the prompt tempe rature coefficient. )

The accident occurs as the result of a human failure in that changing the configuration of the core when the reactor is critical is a violation of good operating practice. Either of the two power level trips could fail with-out changing the results of the accident. If both power level scrams failed the temperature level scram would occur about 20 msec later than the powe r level scram.

The consequences of the accident are:

1. An increase in reactor power from 1 W b'efore the experiment withdrawal to a maximum of about 2000 MW(See Fig. 8-2).

2. An energy release within 0. 355 seconds of 25. 2 MW-sec at which time the maximum fuel temperature occurring at the periphe ry of an element next to the wate r filled channel is 1005"C (See Fig. 8-3) and the maximum clad temperature is ,

67*C. A stress of 10,000 psi is produced in the clad as a I result of the expansion of the air and fission product gases and the release of hydrogen from the fuel material. At this time l the strength of the clad 1.s 28,000 psi (See Fig. 8-4). This margin of 2.8 between the yield strength and the induced stress is the minimum that occurs and, therefore, the clad integrity is maintained. ) ,

8-9

y ~ . v

_ :y o

.. ; . e .

l{. Dh

. nG -

\-

t

.k --

10,000 1

~

I 1 1 1 I i -

bf

e. e ::.; p

j.- + < 4 ._ -*

l

u. * - ;t

3. w S "I 4 s: 4

. :1 y _ -

. L ', ,;

ia !

I000 -- ~~. .

If i _ -

+,e .

y

},

A g 100 - -

~~

f6' .

e w

- i,

?f- h PQ , iP' O.

9

Aq :

a - a; , e -

O

• s &

!k o

? < l0 -

R l. $

E Y Flz h J

&= -

)>e

.n

, j-._

ir l! T f ~; 1.0 - ~~

% p ~

9 d;: .-

@ 'I _

V / r,

. .,s in!

1t! - -

)

iv

, ' ,f 0.1 I I I I I I ~

0.3 0.4 0.5 0.6 0.7 0.8 0.9 l-g j[i ,

TIME AFTER BEGINNING OF RE ACTIVITY INSERTION (SEC)

% l '

- bi Fig. 8-2.

Powe r vs titiie for tlie insertion of 1. 2 % Ak/k l\. th,;

4

. . 6.: ,

8-10 if

i k d

J  !

z \

. 0 0

- 1 n

i .% '

i o

\ t r

i e

s n

i i

) y t

P P C i M M E v g .

E T

E T

S

(

i t

c a

L D N e i E A O r

, U L I r (

y ,

F C T R

E f e

t X X A S a _

M 0 N .,

M e i

_- %A  !

1 I

Y T

I i

t m

s i

V v I

T d

a C l A c E t i- -

R F

e m

n O e i

G l

e _

N l

, I e _

N u f

, N g

I d

, 0 G n

, g I . E a 1 B g, '

l e

R u -

g.

E f g ' T e F h A t x

- g ' E n i c 7 1 M

I s -

}

e

~. ) T r u

t a -

r i

' r

. e .

p

.g m -

e T _

1 O n 0 0 0 3 0 1 -

0 1 8 ,

c 0 .

1 . - .

g '

y i -

mvo- wear <MWoIww F -

7

=

7 -

y me_ -

c x

y,- =

,c . . :_..

>j c

3. A furthe r ene rgy release at a rate varying from 13 MW to '

O. 08 MW at about 3 sec af ter the beginning of the insertion at which time the maximum fuel temperature decreases tc about 750 C and the clad temperature reaches a maximum of 640 C.

8. 2. 2. Analysis The reactor power as a function of time after the initiation of the reactivity insertion was calculated using the combined heat transfer-reactn.

kinetics computer code BLOOST developed by Gulf General Atomic. The input paramete rs to the program were:

1. The reactivity insertion in the pulse 6k/k = [1. 2%] [(t/0. 3)2 ] ; t in seconds

2. The pre-insertion reactor power P o = 1 W in 109 fuel elements

3. The prompt fuel temperature coefficient (end of life)

-D er = {Z. 73 + 0. 012 5T) x 10 6 k/k"C

4. The delayed neutron fraction p = 0. 71 %

5. The prompt neutron lifetime f = 20 usec

6. The reactivity inse rtion in the scram

( 6k/k)s  :

E9 I%3 f(ts /1.0)2]

7. The specific heat of the fuel and wate r c pF =

(785 + 1. 6 IT) watt-sec/ C-elem cp w = 86 0 watt -sec /"C-elem The transfer of heat from the fuel to the coolant channel and from the coolant channel to the pool water was approximated by the use of the data calculated to estimate the thermal resistances required as input to B LOOST. These thermal resistances were:

From the fuel to the cooling channel 1. 0 x 104 C/MW and f rom the coolant to the pool 1. 42 x 10 3 C/MW The fuel the rmal resistance as calculated is predicated on subcooled boiling in the channel. Ilowe ve r, shortly afte r the conclusion of the pulse, film boiling begins, and this resistance increases drastically (by an order of magnitude). As this inability to remove the after pulse heat efficiently is not accounted for in the BLOOST calculation, the tempe rature s calcu-lated are lowe r and the reactor power level decline is slower.

8-12

. ;,.;~ ag:n.:p 100 ;- - 3 i i i i ,

1 i i , =

g i i

~

ULTIMATE STRENGTH y -

m YlELD STRENGTH x _

m 10 -

~ _

.a

?

m

~

~

{0 _

INDUCED STRESS c: _

CL

i e i_, t I i I

t. t i i r i i l 0.1 1.0 10 100 TIME AFTER BECINNING OF REACTIVITY JNSERTION (SEC)

F' ig . 8-4.

Stres s induced in clad by i. sternal gas pres sure and' clad strength vs time af ter reactivity inse rtion xgr; ~~a._;_.  :-=  %=r z -

m? mu _- w ._--w. _ w . _ , ,_._ w . w _ ,_ gn.-,,- ., ,_ .

7 , .-J

n .

N t.(./ ! I i

)

A If[Y h!

t.'T M: >

la Gr

,E[ *p h: ' q:i.; In Fig. 8-2 is shown the reactor power as a function of time after j p ~j

[ ih the in.;tiation of the pulse. Within 0 33 sec the reactor power reaches .

j, a peak value of 2000 MW and by O. 355 see the prompt i aler.se of 25.2 MW- '

] ;5 sec ha d cor.cludec'. The powe r level f rom that poir.t in time de;reasec 1, ;} ' for about 0. 08 wec to 8. 8 MW and then dec reases to abo it 0.03 MW at YI! 3 sec. The power level sc ram occur red at the 2.2 MW level vehich was y _{ at 0. 305 sec a fter the beginning of the ins ertion. T he rods dropped in 1 ec and at 1. 31 see were fully inserted.

jl The temporal power relationship shown in Fig. 3-Z weis approxi- -

mated by analytic expression to use as input to the two-diraensional hat

[;

F tr nsfer program TAC 2D devolaped by Gulf General Atomic. The pc.ver k! de..aity in the fuel cleraent adjacent to the wate r-filled channel f rom which * "

the experiment was rermved was gi,ren by f

- {} t i_ ji g

Q = pCA a P!t) x r M -

where Q is the local power density in I\TU/hr ft 3 7 y:,

U f j.

( p is th. ratio of the c.ve rage power density in the element uext to '

the te r channel to the ave rage pawe r censity in the core.

l

"' 3 -4 i C in a conve rsion factor with the units IWIC/hr ft ) / (MW/cora)

@' A is the axial power distribution in the elemont x

4 ,(, A is the r.viial pcwe r distribution in the element

, 1 >

- O and P(t) is the reactor power ac a function of time. ,

.s '

, t '

The values of the parameter used were [

.'j p = 2.25 j

.  ;,.f' l C = 2. ?..:6 x 10 (13TU/hr ft ) / (MW/ core)  ;

1. 25 cos (0,78ex/I,)

e di)

+,

A x

=

A = 0. 66 (0. 04 e

!# 0 & 1) -

LI 2 t G.331

"- and I'( t) = 2008 sech 0 < t . 0,351 '

-\,96 (6.25 x 10"3)

! = 1, 7 2 t 0. 3 51 < t < 0.48 .

a .

! -4.042  ;

j; = 0. ;8 t .t > 0.48

t i

}

', whe t 3  :< is the a>:ial dit.iance f rom the cente r of the fuel 8

4

' if L is the length of the fuel rnant

\* j; r is the rulial distance from the certer of *he fuel >

'It ,

8-14 ,{

.i;

.l e t .

(

L -

• b!

.'. - . L, E Il g

0

- [' li j r-ois the ridius of the fuel meat 'n M

and t is the time after t'.e beginning of the insertion (sec), .

d' i,

!' The volumetric . heat capacity and the thermal conductivity of the 1

fuel niaterial were h 1

( c p) = (3 7. 58 + 0. 7A ST) BTU /ft - F i

~4 k = (10. 42 - 6. 24 x 10 T) BTU /hr ft- F he f

where T is the local temperature in F. The clad was assumed to be ie f

comart with the fuel with a thermal resistance of 50D BTU /ft2 F betweer.

the two.

• The thermal properties of the stainless steel clad were taken

[

-i to be '

j

)-

( c. p)c P

a 59.4 BTU /ft - F 1

k =. 10 BTU /hr it. F. '

1 c '

The properties of the coolant stream were obtained from the work den ribed in Section 3.1. 3. Below the coolant temperature at which boiling j begins a constant her.t transfer coefficient of 100 Btu /hr ft3 F wa s assume 4 L In the subcooled boiling regime the heat transfer coefficient was derived from McAdams and expressed as a function of the tergerature difference .

! between the inel element surface and the saturation temperature. The film j boiliag heat trtnsfer was taken from the tvork of Ellion.t The heat flux a.m a . I fune.tlon (,f the wall.;aluration temperature difference used to derive an explicit ,

heet tiansfe r coeffident is shown in Fig. 8-5 (the dashed curve was used to ,'

l approximate Ellion's values). The mass flow rate was derived from steady . j state calcula(lons a,nd is given .r, W= [6.461 x 10 ~

- 2. 48 5 x 17

~

(To - T.)] (To - T.) lb/hr "I t 1 ,{

whe re T is the outlet tvater temperature, F 'O o

e and Ti is the inl4t water terapr rature, F, 'i .

! I.

i u i

SSea Section 3.1.1. O, t Ellion, M. E. , "A Study of the Mechanism of boiMrig Mat Transfer, '

.;PL Memorandum No. 20-88, Ma rch 1,1954. j

(.

8-15 i I (-

a.

.e

• • 3

,f i 2

1!jj; yt g.

$1;:l R N-

,3 ; .:,i !i bl 'i hII}j ' 10 6

6

,e 4

h{!

l1

!;;j .

FROM ELLION N (JPL MEMO #20-88, 3/1/5f+)

9l

9. g "i N

!E i'; APPROX TO L fr \ ELL 10N h a, 10 5 -

N f .

s u- -  % 7 e \ /

}- 1 N

1' s

. I;! o r _ V

.I m in i v

- ffI itj x

o .-

FROM STEADY j!~ d STATE ANALYSIS

~

l s Odl- $ 10' -

cv c ~

i:h s} }j <

R ,

y s .

, 50 i k 4 -

4 k -

3.

hj ' ' ' ' ' ' ' ' ' '

, 103 k

[ l;)l 10 100 1000 10,6j

[ ] T g -T SAT ( )

T

. n

-l, y

.t

a 1 -r Jf. .

- . Fig. 8-5. Ilcat flux vs wc.11 supe rheat used in transient heat rernoval calculation

- y. s [1 a

i t. '..

h 4, <$

'j. j . 8-16 .. I V

k i h3

_A hi

w a , 3 j'

3 .

jI i

tt '

S l!,

t The results of this calculation are plotted in Fig. 8-3 and show j the maximum fuel tempe rature and the maximum clad tempe rature. It can "

be seen that the maximum fuel tempe rature (1005 C) occurred as a result of the prompt ene rgy release in the transient at about 0. 33sec after the beginning of the pulse. The maximum clad temperature (640 C) did not occur until a considerably later time ( 3 sec) as it was necessary for heat .

b; to be transferred into the clad from the fuel across the contact resistance between the fuel-clad surfaces.

It is necessary, now, to show that the stress on the clad produced by the heated air and fission product gases and the hydrogen released from the fuel is less than the strength of the clad at all times after the transient. 1; It will be assumed that the temperature of the entrapped gases is equal to j the maximum fuel temperature and that the strength of the clad is character-istic of the strength at the maximum clad temperature.

Calculation of the fission product inventory in the element with the highe st powe r density, i.e., 1. 6 x 2000 kW/100 elements = 32 kW/elem, indicate a total of 4.8 x 10-2 moles of stable gases produced in that element in 8 MW years of operation. The fission product gases that appear in the gap between the fuel and clad comprise the total number of moles produced times the release fraction (see Section 8.1) or N 4. 8 x 10

-2 -5 -6 fp

= x 4. 6 x 10 = 2. 2 x 10 mole s .

i The partial pressure of these gases is 1

-6 RT P = 2. 2 x 10 fp V j

• a

'I whe re R is the universal gas constant, and the free volume in the element j is assumed to be the 1/8-in. space between the fuel and the top reflector and piece, that is V = n(1. 8 2) x 0. 317 = 3. 3 cm i

7:om this we find '

-6 -

P = 0. 67 x 10 RT.

fp 8-17 4 .

l

R"T! Oj; d M .T. .

• d4 1; i P *f . .'F 4 MW s p I

k)k  :

{'n; The pressure exerted by the air trapped inside the cladding is q'

j E RT -5

• gj- i P.

air

=

3

= 4. 46 x 10 RT. j hp }' 22. 4 x 10 ,

Ml I ,

p? The total pressure exerted by the air and fission products is -

i j'i _,

4 .

l Pg = (4. 46 + 0. 067) x 10 ~ RT .

(;q 4 >

' 5 2 = 4. 53 x 10 RT = 1.016 P ,

, air.

M >

Also, we have

. . [f

- 3+ ,

P, = 14.7 p si, j -! air 273

,7

.y 4 i,

m ,

and T T

-  ?! P = 1. 016 x 14. 7 = 14.93 psi, '

g3 4 1 273 273

., *g o

where T is the gas temperature in K.

$1 ]a{;

)

The equilibrium hydrogen pressure ove r Zril l. 6 is shown in Fig.

7 3.2 as a function of fuel temperature. If we assume that all of the gas in

.[ j the gap is at the maximum fuel temperature the total internal pressure in-(j - -

j side the fuel clad is P

T

= P Il (T f) + PI (T f)

{'

T

$ assuming no expansion of the clad to accommodate the entrapped gases. ,

. ;., The stress produced in the clad by this pressure is pl >

S=

0.735 l' = P 36.75 P h PT() 0.020 T

=

T.

n l

l The ultimate strength of 304 stainless steel is shown in Fig. 3 3 as

]i a function of steel tempe rature. In Fig. 84 there is plotted as a function

, j of time after the excursion the stress produced in the clad as a result of h the maximum fuel tempe rature and the strength of the clad determined by 3

! the maximum clad tempe rature, which tempe ratures are shown in Fig. 8-3. '

, j It can be seen that at no time does the imposed stress on the clad exceed the strength of the clad. Thus, it can be concluded that, as a result of the postulated reactivity excursion, the clad integrity would bc

.p!

< maintained and there would be no release of radioactive gases to the atmos-ps phere as a consequence of such an accident.

,s N 8-18

+m: y

h Y k-..--.----..-.-----..-- .fi

- , pr e -

9 It should be noted that kinetics calculation show that had the inse rtion occurred while the reactor was at full power the resulting maximum fuel tempe ratures would have been much lower.

4

7

. , l; -e NUCLEAR SCIENCE CENTER REACTOR SAFETY ANALYSIS REPORT G. Accidental Pulse at Full Power for Mixed and FLIP TRIGA Cores it is necessary to examine this situation in spite of the interlocks which will prevent this from happening. The calculations were performed by Cencral Atomic using the BLOOST 2 code assuming adiabatic processes.

The details of the calculations are presented in Appendix I. The experimental parameters and . core power distributions were supplied by Texas A&M. The results desired were the reactivigy insertion from power that would produce a peak core temperature of 950 C. For valid com-paricons the valueu for pulsing fcom 300 watts were also calculated.

Figure 11-5 shows the renult a obtained as a function of number of FLIP clements in the core at the beginning-of-life. The end-of-life case was calculated only for a full-FLIP core. As can be seen.for every N case, if the reactor is pulsed f rom 1 Mw considerably more reactivity is required to obtain 950"C. The re fo re , it is pulsing from low power that limits the amount of insertion and not the " accident" situation.

f O

Appendix I THE PULSING ACCIDENT IN MIXED AND FLIP CORES The scope of this study performed by General Atomic was defined as:

1) Determining the size of the pulse producing 950 C peak temperature as a function of the number of FLIP elements in the core and the burnup by assuming adiabatic processes, and

2) If the results of 1) do not allow pulses of an acceptable magnitude, to refine the calculations by including the effects of heat transfer which would give a more realistic assessment of the effects of the reactivity addition.

The results of the first part indicated that the second was not required.

Figure I shows the size of the pulse that would produce maximum temperatures of 950 C as a function of the number of FLIP elements in the core. TheThe results are shown for pulses from IMw steady state power and from 300w power.

curves represent beginning-of-life (BOL) conditions for the prompt neutron lifetime and the temperature coefficient. The end-of-life (EOL) values for g' j these parameters were used to calculate the two points for the full FLIP core.

Following is a summary of the process involved in acquiring this information:

The calculations were made using BLOOST 2. The input parameters that were common to all problems were:

No. of elements a 98 Delayed neutron fraction = 0.007 Fuel specific heat = 720.0 + 1.48 T (w-sec/ C-element)

Water specific heat = 860 (w-sec/ C-element)

Fuel thermal resistance = 10000 ( C/Mw/ element)

Coolant thermal resistance = 1175 ( C/Mw/ clement)

Initial average fuel temperature at IMw = 238 C Initial average coolant temperature at IMw - 4 5 C Pulse insertion time = 100 msec

)

Scram delay time = 15 msec Rod drop time = 0.985 sec For the pulses from 300w it was assumed that the system had an initial temperature of 25 C and that there was no scram.

I' 3.80 -

O 3.60 -

'1

%Ok 3.40 -

m.

3.20 -

1 3.00 -

O PULSE FROM i Mw

• 2.80 -

~

- 2.60 -

O

<t w

• EOL (8.'d. Mw-Yrs.

~

\ N  ; BOL .:

2.20 -

EOL (8.2 Mw-Yr.

o 200 - -

1.80 -

I I I I i i t 30 40 50 60 70 80 90 100 N O. OF FLIP ELEMENTS IN 98 ELEMENT CORE FIGURE I PULSE TO PRODUCE 950*C PEAK TEMPERATURE

_ 'w

+o

'() In Figure 2 there is shovn the transient rod integral worth as measured.

The " ramp" table was constructed by assuming this rod was uniformly accel-erated from its initial position to 605 units (i.e., $3.25). The initial position was chosen to provide the desired worth of the pulse. The $3.25 position for the upper end of the insertion was chosen simply to sharpen the pulse as the integral worth curve flattens out drastically above that point. Using the full out position as the final point would tend to clip the pulse although that is a trivial consideration here. The ramp as a function of time is shown in Figure 3 for $2.00, $2.25, and $2.75 pulse.

These curves are plots of the ramp table.

The scram table was constructed by assuming that the total excess available was $7.00, that $3.50 was held down by the 1 Mw temperature, that the total worth i f the rods availabic for the scram was $12.00, that the shape of the rod worth is represented by the General Atomic " standard" shape, and that the rods fall with uniform acceleration. A plot of the scram table is given in Figure 4.

4 Three different core configurations were used to study the effect of adding FLIP fuel. All assumed the FLIP fuel was in the central region surrounded by standard fuel (if any). All the cores consisted of 98 fuel elements.

The values of the several parameters that are dependent on configurations were estimated by interpolation between data points already acquired.

Peaking factors for the power distribution were:

35 FLIP clements: axial = 1.36; radial = 2.00; cell = 1.95 59 FLIP clements: axial = 1.36; radial = 1.91; cell = 1.95 98 FLlP elements: axial = 1.36; radial = 1.56; cell = 1.95 Prompt neutron lifetimes were estimated from the following calculated data:

BOL - 18 FLIP /54 standard t = 28.0 usec BOL - All FLIP t = 17. 5 u s ec l

' L = 21.0 p sec BOL - All FLIP Linear interpolation based on the fraction of FLIP elements in the core gives j for the cores studied:

35 FLIP (35.7% FLIP) BOL L = 26.5 usec 59 FLIP (60.2% FLIP) BOL L = 23.1 usec 98 FLIP (100% FLIP) BOL L = 17.5 usec

! 98 FLIP (100% FLIP) EOL L = 21.0 usec Figure 5 shows the calculated prompt negative temperature coefficients for the 18 rod FLIP at BOL and the full FLIP at BOL and EOL. The BOL coefficient for the mixed cores between 18 rod (which corresponds to 25% FLIP) and full FLIP was estimated by making a linear interpolation between the two curves In Figure with the ratio of FLIP to total as the porportionality constant.

6 the integral temperature coefficients, as used in BLOOST 2, are plotted for the four core configurations considered.

-,--.,-----r- - . . . , _ . , , _ _ . ......-,.% . . , - _ . , --c., - - , - - - , - - . - - ., .---.._--_-.,7 m

g - _

O e

f

=

0 .

B

• s l

N O

l

~ q$

l '

> - O O Z.

h I--

x

' O 3

0

- O e J 4

D E H O y

'}h H

o z

+

9 o O t- x Om M o.

z W

, O O OO 5

7

! M E q W

1--

- O

- O N

. .- N O &

,' = 0 D

e 4

i I l' O O O ' O  !

O O O

$ 'H1HO5\ 00U 7VWD31NI

O .

2.O -

I .5 -

s [/

s'

/

b //

~

s' /,

Q x l .O ,'

/ <7'  % 2.75 PULSE

//

f' f

%2 .25 PULSE O.5 - e',' --

%2 .00 PULSE s'

,V,/

./

'O ' '

O .02 .04 .06 .08 .I TIME, SECONDS FIGURE 3 TIME DEPENDENT PULSE REACTIVITY INSERTION USED TO OBTAIN RAMP TABLE

O .'

O 6.0 -

5.0 -

4.0 -

O X

E 3 .0 -

E M

x

<3 2.0 -

i 1.0 -

O ' ' ' ' '

O. .2 .4 .6 .8 1.0 TIME , SECONDS FIGURE 4 TIME DEPENDENT REACTIVITY INSERTION USED TO GENERATE SCRAM TABLE t -

,o .

20 -

O 18 -

U

's 18 rod FLIP x 16 -

(GA 9350)

<a O

C

$14 FLIP-CLEAN w (PRN C- 123) 9 It w 12 -

O STANDARD w (G A 7882) m h to -

a

?

( 2 y8 -

s F

$ 6 -

FLIP-3OOO Mw-DAYS

$ (PRNC-123) s g4 -

n.

2 -

I I o I I I I I o 10 0 200 300 400 soo Goo 7oo TEMPERATURE, *C FIGURE 5 TEMPERATURE COEFFICIENTS OF TRIGA FUELS

. 2.0 -

' a O,., . 3 -

v

• 2 -

ti 1.6 l ~

g w

QO S O 10 0 200 1.4 -

TEMPERATURE water C 1.2

/

/

[

/

t' BOL, 35 FLIP (35.7%) e e' 1.0

,e j -

BOL,59 FLIP (60.2%)

i

---

B O L , 9 8 F LIP ( 10 0 %) e'

• - O.8

^

i f

EOL,9 8 FLIP (100 %)

i

,' , ,/

/ s l

0.6 -

/ s s' /

s l

s

/ /

/

O.4 -

s f ,l s

, l l ,/ '

/-

j/,,/

O.2 -

p,, '/*,,/,

O ' ' ' ' ' ' ' ' '

O 10 0 200 300 400 500 600 700 800 900 T E MP E R ATURE fuel , C FIGURE 6 INTEGRAL TEMPERATURE COEFFICIENT

.O*

4

. h- The integral water temperature coefficient is the standard FLIP coefficient and is shown in the inset to Figure 6. Since the BLOOST 2 does not. calculate 4

the' peak temperature properly when the pulse is from power, the peak tempera-ture for these cases was calculated by hand. This was done by the following process:

1) Determine the average power density in the " hottest" element at IMw.

-2) From.this power density determine the steady state temperature at the periphery of the fuel at the axial centerline.

3) Calculate-the. energy content at the point at 1Mw.

4). Calculate the energy added at that point in the pulse.

i 5) Determine the adiabatic temperature that would result from that energy i content.

V The peak edge temperature in the fuel before the pulse is:

, 35 FLIP Max SS. edge temp = 165 C Energy content = 0.139 59 FLIP Max SS edge temp = 170 C Energy content = 0.144 98 FLIP Max SS edge temp = 180 C Energy content = 0.154 elem Th'e following; tables give the principle results of the BLOOST 2 calculations.

PUI.SE FROM 1 Mw No. p Prompt _

Prob # $ FLIP llurnup max Energy T T

_11 .2.00 35 BOI. 677Mw 12.4Mw 345 667.3 12 '2.75 35 BOL 1383Mw 18.4Mw 394 .844.1 l 13 2.00 59 B01. 817Mw 13.5Mw 354 685.9 14' 2.75 59 BOL 1579Mw 19.6Mw 403 855.2 15 2.00 98 BOL 1110Mw 15.8Mw 373 -673.7.

16 2.25 98 BOL 1420Mw 18.1Mw 392 728.9 I

17 2.00 98 EOL 1502Mw 22.4Mw 425- 826.2 18 2.25 98 E01. 1951Mw 26.4Mw 455 910.0

ee D .,

-e Pl!!.SE FROM.300W

~

e No. p Prompt- Energy _

Prob #- $ FLIP llurnup max EnerP,y to 2s T T

' 21 ~ 2.00 35 1101, 1263Mw 17.7Mw 24.0 .225 767

'22 .2.75 35 1101, 3625Mw 28.5Mw 34.4 324 1057 23 2.00 -59 11 0 ! . 1619Mw 19.6Mw 26.5 244' 802 24 2.75 59 1501, 4 631Mw 31.6Mw 38.0 349 1094 25 2.00 98 1501. 2645Mw 24.lMw 32.5 286 802 26 2.25 98 BOL 4016Mw 28.9Mw 36.8 328 907 27 2.00 98 E01. 2561Mw 29.0Mw 39.4 329 911 28 2.25 98 Eol. 3905Mw 35.0Mw . '45.7 379 1031 To find the values of the reactivity insertion that would result in peak fuel temperatures of 950 C, an interpolated value was found from the data in the tables above. These values are given in the next table.

REACTIVITY TO CIVE 950 C PEAK TEMPERATilRE

,_ No. li LIP ' llu rnu p Pulse from 300w Pulse from 1Mw

\ 35 Il0L $2.45 $3.32 59 1501. 2.35 3.29 98 -1501. 2.36 3.65 98 EOL- 2.08 2.42

~

_It.was felt that.these results were sufficient to allow effective operation so it was not necessary to do the second part of the program which would consider

! . heat flow from the fuel.after the pulse, i

i

(

I t

i I

l I

L_