ML20238E962
| ML20238E962 | |
| Person / Time | |
|---|---|
| Site: | University of Missouri-Columbia |
| Issue date: | 09/11/1987 |
| From: | Alger D, Mckibben J MISSOURI, UNIV. OF, COLUMBIA, MO |
| To: | Office of Nuclear Reactor Regulation |
| References | |
| NUDOCS 8709150310 | |
| Download: ML20238E962 (94) | |
Text
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-J Research Reactor Facility UNIVERSITY OF MISSOURI 1 g,,,,,ch p,,k Columbia, Missouri 65211 )
Telephone (314) 882-4211 September 11, 1987 l
I i
Office of Nuclear Reactor Regulation U. S. Nuclear Regulatory Commission Washington, D. C. 20555 l
Attention: Director ;
Standardization and Non-Power Reactor Project Directorate l
REFERENCE:
Docket 50-186; License R-103 University of Missouri
SUBJECT:
Response to NRC Request for Additional Information
Dear Sir:
We have completed our response to the questions in your letter dated May 7,1987 concerning our application dated September 12, 1986 for an amendment to our R-103 operating license. This amendment would authorize the use in our reactor of a newly developed extended life aluminide fuel (ELAF) element containing higher densities of uranium and a burnable poison.
The twelve questions and our answers are attached, also enclosed are copies of oxide measurement data sheets which are referenced in Question 1.
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Director September 11, 1987 Page 2 If you have any questions concerning our responses, please contact me at (314)-882-5204 or Dr. Soon Sam Kim at (314)-882-5244.
Respectfully submitted, J. C. McKibben Reactor Manager Reviewed and Approved:
'DCA bk, 04 !
Don M. Alger Associate Director xc w/ attach /encls: A. Adams R. Ambrosek C. Cooper .,
R. Carter K. Brown M _
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QUESTION: NUMBER 1 O '
h Theffue11 plate oxide thickness is a function of operating time. What is the' u
b- , projected oxide thickness for the increased ' operating time?
n l', ,
ANSWER:
The1Griess~ correlation from the ORNL-3541 report (Reference 1) is the accepted method to project oxide _ thickness for. 6061 aluminum fuel plates. The. oxide thick-1 ness:is a function of operating time and fuel plate surface temperature; also,- the coefficient is dependent on coolant pH and heat flux. The: standard form of-the:
correlation.is based on water with a pH of 5.0 and with heat fluxes between
- 1f to 2'x 106 BTU hr-1 ft-2, The Griess correlation is:
x = a 80 778 ,xp ( 82 g) where E a = 443 for the standard conditions x E 0xide thickness (mils = .001 inch).
0 E Exposure time (hr)
- R E Temperature of surface in contact with coolant (*R)
Reference i states "Within the range 1 to 2 x 10 6. BTU hr-1 f t-2, heat flux was unimportant except in the manner in which it influenced surface temperatures.
-Below 1 x 10 6 BTU hr-1 ft-2, heat flux per so was a'significant variable;
. considerably lower oxide thicknesses were observed than calculated by the above correlation." In 'the conclusions of Reference 1 the following-dependencies.on heat flux and pH are given:
%V
- QUESTION NUMBER l' (cont'd)
.At 0.5'x 106 BTU hr~1 f t-2 the rate of oxide formation was about half 'that observed at heat fluxes of 1 to 2 x 10 6 BTU hr~1 ft-2, other conditions being the same.
Under the same conditions except with the coolant pH in the range of 5.7 to 7.0, all constants were the same except the constant a which was 1200 instead of 443. In both pH ranges the above correlation predicted oxide thicknesses considerably higher than were observed when the heat flux was less than 1 x los BTU hr~1 f t- 2, The maximum heat flux for a 775 gm element is 0.45 x 10 6 BTU hr~1 f t-2. - The primary coolant chemistry records over the past 10 years show weekly pH measure-ments of 5.5, 6.0 or 6.5, with 5.5 being the most frequent measurement. Our sensitivity in measuring the primary coolant pH was improved in July 1987. Since the improvement, the pH has ranged between 5.3 and 5.5. To more accurately predict the oxide buildup on MURR fuel elements, a correction factor to the Griess correla-tion coefficient was determined to account for the lower heat flux and higher pH range.
To determine the proper correction factor to the Griess coefficient (a) for use with MURR fuel elements and water chemistry, the typical maximum oxide thickness must be compared to the value predicted by the standard Griess correlation. To do this, oxide thickness measurements were made on July 23 and 24,1987 by Idaho National Engineering Laboratory (INEL) and the University of Missouri Research Reactor Facility (MURR) staffs at MURR. Oxide thicknesses on seventeen different fuel elements were measured. The fully burned up elements with the highest and l lowest oxide thickness were measured twice. Only the outside convex fuel plate f surface (plate 24) can be measured. The data were recorded on "ATR 0xide Thickness l
f
QUESTION NUMBER 1 (cont'd)
Measure ments" forms and are enclosed. The measurements were made using INEL equip-ment and procedures. The technique is based on eddy current measureitents. The oxide thickness values obtained are based on a 0.62 mil oxide standard supplied by INEL. The standard was measured before taking measurement on each element. Each element was measured in at least three places. On the data sheets, "6" line (under l the " convex Plate #19" section) corresponds to a point down one-fourth of the length of the plate (25.5 inches) from the coolant inlet, and "12" line is the plate vertical centerline. "18" line is three-fourths of the way down the plate. All but one of the fuel elements had the highest oxide measurement at either centerline or "18" measurement point. M0-143 had the highest value at the "6" measurement point, which when remeasured yielded a lower measurement. The correction factor used to con-vert the readings to mils of oxide was based on the average of all the 0.62 mil standard measurements. The fuel elements with the maximum measured plate 24 oxide thickness are given in Table 1.1, with the exception of the one high measurement on element MD-143. The elements numbered 133 through 148 were manufactured at Atomics International (California); the rest were manufactured at Babcock and Wilcox (B&W) I in Lynchburg, Virginia. The highest oxide measurements were on elements M0-200, MO-201, and M0-203 which are three of the first four MURR elements made at B&W.
None of the other B&W elements shows the same tendency for high oxide growth. To attempt to determine if in-pool storage time contributes noticeably to the amount of oxide buildup, some alements with approximately the same power history but signi-ficantly different in-pool storage time were measured. By looking at the oxide thickness versus the in pool storage time (from first irradiation date up to July 1987) for elements with similar MWD, it can be concluded that the variation in thickness does not correlate well with storage time.
QUESTION NUMBER 1 (cont'd)
The data were analyzed to obtain a conservative non-temperature dependent I
correlation similar to the Griess correlation but based on MURR operating history, from which the typical maximum oxide thickness could be determined for plate 24 on the 775 gm eierent with a 150 MWD power history. To factor in all the elements measured, a linear regression analysis on the data for the highest' values for the seventeed ehments was performed assuming the following form:
s" Oxide Thickr.ess = C1 (MWD)0.77'S + b (mils) s 3 '
,' i where E slope of the regression line
(*,1 s
. 1 b I 0.0 to assum neioxide on new fuel element, sinilar 4
^
N' to Griess correlation assumption ,
This assumes the same time,$ependenc'y as the Griess correlation,'but assumes thh A: q
'1 temperature dependent term to be cor?stant. The regression analysis yiel_ds a slope
> ~ ,
(C1 ) equal to 0.00961 with a staudard error of 0.000837 in the slope, and is plotted <
. - g y j in Figure 1.1.. '.So the dmived cderelation of the average value far the maximum plate 24 oxide thickne:8'd.easurement for our current 775 gm fuel element is:
Oxide Thickng s = 0.0096). (MWD)0.778 l
- ^(mils) I
\ j MWD = the powcrhistory on toe fuel element
, x' T91s gives thr< Nximm oxide thickness of 0.474 mils for plate '.4 i on a 175 fuel
' i e'ement with 1S0 MWD on ii.' The standard deviation in the slope can tc, derived from the standard error.
. Standrird Deviation (c) = -
begreesofxStandardError Freedom e
\
~
e .- (17-1) x 0.0008'i7 r 0.003?S 775 i
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QUESTION NUMBER 1 (cont'd)
If the temperature dependency had been factored into the regression analysis, it would have yielded a value less than 0.474 mils. This is due to the highest fuel plate surface. temperatures being on new elements with very few MWD of power his tory. This "more accurate non-linear regression" would have a steeper initial slope because of the higher surface temperatures of new elements and a flatter 4 slope towards the end when the fuel element is almost depleted. Looking at Figure 1.1, it can be seen this sort of regression curve passing through the same data point for 150 MWD would increase the error between the regression " fit" and the measurements for power histories less than 140 MWD.
The experimentally determined maximum plate 24 oxide thickness should be compared to the predicted value of the Griess correlation, which requires knowing the fuel plate surface temperature. The fuel surface temperature varies due to control rod height, burnup on the element and relative fuel loading as compared to the other elements in the core. The following assumptions are made to simplify the calculation of predicted oxide growth:
- 1. The 150 MWD power history on a 775 gm fuel element is all from full power operation, i.e.,10 MW with an eight element core; therefore the fuel element is in the core for a total of 2880 hours0.0333 days <br />0.8 hours <br />0.00476 weeks <br />0.0011 months <br />.
}. 24 h 150 MWD x * = 2880 hr If2 WI) da
- 2. The operating history can be approximated by dividing the total operation time into a series of four time intervals, based on assumptions 3 and 4, each at a dif ferent constant hot spot surface temperature. The four intervals are assumed to be:
- o. ,
- y. '?o
.QUESTIONlsVMBER'1' (cont'd)!
~%
"G a) New element operating' at a no xenon rod height for first
! 5.33% ofI the ' time.
b) New element operating at an equilibrium xenontrod height for the1next 44267% of the time.
c)' Depleted element operating at a 'no' xenon rod height for the third time interval of 5.33% of' the total- time.
d) Depleted element operating.at a equilibrium xenon rod height for the last 44.67% of the time.
- 3. With the 150 hours0.00174 days <br />0.0417 hours <br />2.480159e-4 weeks <br />5.7075e-5 months <br /> of operation per week, assume the core operates' for 16 hours1.851852e-4 days <br />0.00444 hours <br />2.645503e-5 weeks <br />6.088e-6 months <br /> at cold clean' rod height (17 inche's withdrawn) and the rest of the time at an equilibrium xenon rod height of 22 inches withdrawn. Actual.
. equilibrium xenon rod heights vary between 22 to 25 inches depending on q total MWD on the core. . Therefore,10.67% of the time the reactor operates at no xenon. rod height..
4.. Fuel elements have the. peaking factors and temperatures of new elements .
for half their operating time and those of depleted fuel for the other j half.
Two computer codes were used to determine the fuel plate hot spot surface
. temperatures. First, the nuclear peaking factors were determined using the BOLD
. VENTURE code s'ystem. The fuel elements were looked at in mixed cores of fresh and depleted fuel elements. Second, the derived nuclear peaking factors are combined with engineering peaking factors which are then used in the COBRA-3C/RERTR code to determine the fuel plate surface temperature. Table 1.2 gives the results from
.these computer codes along with the calculated oxide- thickness using the methodology
. given in Reference 1 on pages 24-28 to predict oxide growth for a series of oper-ating times at different temperatures. Oxide growth during in pool storage was not computed since no significant correlation was found in the oxide measurements. <
pm 1
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- ; QUEST 10ll NUMBER 1 (cont'd)
- The appropriate MURR= correction factor to the Griess correlation is the ratio q,~, .of:the experimentally measured . oxide thickness to the standard Griess' predicted oxide thickness for 150 MWD on plate 24.-
"MURR Griess = { ) x 443; 00 778 ,x g )
L. - Correla tion
= 0.697 :x 443- 00 778 exp ( )
- s 0.7 x 443 00 778 ,xp-( )
Using the same assumptions as before, the predicted maximum oxide thickness for.
- the " hot pla'te"' (Plate 1) in the 775 gm element can be calculated' using the MURR Griess correlation and is given in Table 1.3.- Now modifying the assumptions with the following two changes the predicted maximam oxide' thickness for the new t '1270 gm U-235 fuel element can be calculated using the MURR Griess' correlation:
- 1. A.300 MWD power history -is assumed for a depleted element; therefore, the fuel element will be in the core for a total of 5750' hours.
- 2. The col'd clean critical rod height is 16 inches.
The calculated values .are. given in Table 1.4. Ccmparing the peak oxide thick-ness of 0.854 mils for the 1270 gm element 'with the 715 go element peak oxide of 0.631 mils shows a factor of 1.35 increase. The MURR Griess correlation should be valid for predicting the oxide buildup on both the 775 gm element Plate 1 and the 1270 gm element hot spot since the heat fluxes are not significantly different and
- the chemistry would be the same as for the 775 gm element plate 24 L
u l
L
j s QUESTION NUMBER 1 (cont'd) i The predicted 0.854 mils is the average maximum value expected if several 1270 1 fuel elements were measured. To predict the " worst case" maxinm oxide on a 1270 fuel element, the following assumptions are made:
- 1. A linear correlation can be made between the oxide thickness in mils and the power history on the fuel element in MWD to the 0.778 power.
- 2. For the 775 gm and 1270 gm fuel elements, the ratio of their standard deviations for the slope of the linear correlation is equal to the ratio of the linear correlation's slopes.
I y c 1270 1270 = 1
- 775 C1 'O .
- 3. The " worst case" oxide thickness corresponds to the 300 MWD value for the linear correlation using the plus three standard deviations slope.
" Worst case" #
oxide
=(Cf270 3"1270)(300 MWD [**
=C 3g 270 (300 MWD)0.778 The slope of the linear correlation for the 1270 gm element can be derived from the predicted oxide thickness.
0.854 mils = Cf 270 (300 MWD)0.778
= 0.0101 Cf0 l
-9 L
f,-. '
QUESTION NUMBER 1 (cont'd)
Based on assumption 2 the predicted standard deviation for the 1270 fuel elements !
i can be calculated.
1270 Cy 1270 *y# 1 775 l 0 "
Ob 0.00335 = 0.00352 1270 1 The slope of the + 3 standard deviations can now be calculated.
C 1270 = C1270 + 3 o1270 3a 3
= 0.0101 + 3 x 0.00352
= 0.0207 i
'Now the " worst case" oxide thickness can be calculated for a fully depleted 1270 fuel element using the linear correlation.
1270 Oxide thickness = C 3a (MWD)0.778 (mils)
= 0.0207 (300)0.778
= 1.75 mils
4 QUESTION NUMBER 1 In summary, the predicted maximum oxide thickness on a new 1270 gm fuel element with 300 MWD power history at 10 MW is 0.854 mils. The 300 MWD power history corresponds to a peak burnup of slightly greater than 2.3 x 10 21 fissions per cm3 . The three standard deviations value for the oxide thickness at 300 MWD is 1.75 mils. The probability of having an oxide thickness greater than 1.75 mils is less than 0.14%.
i Reference (1) J. C. Griess, H. C. Savage, and J. L. English, Ef fect of Heat Flux on the Corrosion of Aluminum by Water, Part IV, ORNL-3541, Union Carbide Corp. Nuclear Division, Oak Ridge National Laboratory,1964. l l
- - - - - - - - - - ___m_ _____ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ ,
l i
. i Table 1.1 1
4 0xide Measurements of 775 Elements 1
Fuel MWD 0.778 Maximum 0xide Irradiation Period .
Element (MWD) , Plate 24 (Mils)_ First Last 133 148 48.8 0.36 11/84 9/11/86 135 148 48.8 0.47 11/84 9/11/86 )
143 100 36.0 0.20 8/86 N/A I 145 148 48.8 0.44 8/85 12/11/86 146 149 49.1 0.27 8/85 12/24/86 146 149 49.1 0.34 8/85 12/24/86 147 148 48.8 0.46 8/85 12/11/86 148 149 49.1 0.49 8/85 12/24/86 200 150 49.3 0.61 6/85 11/05/86 201 148 48.8 0.71 6/85 7/24/86 203 148 48.8 0.87 6/85 7/24/86 203 148 48.8 0.85 6/85 7/24/86 206 155 50.6 0.44 10/85 5/21/87 216 110 38.7 0.28 2/86 N/A 218 110 38.7 0.30 2/86 N/A i 226 100 36.0 0.19 8/86 N/A 228 74 28.5 0.17 10/86 N/A 239 70 27.3 0.16 3/87 N/A 246 9 5.53 0.06 7/87 N/A Reference
- 0 0 0.02
element which was pickled for 10 minutes in 30 volume " HNO3 acid at 160*F.
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- t QUESTION NUMBER 2 Reported results indicate that the peak power density will be on inner fuel plates for the new elements. What are the power profiles on the hot fuel plates?
Is the 1.04 azimuthal peaking factor still valid?
QUESTION NUMBER 10 The proposed fuel elements will be loaded with a variable and higher uranium density and some plates have boron added. The hot spot is on a different plate and the azimuthal peaking is higher. What calculations have been performed to demonstrate that the hot channel factors used for analysis of the safety limits are applicable to the proposed core?
i ANSWER:
l A current MURR fuel element has a uniform fuel density in the plates and the peak power density always occurs at Plate 1. However, the proposed fuel element has a variable fuel density between plates and the peak power density is on plates other than Plate 1. To investigate the azimuthal power profiles of hot plates of the current (775 gm) and/or proposed (1270 gm) fuel, detailed 0-R raodels were set up using the BOLO VENTURE computation system.
Cores containing only one kind (some power history and fuel loading) of elements were modeled as half of a fuel element (22.5*) with reflected boundaries at the fuel plate centerline and the centerline of the water gap between fuel elements. Cores containing two kinds of elements were modeled as two half elements (45*) next to each other with reflected boundaries at fuel plate center 11nes of both elements. The model included the water gap between the side plates, the alu-minum side plates, and the homogenized nonfueled fuel plate edge and coolant chan-nel region. Four cases were modeled: uniform loaded cores of each fuel element type (all 775 gm and all 1270 gm), a core loaded with four new 775 gm and four de-pleted 1270 gm fuel elements, and a core loaded witn four new 1270 gm and four
- _ = _ _ _ _ _
f d
QUESTIONS NUMBER 2 AND 10 (cont'd) depleted 1270 gm fuel elements. It is to be noted that the core with the fresh 775 gm plus depleted 1270 gm elements was calculated to have the highest overall peaking factors. Table 10.1 gives the azimuthal power densitics for a core of 775 gm elements with no power history. Table 10.2 is for a core of 1270 gm elements with no power history. Tables 10.3 and 10.4 are for the " worst case" peaking. Table 10.3 gives the azimuthal power densities for the 775 gm elements with no power history in the core. Table 10.4 gives the azimuthal power density for the depleted 1270 gm element. Table 10.5 is the a2.imuthal power densities for the fresh 1270 gm element when combined with depleted 1270 gm elements. It j should be noted the power densities given in these tables are from a BOLO VENTURE 0-R model which gives peak axial values based on buckling and has the control rods rods removed. Inserting the control rods has very little effect on the azimuthal peaking.
According to the 0-R calculations, when the fresh 775 gm elements are combined with either 775 gm or 1270 gm elements, peak power density always occurs at Plate 1, showing the 1.04 azimuthal peaking factor currently used in the MURR analysis is a conservative value. When the fresh 1270 gm elements are loaded with the depleted 1270 gm elements, the local peak power density is on Plate 7 in the 1270 gm elements with no power history. The azimuthal peaking factor within the element for the peak power density is shown to be about 1.33 at Plate 7 in the new fuel element.
1 The MURR safety limits are derived using nuclear and engineering hot channel factors and flow related factors. The proposed core with the new fuel elements will I
have a much flatter radial power distribution than that of the current core because of the variable uranium density in the new fuel element plates. Figure 10.1 l shows the radial peak power distributions in the core regions of the current and the proposed cores when all the control rods are fully withdrawn. The power
- ore is distributions were obtained from AMPX-II/ BOLD VENTURE R-Z runs and tt i
QUESTIONS NUMBER 2 AND 10 (cont'd) modeled as 24 homogenized fuel cells. Nuclear peaking factors of four different cases are compared in Table 10.6. The first case is the nuclear peaking factors currently used in the MURR 10 MW safety limits. The second is the " worst case" mixture of new and old fuel elements consisting of four 775 gm elements with no power history mixed with four 1270 elements that have produced 300 MWD of energy from each element (2.4 x 10 21 fission /cm3 peak burnup). The third is the calcu-lated peaking for four 775 gm elements with no power history mixed with four 775 gm elements that have produced 200 MWD of energy from each element (2.38 x 10 21 fission /cm3 peak burnup). The last case is the peaking factors for four fresh 1270 gm elements mixed with four 1270 elements that have produced 300 MWD of energy from each element.
Accurate evaluation of the engineering hot channel factors is not currently possible since there is no approved fuel specification for the newly designed fuel element. Oxide thickness can influence the flow related factors by reducing the water channel width. However, when the MURR is loaded with mixed cores of new 1270 gm or 775 gm elements plus depleted 775 gm or depleted 1270 gm elements, the hot spot always is located in a 775 gm fuel element with the lowest power history or in a 1270 gm element with the lowest power history in an all 1270 gm element core. A new 775 gm or 1270 gm element will have very little oxide growth, which means no change to the flow related factors is required. The fuel content and fuel thickness / width variations in the new fu21 element may be different from those applied to the current fuel element, which will affect engineering hot channel factors. The changes in the hot channel engineering factors due to the changes in increased fuel density and addition of boron will be evaluated when the new fuel specification is available. However, as shown in Case 4 in Table 10.6, as a result of a flatter radial power shape in the new fuel element, the nuclear peaking factors in the Safety Limits correspond to an overall nuclear peaking
QUESTIONS NUMBER 2 AND 10 '(cont'd) factor 1.45 times the worst case nuclear peaking in a 1270 gm element. If neces- l sary, this provides a sufficient margin in the current Safety Limits to allow for increases in the engineering hot channal factors for the 1270 gm element. The i
1270 gm element will be used only after it is documented that the engineering factors I based on the fabrication specifications yield overall peaking factors less than the assumed overall peaking factors in the Safety Limits. Therefore, the Safety Limits 1 of the current core can still be applied to the proposed core, or a core which has a mixed loading of the 1270 gm and 775 gm elements. This is documented in the following revisions to HSR Addendum 4:
Safety Limits Add the following note to the bottom of page 13 of HSR Addendum 4:
NOTE: The safety limits given in Appendix F are conservative values based on the 775 gm U-235 fuel element hot channel factors given in Table F.2.
This revision adds Table F.2a to give the maximum values for three worst case cores: a core of mixed 775 gm and 1270 gm U-235 fuel elements, a core of all 775 gm elements, and a core of all 1270 gm elements. The mixed core consists of four unirradiated 775 gm elements mixed with four irradiated 1270 gm elements (300 MWD per element). This combination of 775 gm and 1270 gm elements has the highest overall hot channel factor.
The second case has four unirradiated and four irradiated (200 MWD per element) 775 gm fuel elements. The third case has four unirradiated and four irradiated (300 MWD per element) 1270 gm fuel elements. The safety limits are based on overall peaking factors on enthalpy rise and hot spot heat flux that are greater than any of the " worst case" cores.
Therefore, the safety limits given in Appendix F and extended in HSR Addendum 5 are conservative and can be used for any possible combination of eight 775 gm and/or 1270 gm fuel elements.
Add Table F.2a to Appendix F of HSR Addendum 4 as page F-Sa.
I i
. I
! l L___-_____ l
7 Table F.2a l
MURR hot channel factors for worst case core of ,
mixed, all 775, or all 1270 fuel elements.
Corea 4F775 & 4F775 & 4F1270 &
4D1270 4D775 4D1270 Control Rod Height (withdrawn)b 13 inches 17 inches 16 inches Locationc . F775 F775 F1270
' Plate 1 Plate 1 Plate 4e On Enthalpy Rised Power-Related Factors Nuclear Peaking Factors ,
Radial x Non-uniform Loading 2.337 2.301 1.416 Local (circumferential within a plate) 1.040 1.040 1.3300 l i
Engineering Hot Channel Factors l Fuel Content Variation 1.030 1.030 1.030 Fuel Thickness / Width Variation 1.030 1.030 1.030 0verall Product 2.58 2.54 2.00 On Heat Flux Power-Related Factors Nuclear Peaking Factors Radial x Non-uniform Loading 2.337 2.301 1.416 Local (circumferential within a plate) 1.040 1.040 1.3300 Axial 1.327 1.315 -1.347 Engineering Hot Channel Factors
. Fuel Content Variation 1.030 1.030 1.030 fuel Thickness / Width Variation 1.150 1.150 1.150 Overall Product 3.82 3.73 3.00 I
aF775 is an unirradiated 775 gm element and D775 is one irradiated to 200 MWD. F1270 is a unirradiated 1270 gm element and D1270 is one irradiated to 300 MWD.
t bA ctual cold clean critical rod heights are higher but these values were used to generate conservative peaking factors.
cThe total heat flux of the hot plate is assumed to be transferred to the
" hot channel."
dDimensionally the 775 gm and 1270 gm elements are identical; therefo e, the flow-related factors are as given in Table F.2.
eActual Local Nuclear Peaking Factor for Plate 4 is 1.19.
l I
? .
Table 10.1 l
I l' Power Distribution in a Fresh 775 Fuel Element obtained from a 22.5' 0-R MURR Model Azimuthal Power Densities for mesh intervals of 0.0672 radians Channel from the outside edge of fuel meat to fuel plate centerline Max-Avg. x 100 Number 1 2 3 4 5 Avg. Avg.
. ~
1 608.2 611.8 609.5 607.6 606.7 608.7 0.5 2 521.4 507.4 499.1 494.7 492.8 503.1 3.6 3 469.3 439.9 425.7 418.9 416.1 434.0 8.1 4 434.8 395.2 376.4 367.7 364.2 387.7 12.2 5 410.4 364.8 342.9 332.9 328.9 356.0 15.3 6 392.5 343.5 319.9 308.9 304.7 333.9 17.6 7 379.1 328.2 303.7 292.4 288.0 318.3 19.1 8 368.9 317.0 292.2 280.8 276.4 307.1 20.1 9 361.2 308.7 283.8 272.5 268.2 298.9 20.9 10 355.5 302.5 277.7 266.6 262.4 292.9 21.4 11 351.4 297.9 273.3 262.4 258.3 288.7 21.7 12 348.7 '294.5 270.1 259.6 255.6 285.7 22.1 13 347.2 292.3 268.2 257.9 254.2 284.0 22.3 14 346.9 291.2 267.4 257.5 253.9 283.4 22.4 15 347.8 291.3 267.9 258.4 255.1 284.1 22.4 16 350.1 293.0 270.1 261.2 258.0 286.5 22.2 17 353.9 296.8 274.7 266.3 263.4 291.0 21.6 18 360.0 303.5 282.6 274.8 272.2 298.6 20.6 19 368.8 314.4 295.0 288.0 285.8 310.4 18.8 20 382.1 331.4 314.1 308.0 306.1 328.3 16.4 l 21 401.9 357.4 342.7 337.7 336.1 355.2 13.2 l
l 22 432.2 396.3 384.8 381.0 379.8 394.8 9.5 f
23 478.7 454.3 446.6 443.9 443.2 453.3 5.6 24 551.6 540.1 536.2 534.9 534.9 539.5 2.3 ;
l 1
o 1
Table 10.2 Power Distribution in a-Fresh '1270 Fuel Element obtained from a 22.5* 0+R MURR Model m
Azimuthal Power. Densities for mesh intervals of. 0.0672 radians
< Channel from the' outside' edge of fuel meat to fuel plate centerline Max-Avg. x 100 Number: 1 2 3 4 5 Avg. Avg.
1 310.6 319.3: 320.82 320.9 320.8 318.5 2.5 2 329.6 329.4- 327.8- 326.6' 326.1 327 9 0.5 3 344.2 328.5 321.3 318.1 316.9 325.8 5.6 4 '421.8 382.4- 366.2 359.6 357.2 377.4 11.8 5 436.6 376.7 353.1 343.8' .340.6 370.2 17.9 6 476.3 394.0 362.5 350.5 346.5 386.0 23.4.
7, 494.4 395.8 359.2 345.7 341.2 387.3 27.7-8 467.3 365.9 329.1 315.8 311.6 357.9 30.6
-9 449.1 346.9 310.6 297.9 293.8 339.7 32.2 10 436.5 334.3 298.9 286.7 282.9 .327.9' 33.1 11 427.8 325.8 291.4 279.91 276.3 320.2 33.6 12 422.0 319.9 286.5 275.6 272.3 315.3- 33.9
.)
13 418.6 315.9 283.5 273.2 270.2 312.3 34.0 i 14 417.4 313.9 282.2 272.5 269.7 311.1 34.2 15 418.2 313.6 282.7- 273.6 271.0 311.8 34.1 16 421.5 315.8 285.9 277.3 275.0 315.1 33.8 17 428.3 322.1 293.4 285.4 283.3 322.5 32.8 18 440.3 335.1 307.8 300.5 298.6 336.5 30.9 19 461.6 359.7 334.2 327.7 326.0 361.8 27.6 l 20 473.3 383.0 361.1 355.7 354.4 385.5 22.8 21 459.8 391.2 374.9 371.0 370.1 393.4 16.9 22 441.1 397.1 386.6 384.1 383.6 398.6 10.7 23 426.0 403.3 397.9 396.7 396.6 404.1 5.4 24 392.2 385.7 384.2 384.2 384.3 386.1 1.6 i
~22-
Table 10.3 Power Distribution in a Fresh 775 Fuel Element when loaded I with Depleted 1270 obtained from a 45' 0-R MURR Model r
Azimuthal Power Densities for mesh intervals of 0.0672 radians Channel from the outside edge of fuel meat to fuel plate centerline Max-Avg. x 100 Number 2 3 4 5 Avg. Avg.
1 1 716.9 714.3 708.2 704.1 702.2 709.1 1.1 2 612.9 588.7 575.3 568.3 565.3 582.1 5.3 3 549.2 507.0 486.7 476.9 473.0 498.6 10.2 4 505.9 452.4 427.0 415.2 410.6 442.2 14.4 5 474.4 414.8 386.4 373.2 368.1 403.4 17.6 6 450.6 388.1 358.2 344.4 339.1 376.1 19.8 7 432.1 368.7 338.4 324.4 319.0 356.5 21.2 8 417.7 354.4 324.2 310.4 305.1 342.4 22.0 9 406.6 343.7 314.0 300.5 295.4 332.0 22.5 10 398.2 335.6 306.5 293.5 288.6 324.5 22.7 ;
11 391.9 329.5 301.0 288.6 283.9 319.0 22.9 !
12 387.7 325.2 297.2 285.3 280.8 315.2 23.0 13 385.2 322.3 294.9 283.4 279.2 313.0 23.1 14 384.3 320.8 293.9 282.9 279.0 312.2 23.1 15 385.1 320.9 294.5 284.1 280.4 313.0 23.0 16 387.6 322.8 297.2 287.3 283.9 315.8 22.8 ;
17 392.1 327.2 302.4 293.2 290.0 321.0 22.2 i 18 399.2 335.0 311.4 302.8 300.0 329.7 21.1 19 409.7 347.5 325.6 317.9 315.4 343.2 19.4 20 425.2 367.0 347.2 340.5 338.4 363.7 16.9 21 448.3 396.6 379.7 373.9 372.2 394.1 13.7 22 483.2 441.0 427.4 422.8 421.4 439.2 10.0 23 536.8 507.1 497.2 493.7 492.6 505.5 6.2 24 620.4 604.6 598.5 596.0 595.1 602.9 2.9 !
l
.__ -- --- - i
o .. .
Table 10.4 Power Distribution in a Depleted 1270 Fuel Element when loaded
~
with Fresh 775 obtained from a 45' 0-R MURR Model l
l Azimuthal Power Densities for mesh intervals of 0.0672 radians l- Channel frem the' outside edge of fuel meat to fuel plate centerline Max-Avg. x 100 Number. 2 3 4 5 Avg. Avg.
L 1 1
1 91.1 95.5 97.1 97.7 97.9 95.9 5.0 2- 130.0 135.2 137.2 138.0 138.2 135.7 4.2 3 174.7 177.6 178.6 178.9 179.0 177.8 1.7 4 266.3 262.3 260.2 259.1 258.6 261.3- l.9 5 326.3 309.9 302.0 298.4 296.9 306.7 6.4 6 401.7 366.6 350.4 343.1 340.3 360.4 11.5 7 452.9 397.3 372.2 -361.3 357.1- 388.2 16.7 ;
8 474.8 402.0. 369.8 355.9 350.7 390.6 21.5 9- 452.1 372.0 337.1 322.1' 316.6 360.0 25.6 10 436.6 351.7 315.0 299.5- 293.9- 339.3 28.7 11 426.0 337.7 .300.1- 284.5 278.9 325.4 30.9 12 419.2 '328.4 290.6 275.1 269.6 316.6 32.4 13 415.1 322.6 284.9 269.9 264.6 311.4 33.3 14 413.9 320.0 283.1 268.7 263.'7 309.9 33.6 15 414.2 320.2 284.6 271.1 266.5 311.3 33.0 16 416.3 324.0 290.4 278.1 274.0 316.6 31.5 17' 432.6 342.3 310.4 299.0 295.3 335.9 28.8 18 454.0 368.7 339.4 329.1 325.8 363.4 24.9 19 485.6 408.8 382.9 373.9 371.1 404.5 20.1 20 449.4 394.7 376.3 370.1 368.1 391.7 14.7 21 397.3 364.4 353.1 349.4 348.2 362.5 9.6 22 326.7 311.5 306.1 304.3 303.8 310.5 S.2 23 255.1 250.9 249.1 248.7 248,7 250.5 1.8 24 180.8 181.3 181.S 181.8 182.0 181.5 0.4 l
_=___1___.____ _ . _ . _ _ _ _ _ _ .
j
Table 10 05 Power Distribution in a Fresh 1270 Fuel Element when loaded with Depleted 1270 obtained from a 45' 0-R MURR Model Azimuthal Power Densities for mesh intervals of 0.0672 radians Channel f rom the outside edge of fuel meat to fuel plate centerline Max-Avg. x 100 Number 3 4 5 Avg. Avg.
1 2 1 392.2 393.8 391.5 389.9 389.2 391.3 0.6 2 408.9 393.3 385.3 381.6 380.2 389.9 4.8 3 422.1 383.4 367.6 361.3 359.0 378.7 11.5 j 4 512.3 438.7 411.1 400.7 397.3 432.0 18,6 4
5- 527.7 428.9 393.6 381.0 376.9 421.6 25.2 6 573.8 448.1 405.2 390.2 385.6 440.6 30.2 i 7 595.2 454.1 406.9 389.4 384.9 446.1 33.4 8 563.9 421.3 375.9 361.3 356.9 415.9 35.6 9 543.2 402.8 359.1 345.5 341.6 398.4 36.3 .
I 10 528.9 390.8 349.1 336.4 332.8 387.6 36.4 11 518.6 382.6 342.7 330.9 327.8 380.5 36.3 12 511.6 376.9 338.6 327.7 324.8 375.9 36.1 13 506.9 372.7 335.9 325.8 323.2 372.9 35.9 14 504.5 370.0 334.5 325.1 322.8 371.4 35.8 15 504.1 368.8 334.5 325.8 323.7 371.4 35.7 16 505.9 369.7 336.6 328.6 326.7 373.5 35.4 17 511.1 374.3 342.6 335.2 333.5 379.3 34.7 18 521.5 385.3 355.2 348.5 347.0 391.5 33.4 19 541.8 408.5 380.3 374.2 372.9 415.5 30.4 20 551.5 430.5 405.8 400.6 399.5 437.6 26.0 ,
21 534.6 438.8 419.4 415.3 414.4 444.5 20.3 22 516.4 449.8 436.0 432.9 432.1 453.4 13.9 23 507.5 468.5 459.6 457.2 456.5 469.9 8.0 24 483.4 466.0 461.1 459.3 458.6 465.5 3.6
' Table 10.6 Description Case 1 Case 2 Case 3 Case 4 Safety Limits' F775+D1270' F775+D775 F1270+D1270-Control-Rod Height N/A> 13" 17" 16" (withdrawn)
' K-ef fective: .
' *Calcula ted N/A 1.005- 1.027 1.023
- Correcteda ---
0.96 0.98 0.98-Peaking. Factors:
- Azimu thal Within Elements 1.04 1.04 1.04 1.33b Between Elements 1.112
- Radial- 2.22 2.337c 2.301c- 1.416c-
- Axial 1.432 1.327 '1.315 1.347 q
- 0verall 3.676 3.225 3.147 2.537 Position of Hot Spot Plate 1 Plate 1 Plate l' Plate 4 aThe corrected F775+D775 is based on the.. empirically derived estimated critical-position and rod worth curves. The same' correction was assumed .to apply to the F775+D1270.
.bThis value is an azimuthal peaking factor at Plate 7 obtained. from a detailed 45' e-R BOLD' VENTURE run for a combination of F1270+01270; this value was used .to be conservative. .The calculated azimuthal peaking factor for Plate 4 is 1.19.
cThis peaking factor includes radial factor and azimutha1' peaking factor between elements. This is the average power density in 'the fuel plate with the hot spot f divided by the core average power density. It was obtained from a BOLD VENTURE 0-R-Z model.
l
800, 700f .
600 M
~
500-o - ' , %
N.
O p .g O -
p Proposed Core "
D
- w. 400i . .y .- % s / 'g + .\ ~..,_'+ ,,
g .
Ci : 7
~
o .
e 300--
w .
N =
o -
Current Core
- o. -
x -
w
- n. 200f .
100-'
0-. - ---
e 7 8 9 10 11 12 13 14 DISTANCE FROM CENTER LINE (CM)
Ticure 10.1 The Peak Power Density Distribution
Li
- QUESTION.. NUMBER 3 1
The increased element operating time-above"your current practice will result'
. in increased likelihood of fuel plate' corrosion and in pit depth. What is the pro-
' jected primary coolant radioactivity level .if a pit results in release of fission products, and what are the projected consequences to.the public?
l AWSWER:
The University of Missouri Research Reactor-(MURR) has had no problems or failures of aluminide fuel' caused by corrosion pitting. With this excellent per-formance, there' is no empirical basis .for calculating pitting rate at MURR. Thus, the best estimate of pitting rates would be based on the " Extended Life Aluminide -
' Fuel Final Report" by L. G.' Miller and J. M. Beeston, EGG-2441 June 1986 (Reference.
2). Pit replication was done on 15 plates to determine pit sizes and results are.
given in Table 9 of Reference 2. Only two plates had _ pit depths great'er than 3.0 mils. and only.one (depth 16.0 mils) was greater than 7.0 mils.. .The typical pit diameter is stated' to be about six times the depth of the pit. The following equation for pitting corrosion rates from Reference 2 was used to ' calculate the maximum predicted corrosion pit depth given in the referenced Table 9'and was conservative for all but the 16 mil pit.
CR,
= 7. 6 x 10- 2 4 T* in./ day where maximum corrosion rate CR max TE fuel plate surface temperature, *K.
I 1
I
1
. QUESTION NUMBER 3 (cont'd) f' The 1270 gm fuel element will have twice the operating time of the present
.775 gm elements, but the hot spot will have a lower surface temperature due to the more uniform power density. Table 3.1 compares the effect this has on the pitting rate, using the same four time intervals and hot spot surface temperatures as used in Table 1.3 and 1.4 in Question 1. The 13.09 mil pit depth for a 1270 gm element .
indicates a factor of 1.69 increase over the 7.73 mil pit depth for the 775 gm element. The 13.09 mils is a conservative overestimate since it was based on a ;
higher overall peaking factor than actually will occur in the 1270 gm element.
The increase in pitting depth may be predicted more accurately by comparing the naximum oxide thickness. As calculated in Answer 1, the maximum oxide thickness for the 1270 gm element is 1.35 times the maximum for the 775 gm element. The 1.35 increase in hot spot oxide thickness also indicates a greater potential for pitting corrosion in the new fuel. However, this should not cause any safety problem for MURR, since this kind of failure is not catastrophic and has been experiences by reactors such as ATR. A corrosion pit develops slowly and penetration of the ,
cladding is easily detected due to the installed on line fission product monitor and the minimum detectable activity levels for most of the iodine fission products is R approximately 1 x 10-7 to 1 x 10-6 pci/ml. f The following are typical alues for radioactive iodine activities in' the ~ ' f primary coolant: ,
Nuclide 1/2 Radioactive ty I-131 8.04d 1 x 10-6 pc/ml*
I-132 -
2.29h 8 x 10-6 pc/ml I-133 / 20.8h 3 x 10-6 pc/ml I-134 52.6h 4 x 10-5 pc/ml ,
I-135 6.58h 1 x 10-5 pc/ml
- Minimum detectable activity for I-131.
_ _ _ - 1
^
'~
QUESTION NUMBER 3 (cont'd)
The actual primary coolant radioactivity levels that would result from a pit releasing fission product cannot be predicted accurately. With no previous experi-ence with this mode of failure, there are too many unknowns. The following paragraph of Technical Specification 3.9 places a limiting condition for operation on the maximum I-131 radioactivity in the primary coolant:
3.9c) The reactor shall not be operated when radiochemical analysis shows that a concentration of the radioisotope I-131 exceeds 5 x 10-3 pc/ml in the primary coolant.
To put an upper limit on the problem, a " worse than possible" pit release will be calculated and consequences evaluated.
The Design Basis Accident (DBA) already addresses a release of fission products with acceptable consequences to the public. The DBA assumes the melting of the #1 plate in four 775 gm fuel elements, which releases 2.08 percent of the fission products. Tr,e,nower density and fission product iodine concentrations are greater in Plate 1 ofi the 775 gm element than in the hot plate of the 1270 gm element. So, it is conservative [ to assume the pit occurs on Plate 1 of the 775 gm element instead of the 1270 element to quantify the consequence to the public.
Based on the statement in Reference 1 that the diameter is six times the depth, the surface area of a pit whed it penetrates the 15 mil clad of a fuci element would be 0.090 inches in diameter. To be conservative, assume the pit is 0.2 inches in '
dineter and instantly release the fission froducts from a fuel meat volume equal lta, the pit area times the fuel meat thickness (20 mils). This volume contains e />
0.015 gm of U-235 compared to the 78.58 gm of U-235 contained in the Plate 1 of (cur 77o gm elements. So it would be equivalent to 1.91 x 10-4 times the release osped in the DBA which would not endanger the health and safety of the general public. Assuming instantaneous dilution in the 2000 gallons of primary coolant anti equilibrium I-131 activity in the . core for 10 MW operation of 2.5 x 105 curies, the primary coolant I-131 radioactivity cm be calculated:
c ,'
k_._,.m_ ___ _ _
-QUESTION' NUMBER 3 (cont'd)
=
2.5 x 10M pc x 0.0208 x 1.91 x 10-4 1-131 activi ty 2000 gal x 3,785
= 0.131 pc/ml where 2.5 x 101 luc E I-131 activity in the core
' O.0208 E the fraction of activity released in the DBA 1.91 x 10-4 E the ratio of activity released by "the pit" compared to the DBA release.
The MURR primary coolant system has a demineralized cleanup system with a 50 gallon per minute flow rate. The 0.131 pc/ml of I *231 activity could be reduced to 5 x 10-3 pc/ml in approximately 3 hours3.472222e-5 days <br />8.333333e-4 hours <br />4.960317e-6 weeks <br />1.1415e-6 months <br />.
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[ -
-QUESTION NUMBER 4
! Projected oxide thickness exceeds the minimum that has been shown to result i
)
in spontaneous spallation. What data is available for validation of oxide thick- !
I ness projections? What are the consequences to fuel clad as a result of spal-- ')
1ation, e.g. , subsurface voids? What is the impact to safety limits as a result I
of changes in heat transfer from the spallation zone? Is the oxidation rate accelerated on the spallation zone such that multiple spallation at that location 'l l
1s more likely? What would be the pocential consequences to the public? l ANSWER:
The initial projected oxide thickness was greater than 2 mils, which cor-responds to the range where spontaneous spallation can occur. The initial projected oxide thickness was based on the Griess correlation for a pH of 5.7 or ebove since the methodology used at MURR over the last 10 years to determine the primary coolant pH indicated it was between 5 and 7. In July 1987, the pH measuring sensitivity was improved and the oxide thicknesses on MURR fuel elements were measured for the first time. Baced on these measurements the projected max-t imum oxide thickness was calculated to be 0.854 mils for 300 MWD on the new 1270 gm fuel element (see answer to Question 1). The projected value for " worst case" maximum oxide thickness (3a) is 1.75 mils. Neither the more accurately projected oxide thickness nor the projected " worst case" oxide thickness exceeds the minimum that has been shown to result in spontaneous spallation (Reference 1.)
Therefore, the rest of Question 4 is no longer applicable to the 1270 gm element and will not be addressed.
i
QUESTION NUMBER 5:
The predicted increased oxide thickness above that for current operating con-ditions will result in changes to the reactor operating characteristics. What are the changes in fuel plate (steady state) operating temperatures? What are the changes in reactor response to reactivity perturbations as a result of increased oxide thickness on fuel plates?
ANSWER:
Analytical calculations were performed to determine the change in fuel plate steady state operating temperatures due to the predicted increased oxide layer.
The PeMimum predicted oxide thicknesses for 775 gm and 1270 gm fuel elements were calculated based on the MURR Griess correlation and discussed in the answer to Question 1. The values are: 0.631 mils for the average depleted 775 gm element and 0.854 mils for the average depleted 1270 gm element. The analytical calcula-tions in this answer were performed using earlier estimates before these Question 1 values were finalized. For the 775 gm element, 0.61 mils was used to represent the.
maximum oxide on the average deplete element while 1.27 mils was used to represent
" worst case" oxide thickness. For the 1270 gm element, 0.85 mils was used for the average maximum oxide and 1.74 mils for the " worst case."
For each fuel element type, the maximum heat flux and the corresponding heat transfer coefficient were obtained for their respective hot spot using the thermalhydraulics code, COBRA /RERTR. The 775 gm element has a maximum heat flux of 5.57 x 10 5 BTU ft-2hr-1 and a heat transfer coefficient of 7636.7 BTU ft-2hr-l*F-I. The 1270 gm element has a maximum neat flux of 4.94 x 10 E BTU 1
f t- 2hr-1 and a heat transfer coefficient of 7575.5 B10 f t-2hr-l'F-1 The following one-dimensional hest conduction and transfer equation (see Reference 1, p.117)
I was solved adding an oxide layer adjacent to the cladding to calculate the maximum fuel and cladding temperature changes due to the three assumed oxide thick-nesses (zero, maximum on average element, worst case).
1
L . QUESTION HlMBER 5 '(cont'd) c;t o=
' t t t q = 2k A f t*[t 8=kA c s
c 3=kAg hA(tg -t)f L where o, ko and to are respectively the oxide layer thickness, the oxide heat con-
.ductive constant and surface temperature of the layer, and other notations are same as those in Reference 1.
The calculated fuel operating temperatures for three oxide thicknesses for .
I each fuel type are compared in the following table.
' Fuel 0xide Thick- Fuel Fuel Plate Coolant Type ness (mils) Centerline (*F) Surface (*F) Surface (*F) _ (*F)
COBRA COBRA 1270 gm 0.0 229,7 229.4 224.9 224.8 218.5 153.3 ,
- 0.85 256.2 --
261.4 --
218.5 153.3 !
1.74 284.1 --
279.3 --
218.5 153.3 775 gm 0.0 240.3 240.1 234.9 234.8 227.7 154.7 0.61 261.8 --
256.4 --
227.7 154.7 1.27 285.1 --
279.7 --
227.7 154.7 1
As the table shows, comparison of analytical results with those of COBRA for the fuel d without oxide layer gives excellent agreement. The results indicate the 1270 gm
'I fuel elements will operate with a slightly lower fuel meat temperature than the !
775 gm element. This is true even in the unrealistic case of maximum oxide thick- ,
ness of a depleted element applied to the hot spot of a new element.
I Also shown is the fact that there would be an increase of 54.4*F in the steady state fuel centerline temperature if a " worst case" 1.74 mils oxide layer potentially caused by the longer operating cycle of the proposed fuel elements is ,
i applied to a new 1270 gm element hot spot heat flux. A 44.8'F increase in the fuel centerline temperature is calculated for the corresponding " worst case" 1.27 mils oxide layer buildup in the current fuel element.
l
i QUESTION NUMBER 5 (conf d)
Reactor response to reactivity perturbation was also studied using a point kinetics code, PARET/ANL. PARET/ANL is a modified version of the original PARET code and includes heat transfer correlations and models typical of research reactors. Benchmark calculations against SPERT-I (see Reference 2) cores which are similar to a plate-type research reactor, show that good agreement was obtained between the ANL version and the experimental results. It was reported that a core with the fresh 775 gm elements mixed with the depleted 1270 gm elements showed the highest power peaking among various combinations having the hottest channel in Plate 1 in the fresh 775 gm elements. To be conservative, MURR was modeled using two channels. The axial power profile in Plate 1 of the case was input to the first charael for the Plate 1 and the average distribution was input into the second channel for the rest of the core. Also used was an assumed coolant inlet tempera-ture of 80*F with a coolant flow rate of 3800 GPM. This will cause the reactivity I feedback to be delayed because of the greater sub-cooling when compared to the normal operating condition. Cince the MURR has highly enriched uranium fuel ele-ments, Doppler feedback was not considered, and in the analysis the conservative
-2.0 x 10-3 tA/% void from Technical Specification 3.1 was used in the model to take into account the void reactisity feedback. No temperature coefficient was used. It was also assumed that 3.5% of the energy generated in the fuel is deposited promptly in the moderator. The effect of the oxide film layer on the reactivity transient was analyzed assuming that the thickness of the layer is 2.5 mils and the oxide layer thermal conductivity is 1.3 BTU f t-lhr-l*F-1 The MURR Technical Specification 3.1, Reactivity, limits the maximum reactivity worth of the regulating blade, each secured removable experiment, all experiments in the center test hole, and all unsecured experiments not to exceed 0.006 tk. The basis is to limit these items to a value (0.006 tA) to which an inadvertent step insertion of reactivity would not result in fuel plate damage. To be conserva- g
(
tive, a step insertion of 0.007 Ak($0.9485) was modeled. The only scram modeled j
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j- QUUTION NUMBER S (cont'd) l was the 125% of full power with an assumed 150 millisecond (ms) time delay between trip point and start of control rod insertion. The control rod insertion was b
- modeled as -$11.26/sec for a total of 750 ms of insertion time. This gives a total control insertion of $9.5308 of reactivity, equal to .0566 Ak. The measured total control blade worth is -0.1655 Ak. The MURR Technical Specifications require that the control blades must be able to be inserted from fully withdrawn (26 inches) to the 20% withdrawn position in less than 700 ms and are verified to comply with this requirement monthly. This makes the model very conservative and is equivalent to having the control rods at 11 inches withdrawn at the start of the accident and taking 750 ms to go to the 20% withdrawn position ar,d stopping there. Figures 1 and 2 show the results of the step reactivity insertion.
For the model with an oxide layer, it was found that during the transient, the peak power reached 57.9 MW at 65.7 ms, the maximum fuel centerline temperature being 765.6 "F at 100 ms and the cladding temperature increased up to 396.0 *F. For the model without an oxide layer, the peak power reaches 51.48 MW earlier at 54.8 ms and then begins to decrease because of better heat transport from fuel to coolant. 1 The temperature transients in the tables also show that the peak fuel centerline temperature (437.5 'F) occurs at 65.2 ms and that the cladding temperature reaches its maximum 378 *F at 69.2 ms. It was found that the peak power increased from 51.48 to 57.94 MW and the maximum fuel centerline temperature went up from 437.5 to 765.6*F because of the oxide buildup. It is to be noted that the above calcula-tions are based on the conservative assumptions that the maximum oxide layer of 2.5 mils is deposited on the surface of the hottest fuel plate and that the 0.007 Ak step insertion is used instead of Technical Specification limit of 0.006 Ak.
REFERENCES
- 1. M.M. El-Wakil, " Nuclear Heat Transport," International Textbook Co. ,1971.
- 2. W.L Woodruf f, "The PARET Code and the Analysis of the SPERT-I Transients,"
Proceedings of the International Meeting on Research and Test Reactor Core Conversions from HEU to LEU Fuels, ANL/RERTR/TM-4 (September 1983)
~~ '
-QUESTION NUMBER 6-The Group 3 and Group 4 control rod constants (cross sections) were adjusted
'to force control blade worth agreement with measured worths.in the-present core.
The proposed core is expected to have a-harder neutron spectrum, resulting in changes to the core leakage and leakage' energy ' spectrum. . What evaluations have been completed to demonstrate that the cross sections for- the current core are sufficiently. valid for, the proposed- cores that control. binde worths and reactivity '
conditions will be ' consistent with Technical Specifications?
ANSWER:
The control rod cross sections have been evaluated using AMPX-II modules.
First, the cell averaged cross sections were evaluated for the control blade materials using NITAWL and XSRDNPM (transport code) of the AMPX code system (see Reference 1). This set was then spatially weighted using zone dependent neutron spectrum. .Using this data in the form of IS0TXS file, the reactivity worth of the control blades in the BOLD VENTURE model was checked against their measured worth.
The control blade cross section set was found to give a total reactivity worth for the control blades- that was too high. BOLD VENTURE can Nse " rod constants" (internal extrapolated boundaries) to model black absorbers, where the constant Cg defines the following boundary condition in group g at the face of the black absorber zone:
-D 64
= 9 9 C
g 6x
&g where D E diffusion coefficient in energy group g for the 9 material adjoining the black absorber i.
l l $ E flux in energy group g 9
l L
C i
QUESTION-NUMBER 6 (cont'd)
The control blades were then modeled using the cross sections data on the ISOTXS file for energy groups 1 and 2. A rod constant of 0.08406 was used for energy group 3, which was calculated using equation 11.5 of Reference 2. The !
energy group 4 (thermal) rod constant was adjusted to give good agreement between calculated and measured values, with a value of 0.2. These cross sections and rod constants are used in the rest of the BOLD VENTURE diffusion calculations.
i Since the measured blade worths for the proposed core with the new elements are not available, corresponding new rod constants are not determined yet. The proposed core with the newly designed fuel elements will have a higher U-235 loading in the middle portion of the core and a lower loading in the inner and outer portion of the core. The burnable poison will be loaded in <., elected inner and outer plates. To see the effect of spectrum change of the proposed core on the control blade reactivity worth, several XSRDNPM runs were made and the zone weighted B-10 (infinite dilution) microscopic cross section data in the control rod gap (space between outer pressure vessel and beryllium reflector) for the two different spectra were obtained. Table 6.1 compares the generated B-10 micro-scopic absorption cross sections (in barns) for the current and the proposed cores. It was found that due to the more hardened neutron spectrum in the proposed core, B-10 microscopic absorption cross sections at the control blade gap become lower than those of the current core by a maximum change of 5.4%. Due to j the B-10 self shielding the effect would be even less if the B-10 was modeled at the number density corresponding to control blades.
Four group neutron flux levels (in neutrons /cm2 sec) in the control blade gap I I
were also calculated (see Table 6.2) using BOLD VENTURE. With the slightly harder )
) spectrum and higher U-235 loading, BOLD VENTURE calculates a total rod worth 6%
less for a core of 1270 gm elements compared to the current core. Based on Table {
6.1 results, using the same rod constants should cause less than a 4% additional )
1 l
1 l
I 1
QUESTION NUMBER 6 (cont'd)
. error in the modeling of the control rods. This error will result in slightly overpredicting the rod worth in the proposed core model.
Technical Specification 3.1 has three requirements that either directly or indirectly apply to the reactivity of the control blades. Technical Specifica-tions 3.1d, 3.le, and 3.lf place limits on maximum reactivity insertion rate, .
l shutdown margin, and excess reactivity. Meeting these requirements has to be i based on measured rod worth. The reduction in total rod worth will cause a cor-responding change in the reactivity insertion rate. This is based on no change being planned in the speed of the control blade drive mechanisms. A change to the shutdown margin Technical Specification has been requested to base it on a value (0.005 Ak). that provides a sufficient safety margin. The excess reactivity of every core is verified to be within Technical Specifications based on the measured rod worth curve and cold clean critical rod height.
REFERENCES
- 1. Soon S. Kim and Charlie McKibben, "MURR Upgrade Neutronics Analysis Using AMPX-II/ BOLD VENTURE IV Computation System Benchmarked
, to the Destructive Analysis of Fuel Element 775F3," Internal Report, University of Missouri Research Reactor Facility (September 1986).
- 2. G. H. Kear and M. H. Ruderman, "An Analysis of Methods in Control Rod Theory and Comparisons with Experiment," GEAP-3937, Vallecitos Atomic Laboratory, General Electric Company, San Jose, California (May 1962).
i
Table 6.1 Energy Group Current Core Proposed Core % Difference 1 0.282 0.282 - 0.0 2 2.66 2.59 - 2.63 3 160.80 152.08 - 5.4 4 3003.70 2896.60 - 3.6 Table 6.2 Neutron Flux at Control Rod Gap Energy Group Current Core Proposed Core % Difference 1 6.22 x 10+13 6.76 x 10+13 + 8.7 2 9.68 x 10+13 1.05 x 10+14 + 8.5 3 8.35 x 10+13 8.70 x 10+13 + 4.2 4 1.69 x 10+14 1,69 x 10+14 + 0.0 i
Keff 1.112 1.093 Total Loss Rate 6.82x1017 6.94x1017 Production Rate 7.58x1017 7.59x10 17 l
I o - _ _ _ _ _ _ _ . _ _ _ _ _ _ _ _ _
QUESTION '!! UMBER 7 The nuclear peaking for all fuel combinations was evaluated with a water-filled flux' trap. Maximum peaking does not occur on Plate 1 for the. proposed ele-ments. What are the peaking factors in a core composed of the proposed fuel with experiments allowed within technical specification limits?
ANSWER:
The nuclear peaking factors for all fuel combinations were evaluated using a BOLD VENTURE (Reference 1) model with a water-filled flux trap to be conservative.
The highest overall peaking considering all possible cores is Plate 1 of a new 775 gm element. This peaking factor is highest when there is only water inside the fi.ner wall of the pressure vessel. When a sample-filled flux trap is located at the core center, the MURR core will have a more hardened neutron spectrum.
Starting with a MURR R-Z model of the core composed of the proposed fuel, the flux trap with samples was added to the model. In the model, the three barrel flux trap was modeled as one single tube with volumes of the three tubes combined into one big tube in such a way that a single ring of aluminium has the same volume as that in the three Aluminum tubes of the flux trap. To represent the typical reac-tivity worth of a flux trap sample loading, the volume inside the ring was modeled as aluminum with a B-10 number density adjusted to give an overall flux trap reac-tivity worth, which is greater than the maximum or minimum reactivity worth (0.006 Ak) specified in the MURR Technical Specifications. Table 7.1 summarizes the R-Z results for the rods out and the rods half-in cases. In the table," Ref."
represents a case where there is no flux trap in the core model, and is the refer-enced kef f corresponding to the zero reactivity worth for the flux trap. '
Based on the BOLD VENTURE R-Z results, it was found that the addition of experiments in the flux trap tends to depress the power on the inner plates and l raise the power on the outer plates. However, the magnitude of the change is not significant. With all rods out, Plate 23 has the highest Average Power Density
3
-o QUESTION NUMBER 7 (cont'd)'
~
( APD) and the- maximum change is plus 1.8% when e conservative negative 0.0065 ok is applied. ' Local hot spot Maximum Power Density (MPD) at Plate 23 also increased by 2.94% due to the negative . insertion of 0.065 Ak. .Due to the power shif t, when the
~ control rods are-inserted half way into the core, Plate 4 is shown to have the maximum APD. Insertion of either a positive or negative reactivity flux trap simply . lowered the APD of Plate.4,: with.the negative flux trap causing the largest decrease in APD.
As shown in table-.7.1, a maximum 2.9% increase in the local hot spot peak power .
density (543.4 ' watts /cm 3
) was also shown in Plate 23 when a a flux trap of -0.0061 ok reactivity worth is inserted.
. It was reported that in the detailed 0-R calculations, a peak power occurred at Plate 7 in the 1270 gm fuel element, with Plates 6 and 20 the next highest,
- respectively. Including the flux trap will reduce the power on Plate 7 and will increase the power at Plate 20L(see Table 7.2). In the table, "Ref." case represents the model with water instead of flux trap and samples inside the inner pressure vessel region. According to the table, the local hot spot power peak is reduced when the flux trap experiment is introduced; however, maximum APD is moved from Plate 7 to 20 and increases about 2.2%. The maximum power density in Tables 7.1 and 7.2 was found to be 543.4 watts /cm 3 , which is far less than the APD of Plate 1. in the fresh 775 gm fuel element (see answer to Questions 2 and 10). The presence of the flux trap with samples will depress the power on Plate 1.and increase the power on outer plates', but the change is not significant. The peaking factors for this change are covered by the envelope of the worst power peaking factors derived in the answer to Questions 2 and 10.
REFERENCES
- 1. Soon S. Kim and Charlie McKibben, "MURR Upgrade Neutronics Analysis Using j
AMPX-II/ BOLD VENTURE IV Computation System Benchmarked to the Destructive Analysis of Fuel Element 775F3," Internal Report, University of Missouri Research Reactor Facility, September (1986).
N,S '
a
- ^
- Q'JESTION NUMBER 7 (cont'd) . . .
Table.7.1 R-Z BOLO VENTURE Result (1270 gm Element Core)
All-Rods-Out Rods Half-In-Tech Spec Limit (6k) .+0.006 0- -0.006 +0.006 .0 -0.006 l
L Equiv.1 B-10 1.48x1019 Ref. 7.15x1019 1.17x1019 Ref. 5.085x1019 Number Density, k-effective 1.0995 1.0929' 1.0864 '1.0253 1.0191< 1.0130' Model Ak 0.0066 0 -0.0065 10.0062 0 -0.0061 a.
APD0 4t.
Plate 4 '323.2. -343.1 310.3 357.2 377.2 344.2-Plate 23 372.1 370.7 377.4' '313.0 308.4 318.1 Hot: Spot:MPDb 468.1 461.8 475.4 532.5 528.0 543.4 at Plate 23 23 23 23 23 23 aunit of L the Average Power Density (APD) is watts /cm3.
bunit of the Maximum Power Density (MPD) is watts /cm3.
Table 7.2 0-R BOLD VENTURE Result
-(1270'gm Element Core)
-Tech Spec Limit .
( Ak) +0.006 0 -0.006 B-10 0.62x1019 Ref. 5.1x10 19 k eff' 1.0685- 1.0621 1.0557 Model ok 0.0064 0 -0.0064 APD at Plate 4 309.0 322.7 298.8 Plate 7 329.4 331.2 325.9 Plate 20 334.5 329.9 338.5 Hot Spot-MPD 483.9 494.4 482.8 at Plate 7 7 20
t
.. b,
. QUESTION NUMBERL8 L
'l The cross sections for the reactor materials 'are. dependent on temperature and
- fuel _burnup. How were corrections made to the cross' sections used for the' bench -
t- mark to represent spectra 1' changes and different operating temperatures as a con-
. sequence of 10'MW' operation and increased oxide thickness?. Were calculations for mixed" cores performed with cross' sections' representative of: burnup? What.are the cross /section . sensitivities to temperature and fuel. and boron' depletion, and.
what are' the effect on calculated control blade worths and power peaking factors?
Are these consistent with your; Technical . Specifications? .
- ANSWER:
The cross sections for the MURR materials were generated using the modules
- in AMPX-II. The source is a collapsed 27 neutron group cross:section library The library is originally developed at ORNL for the criti-
' derived from CSRL-4 cality safety neutronics analysis and is based 'on ENDF/B-IV data. For the. thermal data, it uses Maxwellian weighting for a temperature of 293*F. To briefly describe the ' procedure (referring to Figure 1 in Reference 1), . AIM and AJAX were first pro- ;
' cessed to read and select the cross section. data of the MURR materials. Resonance t
. self-shielding calculations- were taken care of by NITAWL. The module.uses Nordheim
)
Integral method to treat the resolved resonances. The cell / zone weightings were
'l performed using one-dimensional neutron transport theory code, XSDRNPM. Hetero-geneous fuel cell including fuel meat, aluminum cladding and water moderator was set up.. and'the equivalent reaction rates when the cell is modeled homogenized were evaluated. Uranium content in the fuel meat is based on 1.55 gm uranium per cm3'. loading with a 93.14% enrichment. In zone weighting, the MURR reactor was I
modeled as a set of homogeneous zones where the cell weighted cross sections were used, e.g., fuel cell cross sections for the core zone. A four energy group nuclide I
ordered IS0TXS file that was collapsed from the 27 neutron group cross section j library is the final product of the XSDRNPM runs.
QUESTION NUMBER 8 (cont'd)
The HURR' depletion calculations have been performed using a member of the BOLD VENTURE modules, BURNER. The user provides primary data required to deplete core materials and the module solves nuclear chain equations to estimate nuclide concentrations at the end of each exposure time. In the MURR depletion model, eight heavy metals and seven fission product materials were used.- These are U-234, U-235, ')-236, U-238, Pu-239, Pu-240, Pu-241, Pu-242, and Xe-135, Pm-147, Pm-148, Pm-146m, Sm-149, NSFP (Non-Saturating Fission Product), and SSFP (Slowly Saturating Fission Product). Also added are B-10, B-11, Be-9, and buildup of Li-6 to allow one to treat neutron absorber in the fuel and/or control rods, and to simulate (n, .2n) reaction in the beryllium reflector. Nuclides at fuel zones in the core are exposed to the zone-average flux. The MURR depiction calculations used a R-Z model consisting of three different fuel zones in Z-direction and multiple fuel zones in R-direction. Radially, several of the middle fuel cells which have the same loading densities and approximate power density were grouped into one fuel zone. The 775 gm element core model has 7 radial fuel zones and the 1270 gm element core model has 14 radial fuel zones. Therefore, MURR depletion calcula-tions are based on one IS0TXS cross section file based on 293*K and the burnup :
dependent nuclide number densities calculated by BURNER. The MURR depletion model l
I was tested against the burnup measurements obtained by post-irradiation examination of the 775-F3 fuel element (see Reference 1). The BOLD VENTURE result showed good agreement with the measured data, giving confidence to the accuracy of the core 1
model and the cross section data. {
No corrections were made to the cross sections used for the benchmark to ]
represent spectral changes and different operating temperatures as a consequence l of 10 MW operation and increased oxide thickness. As a consequence of the power i
increase from 5 MW to 10 MW, there is an increase in operating fuel temperature, l
If there exists an error in evaluation of the core reactivity and other parameters j J
l.
l QUESTION NUMBER 8 (cont'd) due to this change, the magnitude is very small. Using a cross section set based on 293*K for all fuel zones is more conservative than using actual temperature
! dependent cross sections. The highest power density (hot spot) is overestimated by assuming no variation in fuel meat temperature. To perform a sensitivity analy-l sis, we have obtained a more flexible cell code, EPRI-CELL, from ANL. It has cross section libraries for different temperatures so that unlike the CSRL library in AMPX, the user can select different temperatures for fuel, cladding, and cool-ant. The original version of EPRI-CELL was designed for power reactor analysis, but ANL intensively modified the code to apply it for use in the plate type fuel analysis in research reactors like MURR. Validation of the code and benchmark calculations are well described in Reference 2. A new seven neutron energy group structure was set up for EPRI-CELL as shown in Table 8.1. The new set has four thermal groups which describe in more detail the neutron behavior in the thermal range. This seven group structure was used for the core temperature coefficient calculations discussed in the answer to Question 9.
Using EPRI-CELL, two sets of fuel cross section data at 293*K and 350*K were generated to see the effect of temperature change on cross section data, core reactivity, and peak power density. It is to be noted that the thermal neutron group libraries for major materials in EPRI-CELL are available at 293*, 350*,
400*K, etc., and 350*K is approximately the average fuel meat temperature, Also noted is that Hydrogen cross section library for EPRI-CELL was at 296*K and not 293*K and this was used in the temperature coefficient calculations. Tables 8.2 through 8.6 compare the data for the two different temperatures for the core materials, U-235, U-238, aluminum, hydrogen, and oxygen. As shown in the tables, as a con-sequence of the temperature increase, thermal absorption cross sections of i fuel, cladding, and coolant materials decrease. Also decreased is the thermal }
I '
1
QUESTION NUMBER 8 (cont'd) fission cross section in U-235. Using BOLD VENTURE, the seven group ISOTXS file was input in the current 6.2 kg core model to calculate the effect on the power peaking and core reactivity (see Table 8.7).- As the result shows, the effect of the temperature increase is that due to the spectrum hardening, the keff is
.slightly lowered by 0.35%, and the peak power density is also reduced by 0.71%.
Therefore, it can be concluded that the core analysis based on room temperature is more conservative than that based on core operating temperatures.
Normally oxide buildup at the aluminum cladding is not considered in the fuel cell cross section generation because' initial fresh fuel does not contain an oxide layer. However, the oxide layer during fuel depletion must be considered in thermalhydraulic and transient analysis and the related answers are provided in other parts of this report.
The depletion calculations for " mixed" cores were not performed with cross sections representative of specific fuel burnup. Calculations for " mixed" cores were performed with cross sections produced at conservative cold clean core conditions from AMPX modules and the number densities representative of burnup evaluated from the BURNER depletion option in the BOLD VENTURE calculations.
The effect of fuel burnup on power density was calculated from the BOLD VENTURE depletion runs. Table 8.8 shows the peak power density vs. burnup for the current and the proposed MURR cores with 10 MW operation. It should be noted that the proposed core contains boron in the front five and back four fuel plates.
The core model and control rod movement scheme for the calculations are described in Reference 1. For both cores, the trend is that the power peaking decreases as {
fuel burnup increases, which indicates that use of power peaking factor at cold clean condition is conservative. Table 8.9 shows U-235 content in the current 6.2 )
i kg core for different fuel burnup and the U-235 and B-10 contents for the proposed j
) core fuel burnup are given in Table 8.10. )
t 1
l l
_ - _ _ _ _ _ _ - _ - - - - _ _ - -- 1
. QUESTION NUMBER ~ 8 < (cont'd)'
The effect of- fuel- burnup on' control blade worths was also evaluated. Forlthe.
~
. current' core, the worth of all' four shim blades was -0.175 Ap (-0.1814K), whereas at 1600 MWD, -0.202 Ap (-0.171'4K) was obtained. :In-the proposed core, total worth was calculated to be -0.168- Ap (-0.170aK) at beginning of life and -0.191 Ap
(-0.169 ' ' AK). at the end of 2400 MWD .burnup. This trend is consistent with the MURR-Technical Specifications.
Neutron spectra in the proposed core will L be'different from those of the current core because of the different fuel loading. To see the effect of this, a sensitivity analysis was performed using AMPX/ BOLD VENTURE. Using modules in AMPX, l
two cross section libraries were produced. The first one is the one used for the.
current core analysis, but the other one consisted of the heavily loaded fuel plates in the middle- portion of the proposed core. In generating the cross section library (1st set) for the current core, uniform distribution of the current 775 gm U-235 fuel was assumed and one set of U-235 cross section was obtained. For the' proposed core, a cross section library (2nd set) that has three sets of U-235 cross section data was generated, one each for the inner region, the middle, and the outer region of the core. For inner and outer regions, the same loading density of the current core was' applied and a loading of 3.0 gm uranium per em3 was used in the middle portion of the core. Table 8.11 compares result of the proposed core when either of the two cross section libraries was used. As shown in the table, when the cross section library for the current core was employed in the proposed core calculation, a slightly increased keff (+1%) was obtainbd along with an increased power density, showing that use of the first cross section library for the proposed core is more conservative.
Cross section sensitivities to fuel and boron depletion were also performed using EPRI-CELL with the AMPX four group structure. The cases for with and without boron in fuel cell were considered in the depletion calculations. Microscopic L______-___
~
l: ; .
QUESTION NUMBER 8 (cont'd) cross. sections of the core materials ice two burnup steps of 0 and 400 MWD with and without boron cases are presented in lebles 8.12 through 8.17. According to the EPRI-CELL results, changes in microscopic coss sections are much less than the changes. in material number densities during fuel deoletion. The effect of boron in the fuel is that the harder spectrum in the fuel cell with the boron causes i
less mbsorption in the core materials when compared to the fuel cell without boron.
1 i
REFERENCES (
i
- 1. Soon S. Kim and Charlie McKibben, "MURR Upgrade Neutronics Analysis Using AMPX-II/ BOLD VENTURE IV Computation System Benchmarked to the Destructive i Analysis of Fuel Element 775F3," Internal Report, University of Missouri Research Reactor Facility, September (1986).
- 2. Research Reactor Core Conversion from the Use of Highly Enriched Uranium to the Use of Low Enriched Uranium Fuels, I AEA-TECDOC-233, IAEA,1980.
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- ~ - - - - - - - - - - ----
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I < Table 8.1 p Neutron Group Structure
, y Group' Upper Energy Lower Energy Bound-(ev)- Bound-(ev) 1 '10 x 10 7- 8.2085 x 10 5
- 2 8.2085 x 105 5.5309 x 10 3 3- 5.5309 x 10 3- .1.855 4 1.8550 6.2493 x 10-1 V
5 6.2493 x 10-1 '2.5104 x 10-1 ,
6 2.5104 x 10-1 5.6925 x 10-2 7 5.925 x 10-2 1.0 x 10-6 H
+<
Table 8.2 0-235 Microscopic Cross Sections -(bar ns)
' Absorption Nu-Fi ssion Scattering . Transport Energy Group 293*K 350*K 293*K 350*K 293*K' 350*K 293*K 350*K 1 1.3067 1.3067 3.3989 3.3989 5.9630 5.9630 4.5539 4.5539 2 2.1124 2.1124 4.0229 4.0229 8.8610 8.8610'. 9.0679 9.0679 3- 35.33 35.46 53.94 54.07 12.138 12.138 46.773 46.893 4 55.98 55.99 115.14 115.16 12.023 12.023 54.396 54.408 5- 146.82 147.59 298.91 300.32 11.862 11 860 137.93 138.56 6 277.31 276.10 566.74 564.18 11.607 11.608 278.98 277.96
'7 608.67 598.81 1260.4 1240.1 10.969 10.984 566.19 , 559.59
)
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= _ _ _ _ _ _ . _ _ _ _ _ _ _ _ _ _ _
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Table 8.3 U-238 Microscopic Cross Sections (barns)
Absorption Nu-Fission Scattering Transport Energy Group 293*K 350*K 293*K 350*K 293*K 350*K 293*K 350*K 1 0.4463 0.4463 0.3877 0.3877 6.9524 6.9524 4.6504 4.6504 2 0.2623 0.2623 - 3.6x10-4 3.6x10-4 10.886 10.886 9.1332 9.1332
~3 29.79 29.88 7.11x10- 5 7.11x10-5 29.59 29.594 58.17 58.26 4 0.5175 0.5175 3.91x10-8 3.91x10-5 8.5740 8.5)40 9.0673 9.0672 ,
5 0.7175 0.7185 0.0 0.0 8.4590 8.4590 9.15?9 9.1528 6 1.2596 1.2555 0.0 0.0 8.2770 8.2780 9.5114 9.5081 7 2.4182 2.3839 0.0 0.0 7.8226 7.8330 , 10.19 10.17 Table 8.4 Aluminum Microscopic Cross Sections (barns)
Absorption Nu-Fi s sion Scattering Transport Energy Group 293*K 350*K 293'K 350*K 293*K 350*K 293*K 350*K 1 6.21x10-3 6.21x10-3 0.0 0.0 2.891 2.891 1.966 1.966 2 2.71x10-a 2.71x10-3 0.0 0.0 4.023 4.023 3.549 3.549 3 6.49x10-3 6.49x10-3 0.0 0.0 1.399 1.399 1.382 1.382 ,
4 3.57x10-2 3.57x10-3 0.0 0.0 1.398 1.398 1.399 1.399 5 5.78x10-2 5,79x10-2 0.0 0.0 1.385 1.385 1.408 1.408 6 1.07x10-1 1.07x10-1 0.0 0.0 1.366 1.366 1.439 1.439 7 2.12x10-1 2.09x10-1 0.0 0.0 1.327 1.328 1.505 1.502 l
1
Table 8.5 Hyrodgen Microscopic Cross Sections (barns)
Absorption Nu-Fission . Scattering Transport Energy Group 296*K 350*K 296*K 350*K 296*K 350*K 296*K 350*K
'1 3.48x10 3.48x10-5 0.0 0.0 2.951 2.951 1.573 '1.573
- 2. 1.44x10-4 1.'44x10 0.0 0.0 11.24 11.24 3.061 3.061 3' 8.82x10- 3 ' 8.81x10-3' O.0 0.0- 17.91 20.21 6.405 6.404
'4 5.2x10-2 5.20x10-2 0.0 0.0 9.263 9.363 9.169 9.315.
'5 8.57x10-2 8.59x10-2 0.0 ' O.0 12.66 12.78 12.6- 12.64 6 1.645x10-1 1.64x10-1 0.0 0.0 22.51 22.67 21.86 21.86 7 3.59x10-1 3.52x10-1 '0.0 0.0 45.45 45.89 42.63 42.61 1
l Table 8.6 0xygen Microscopic Cross Sections (barns)
Absorption Nu-Fission Scattering Transport Energy Group 293*K 350*K- 293*K 350*K 293*K 350*K 293*K 350*K 1 9.2x10-3 9.2x10-3 0.0 0.0 2.357 2.357 1.879 1.879 2 0.0 0.0 0.0 0.0 3.905 3.905 3.569 3.569
.3 6.7x10- 6 6.7x10-6 0.0 0.0 3.642 3.643 3.539 3.539 4 2.8x10-5 2.8x10-5 0.0 0.0 3.780 3.780 3.621 3.621 1
5 4.6x10-5 4.6x10- 5 0.0 0.0 3.804 3.805 3.645 3.645 6 8. 9x10- 5 8.9x10-5 0.0 0.0 3.858 3.858 3.696 3.696 i
7 1.91x10-4 1.88x10-4 0.0 0.0 4.065 4.057 3.881 3.875 i
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>4 ' Table 8.7.
i' ' : Comparison .of Keff and Peak Power Density ' Lj f
, 293'K 350*K .% Change.
-Keff' 1.09027 1.08639. ,
-0.35
' Peak' Power
'! Density ,(w/cm3 ) 700 695- -0.71 I
" Table 8.8 Peak Power Density vs. Fuel Burnup
' Burnup (MWD) Current Core (w/cm3) Proposed Core (w/cm3) s ,0 799.8 518.3
~
400 707.8 --
.. 600 --
500.8 :!
800 607.9 --
1200 --
492.9 o
1600 564.5 --
l 1
1800 --
481.1
- - 4 2400 --
485.0
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Table fl.9 e /
4 U-235 Content vs Fuel Burnup for The Current Core - 1' s 7 a' d' Incremer.tal Burnup (MWD) U-235 (Kg) . ,.f!/ ' Loss in U-235 (g) ti 4 .
'O 6.2 0
L p,
o q
400 5.7165 .' 483.5?' f, /g '
-800 5.2352 a 481;3 /'
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1200 4.7569 478.3 1600 4.,2819 / 475.0 lI f ,; ,1 1
4' Table 8.10 .i
,i U-235 and B-10 Contents vs., Fuel Burnup for The Proposed Core
.i y ,.
Incremental i Loss inU-235-(g) /
Burnup (MWD) U4, .!3 (Kg) B-10 (g) 0 10,#f6 1.583 0 !
600 9,A211 O.714 738.9 1200 8.6898 0.342 734.3 1800 759583 0.167 728.5 2400 7.2357 0.078 722.6 1 ,
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1 kery 1.092ti8 , 'l.08158- )
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Peak Power 7 ./ Density (w/cm3) 46:!.4 464.0 y
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Table 8.12 U-235 Microscopic Cross Sections (barns)
Depletion with/without Boron in Fuel Energy Case Group Absorption Nu-Fi ssion Scattering Transport 0 MWD 1 1.3068 3.3988 5.9629 4.5532 with Boron 2 2.1170 4.0219 8.8601 9.0649 3 36.973 59.301 12.126 46.724 4 310.09 637.80 11.359 217.49 400 MWD 1 1.3068 3.3988 5.9630 4.5534 with Boron 2 2.1119 4.0223 8.8604 9.0658 3 36.711 59.075 12.126 46.478 4 311.47 640.81 11.368 218.11 0 MWD 1 1.3068 3.3988 5.9630 4.5533 without Boron 2 2.1118 4.0220 8.8601 9.0652 3 37.036 59.431 12.127 46.772 4 317.75 653.89 11.395 222.20 400 MWD 1 1.3068 3.3988 5.9630 4.5534 without Boron 2 2.1120 4.0223 8.8605 9.0660 3 36.756 59.168 12.127 46.511 4 316.97 652.22 11.393 221.45
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4 U-238'MicroscopicCrossSecNni(barns)
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f s Dipleticiii6dwithout Berwi f n Fuel
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'!' Case hprption St.attering Transport lg Nu-fission (
- 0 ' MD ' 1 ' 'i 0.4462 1.0851 6.9524 4.6497
.yp
? 0.2622 8. Wx'10- 4 .!0.834 9.1305
>Eoron 26.142 1.4503.)-4 I 27.283 36.228 A.(.n 3 g
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4 1.3528
)
0.0 8.1009 9.4177 4 3 ..
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3 400 IND 1. O. 462' 1.0005 7 5.9525 4.6498 with 4 I g4 8.693r10-4 10.885 9.1313
' nt<:n c 2 0. 2C!2 i
s
\ 3 25.454 .
1.483:10-4' 27.267 35.894 g
, '4 1.35MI. h 0.0 -
t 1069 9.4277
, .. ._ L . . ,
4 c,:
.' 0 :IWD ' 1 0.4462 -
1.986.i' 6.9524 4.6497 w Mheut Bt.rco ' 2 40.2622 8.6940r.10-4 10.885 9.1307
'I 3 26.143 '
l 4882x10-4 27.2450 36.095 J .J 4 1.3816 0.0. 8.1263 9.4693 i
i eAAu M n 1 0.a462 i 1.0866 6 s9525 4.6499 without I Boron 2 0.2622 8.6929x10-4 10.885 9.1314 3 2'i,492 1.1789x10-4 27.240 36.801 l
4 1.3783 0.r 8.1248 9.4644 t
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-Table 8.14 Aluminum Microscopic Cross Sections (barns):
Depletion with/without Boron Lin Fuel .
Energy .. .
Case Group: ~ Absorption Nu-Fi ssion Sca ttering - ' Transport 0 MWD 1 1 6.203x10-3 0.0 2.8915 1.9655-
, with-Boron 2 2.714x10-3 0.0- 4.0229- 3.5492
- 3. 9.347x10-3 0.0 1.3988 .1.3832 4 1.166x10-1 0.0 1.3493 1.4312 x
400 MWD 1 6.203x10-3 0.0 2.8915 1.9656 with Boron 2 2.714x10-3 0.0 4.0229 3.5491 3 19.402x10 0.0 1.4181 1.3833
-4. 1.169x10-l' O.0 1.4669 1.4322 0 MWD. 1- ' 6.2038x10-3 0.0 2.8915 1.9656 without Boron 2~ 2.7141x10-3 0.0 4.0229 3.5492 3 .9.4057x10-3 0.0 1.3989 1.3834 4 0.1190 0.0 1.3521 1.4362-400 MWD 1 6.2046x10-3 0.0 2.8915 1.9656
,without Boron 2 2.7143x10-3 0.0 4.0228 3.5491 3 9.4432x10-3 0.0 1.3988 1.3834
- 4. 0.1187 0.0 1.3519 1.4357.
1 i
8 e_________-_-_-.-_
Table 8.15 Hydrogen Microscopic Cross Sections (barns)
Depletion with/without Boron in Fuel Energy Case Group Absorption Nu-Fission Sca ttering Transport 0 MWD 1 3.4807x10-5 0.0 2.9512 1.5733 with Boron 2 1.4390x10-4 0.0 11.233 3.0585 3 1.3053x10-2 0.0 16.986 6.5485 4 0.1901 0.0 25.386 19.618 400 MWD ' 1 3.4807x10-4 0.0 2.9512 1.5732 with Boron 2 1.4394x10-4 0.0 11.233 3.0594 3 1.3129x10-2 0.0 16.965 6.5635 4 0.1905 0.0 25.432 19.635 0 MWD 1 3.4808x10-5 0.0 2.9512 1.5733 without Boron 2 1.4391x10-4 0.0 11.233 3.0586 3 1.3132x10-2 0.0 16.964 6.5625 4 0.1932 0.0 25.755 19.887 400 MWD 1 3.4808x10-5 0.0 2.9511 1.5732 without Boron 2 1.4394x10-4 0.0 11.234 3.0595 3 1.3186x10-2 0.0 16.949 6.5733 4 0.1928 0.0 25.697 19.827 l
l j- .:
- Table 8.16 0xygen Microscopic Cross Sections (barns)
Depletion with/without Boron in Fuel
~~
Energy l Case Group i,,7 6bsorption Nu-Fission Scattering Transport 0 MWD 1 9.213x10- 3 0.0 2.3578 1.8799 with Boron 2 0.C 0.0 3.9054 3.5696 3 8.868x10-6 0.0 3.6590 3.5481 4 1.022x10-4 0.0 3.9118 3.7412 400 MWD 1 9.214x10-3 0.0 2.3577 1.8798 wi th Boron 2 0.0 0.0 3.9053 3.5696 3 8.906x10-6 0.C 3.6588 3.5483 4 1.023x10-4 0.0 3.9113 3.7407 0 MWD 1 9.2140x10-3 0.0 2.3578 1.8799 without Boron 2 0.0 0.0 3.9054 3.5697 3 8.9073x10- 6 0.0 3.6587 3.5481 4 1.0381x10-4 0.0 3.9101 3.7395 400 MWD 1 9.2154x10-3 0.0 2.3576 1.8798 without Boron 2 0.0 0.0 3.9053 3.5696 3 8.9341x10-6 0.0 3.6586 3.5483 4 1.0357x10-4 0.0 3.9101 3.7396
)
i l
i
Table 8.17 4 Boron-10 Microscopic Cross Sections-(barns)
Depletion with Boron in F el Energy.
Group- Absorption Nu-Fi ssion Scattering Transport Case 0 MWD- 1 0.3820' 0.0 1.7241 11.7566 with 4.7455 Boron. 2 2.2070 0.0 2.9761
- 3 149.33 0.0 1.96/5 108.71-4 1883.9 0.0 1.9012 1453.6-400 MWD l' '0.3820 0.0 1.7240 1.7566 with Boron .2 2.2074 0.0 2.9760 4.7458 3 150.26 0.0 1.9673 109.38 4 1890.6 0.0 1.9028- 1456.6.
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QUESTION NUMBER.9L The proposed core.is expected to have different. neutron spectra different from
'the. current core. What are' the expected changes in~ core temperature and voiding reactivity coefficients? What impact will these changes have on conditions and !
postulated accident scenarios? j ANSWER:
The current and the proposed cores were modeled using the R-Z option in the l
BOLD VENTURE system to calculate the core temperature coefficient (T/C) and void reactivity coefficient (V/C). In calculating'V/C, the AMPX four group cross' I section set was used and a 10% reduction in water number density in the core was assumed to get in units of 2/% void. Two BOLD VENTURE runs were made for each core -
l using two different water densities in the coolant. The V/C of the proposed core ]
was calculated-to be -2.83 x 10-3 &/% void, which was lower in magnitude than the I V/C calculated for' the current core (-2.91 x 10-3 &/% void) by about 2.75%. Based on the measured value of the current core, -2.51 x 10-3 &/% ~ void, it is predicted that the V/C of-the proposed core will be -2.44 x 10-3 &/% void. This value is still more negative than the Technical Specification value of -2.0 x 10-3 &/t void.
In calculating T/C, the seven group EPRI-CELL cross section library, .which l 1
consists of four thermal and three fast groups, was used to better de3cribe the i behavior of the thermal neutrons. Hydrogen temperature change in the coolant was from 296*K to 350*K. Four BOLD VENTURE runs were made for both cores using four ;
'different cross section sets and the calculated T/C's for the current and proposed j cores were, respectively, -9.85 x 10-5 and -7.94 x 10-5 &/*F. In contrast to the slight change in the V/C, the proposed core has a 19.4% less negative T/C value than theicurrent core. A. low power testing program in 1971 (see Reference 1) determined j the measured value of T/C of the current ce*a is -7.00 x 10-5gf.F. This is an average of the 15 observed coefficients that range from -4.4 x 10-5 to l
-11.9 x 10-5 &/*F over a temperature range from 304*K to 334*K with an aver. age 1
u_._. _ _ _ _ _ _ _ _ _ _ . _ _
QUESTION NUMBER 9 (contd) temperature of 315.8'K. When the average ineasured value is reduced by 19.4%, the average core T/C' of on eight 1270 gm element core is predicted to be -5.64 x 10-5
&/*F, which is less negative than the Technical Specification value, -6.00 x 10-5
&/*F. A core of four 775 gm elements and four 1270 gm elements should have a T/C of approximately -6.32 x 10-5 &/ *F. Based on this, the V/C of the proposed core is consistent with the current MURR Technical Specifications, but the less negative T/C suggest a possible need for a Technical Specification revision before a core of more than four 1270 elements can be used. The basis for the T/C Technical Specification 3 la is section 3.5 of Addendum 3 to Hazards Summary Report (HSR). Analysis of Rapid Step Reactivity Insertions from Full Power in the MURR. The analysis con-cludes the MURR reactor can withstand a positive step insertion of +0.006 & reac-i
~
tivity without core dmage. This analysis has been redone for a +0.007 & step insertion using 2.5 mils of oxide and a V/C of -2.0 x 10-3 &/% void with no T/C. 1 The results show no core damage and are discussed in the answer to Question 5. In the future, a request to revise the Technical Specification on core temperature i coefficient of reactivity will be submitted to the Nuclear Regulatory Commission, which will include an expanded version of the reactivity transient analysis to replace all the reactivity transient analyses in the HSR.
REFERENCES j 1 Low Power Testing Program for the Missouri University Research Reactor j
6.2 Kg Core, MURR Internal Report, 1971.
l l I
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_ _ _ _ - _ _ _ __ _ _ - _ _ _ _ . 1
i i
% w:s QUESTION NUMBER'11.
The proposed core has different' nuclear characteristics, i.e., power peaking,
' fuel temperature, oxide thickness,' neutron spectra. What is. the reactor response to accident conditions with the proposed core that have not been analyzed and discussed previously?
lANSWER:
A thorough review of Hazards Summary Report (HSR) analyses was performed.
Several descriptive changes need to be made in the next annual report in accordance with 10 CFR 50.59. Changes to two of the most significant safety analyses are
.: included in the answers to other questions. The combined answer to Questions 2 and 10 contains the revisions to the' HSR Safety Limits ' analysis. The answer to Question 12 contains the revisions to the' Design. Basis Accident (DBA)' analysis to cover the.
1270 gm element.
The Loss of Flow Accident (HSR Addendum 4 Appendix D) and the Loss of Coolant Accident (HSR Addendum 4 Appendix E) do not need to be revised to cover the 1270 gm fuel element.. Since the dimensions for the 1270 gm element are the same as the 775 gm element,' thus there is no change in the hydraulics response of the core. The
'" hottest"< possible hot plate is Plate 1 of the 775 gm element on which both analy-ses are based. Neither analysis assumed any oxide thickness on the fuel plates.
The addition of- an oxide film would be insignificant in either analysis due to the relative low. heat flux of decay heat compared to full power. As given in Answer 5, the hot spot in the 1270 gm element at full power would have a 60.8'F differential i temperature across a " worst case" oxide thickness of 1,74 mils. Therefore the dif.ferential temperature due to decay heat would start out at less than 4*F and
- would decrease with time. This value is extremely conservative since it is based p
on a 1.74 mil maximum oxide located on the hot spot. The maximum oxide is really found on the almost depleted element.
. . o QUESTION NUMBER 11 (cont'd)
The results of a new reactivity transient analysis, given in answer to Question 5, show that the Technical Specifications limiting potential reactivity insertions to 0.006 Ak are still conservative. In the future a request to revise the Technical Specification on core temperature coefficient of reactivity will be submitted to the Nuclear Regulatory Commission. An expanded version of the reac-tivity transient analysis will be included with the submittal to replace the reactivity transient analyses in the HSR.
The following revision to the " Justification of the Acceptability of the 125%
l Power Level Scram for 10 MW Operation" is made to document applicabil1+.y to 1270 gm 1
l elements.
Add a note to the bottom of Table 3.2.1 on page 49 of HSR Addendum 3.
NOTE: The peaking factors from Reference 1 are conservative values for cores using new and depleted 775 gm U-235 fuel elements. The overall peaking factor (PrPaPOPn) for mixed cores of new and/or depleted 775 gm and 1270 gm elements, or new and depleted 1270 gm elements is less than the values given in the table. The values are given in Table F.2a of HSR Addendum 4 page F-Sa, which is included in this revision to the HSR.
l
F , J. e -
QUESTION NUMBER 12~
u The proposed elements will operate with peak power on' a. different plate that has'an increased fuel loading. What effect'does this.have on the DBA environmental
! ' considerations related to accident scenarios?
ANSWER:
y
.The. Design Basis Accident (Hazards Summary Report, Addendum 4) is a' sot of postulated conditions which lead to consequences worse than those resulting from any. other_ anticipated accident. The arbitrariness of what causes the accident is l
given in section C.2 of Appendix C to Hazards Summary Report (HSR), Addendum 4:
As a design. basis accident it is assumed that an accident condition has -led to the melting ~ of four, fuel plates in the reactor core. Because this accident is considered worse than any anticipated accident, the. conditions that lead to this cir-
- cumstance are immaterial to the analysis. The DBA may be postulated to result from partial flow blockage to the fuel.
However, the coolant strainer, the fuel end fitting and the pre-operational inspection of the pressure vessel and core following any fuel handling,' all . prevent an accident of this type. In addition, it has been shown that a 75% blockage of flow to the hot channel is insufficient to cause a clad failure (1).
It is assumed that the fuel plate melting occurs in the peak power region of the core; that is, four plates are chosen which represent the area of peak fission product inventory.
The assumption is melting of the #1 plate on four 775 gm U-23S fuel elements.
With an allowance for peaking, this is determined to be 2.08 percent of the fission products in a 6.2 kg U-235 core that has operated at 10 MW for 80 days (page C-5 of.HSR Addendum 4). The rest of the DBA analysis concerns the health effects of a release of this quantity of fission products, ;
The proposed fuel elements have a more uniform power density. This means a I
greater percentage of the proposed core has to melt to release 2.08 percent of the 1 fission products as compared to the 6.2 kg U-235 core. A core containing all $
proposed fuel elements would have more U-235 (10.16 kg); but this is not signifi-cant to the concentration of radioactive iodines., which are the fission products )
l of greatest health concern. Since the fission product inventory is based on 80 1 days of continuous operation, the fission product iodines are at equilibrium
\
)
,,,o QUESTION NUMBER .12 (cont'd) values, which are only a function of the power level. Therefore the DBA is even more conservative when applied to the proposed core than the 6.2 kg U-235 core it is based on. The following revisions are made to HSR Addendum 4 to account for both the 775 gm and 1270 gm fuel elements in accordance with 10 CFR 50.59.
- 1) Change the last sentence on page 6 to:
"The accident that is considered to constitute the most severe consequences to the health and safety of the public is the release of 2.08 percent of the fission products from the accidental melting of the #1 fuel plate in four 775 gm U-235 elements postulated to occur as a result of blockage of flow."
- 2) Add the following note af ter the " Design Basis Accident" heading and before the " Introduction" heading on page C-2:
" NOTE: The Design Basis Accident is based on the 6.2 kg U-235 core (eight 775 gm U-235 fuel elements). The analysis is valid for any pos-sible combination of 775 gm or 1270 gm U-235 fuel elements totaling eight elements, due to the highest possible power density occurring on the #1 plate of the 775 gm element."
- 3) Change second paragraph on page C-3 to read:
". . . four plates are chosen which represent the area of peak fission product iodine inventory."
- 4) Add the following note af ter the "C.4 Analysis" heading on page C-5:
"N OTE,: The fission product release is based on the melting of the #1 plate in four 775 gm fuel elements (1.3 percent meltdown).
The #1 plate of a 775 gm fuel element has the highest possible l
power density in cores made of any possible combination of 775 gm or 1270 gm U-235 fuel elements containing a total of eight elements. Due to the more uniform power density, a core
1
,elO QUESTION NUMBER '12 (cont'd) containing only 1270 gm U-235 elements would have less release of fission product iodines than the referenced 6.2 kg core from a 1.3 percent meltdown."
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