ML20002D107

Forwards Proposed Change 13 Request for Change to Tech Specs of License DPR-6,to Insert Six High Performance Developmental Fuel Bundles Designed to Explore Central Fuel Melting Regime

ML20002D107
Person / Time
Site:	Big Rock Point File:Consumers Energy icon.png
Issue date:	05/26/1967
From:	Haueter R, Wall H CONSUMERS ENERGY CO. (FORMERLY CONSUMERS POWER CO.)
To:	Morris P, Skovholt D US ATOMIC ENERGY COMMISSION (AEC)
References
NUDOCS 8101190449
Download: ML20002D107 (64)

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%.!.y V; May 26,1967 Dr. P. A. Morris, Director Re: Docket 50-155 Division of Reactor Licensing United States Atomic Energy Commission p, _, a s _ _ _, __ q, i:0 Cy*

Washington, D. C.

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Dear Dr. Morris:

Attention:

Mr. D. J. Skovholt Transmitted herewith are three (3) executed and nine-teen (19) conformed copies of a request for a change to the Technical Specifications of License LPB-6, Docket No. 50-155, issued to Consumers Power Company on May 1, 1964, for the Big Rock Point Nuclear Plant.

The proposed change (No.13) will enable Consumers Power Company to insert into the reactor at Big Rock Point six (6) high per-formance develaomental fuel burdles designed to explore the central fuel melting regime. These bundles are an essential part of the " General Electric-Atomic Energy Commission-Euratom Project on UO2 Fuel Operation With Central Melting in a Large Power Reactor (Contract AT(Oh-3)-189, PA 50)."

To allow insertion of these bundles, during the present refueling outage at Big Rock Point, requires approval of this request by June 12, 1967 The bundles are presently in the final stages of fabrication at the General Electric Company's facilities in San Jose-and can be shipped to meet the above date.

It is recognized that the time interval is extremely short and that we are acking for special handling of this request. To assist you in this matter, we are prepared to come to Washington and spend whatever time is necescary with your staff after they have had an opportunity to review this submittal.

We will appreciate eny special attention that you can give to this request. The AEC Reactor Development Division, Euratom and General Electric Company are extremely anxious to get this program under way.

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-We must apologize for not getting this request submitted earlier; however,'the physics, thermal-hydraulics and safety. analyses for this. design were very time-consumin6 -

Yours very truly, fa

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RLH/dmb Robert L. Haueter a

Attach.

Assistant Electric Production ~

Superintendent - Nuclear

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CONSUMERS POWEB COMPANY U

Docket No. 50-155 k"' t:rf N -.!.T;;;

Reque:st for Authorization of Change in Technic'al Spec fications License No. DPR-6 I.

Changes in Technical rpecifications For the reasons hereinaft. set forth, it is requested that the Technical Specifications Appended to Operating License No. DPR-6 issued to Consumers Power Company on May 1, 1964, for the Big Rock Point Nuclear Plant, be changed as follows:

A.

Replace paragraph at beginning of Section 515 (c) to read as follevs:

"The general dimensions and configuration of the types of fuel bundles shall be as shown in Figures 5.2, 5.3, 5.h, 5.5, 5.6 and 8.1 of these specifications. Principal design features shall be essentially as follows:"

B.

Add Figures 5 5 and 5.6 (8 x 8 and 7 x 7 Fuel Drawings).

C.

Add a column to the table in Section 5 1 5 (c) as follows:

General Research and Development Geometry, Fuel Rod Array 8x8 7x7 Rod Pitch, Inches 0.807 0 921 Ctandard Fuel Rods per Bendle 36 29 Special Fuel Rods per Bundle 28***

20***

Spacers per Bundle 5

5 Material Zr-2 Zr-2 Standard Rod Tube Wall, Inches

.035

.0ho Special Rod Tube Wall, Inches

.035

. oho

• • • Special rods have depleted vranium.

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2

' Fuel Rods Research and Development Standard Rod Diameter,' Inches-

-0 570-0 700 Special Rod Diameter,-: Inches.

0 570 0.700.

UO Density Percent Theoretical 9h Pellet

.t 2

85 Powder Active PueloLength, Inches 66-6713 65-66.3 Fill' Gas Helium Helium"

.D.

Add a new section as follows:

"5 1 9 Centermelt Test Fuel Bundles Six. fuel bundles may be operated at increased

-thermal output, up to and including various-

. amounts of center melting of the UO. The fuel 2

.has.bcen spect. 11y designed for this operation and is permitted to exceed the general core

-operating limitations of Section 5.2.1.(b) but

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vill be limited to the most conservative of the following values:

Fuel Type 8x8 7x7 Number of Bundles Pellet UOp 1

2' Powder UO -

1 2

2 Maximum Steady State 2

-Heat Flux, Btu /Hr-Ft 500,000 500,000 Maximum Steady State Fuel Rod Power, Kw/Ft 21.8

-l26. 8 Funim', Nec Burnout Pi>-

t Ge:erpower h

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1.5*

1.5" Rate of caange... Reactor Power During Reactor Operation:

" Control rod withdravol vill be limited as in Section-5 2.1.

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In addition, when the centermelt fuel is in the core, the following restriction shall be observed whenever (a) a i

centermelt fuel-bundle is being icought to full power for

. Based-upon new Critical Heat Flux Correlation, APED-5286 (see Proposed Change #12).

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the'first: time, or'(b) a scram recovery is being made at a time-in the xenon transient.such that the pe'k'of a

the' axial power distribution is lower in the core t'han the peak _ existing at1the time of the last chutdown:

When power is between-170 Mwt and 2h0 Mwt,

.the rate of~ power increase vill be limited-to an average _of 1/2 Mwt per minute."

II. -Centermelt Fuel Mechanical and Thermal Design

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A total'of 32h fuel rods vill be incorporated into six fuel assemblies for irradiation in the Big Rock Point reactor. Two of the assemblies, comprised of 64 rods each in an 8 x 8 configuration, vill be designed for incipient melting of UO at rated power conditions. These 2

assemblies have teen designated tlie intermediate performance fuel and vill

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cortain both powder and pellet type fuel. The remaining four assemblies, comprised of h9-rods each in a 7 x 7 configuration, vill be designed for

' definite but moderate melting of the UO at rated power conditions. These 2

assemblies have been designated the advanced performance fuel and contain i

both powder and' pellet type fuel.

'The basic fuel bundle design will be similar to that used previously.in the high power density research and development program.

The ability to remove and replace individual rods during various phases of the irradiation vill be of value during the operation of the program.

Individual rods can be examined visually, dimensional measurements can be made and selected rods can be shipped to the RML (Radioactive Materials Laboratory) of_ General-Electric Company for destructive examination at various levels of exposure without destroying the integrity of the assembly.

i The fuel vill be limited to the same critical heat flux ratio as the remainder of the core (MCHFR >l.5 at an overpower of 1.22).

The new critical. heat flux correlation, APED-5286, will be used as the I

basis for the calculations.

Fuel rod size vill be such that when operated at a rated

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surface heat flux of h50,000 250,000 Btu /Hr-Ft the intermediate performance w

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fuel vill operate with incipient central melting while the advanced per-formance; fuel vill operate.with definite but moderate centr'al melting.

A.

Fuel' Bundle Design-Both the advanced performance fuel (0 700" diameter) and t

the intermediate performance fuel (0.570" diameter) vill use the remov-able rod design used previously in the high-power density fuel development program. The 0 700" diameter rods are arranged in a 7 x T array.. Figure-5 5 describes the fbel assembly in detail. The 0 570" diameter rods are:

arranged in an 8 x 8 array. Figure 5.6 describes the assembly in detail.

~ Key components of the support structure are as follows:

1.

Handle - Completely removable to allow rod removal.

Handle has notches to permit visual identification of these bundles after loading into the reactor.

2.

Pin - Captures the handle to the support structure.

3.

Spacer - Double layer wire, constant pressure spring type.

Provides maximum protection against wear of the Zircaloy-2 cladding.

.h

, Angle - Acts as axial support member of bundle. Prcvides p css. t.,.e., g a means of.periti spacers axially. Also minimizes the possibility of dam-age to rods and spacers during insertion and removal from in-core channel locations.

5 Base - A grid bar arrangement velded to the lower end of 2 -

each corner angle. The fuel rods rest on the grid.

The support structure is fully capable of supporting the rods when resting on the base and when hung from the handle. However, when rest-ing on a side with a full load of rods, adequate support must be provided for individual rods to prevent damage to the spacers. This will be provided

'during the shipment of unirradiated bundles, and the method of packing and shipment has been proven to be adequate to prevent damage. Other pertinent design data are shown in Table 1.

I The fuel bundles in the fully loaded condition will be of the following approximate weights: (Includes weight of cage.)

I-1.

0.700 OD Pellet UO R ds - 385 Lb 2

2.

0 700 OD Powder UO Rods - 358 Lb f

2 3

0 570 OD Pellet Uo R ds - 330 Lb 2

h.

0 570 OD Powder UO R ds - 309 Lb 2

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5-s B.

Puel Rod' Design 1.

Intermediate Performance Fuel There are two intermediate performance fuel assemblies.

Each contains 6h' rods in an 8 x 8 configuration. Four U-235 enrichments are used-in Msch assembly and one assembly will contain pellet type fuel:

while the other will contain powder type fuel.

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The rods in the assemblies are clad with 0 570" OD.x 0.035" vall Zircaloy-2 tubing-and contain a compression spring in the plenum region to minimize axial fuel movement during handling and shipping.

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' depleted UO pellet is to be placed.at each end of the active-fuel column 2

.to minimize temperature effects at the lower end plug and'the plenum spring.

~The UO2 pellets will be dished to provide a 5% void volume to accommodate the phase change volume expansion.of the UO n melting. The UO powder 2

willibe compacted to an apparent' density of 85%-of theoretical. The 15%

void volume is adequate to accommodate the volume expansion which will occur in melting for these rods.

Fuel-clad mechanical interaction resulting from differential thermal expansion is considered as one of the' largest potential contributors to high. cladding strain. In an attempt to minimize this on a gross scale, a rather~1arge cold gap of 0.012" is built into the intermediate performance

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pellet' fuel. At the beginning of life at normal rated power, at least two mils of diametral clearance will be present in the rods.

A's burnup prc-ceeds, this clearance will decrease to about one mil at 15,000 Mwd /T.

2.

Advanced Performance Fuel There are four advanced performance fuel assemblies, each containing h9 rods in a 7 x 7 configuration. Two of the assemblies will contain UO in the form of sintered pellets and the other two assemblies 2

will contain UO in the form of compactible grade powder.

2 In the advanced performance fuel, there are three distinct

. designs. Two differ only in the form of the UO2 (p vder or pellet). The

-third design is the same except tungsten wafers are placed in the rods at approximately 18" intervals in the axial direction. The purpose of the L

' tungsten wafer-is to minimize axial movement of the molten UO during 2

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6 Joperation. 'It is tentatively planned to place tungsten vafers in eight of?theihigh: performance rods - four powder rods and four pellet rods.

LTungsten was chosen as the material for the'vafers because

of its compatibility with UO at-high temperatures. The melting point of 2

tungsten is 3370 C as compared to 2800 C for UO. The thermal conductivity 2

of.the tungsten wafer plays an important part in maintaining the adjacent UO2.below its melting point.

The.effect of the tungsten wafere inserted into some. of the 0.700" OD 7 x 7 array centermelt bundles has been calculated. The high neutron. absorption cross section of tungsten causes a calculated power de-pression-of 2 5% from the maximum value between wafers to the point adjacent '

to the wafer. The calculation performed with two-dimensional, r-z geometry diffusion theory'necessarily assumes symmetric boundary conditions which then assumes that all fuel rods contain wafers. This is not the case, and the actual power depression in the segmented rods will be negligible and lower than' the calculated value.

All rods in the advanced performance fuel are clad with 0 700" OD x 0.0h0" vall Zircaloy-2 tubing and contain a compression spring in the plenum region to minimize axial fuel movement during handling and shipping. The pellet fuel vill be fabricated with a 0.100" diameter central n melting.

hole to accommodate the phase change volume expansion of the UO2 F

The 15% void space is considered sufficient to accommodate the volume ex-pansion on melting in the powder fuel.

Fuel-clad interaction vill be minimized in the pellet rods by use of a 0.013" nominal cold gap.

At 450,000 Btu /hr-ft heat flux and i

beginning-of-life conditions, no gross interaction vill occu., It is,

. nterac+e a o o

possible, however, that.at 15,000 Mwd /T, a very slight degree could occur.

3

- The combined effects of increased fuel temperature and fission swelling are the cause of this.

3 Uo and cladding Temperature calculations 2

Temperature 5of the UO and cladding were calculated using 2

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the following input-information:

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.7 thermal conductivities UO2. Thermal Conductivity - The UO2

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use'd in this analysis are as follows:

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=

. Pellet Fuel:.K(T) = h02

+T This equation results in.an /

d

• 9

""0 '2800 C" 2800 C 0

500 C K(T) = 3 5g

,y + 5 2922 x 10-13 (T + 273)3 (2) 9

Powder Fuel:

2800 C 2800 C This equation results in an:/

= 63 W/cm, and 1 KdT = h9'W/cm.

0 500 C

.The' integral KdT for pellets and powder versus temperature is The pellet integral KdT curves are from Lyons, et al.(1) plotted in Figure 1.

The powder KdT meets the constraints as suggested by D. H. Coplin in Refer-ence (2), Pages h-8, and agrees with post-irradiation measurement data points in Reference'(1), Figure 7-3.

The slope of the powder. curves has a K(T) for powder in the range 500

'C to 1500 C approximately 25% less than for' pellet fuel and, in the range of 1500 C and 2800 C, K(T) is appioximately equal.to that for pellet fuel.

Specific Heat Generation Required for Melting - The specific heat generation required to cause melting for three conditions is:

8 x 8 Bundle 7 x 7 Eundle Pellet Powder

• Pellet Powder Beginning of Life 21.h Kw/Ft 19 1 Kw/Ft 2h.2 Kw/Ft 19.h Kw/Ft 10,000 Mwd /T 20.1 Kw/Ft 18.5 Kw/Ft 22.9 Kw/Ft 19 0 Kw/Ft 20,000 Med/T 20 5 Kw/Ft

-18.7 Kv/Ft 23 3 Kw/Ft 19 2 Kw/Ft e er e ra s

• Value for a = 0.16h where a =

= 0.16L pellet radius.

Pellet radius 1

Results - The computer results of the calculation of clad and UO temperature and thermal expansion are shown in the following figures. The 2

clad average temperature at any location on the rod is shown as a function of heat flux in Figures 2, 5, 8 and 11.

The fuel center line temperature as a function of heat flux is shown in Figures 3, 6, 9 and 12, and the UO2 **lt fraction as a function of heat flux is shown in Figures h, 7, 10 and 13.

Thermal Conductivity From Irradiation With Central Lyons, M.

F., et al, "UO2 Melting," GEAP h62h, July 196h.

Lyons,-M.

F., et al, "UOp Powder and Pellet Thermal Conductivity During Irradiation," GEAP-5100-1, March 1966.

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III.. Physics Design Analysis The centermelt bundle designs include four d'irferent U-235 enrichments which allow a statistically significant number of' fuel rods

-operating at 450,000 250,000 Btu /hr-ft without exceeding critical heat

- flux limits. The design goals were the following:

a. ' Adherence to the same MCHFR limit-as the rest of the Core.

.b.

Reactivity ~ coefficients which are as r.cgative as

.possible.

Bundle average exposure capability of 15,000 Mwd /T c.

vithout bundle modifications (burnable poison in some form).

A.

General Features Common to Both Designs Both the 7 x 7 and the 8 x 8 designs have four enrichments:

4.3, 5.0, 5.6 and 0.22 (depleted) wt percent. The enrichments and placement of the rods.were specified such-as to provide the maximum number of fuel pins operating at or very close to the desired heat fluxes. The general arrangement of the' various enriched rods is shown in Figures 1h and 15 The number of depleted UO r ds was decided from a thermal 2

limit MCHFR consideration,.where their low power production effectively

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increases the allowable power of the remaining enriched fuel' rods at the same overall bundle power output. The depleted rods were arrayed in a roughly annular zone inside the corner and side rods above. Then, the highest enrichment rods were placed in a sycmetrical arrangement near the center of the bundle. Their enrichment is sufficient to produce rod powers similar to the other outer power producing rods.

The bundles have adequate thermal margins throughout their lifetime based on typical "beginning" and "end of cycle" power shapes.

These power shapes are very important to the design and were deter =ined from typical recent operating shapes experienced at Big Rock Point and

'are the worst expected during normal operation.

Both bundle designs have been enriched to attain approxi-l mately 15,000 Mwd /T. average (20,000 Mwd /T average for the power producing l

rods) in a core of the same type. This is the standard method of

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9 specifying fuel bundle exposure capability, and it is possible to " push"'

any' desired bundle or bundles past the specified exposure cahability at'

.some expense in core reactivity.

" B.

Calculated Local Peaking Factors

.The-calculated local peaking factors for both the 7 x 7 and the 8 x 8 bundles are.shown in Figures 14 and 15 The local peaking factors'shown are for the new bundle, and the average rod power in each-case is'1.0.

Because: of plutonium production in each of the fuel rods, the peaking, factors of the depleted rods will increase with time and, because of normalization, the peaking factors for the enriched higher 1 power producing rods will decrease.

The relative power of the three high enrichment groups decreases with exposure, and the depleted. group increases. The fuel will be moved to positions with higher radial power factors as the irradiation proceeds. In this way, the design heat flux for the high enrichment roda.

can be maintained throughout the exposure lifetime. The bundles are capable of the' target' heat fluxes at 15,000 Mwd /T average exposure and

. typical-beginning of cycle power shapes without exceeding the critical ~

heat flux limits.

C.

Doppler Coefficient The Doppler. coefficient'for the 8 x 8-bundle design has

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.bcen calculated to be -0.86 x 10 Ak/k/ F at %1000 K.

This compares l

to the standard fuel designs (A, B and C) for Big Rock which-have values

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in the range of -1.0 to -1.1 x 10 ' Ak/k/*F. The 7 x 7 design is about

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-0.96 x 10 Ak/k/ F.

D.

Temperature and Void Coefficients

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The temperature and void coefficients for the cold center-melt bundles have been calculated. The coefficients are noticeably more positive than standard BRP fuel, as can be seen in the following com-i parison with BRP "B" reload" fuel presently loaded into the reactor.

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10 TemperatureL Void Coefficient

' Coefficient at 25* C at 20

'C (ak/k)/ C (ak/k)/ Unit Void Beginning of Beginning Life End of Life of Life

_End of Life "B" Reload With-

-6

~5

-2 out Cobalt

+3.20 x 10

+5 5 x 10

-0.28.

-3.8 x-10

-3

-Centermelt 8 x 8

-+1.0h x'10

+1.8 x -0.2h

+5.3* x 10

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-2 Centermelt 7 x 7

+5.60 x 10-

+1 7 x 10"

-0.27

-1.3 x 10

1. For 0 to 10% voids; the corresponding number for 0 to 20% voids is

-6.8 x 10-3, The small number of centermelt bundles in the. reactor core (six of an 84 total) vill produce a very small, if.even detectable,.effect i

on overa11' reactor behavior with temperature and void. The. total addition of reactivity due to any possible cold voiding of the two 8 x 8 bundles has been calculated to be 0.0000126 (ak/k). The temperature coefficient is negative at the operating temperature. It turns from positive to negative at about 130 C.

E.

Burnup Behavior Although the bundles do not resemble other designs in the core to any great extent, the behavior with lifetime is very -similar. The burnup slope for the 7 x 7 bundle is 0.0113 Ak/1000 Mwd /T and 0.0116 ak/

1000 Mwd /T for 8 x 8.

This compares to the last reload ("C" fuel with cobalt) value of 0.0105 Ak/k/1000 Mwd /T. The beginning reactivities are comparable to the "B" and "C" reload des'gns, and no special power shape or lifetime behavior problems are expected.

F.

Location in Core The six centermelt test fuel bundles are to be loaded in the core in a dispersed array with a minimum center-to-center distance of 42 cm (16.5"). The fuel bundle plus water gap is a square 7.h" on a side.

This restriction means that the closest centermelt bundle spacing vill be no closer than two bundles in the x-direction and one in the y-direction.

Also, the centermelt bundles will not be placed in the outer row at the periphery of the core. Notches have been placed on the handles of the centermelt bundles to permit visual verification of their location after i

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1 they have.been loaded'into the core. The centermelt fuel-i,s designed to-have albundle power comparable to the adjacent' fuel and will,not cause

. excessive flux:or power in' adjacent fuel bundles.

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IV. ~ Experience With Central Melting 1.

'This irradiation program can be considered an extension-of the hi h performance UO pr gram which vas. proposed jointly;by the S

2

.U.S. Atomic. Energy.ComLission and Euratom. This program established the basic. feasibility of operation with-UO central melting. Forty-three fuel.

2 roda were irradiated during the high performance UO2 pr gram as summarized'

.in Table 2.

A brief synopsis of the irradiations is given below. More

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idetailed'descriptiens of the irradiations are availab:,e.(3,4,5)

The.cor-respondence between individual fuel rods and the assemblies in which they were irradiated can.be~ determined from Table 2.

The first 'three fuel assemblies irradiated contained rela-tively high-density, solid pellet, fuel rods.and were representative of the best design practice at the time. These three assemblies,'EPT-6, EPT-8 and' EPT-10, were-irradiated in consecutive:GETR cycles. The third assembly irradiation with EPT-10 was terminated prematurely..by indications of a fuel rod failure and all three assemblies showed severe cladding swelling, suffi-cient to preclude continuation of the irradiations to higher thermal per-formance levels.

Evaluation of the problem, and especially supplementary.

. capsule irradiation evidence, led to the conclusion that the most probable

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mechanism for the swelling was the occurrence of a significant UO # 1"**

2 increase: accompanying the solid-liquid phase change. The primary evidence for this conclusion was:

i Lyons, M.-F., et al, "UO2 Fuel Rod Operation With Gross Central

- Melting," GEAP-426h, 0ctober 1963 (h)Iyons, M. F., et al, " Molten Fuel Rod Operation to High Burnup," GEAP-5100-2.

Lyons, M.

F., "High Performance UC Program - Irradiation og Btu /Hr-Ft 2

0 5 Inch 2

' Diameter Vibratory Compacted Powder Fuel Rods at 1.3 x 10 Peak Heat Flux to 10,000 MWD /Te Burnup," GEAP-5100J, September 1966.

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.1.

-The correspondence of the maximum swelling zone to.

ithe. peak heat flux zone at' start-up.

3 P. ' The correspondence of the threshold thermal performance -

level lfor swelling with that required to initiate melting.

3.. The apparent increase in. swelling with thermal per-formance level.

k'. The distinct axial expans' ion-of.the molten UO fuel 2

column in short-length capsule irradiations.

Almost simultaneously.with the occurrence of fuel' rod clad swelling during lume hange on melting-these irradiations, new measurements of the UO2 'y

. were. reported (6*7) and they indicated a greater volume increase than pre-Lviously suspected.. Preliminary results suggested a valueLof 7% which was

' subsequently refined to 9.6%. Based on these results, it became apparent h

that higher performance conditions could only be attained by devising'a.

means to accommodate the increased fuel volume.

Two methods were given serious consideration for reducing fuel density to allow for the phase change volume increase:

1.

Divide the present through-rod design into capsule

1ength segments and provide axial expansion areas within the molten fuel region, and 1

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2.

Use hollow or dished pellets to provide the increased l

1 volume.

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i The use of hollow pellets was finally selected as the most straigntforward approach. To evaluate this idea, a special assembly irradiation, desig-nated EPT-COM, was performed with pellets cored with an 85-mil' diameter

hole', equivalent to 43% of the fuel volume. The pellets'of these four

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' rods were enriched such that a range of performance levelE would be l

'obtained. The irradiation of this assembly was performed in two steps:

-(6)Christianson, J.

A., " Specific Volume of Molten Uranium Dioxide," Paper-

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.for'American Ceramic Society, 15th Pacific Coast N. W. Regional Meeting, Seattle, Washington, October 1962.

Christianson, J. A., " Thermal Expansion of UO," HW-75148, October 1962.

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-(l) A single short cycle of full power' operation.followed by examinatien, and (2) A full' cycle of irradiation with normal reactor power cycling.

Post-irradiation examination revealed only slight swelling of one rod, Rod 10E,_which occurred during the short cycle start-up, with no indi-cations of a ratcheting mechanism during the subsequent continued opera-tion. However, because of the slight swelling of Rod 10E, it was conclu'ded that the free volume provided by the small cored pellets was marginal for a' performance level of 1.0 x 10 Btu /hr-ft and above, and a larger core was provided for the maximum thermal pe.-formance'irradia-tion in the sintered pellet series, Assembly EPT-12C.

Assembly EPT-12C, to be irradiated at a nominal performance 6

level of 1.2 x 10 Btu /hr-ft, was provided with sintered pellets cored-

-to an inside diameter of 0.140".

This larger core size was selected to provide about 8% (of

.e UO ) free v lume for.the phase change expansion.

2 To protect the' rods further, a programmed start-up procedure was, introduced in the irradiations. This procedure was formulated in recognition of the 1

significant axial fuel relocation observed in the previously iriadiated rods, and the possibility that the relocation could undo the initial benefit of the de:iberately included central hole. The objective of the procedure was to prevent any sudden increases in internal volume without allowing time for axial redistribution. The reactor power rise rate was limited to less than 1 Mw (or 3%) per 15-minute time increment and-the restriction initiated at a power level below that at which the onset of central melting would be expected. This start-up procedure was specified l

for all start-ups following a reactor refueling, provided the predicted flux peak location was three inches or more below that existing at the previous shutdown.*' In the absence of a refueling resulting in a sig-nificant downward flux shift, no restriction on-start-up rate was made.

Irradiation of Assembly EPT-12C was successful through the first two reactor cycles to 5300 Mwd /T burnup with no detectable swelling of the rods at performance levels as high as 1.h x 10 Btu /hr-ft "GETR operation results in an axial flux peak shift as control rods are retrmeted to compensate for reactivity depletion during a reactor cycle. Refueling then results in an abrupt chift of the peak heat flux axial location.

39

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[However,onthethird' cycle. start-up,rodswellingandonerodfailure e occurred. The unexpected result was subsequently traced to an ina'd -

fvertent4 Mw'(13.3%): step power increase before failure. Two of the

' rods ~from the EPT-12C assembly were only slightly swollen-and continua-

tion'of their irradiation was attempted in two subsequent assemblies, EPT--12C-A and EPT-12C-B.

However, both rods experienced pinhole type failures during their initial start-up or very shortly thereafter. No satisfactory explanation for the pinhole failures was obtained, but some connection with tlie same. event that caused the swelling anc failure. in the.other two rods seems inescapable. New cored pellet rods,-substan-

-tially identical to the EPT-12C rods, were fabricated later and operated satisfactorily to more than twice the EPT-12C assembly burnup in Assembly EX-12A.

. Irradiations in the compacted powder UO series were begun 2

with 90,5% dense, vibratory-compacted and svaged rods in-Assembly EPW-6/8.

1:

Because of the lower as-built density of th'se rods compared to solid pellets, they were expected to be less susceptible to swelling. -Instead, the powder rods showed greater swelling at equivalent thermal performance to the pellets,'even with carefully programmed start-ups. This greater swelling was verified in a second assembly irradiation with 90.5% dense rods, EPW-6/8A. However, this same assembly also contained two 88% dense,

-straight, vibratory-compacted powder rods which demonstrated significantly reduced swelling at higher thermal performance than the two 90.5% dense f

rods. The effectiveness of reduced initial density in preventing swelling was '.onfirmed in a third assembly containing straight, vibratory-compacted 875 to 88% dense rods, EPW-6/8V. The concept demonstration was then ex-tended to much higher thermal performance 3evels with the successful 6

irradiatien of 83% to 85% dense fuel rods at peak heat fluxes of 1 5 x 10 L

in a fourta powder assembly, EPW-10/IPV.

On the basis of the above experience, the extended irradia-p

tion of the maximum thermal performance powder fuel assembly, EPW-12V, was undertaken and carried successfully beyond 10,000 Mwd /T average rod burnup 6

2

'at heat fluxes up to 1.3 x 10 Btu /hr-ft. During the fifth GETR cycle of i

? -

15-

[

irradiation,~.a fuel rod failure occurred which was subsequently traced to an initial fabrication defect. The fail, re we a 1/2" lorig longi-tudinal-. split located on the' low heat flux side of the rod h-1/2" from the bottom.

It is' thought.to have resulted from a tubing defect. intro-E:

.duced by svaging. Very slight.svelling was noted in the rods, which presumably occurred at the time of the initial-start-up. Two of the-unfailed rods were.then' transferred to a new assembly with two'new

~

cored pellet rods. This assembly, EX-12A, was successfully irradiated for five more GETR reactor cycles.at peak heat. fluxes up.to 1.h x 10 Btu /hr-ft which brought the two powder rods to 20,000 Mwd /T average burnup'and the two new pellet rods to about 12,000 Mwd /T. Extensive post-irradiation-examination of these rods showed no further swelling

=in the-powder rods, none-in the pellet rods and no detectable deteriora-tion'of any kind. Coupled with the satisfactory irradiation performance, this assembly is considered to hava demonstrated the intrinsic _ feasibility-of high burnup operation with gross central melting.

V.

Hazards

,iderations A'

Gener.

lensiderations The centermelt fuel bundles described above and shown in f

Figures 5 5 and 5.6 are the same basic mechanical design as the Reload

~"B" and "C" and the Phase I and II R&D bundles. Experience to date with this design has been excellent and there is no reason to expect problems rising out of the design.

B.

Normal-Operating Considerations Except for the highcr internal temperatures of the UO '

2 i

the thermal design of the centermelt fuel is similar to the fuel currently licensed for use in the Big Rock Point reactor. Table 3 below lists the i

critical heat flux ratios (CHFR) and coolant properties for the 122% over-power case. This case is for the 8 x 8 assembly for a typir' ally " worst" I'

case of power distribution. The CHFRs are based on the new GE correlation from APED-5286. The 7 x 7 assemblies show similar quality trends but have higher CHFRs.

it should be noted that the minimum critical heat flux.

i ratio (MCHFR) for this typically " worst" pceer distribution is within the current. Big Rock Point license limit fo-MCHFR. This has been achieved l'

16 1

by ' including very low power fuel rods in the bundle with the high power centermelt rods. This combination yields essentially'the sam,e node-by-node coolant quality profiles as-in the current fuel-license 4for the Big Rock (Point reactor.

TABLE 3' Heat Flux _and Coolant Property Data-for a' Typically " Worst" Power Distribution at 122% Overpower-(8 x 8 Assembly)

" Axial Surface

-Power.

-Heat Flux.

Surface Coolant.

Node Factor Btu /Hr-Ft Temp *F CHFR Quality 1

.0.391 0;1513-+ 006 590.6 5.8 0.

2 0.510 0.1978-+ 006 591.2' h.5 0.

3.

0.630 0.24h3 + 006' 591 7 3,6 0.

h 0 7ho 0.2870 + 006 592.1 31 0.

5 0.830-0.3219 + 006 592.h 27 0.

6-0 910 0.3529 + 006 592.6 25 0.

7 0 970' O.3762 + 006 592.8 23 0.002 8

1.030 0.3995 + 006-592 9 2.2-0.010 9

1.060 0.h111 + 006 593 0 2.1 0.019 10 1.090

.0.4228 +'006 593.1 2.1 0.027 11 1.100 0.h266 + 006 593.1 2.1 0.036 12 1.130 0.4383 + 006 593.2 2.0 0.0h6 13 1.180 0.4577 + 006 593.3 19 0.055 1h

'1.270 0.h926 + 006 593 5 1.8 0.066 15 1.380 0 5352 + 006 593.8 1.6 0.077 16

.1.400' O.5430 + 006 593.8 1.5**

0.088 17 1.310 0 5081 + 006 593.6 1.6 0.099 18 1.180 0.h577 + 006 593.3 17 0.109 19~

1.0h0 0.h03h + 006 593 0 19 0.117 20 0.850-0.3297 + 006 592.h 2.2 0.12h i

6 2

• Mass Flow Rate =-0.75 x 10 Lbm/Hr-Ft, Radial Power Factor = 1.16, l

Loc M Power Factor = 1.60

• • Minimam Critical Heat Flux Ratio (MCHFR) l l

l

n

-~

eq'..

L

+

.. a 17 C.

' Accident Analyses 1.

Loss-of-Coolant' Accident-The implications of the higher power rating of the center-melt fuel rods on emergency cooling system performance as described in.

' Big Rock. Point ' license change No. -8 have been appraised. Two aspecto of

.the higher power rating are considered to'be of significance:

a. - The-increased stored energy in the:centermelt fuel tresulting from its higher average operating temperature level.

'b.

The larger decay heat generation rate; this, of course, is-directly proportional to the higher design power level of the center-melt fuel.

It'is particularly significant to note that t'is change.

. request relates to the introduction of a maximum of six experimental

. assemblies into the Big Rock Point core. Each of these assemblies is constituted partially of higher rating (centermelt) rods, the balanca being of a low rating such that overall the fuel bundle generates a power

-level comparable to that of a standard bundle. Consequently,,the decay heat generation rates for the complete.centermelt assemblies are compurable to the maximum levels computed for previous fuel assembly designs for the

]

Big Rock Point core, even though.the individual centermelt (high heat flux).

. rods within the. core will have proportionately higher decay power levels.

In view of this plemned maximum core loading of six centermelt bundles under the present program, the centermelt rods will constitute only about 2% of the total number of rods making up the core.

Analysis of the effect of increased stored energy in the centermelt rods on fuel temperature following a major system rupture, as discussed in Big Rock Point license change No. 8, indicates the following:

As a consequence of the duration of the postulated blowdown (>h seconds) and of the high heat transfer rates expected in the core during this process, these centermelt fuel rod temperatures will be reduced to a level characteristic of the ensuing decay power generation and are thus virtually independent of the initial stored energy content. This aspect of the center-melt ~ fuel design is therefore not errected to influence the consequences of a' loss-of-coolant accident.

r l

l~

+'

~8 1

The effect of the increased decay heat level of individual'

.-rodswithinanassemblycontaining'a. proportionate.numberof[ underrated

~

rods, such that' on the average the fuel bundle decay power is unchanged, is in this case minimal and is found to be within the limits of accuracy _

~ f the temperature measurements (2100 F at'2000 F) carried out by the

^

o

. General Electric-Company.in the various spray cooling test programs con-ducted inusupport of BWR safety analyses.' Undoubtedly.the general effect

~isone of. increased temperatures as necessary to dissipate the increased decay heat flux. 'However, at the higher temperature levels (>2000 F) of

-prime interest from the viewpoint of zirconium behavior'in sater reactor accidents, a 200 F increase in temperature not only radically increases the evaporative cooling rate for the rods but also increases 'he radiative

-heat transfer by 25%-or more. This trend is well illustrated by Figure attached.

i

.As a result of the proportionate number of low rating rods also present.and dispersed relatively uniformly through the bundle, the net effect will be somewhat less than that implied by Figure 16.

Further-more, as this effect will only apply to some 2% of the rods in the core, its significance on the degree of zirconium water reaction and, indeed, on the whole course cf the loss-of-coolant accident is minimal and is within the uncertainty margin'of ava lable computational techniques.

2.

Control Rod Drop Accidents The Big Rock Point reactor operates with one specified rod withdrawal pattern. The rods are grouped in banks of two.or more; all the rods in a bank are withdrawn together, with a procedurni limit of one notch between any two rods in a bank. This sequencing prevents.large rod worths; however, an operator error or series of errors can result in larger worths.

The possible rod drop situations and rod strengths when the core is critical and at hot standby are:

I Case.1:

In-sequence potential of.008 Ak for drop from-full-in position to drive position.

Case 2:

In-sequence potential of.021 Ak for drop from full-in-to full-out.

l-t-

I

~

s

-19 Case 3:.Out-of-sequence potential of less than.021 Ak for drop from-full-in to full-out.

Case h: LMaximum theoretical worst case of about.045 Ak, Case 1 requires the following equipment malfunctions and oper'ator. error:

a.

Rod becomes uncoupled from drive.

(

'b.. Drive is withdrawn (in-sequence), but blade hangs up

temporarily. Operator does not notice that-blade is

~

not'following.

Rod then unexpectedly releases and drops from full-in c.

to position of the drive due to gravity.

Case 2 requires an additional operator error of withdrawing the. drive completely rather than concurrent with the bank.

Case 3 consequences are less than those for Case 2.

Case h is considered hypothetical, as it requires still further wompour. ding of unrelated errors beyond those enumerated above.

The analyses are performed for the hot standby (HSB) con-dition, i.e., power at neutron source level and normal water level in the vessel. The hot standby rather than cold case is analyzed because:

a.

The-rod strengths in the cold condition for the same potential.

- rod drop situations are no greater than in the hot standby condition.

b.

According to our calculations, primary system integrity is more vulnerable when there is a free surface of the possible water-hammer and subsequent vessel movemerc. This situation exists so long as the reactor is critical only when

.e vessel is~at a temperature sr.fely above the ni.* ductility temperature. Maintenance of this safety margin allows a-strain to rupture, of at least 13%. B16 Rock Point operating procedures provide this margin.

The sero power initial condition accidents are more severe than full power accidents because:

l'

E

-]

~

~

1 20 a.~

The comparable rod strength is less.at full power; e.g.,

the worth for Case 2 is.0095 b.' The core void give. the moderator system more compliance, thus the water-hammer situation is not as severe.

To. prevent a large amount of centermelt fuel from being in~the peak neutron-flux during a reactivity accident, the six centermelt-bundles are toibe loaded in the core in a dispersed array with a minimum-

~

center-to-center distance of h2 cm.

This restriction means that the closest centermelt bundle spacing will be no closer than two bundles in the x-direction and one in the y-direction.

a.

Analytical Methods and Results The latest analytical tools and methods were employed throughout the present analysis. A brief description of the methods and changes from the "C"~ core analysis follows:

1.

Neutron Kinetics Methods - The kinetics analysis analytical model was essentially unchanged from that used in the submittal for change No. 10.

It should, however, be pointed out that the calcula-tions for change No. 10 were among the first done using this kinetics analysis analytical model. Since that time, the input parameters have been refined..The results of this analysis differ from those submitted earlier in change request No. 10.

These refinements of input data are discussed later on in this submittal.

(a) To incorporate reactor spatial effects into the Doppler feedback to the kinetics calculation, a space-time kinetics analysis is synthesized by a marching calculation.

Initial neutron flux distribution associated with accidental reactivity addition is first determined utilizing a three-group, steady state diffusion calculation.

A core averaged Doppler spatial weighting factor is estimated from the flux distributions and utilized in the point kinetics equation to generate a small increment of power.

In this case, the kinetics calcu-lation represents the average reactor condition.

This increment of power, expressed as a fuel temperature change, is then spatially dist 'ibuted across the core according

21

.to initial flux distributicas. New spatially distributed cross sections

~

are computed, reflecting the Doppler effect due to the added. temperature,-

and another diffusion calculation is made. Comparison of the eigenvalue-change in 'this calculation to the eigenvalue change resulting from a uniformly distributed temperature increment provides an accurate estimate of the Doppler weighting factor appropriate.for the next kinetics calcu-lational step. Utilizing this procedure, the calculation is marched through to the termination of the excursion. The space-time dependent calculation includes the inherent power flattening as fuel is heated in the zone of.the dropped control blade.

(b) For excursions commencing at zero power and having relatively large reactivity insertions (.025 Ak or greater) at gravity rod drop rates, the Doppler effect turns the primary power burst over before all the reactivity has been inserted. This additional reactivity will cause another prompt power burst if there is no negative reactivity immediately available in the system. The scram is assumed not to actuate until 200 milliseconds after the overpower condition is reached, so it can do no better than terminate a second burst, not prevent it.

For the centermelt core all reactivity insertions greater than about.02 Ak produce some material of an energy above h25

~

cal /g, the prompt rupture threshold. This material vill disperse in sur-rounding water and create a rapid steam explosion. Dispersal of the highly enriched centermelt fuel into the moderator is a considerable negative reactivity effect. For example, complete homogenization creates a reduc-tion of approximately 16% in k,of a bundle. Proper volume and importance weighting of such an occurrence would indicate the whole core effect, and

it could be as much as a negative.02 Ak.

Removal of moderator and fuel material or just moderator from the core with a steam explosion is also a 1 considerable negative reactivity effect. The effect of these reactivity changes is not only one of reduced reactor power but also a spatial shift of. the peak neut. a flux away from the affected area. Now, even in the end of spectrum reactivity insertion,.045 Ak, the great majority of fuel j

reaching high energy is located in the centermelt bundle adjacent to the l

l

22

. 'droppe6. rod. Thus, the spatial neutron shift to a' lower reactivity. region at'thegreatlyreducedcorepowerdoesnotsignificantlyincheasetheamount of fuel ~containing prompt rupture energy.

Considering that. dispersal, st'eam explosion and

. flux spatial shift occur.in several milliseconde, while a second prompt-power burst would not occur.until tens of milliseconds after the_first burst,'ne second power burst-is. prevented.or at-leart severely restricted.

It is for this reason that the kinetics analysis.'for accidents producing a steam explosion doe-not. include the calculated integrated energy'found in a second power burst.

(c) The excursions are assumed to.be adiabatic with only the~ Doppler effect supplying prompt reactivity. feedback to terminate the primary power burst. In particular, no negative feedback from prompt' moderator heating is assumed. The magnitude of the reactivity effect due-to Doppler broadening in the calculation has a nominal conservative bias of about 20% in predicted reactivity decrement. This inferred bias is due to a number of factors. Experimental data on UO2 :uel against which the

. design model is compared is that of Pettus and Hell trand. This com-

.parison shows an average conservitive bias of the des,gn model on the

+

order of 5% to 10%. In addition, there is about a 3% conservatism intro-3 duced through the assumption of a constant radial fuel temperature. The contribution of Pu-240 to the Doppler reactivity has not been assumed.

e This effect will increase the decrement by 10% to 15% throughout most of the reactor life.

I The temperature dependence of the Doppler coeffi-

)

cient used in all of the analyses is based on the inverse square-root of temperature form. Experimental data for oxide fuel are fit more precisely i

by this form in both differential mcasurements ' and integral measure-ments of lattice parameters. Analysis of SPERT oxide core excursions,

utilising this Doppler model and the space-time kinetics calculation dis-p cussed above, have shown the calculations to give excellent agreement on-I both the measured peak excursion power and total energy release.

i

_,-.,m_

~ ~

-r-23 (d)- The zero power accident has historically been initiated at a power level of 10 of rated. As the accident severity

' decreases monotonically with increasing initial power level in the zero

~

. power range further investigation has been undertaken.

Our analysis indicates a free surface is neces-sary if-a water-hammer-activated vessel' movement is to occur. Normal start-up procedure insures that whenever the reactor vessel has a free water surface, i.e., it is not full r,f water, the core' power is at least t

'10% of rated. -However, it is conceivable that a lower power could be obtained with a forced shutdown. at which time the isolation valves are

. closed, momentarily preventing the steam drum to backfill into the pres-

+

sure vessel. It is this improbable situation that is postulated to embody both a free surface and a zero power condition. iiovever, the power level must.be at least as great as the neutron source level. There are two fp - O c:

8,000 curie % g Fg sources in the Big Rock Point Plant. Conservatively assuming the plant has been previously ' shut down for one source half-i life, six months, and considering their~1ocation for effective neutron produc. tion to the core, the total source level is 3,000 curies. Using this value, the relationship between core effective multiplication con-stant, k ff, and power is calculaMd and presented on Figure 16.

It is

-6 seen that for a power of 10 the cort..a'.002 Ak suberitical. Using the various combinations described by the line on Figureff, it was determined that an-initial power of 2 x 10 times rated and.001 Ak suberitical yielded the most severe results for insertion of.021 Ak or less. For larger insertions, this effect was of little consequence, and the his-torical values of 10 rated power and just critical were used. This vork also indicates the conservativeness of the historical cold accident-j.

l

-8 initial conditions of 10 rated power ant just critical.

(e) For the rod dropping from full in with gravity the power burst occurs in the vicinity of 50% vithdrawn, depending upon the reaJ ivity worth of the rod. This gives a relatively large axial l_

power peaking. For a rod ejection, the rod will be withdrawn completely when the power burst. occurs, thus giving a normal cosine axial power shape.

.a.

-(

-m y

-.7 s.

24 l

In.the' "0"i core analysis, the peaked axial shape was used in the ejet < ion calculation; in'this analysis the cosine shape is used.- In addition, the trod' drop analysis was performed'on a radially. central control rod due to

(

"the11 imitations of'our models.1 This is conservative in'that a greater

total peaking.is obtained than for an off-center reactivity addition in i

a-.small leakageEcontrolled core such as Big Rock Point. As it was not necessary to model the axial effects dy amically'in the ejection situation, 4the actuel situation of an off-center rod was analyzed. An off-center' rod

~

should always be analyzed as they are of highest worth at Big Rock Point, i'

and it'is' improbable that a centermelt bundle will be placed in the-center.

of the core.

ii. Results of' Kinetics' Analysis - The peak energy. density-results of the kinetics analysis are shown on Figure 17 As:vas explained

[

~

r in the methods descriptions, when the energy density exceeds'the prompt rupture threshold ~ of h25 cal /g, the power burst is terminated before any.

2 -

secondary prompt reactivity burst appears. This explains why the slopes

-of tLe free fall curves.are reduced at prompt rupture energies. The "C" core and centermelt core results are essentially separated by the peak 4

Ibundle local power. factors in the two cores, that is the centermelt bundle

~

~ of highestispecific power is assumed to be adjacent to the dropped control rod.

The peak energy density of low reactivity ejecticns

}

'[

'is not as great as would be expected relative to the corresponding free fall urop. This is explained in_the methods discussion by more. realistic analysis in the ejection cases.

As indicated previously, the primary difference be-i-

tween the two cores is the local peaking distribution. This is vividly.

demonstrated on Figure 18, the peaking distribution at peak lower as a

result of a.025 Ak rod drop.

The final energy distribution is illustrated in I

. Figure 19 Prompt rupture is indicated by an energy density of h25 cal /g and the melting range by 220 cal /g to 280 cal /g.

From Figure 19, it is t

seen that the mass of molten fuel is greater in the "C" core for the same

?

+

-,a.

r y,y

-.c

,,e

,r-,

t 25 reactivity addition; however, the amount of fuel undergoing prompt' rupture is less. The reason for more molten mass, as explained in the

' methods discussion, is due to the shutdown provided by the steam explo-sion in the centermeT+ core. As far as steam explosion effects are con-cerned, the - may component is the amount of fuel promptly dispersed into the moderator.

iii. Heat Transfer Methods - The heat transfer mechanism.

can be. visualized in the following steps:

(a) Loss of clad integrity.

(b) Puel dispersal.

(c) Vapor pressure adjustment to ambient.

(d) -Particle heat dissipation.

There is little time lapse between each of the above occurrences for a particular fuel rod. The specific character of the process depends on the accident parameters of peak energy density, reactor period and fuel particle size.

(a) Lons' of Clad Integrity - The shape of a power burst can be reasonably represented by a Gaussian distribution. The total energy density deposited in the fuel is simply proportional to the time integral of the Gaussian. From the kinetics calculations, the shape of the Gaussian, defined by its standard deviation, can be found for each reactivity addition. The standard deviations for the Big Rock Point rod drop accidents vary between 5 and 8 milliseconds corresponding to the.045 Ak insertions, respectively. With this knovledge and the assumption that prcept rupture occurs when the fuel eathalpy reaches h25 cal /g, the relative time of prompt i

1 failures can be found. For final enthalpies less than 425 cal /g, the modes of failure vary from fuel vapor pressure exceeding hot ;1 adding ~ strength to cladding meltdown. Cladding failure is assumed not to occur for final energies less than 150 cal /g. The relationship used in subsequent calcu-lations is shown on Figure 20.

In the prompt rupture domain it primarily represents the short period, large insertion accident which means it is

' conservative for the lesser accidents.

I (b) Puel Dispersal - When the clad is breached by internal vapor pressure, the fuel is assumed to disperse instantaneously M

'w w

m

i 26.

linto the surrounding vater in the form of small spherical particles.

~

6 lThe available transient data with rod-type UO2. I"*1' Powder,cn pellet,

indicates a~ distribution of particle sizes from a few mils in diameter to the ori inal fuel diameter.

.As the data' has been taken' with fuel t

lengths from.5" to 5", an axial power distribution vould not seem to

.be important. Irrespective of the reasons for the vide particle dis-Ltribution, it has been shown to occur;-thus, an effective particle size,.

properly reflecting ths surface area to mass ratio, should give the correct heat transfer calculational results.

(c) Vapor Particle Heat Transfer - When the'het fuel is exposed to a lower pressure atmosphere, 1000 psi, the tendency.

vill be to vaporize more until its vapor pressure is in equilibrium with its temperature. This will be nearly an instantaneous operation. How-ever, particle quenching by the water vill also commence, und this energy transfer is taken into account by the particle time constant. It is thought realistic to instantaneously add a quantity of energy to the water to account for the extremely small vapor particles. Somewhat

,- arbitrarily 30 cal /g is added, immediately bringing the UO to a level 2

of 395 cal /g which has a relatively low vapor pressure.

(d) Particle Heat Dissipation - As explained in the discussion of fuel failure, various fuel pins reach prompt breach energies at various times on the back side of the power burst. If the power burst has not terminated, there vill continue to be fissioning in the dispersed fuel. As discussed pr2viously, this fissioning vill be at a lover rate relative to the core average than when the pins were intact due to the neutron flux shift. The heat generated in this manner must be dissipated at about the same rate as generated because of the proximity of the fuel I

temperature to the vaporization range. Conservatively assuming no flux shift occurs, this heat transfer rate is taken as the lower limit for this analysis. Now the power burst is an exponential decaying function, so an e-folding time constant, T, is used to describe the heat transfer rate.

Figure 21 shows the time constant used to be a minimum of 15 ms in the l-

'high energy density range. As this particular heat transfer analysis-l' I'

i i

i

~

gw

-27 6

does not apply'to any fuel at: energy. density.less than prompt failure, I

425' cal /g/ a constant value. of 100 ms was used for the time c"onstant of-l particles that' vere;not dispersed.during the power burst.

4 The_ upper limit on heat transfer rate is-the

' particle. property-dependent internal conduction.' This rate calculation f

uses an infinite particle surface heat transfer coefficient..- The fission-Egeneration. should.also be-included in this upper limit; however, its time

[

. constant is large compared to-the conduction constants so the effect is-small.' These time constants are shown-on Figure-22 for the various par-I-

ticle sizes. The minimum curve corresponds to a UO heat capacity of

~

2 h

.08 cal /g C.

However, the latest data indicates that the heat capacity.

in the temperature' range just below melting is about.15 cal /g *C.

~Now many materials similar to UO have essentially 2

- the same heat capacity in the molten state as at temperatures just below melting. With this assumption for UO, heat transfer time constants.are 2

larger, as1 indicated on Figure 22.

In addition the ANL-TREAT calculated-

. time' constants used in the previous "C" core analysis are shown as well as a radiation controlled heat transfer rate. 'For~ conservatism, this l7 analymis used the conduction relationship calculated with the lesser heat capacity.. The question of particle size distribution at the' time.of dis--

1:

persal.is not easily answered. It is believed the TREAT tests-are the-most representative of BWR transients that exist.crom the standpoint of l

fuel failure mechanisms and particle sizes. Argonne National Laboratory F

has published a reasonable correlation between final mean diameter and peak enthalpy for these tests. This correlation indicates that the mean diameter.for fuels undergoing prompt rupture is about 20 mils..From i

Figure 22, this diameter gives a time constant of k mr. which is used in

~

Figure 21.

The time constants for less than prompt rupture energy are also direct functions of the TREAT data.

There are now two sets of heat tranc.fer rates with which to bound the problem, fission energy generation and conduction limited. As mentioned above, both are somewhat conservative as in the l

face of uncertainty the time constants are decreased.

.o j

i

. - - ~ -. -- -

2$'

iv Thermodynamic-Hydrodynamic' Analysis - There-are many inputiparameters;to this calculation'that are not well unders'tood. For

" this reason, it is realistic. to show the effect of varying these parameters rather than choosing a particular value. This analysis ~uses extreme values

- of.the.two most important-time variant inputs, the flow path of the wat'er slug and the energy deposition into the water.

The energy deposition extremes were discussed pre-viously and are identified by internal conduction limited and fission heat

. generation limited. A third set of values' corresponding-to the TREAT'

- transient powder data is also used. A mathematical expression derived-from the basic differential equation describes the flow path in terms of-volume velocity and kinetic energy; it is called restraint. The minimum

< restraint to flow is defined as a cylindrical column of water having a cross-sectional area equal to that of the vessel. The maximum restraint

~

is ' defined as a cylindrical colu=n with a cross section equal to that of

.one bundle, e.g., the centermelt bund]e adjacent-to the dropped rod. From theoretical considerations, the flow path should be somewhat conical in nature.

Thus,.a best guess flow restraint is chosen to physically repre-j sent a two-dimensional. flow path with dimensions in between the two extreme

- one-dimensional situations.

The two important primary system rupture modes are' vessel vertical movement with subsequent shearing of pipes and vessel overstrain. The limit of allowable movement on the steam risers and coolant recirculation lines is about 6".

However, the limiting case

. is 'an allowable 1,5" on the small core spray line. The limit on the j

Big Rock Point vessel ductility after 40 years of operat! sn has been determined to be 13% strain.

The calculational results are shown on Figures 23 and 24. Some general trends can be noted:

l l

(a) The higher the restraint, the greater the steam I

. pressure and thus the vessel strain. With the high restraint value, there. is no vessel vertical movement.

I (b) The lower the restraint, the greater the vessel movement. There is no strain with the minimum restraint.

f r ta w

--n-r

.N e m

- ae

-A-

29 (c) There'is no threat to primary system rupture unlessthereisasignificantamountuoffuelinthepromptrkpturedomain-greater than h25 cal /g.

D.

Conclusions Based.upon the.above analyses and comparisons of a whole l core of "C" fuel.and a'. core of."C" fuel containing six centermelt bundles, the following conclusions may be drawn:

1.1 The ~ mechanical design of the ' centermelt fuel bundles

.is a well-proven concept and no problems'should be expected based on the good experience with the design to date.

2.

The-total power generated in each centermelt-fuel bundle matches the power generated by currently licensed fuel for the Big Rock Point reactor. This yields equivalent coolant quality profiles and-MCHFRs. Consequently, currently licensed MCHFR limits can be readily

' met.

3.

The addition of a maximum of six centermelt in2el bundles 4

with their relatively small number of higher performance fuel rods does not significantly change the loss-of-coolant accident. The high performance centermelt rods will constitute only about 2% of the total number of rods.

i' in the reactor core. Even though the centermelt rods would have an in-l creased decay heat level, the ' higher rod temperatures radically increase i

their evaporative cooling rate and also increase their rcdiative heat l

transfer by 25% or more. Consequently, the effect of the six centermelt j

fuel bundles on the loss-of-coolant accident is minimal.

(,,

4. lThe rod drop accident is not expected to cause a ureak i-of the primary system for the following reasons:

a.

The maximum control rod worth obtainable with one procedural error is about.021 Ak.

The theoretical maximum rod vorth is abcut.045'Ak.

The attainment of this high a rod varth requires several procedural errors in succession. Because only one control rod withdrawal pattern is used at any one time, we consider this contingency as being incredible.

+v

i 30 b.

Considering reactivity accidents, the primary

~

difference between the "C" core and the centermelt core is the increased local peaking in the latter. The free fall drop of the.021 Ak rod mentioned above yields peak energy densities of about 330 cal /g and h50 es1/g in the "C" and centermelt cores, respectively. Neither of these values represents a severe threat to the integrity of the pri-mary system, c.

Because of the restriction on location of the centermelt bundles, all of the elevated temperature fuel is located in the centermelt bundle assumed to be adjacent to the dropped control rod.

d.

No threat to primary system integrity was cal-culated to exist for accidents not hsving a significant amount of fuel in the prompt rupture domain (>h25 cal /g).

e.

Using extreme energy addition rates and thermal

'to mechanical conversion efficiency, the calculated vessel strain as a result of a free fall rod drop of maximum worth in the centermelt core did not reach the vessel ductility limit.

f.

Under the most extreme eet of assumptions, the calculated vessel vertical movement as a result of a free fall rod drop of maximum worth in the centermelt core would exceed the possible primary system rupture limit. It should be pointed out, however, that these analyses were all done ignoring the resistance to vessel vertical movement provided by the vessel support system.

Considering this and the other conservatisms in the calculation, along with the extremely low probabilities of the contingent string of events, it is very unlikely that the primary system would be breached, even under the most extreme set of assumptions analyzed.

I i

31 1

Based on the above considerations, we have concluded thattheuseofcentermeltfuelintheBigRockPointreactor[doesnot present a significant change in the hazards considerations described or implicit in the Final Hazards' Summary Report.

L l

CONSUMERS POWER COMPANY J

M d

By Vice President Date : May 26,1967 Sworn and subscribed to before me this 26th day of May 1967 Cu r _h) i Notary Public, Jackson County, Michigan My Commission Expires February 16, 1968 1

i e

i I

i-

i-4

.~

' TABLE 1 CENTERMELT FUEL DESIGN DEPAILS AND ' OPERATING CONDITIONS 1 Cladding Number Oatside.

Wall UO2 Fuel of Diameter Thickness-Puel Assembly Type TD, U-235' Rods Type

-(Inches)

(Inches)

-Condition'-

Group A D-50 Pellet 95 4.3-12 Zire-2 0 570.

0.035 cold Worked and'

'(8 x.8 Lattice)'

Pellet 95 50 16

-(Tube

' Stress Relieved Pellet 95 5.6 8

Reduction Powder.

85 0.22 28 Process) 4 D-51 Powder 85 h.3 12 Zirc-2 0 5T0

'O.035 Cold Worked and (8 x 8 Lattice)

Powder 85 50 16 (Tube Stress Relieved Powder 85 5.6 8

Reduction Pcwder 85 0.22 28 Process)

Group B D-52 Pellet 95 h.3 12 Zire-2 0.700 0.0h0 cold Worked and I..

and

• Pellet 95 50 12 (Tube-Stress Relieved

-D-53 Pellet' 95 5.6 5

Reduction (7 x 7 Lattice)

Powder 85 0.22 20 Process)

D-Sh Powder 85 h.3 12 Zirc-2 0 700 0.0h0 cold worked and and

" Powder 85 5.0 12 (Tube Stress' Relieved i

D-55 Powder 85 5.6 5

Reduction (7 x 7 Lattice)

Powder 85 0.22 20 Process) qLfj; Mc3

$n

?

'f, s

4 Mn &

, f-l' ce n

IE.?,/C ^~ $Ef1 l

• Four rods of pellet fuel'and four rods of powder fuel of this enrichment (eight _ rod total) will be hI segmented by tungsten wafers into lengths'of about 17".

'V t'

s 9IN~j,6' u'$

J a

l-4H j L I.5 \\

. +

TABLE 1 CENTEINELT FUEL DESIGN DETAILS AND OPERATING CUD 1TIONS (Contd)

Design Design Heat Flux Fuel Rod Power Burnup Normal overpower Normal overpower Uop Fuel Fuel Rod (Mwd /TU)

. (Btu /Hr-Ft )

(Btu /Hr-Ft )

Kw/Ft

-Kw/Ft Group A (8'x 8)

Cold Press Standard 20,000 450,000 610,000

-19.65

26.6 and Sinter Dished 5%

250,000 22.18 Dynapak Vibratory 20,000 450,000 610,000 19 65 26.6 Compaction

• 50,000-22.18 I

Group B (7 x 7)

Cold Press

-Standard 20,000 h50,000 610,000 24.15

' 32.8 '

i and Sinter O.100" 250,000

• 2.68~

Central Hole 4

Dynapak vibratory 20,000 450,000 610,000 2h.15 32.8 l

. Compaction

• 50,000 22.68 l

l-i

\\

t PJ l

~

.__m

TABLE 2.

Fabrication and Irradiation Data Fuel Pre - Irradiation Pellet-Clad Fuel Rod Fuel Column Max. Surface C.addir.g Diameter (Inches)

Diametral Gap Isngth Length Heat Flq Assembly No. and Rod GETR Fuel Type Number Cycle Outside Inside (mils)

(Inches)

(Inches)

Btu /h-ft*

EPT-6 6A 33 0.5647-0.5653 0.5058-0.5064 5.5 39.018 33-3/4 630,000 Solid Pellets 6B 0.5642-0.5659 0.5060-0.5067 5.8 39.008' 34-5/16 665,000 6C 0.5611-0.5638 0.5064-0.5068 6.1 39.006 34-1/4 675,000 6D 0.5642-0.5654 0.5051-0.5056 5.9 39.004 34-3/16 732,000 EPT-8 8A 33 0.5642-0.5652 0.5092-0.5106 6.4 39.014 34-1/4 950,000 Solid Pellets 8B 0.5642-0.5668 0.5082-0.5091 5.1 39.020 34-9/32 920,000 BC 0.5645-0.5663 0.5091-0.5100 6.0 39.031 34-3/16 780,000 BD 0.5635-0.5648 0.5092-0.5100 6.1 39.032 34-1/4 915,000 EPT-10 10A 35 0.5671-0.5688 0.5093-0.5101 5.7 39.033 31-1/4 1,030,000 Solid Pellets 10B 0.5675-0.5685 0.5094-0.5107 6.0 39.012 ~

31-3/8 1,040,000 10C 0.5663-0.5680 0.5100-0.5104 6.2 39.026 31-5/16 1,140,000 10D 0.5668-0.5690 0.5097-0.5107 6.2 39.025 31-1/4 1,140,000 EPT-COM 6E 37,39 0.5647-0.5660 0.5046-0.5055 4.5 '

38.965 33-3/4 605,000 Small Cored Pellets 8E 0.5635-0.5657 0.5092-0.5100 6.I 38.960 34 905,000 10E 0.5679-0.5688 0.5098-0.5109 6.3 38.990 31-1/8 1,000.,000 12E 0.5685-0.5690 0.5108-0.5110 9.0 38.990 29-1/2 1,148,000 EPT-12C 12AC 41,42,45 0.5620-0.5630 0.5001-0.5006 4.8 39.032 29-1/2 1,440,000 Large Cored Pe!!ets 12BC 0.5620-0.5630 0.5001-0.5018 5.4 39.036 29-1/2 1,360,000 12CC 0.5610-0.5620 0.5000-0.5009 5.0 39.035 29-1/2 1,210,000 12DC 0.5610 0.5630 0.5002-0.5009 5.0 39.045 29-3/4 1,2.%,000 EPT-12CA 12BC 47 0.5620-0.5640 39.056 1,137,000 Large Cored Pellets

.12CC 0.5610-0.5630 39.069 1,107,000 12GC 0.5610-0.5620 39.054 29-11/16 1,098,000 1211C 0.5590-0.5605 39.051 29-5/8 960,000' 1,177,000 39.069 EPT-12CB 12CC 52 0.5600-0.5640 Large Cored Pellets 12FC 29-11/16

~ I,230,000 12HC 0.5590-0.5605 39.079 1,233,000 12GC 0.5600-0.5630 39.062 1,061,000 i

.I j IO

. ~ts

', / c3 f

p'

.l ;N - q yO m

-I c,

b

' -i d Nl C

'd nf;:

L s

. >N 1

,J C *% '.

s r :3

! /o

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TABLE 2.. (Continued)

Pre - Irradiation Fuel Rod Fuel Column. Max. Surface j

Cladding Diameter (fnches)

Powder Rod.

Length Length Heat Flux

' Assembly No and Rod GETR Fuel Type Number Cycle Outside Inside Densities (Inches)

(Inches)

Btu /h-ft2 EPW-6/8 Vipac.

6AA1 40 0.5654-0.5671 91%

-39.034 34-1/8 531,000 and Swaged Powder 6AA4 0.5671-0.5682 38.924 34-1/8

-536,000 8AA2 0.5670-0.5695 39.067 34-l/8'

'l21,000 8AA4 0.5673-0.5695 39.992 34-1/8 615,000 EPW-6/8A Vipac and 6AAI 40 0.565 -0.567 91%-

39.026 642,000 Sw2ged Powder 90bTD 8AA4 0.5655-0.569 91%

38.990 744,000 Vigac Powder 6AV 88%

34-l/8 648,000 8AV 88%

34-l/8 872,000 EPW-6/8V 6BV 43 0.5605-0.5620 86.33%

39.000 34-l/4 757,000 Vipac Powder 6CV 0.5605-0.5620 86.90b 39.039' 34-3/16 750,000 8BV 0.5600-0.5610 87.39%

39.040 34-1/8 888,000 8CV 0.5610-0.5620 88.86%

39.03 34-l/8 1,000.000 EPW-10/12V 10AV 44 0.5620-0.5625 84.6*/5 39.017 34-1/4 1,210,000 Vipac Powder 10BV 0.5610-0.5625 84.67E 39.010 34-1/4 1,107,000 12AV 0.5600-0.5610 85.06%

38.970-29-1/2 1,482,000 12BV 0.5615-0.5625 83.63%

38.943 29-1/2 1,351,000 EPW-12V 12CV 47-52 0.5610-0.5625 83.83%

39.015 29-7/8 1,222,000 Vipac Powder 12DV 0.5655-0.5665 85.00%

38.995 29-1/2 1,357,000 12GV 0.5655-0.5660 83.0db 39.008 29-7/8-1,123,000 12 FV 0.5640-0.5650 84.975 39.021 29 1/2 1,260,000 EX-12A 12GV 54-59 0.5660-0.5680 39.027 1,070,000 Vigne Powder 12FV 0.5640-0.5670 39.050 1,030,000 large Cored Pellets 12 FC 1.360,000 (Liner) 12KCF 1.170,000

• There is approximately 10 percent variation in the heat flux values reported for the fuct rods from EPW-6/8 and EPW-6/8 A m the various program reports.

These variations arise from differences between physics and gamma scan indic ations of the power sp;it among the rods.uul changes in the PWL asial peaking factors based on measurements performed midway through the prograni.

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l 1

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?

e FIGURE 6. FUEL CENTERLINE TEMPERATURE VER$US HEAT FLUX -INTERMEDIATE PERFORMANCE POWDER FUEL

0.7 l

l 10,000 MWD /T 04 0.5 1:

BEGINNING 5 = 0.4 Or 18FE e

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l FIGURE 7 UO MELT FRACTION VER$US HEAT FLUX -

2 INTERMEDIATE PERFORMANCE FUEL POWDER

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l

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b t

D A

B B

B A

D 0.13 1.60 1.8 1.67 1.68 1.60 0.13 A

A D

D D

A A

1.60 1.42 0.10 0.10 0.10 1.42 1.60 B

D C

D C

D B

1.68 0.10 l.71 0.09 1 71 0.10 1.68 B

D D

C D

D B

1,67 0.10 0.09 1.65 0.09 0.10 1.67 B

D C

D C

D B

1.68 0.10 1 71 0.09 1 71 0.10 1.68 A

A D

D D

A A

1.60 1.k2 0.10 0.10 0.10 1.42 1.60 D

A B

B B

A D

0.13 1.60 1.68 1.67 1.68 1.60 0.13 Type A U-235 Enrichment = 0.0h3 Type B U-235 Enrichment = 0.050 Type C U-235 Enrichment = 0.056 i

Type D U-235 Enrichment = 0.0022 (Depleted)

Figure lh - Individual Fuel Rod Relative Powers in the 7 x 7 Center-melt Bundle. Beginning of Life. Average Rod Power = 1.0 i

9 t

k e

c-e

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.-y.

--S

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D A

B B

B B

A D

o.13 1.65 1 75 1 73 1 73 1.74 1.63 0.13 A

A D

D D

D A

A 1.65 1 53 o.lo o.lo o.lo o.10 1.49 1.63 B

D D

C D

C D

B 1 75 0.10 0.10 1 79 0.10 1.81 0.10 1.7h B

D C

D C

D D

B 1 73 0.10 1 79 o.09 1 72 0.10 0.10 1 73 B

D D

C D

C D

B 1 73 o.lo 0.10 1 72 0.09 1.79 0.10 1 73 B

D C

D C

D D

B 1.Th o.10 1.81 0.10 1.79 0.10 0.10 1 75 A

A D

D D

D A

A 1.63 1.49 0.10 o.lo o.lo 0.10 1 53 1.65 D

A B

B B

B A

D o.13 1.63 1 7h 1 73 1 73 1.75 1.65 0.13 i

l Type A U-235 Enrichment = o.oh3 Type B U-235 Enrichment = 0.050 Type C U-235 Enrichment = 0.056 Type D U-235 Enrichment = o.0022 (Depleted) l Figure 15 - Individual Fuel Rod Relative Powers in the 8 x 8 Center-melt Bandle.

Beginning of Life. Average Rod Power = 1.0

~.

4 i

I

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10-3 4

!3n 2

10-4 Oy cc LL.o

' E 5

10-5 Ei m

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(0.93)

(0.9) 1-k g (keff) e FIGURE 16. BRP SOURCE AND MULTIPLICATION c-.,

f SS3 1

1 1

ji I

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l "C" CORE WITH CENTER'.(ELT FUEL 700 EJECTION N FREE FALL l

/

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~

l O0 0 0.0 -

FIG. 18 PEAKING DISTRIBUTION BRP

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1000.0-o y

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FIGURE 19 ROD DROP FINAL ENERGY DISTRIBUTION

l I

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FIGURE 20.

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PARTICLE DI A'.'ETER (mits) i FIGURE 22.

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FIGURE P3 BRP VESSEL JU?P

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r 4

REPERENCES 1.

BAW-1244, " Resonance Absorption in U-238 Metal and oxide' Rods,"

'W.G.'Pettus(1962).

- 2.

NS&E 8, 497 (1960) E. He11 strand, et al, "The Temperature Coefficient of the Resonance Integral for Uranium Metal and Oxide." -

3.- NS&E 11,- 39 (1961) L. Dresner, "Some Remarks on the Effect of em Non-

~

Uniform Temperature Distribution on'the Temperature Dependence of

Resonance Absorption."

4.

WCAP-1434, ~ " Multi-Region Reactor Iattice Studies Microscopie Lattice Parameters in Single and Multi-Region Cores," June 1961.

5 NS&E 19, 172 (1964) A. H. Spano, " Analysis of Doppler-Limited Excursions in a Water-Moderated Oxide Core."

6.

L. Baker, Jr., R. o. Ivins, A. D.' Tevebaugh, Chemical Engineering Division Summary. Reports:

ANL 7325 - July-December 1966 7225 - January-June 1966 7125 - July-December 1965 7055 - January-June 1965

- 7 Conway, Fincel, Hein, Trans. Amer. Nuclear Society, 6, No. 1 (1963)

Page_153 l

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