ML20003A345

Preliminary SAR for Fuel Reload 2 (Cycle 3).

ML20003A345
Person / Time
Site:	Fort Saint Vrain
Issue date:	07/31/1980
From:	GENERAL ATOMICS (FORMERLY GA TECHNOLOGIES, INC./GENER
To:
Shared Package
ML20003A343	List: ML20003A342 ML20003A345
References
NUDOCS 8102030529
Download: ML20003A345 (65)

Text

l O

f D p{ lt??S t 4 "V s, --. li s a s 's , b i s i

i SAFETY ANALYSIS REPORT FOR FUEL RELOAD 2 (CYCLE 3)

FORT ST. VRAIN NUCLEAR

+ GENERATING STATION J

l i

! GENERAL ATOMIC PROJECT 1900 JULY 1980 MEUM!NSHY 8102eao$23

> a y .

CONTENTS

1. INTRODUCTION AND

SUMMARY

. . . ........ .. . . .... . 1-1

2. REACTOR OPERATING HISTORY .. ............... . . 2-1

3. GENERAL DESCRIPTION .. . . . ............ . . . .. 3-1

4. FUEL SYSTEM DESIGN . . . . . . ................. 4-1 4.1. Fuel Design .. . . . . .... ... ..... .... . 4-1 4.2. Mechanical Design . . ......... .. ... .. . 4-2 4.3. 'Ihe rma l De s i gn . . . . . . . . . . . . . . . . .. . ... 4-3 l

4.4. Fission Product Release ................. 4-4

5. NUCLEAR DESIGN . . . . . . . . ........ . .. ... . .. 5-1 5.1. Segment 8 Fuel Loading . ................. 5-1 5.2. Burnable Poison Loading ................. 5-2 5.3. Control Rod Sequence . . ................. 5-4 5.4. Projected Cycle 3 Operation .. . .. ... . . . . . .. 5-6 5.5. Maximum Control Rod Worth .. ...... .. ... ... 5-6 5.6. Core Shutdown Margin . . ........... . .. ... 5-7 5.7. Kinetics Parameters .. ................. 5-9 5.8. Analytical Input . . . . ....... .. . . .. . . .. 5-9 5.9. Core Operating Procedures ... . .. .. ..... .. . 5-10

6. THERMAL-HYDRAULIC DESIGN . . . ........ ...... .. . 6-1

7. SAFETY ANALYSIS ... . . . . ................. 7-1 7.1. Introduction and Summary . ... . ... . .. . .. .. . 7-1 7.2. Incidents .. . . . . . ................. 7-2 7.3. Loss of Normal Shutdown Cooling ... .. . . .... .. 7-2 7.4. Moisture Inleakage . . . ................. 7-3 7.4.1. Stean-Graphite Reaction .... .. .. .. ... 7-4 7.4.2. Hydrolysis of Failed Fuel ... ... .. .. . . 7-4 7.4.3. Fission Product Release from Oxidized Graphite . . 7-5 7.5. Permanent Loss of Forced Cooling (Design Basis Accident No. 1) ... .. .. .. .. .. . 7-5 111

O 7.6. Rapid Depressurization/ Blowdown (Design Basis Accident No. 2) . . . . . . . . . . . . . . 7-7 7.7. Conclusion .. . . . . . . .. . . . . . . . . . . . . . 7-8

... . . . 8-1

8. PROPOSED MODIFICATIONS TO TECRNICAL SPECIFICATIONS 9-1

9. STARTUP TESTS . . . . . . . . . . . . . . . . . . . . .. . . .

.. . . . . . . . . - . . . . . . . . . . . . . . . 10-1

10. REFERENCES FIGURES 3-1. Core regions refueled in reloads 1 through 6 . . . . . . . . . 3-5 3-2. Refueling region age distribution for the equilibrium cycle . 3-6 5-1. Segment.8 poison rod using surplus Segment 7 LBP rods .. . . 5-18 5-2. Identification of control rod groups . . . . . . . . .. . . . 5-19 5-3. Tilt envelope for cycle 3 .. . . . . . . . . . . . .. . . . 5-20 5-4. Axial power distribution during cycle 3 ... . . .. . . . . 5-24 5-5. Maximum allowable rod pair worth versus core average gas ~

outlet temperature . . . . . ................. 5-26 5-6. Temperature coefficient in the initial core with equilibrium Xe-135 and Sm-149 ... . . . . . . . ... .. . 5-27 5-7. Temperature coefficient at end of equilibrium cycle .. . . . 5-28 5-8. Temperature defect versus average core temperature . . . . . . 5-29 TABLES 4-1. Irradiation data base for fuel rods containing near-isotropic graphite shim particles and petroleum pitch binder matrix .. 4-5 4-2. Summary of irradiation data on fuel rods fired at temperatures <1800*C . . . . ................. 4-6 4-3. Segment 8 calculated peak operating conditions versus 4-7 FSAR initial core peak values .... ..... .. .. . . .-

5-11 5-1. Segment 8 design uranium and thorium loadings .. . . . . ..

i 5-12 5-2.- Projected core loadings at the end of cycle 2 .. . . . . ..

5-13 5-3. Use of burnable poison in Segment 8 .... . . . .. . . . .

5-14 5-4. Control rod sequence for cycle 3 . . . . . . . . . . . . . . .

5-5. Calculated control rod group worth and power peaking factors 5-15 with proposed rod sequence . .................

5-16 5-6. Control rod bank worth ... . . . . . . . . . . . . . . . . . .

iv

4 i

TABl.ES (Continu=f) 5-7. Shutdown margins - cycle 3 . . . . . . . . . . . . . . . . . . 5-16 5-8. Kinetics parameters ..................... 5-17 7-1. Potential effects on FSV accident analysis of petroleum-derived pitch binder and graphite shia particles in fuel rods and the use of whole block buffer elements . . . . . . . 7-10 I

y l

2 l

t

1. IhTRODUCTION AND

SUMMARY

This Safety Analysis Report (SAR) is prepared to obtain concurrence to operate the Fort St. Vrain Noclear Generating Station (FSV) through the forthcoming reload cycle (cycle 3). For this cycle, 6 of the 37 fuel regions in the core will be loaded with fresh fr.el elements fabricated by General Atomic Company. The introduction of new fuel elements is consistent with the fuel management program described in the Final Safety Analysis Report (FSAR). The regions to be refueled contain 204 standard fuel elements, 30 control fuel elements, and six bottom control fuel elements, a total of 240 e!ements to be replaced.

This report contains sections describing the operating history of the reactor through February 29, 1980, the fuel system design, the nuclear design, the thermal-hydraulic design, and the safety aspects of the core during cycle 3. The planned startup program for the refueled core is also briefly described.

The replacement fuel elements feature three minor design changes and one manuf acturing process change relative to the fuel design and process described in the FSAR. The safety evaluation for these changes is presented in this report. No changes to the plant Technical Specifications are necessary for cycle 3, and no unreviewed safety questions, as defined in 10CFR 50.59, are presented. None of the peak operating conditions presented in the FSAR are exceeded.

l-1-1 i

2. REACTOR OPERATING HISTORY Initial criticality of the FSV reactor was achieved on January 31, 1974, with initial generation of el' ectricity on December 11, 1976. Prior to February 1,1979, when the plant was shut down for refueling, the initial core had operated a total of 174 effective full power days (ETPD). Cycle 2 operation began on May 26, 1979, and as of June 30, 1980, had accumulated approximately 70 EFPD.

The nuclear performance of the cycle 1 core was, in general, as 1

predicted. Good agreement between measurements and calculations was obtained for shutdown margins, temperature coefficients, xenon worth, and control rod worth (i.e., measurements were well within the acceptance criteria specified for the tests). Initial cold clean criticality was predicted tthin 0.003 Ak. Analyses overpredicted the end-of-cycle (EOC)

reactivity by a few tenths of a percent; however, the difference between observed and expected reactivity remained within the 0.01 Ak limit of Technics 1 Specification LCO 4.1.8.

Fission product release during cycle 1 was very low. Measured circulating activity was approximately a factor of 60 less than the limit provided in Technical Specification LCO 4.2.8. Measurements of plateout activity must avait removal of the first plateout probe.

The nuclear performance of the core during cycle 2, to date, has been much like that during cycle 1. The cooperisons of measured and computed temperature coefficients, control rod group wortha, and reactivity discrepancy indicate that all measured data are well within the acceptance criteria specified for the cycle 2 startup tests.

l 2-1 l

l

The most unusual occurrence, to date, has been the detection, initially in October 1977, of temp 9rature fluctuations. The effects of the fluctua-tions are seen on the nuclear channels, the region exit temperatures, and the steam generator module temperatures. During fluctuations, however, the total core coolant flow and core thermal power remain essentially constant.

During the first refueling, in early 1979, a comprehensive in-core inspection program verified that no damage to the core components had occurred as a result of the fluctuations. Details regarding fluctuation tests and analyses, postulated fluctuation mechanisms, stable core operating configurations, the in-core inspection program, and possible means to eliminate the fluctuations are contained in the following Public Service Company of Colorado subsittals to the Nuclear Regulatory Commission (NRC):

" Safety Evaluation-Reactor Outlet Temperature Fluctuations," P-78137, August 11, 1978.

" November 4 and December 12 and 13, 1978 Temperature Fluctuation Events," P-79032, January 29, 1979.

" Amendment A-1 to SAR for Instrumented CRDs," P-79062, March 19,1979.

"SAR for Core Region Constraint Devices," P-79068, March 23, 1979.

" Report - Fort St. Vrain In-Core Inspection, Region 35 and Region 13, Core Support Llock," P-79193, August 30, 1979.

" Fort St. Vrain Unit No.1 Fluctuation Testing," P-79293, December 4, 1979.

2-2

i

3. GENERAL DESCRIPTION The FSV fuel management scheme is designed so that approximately one-sixth of the core is reloaded at periodic refueling intervals. This docu-ment describes the second reload segment (also known as Segment 8) to be inserted into the FSV core. It is projected that this reload segment will reside in the core up to the 1800 EFPD limit of Technical Specification LCO 4.1.1. Since cycles 1 and 2 are each less than 300 EFPD in length, Segment 8 is the first segment projected to attain 1800 EFPD of operation. This reload fuel segment is designed so that the core performance with the new fuel added satisfies the reactor Technnical Specificationc. These limita-l tions apply to the total core performance.. That is, not only do the freshly loaded refueling regions meet power distribution limitations, but the per-turbations to the remainder of the core are such that the segments remaining in the core from the previous cycle also satisfy the performance require-ments. These performance requirements include core excess reactivity and shutdown margins, power distribution behavior, and all the core safety f considerations discussed in the FSV FSAR.

About one-sixth of the' core is replaced at each refueling. Figure 3-1*

summarizes the scheduled refueling sequence. As can be seen, six refueling regions are reloaded at each refueling, except for. the fif th reload, at which time the central refueling region is also replaced. Segment 8 con-sists of 204 standard fuel elements, 30 control fuel elements, and 6 bottom control fuel elements. In this reload segment,-three of the refueling regions are located in the central portion of the core (regions 4, 8, and

15) and three are located adjacent to the side core-reflector interface

~

(regions 25, 32, and 36). The refueling region sequence was chosen so that freshly refueled regions are never adjacent to each other (except when region 1, the central region, is reloaded). Therefore, each refueling region is surrounded by regions of varying ages.

• Figures and tables appear at the end of each stStion.

3-1

- - - - , _ _ _ . . ~ . _ _ _ . _ _ . . . - . _ , . _ . , _ _ _ _ _ . . _ _ . . _ _ _ _ _ _ _

Figure 3-2 shows the refueling region age distribution for the equilibrium cycle as given in FSAR, Section 3.5. By comparison of this figure with Fig. 3-1, it can be seen that the reload sequence follows that given in the FSAR.

Segment 8 features three minor design changes and one manufacturing process change relative to the fuel design and manufacturing process described in the FSAR. The first of these is use of a new fuel rod matrix formula. The matrix formula used for FSV Segments 1 through 7 contained a coal tar pitch binder, as described in Appendix A.l.l.6 of the FSAR. In the matrix to be used in Segment 8 manufacture, the binder is petroleum-derived pitch. This petroleum pitch contains fewer impurities and is in more ample supply than the coal tar pitch. The matrix will continue to use graphite filler material.

The second design change is use of near-isotropic' graphite shim particles in the fuel rods. As explained in Appendix A.l.l.5 of the FSAR, adjustments of heavy metal loading within the fuel rod volume have been accomplished by varying the volume fractions in the rod of four different fissile and fertile particle size types (i.e., fissile A and B, and fertile A and B). The different size types are shown in FSAR, Fig. A.l.1-8. The method used to control metal loading in the initial core and Segment 7 involved carefully blending these different fuel particle sizes; the four particle sizes were handled simultaneously, and extensive record keeping was involved. The use of graphite shim particles in the FSV fuel rods permits the use of fewer coated particle sizes and provides greater flexibility in meeting fuel loading requirements. No change to the reference coated fuel particle design is required.

The third design change is use of a thick (i.e., whole block) buffer fuel zone at the radial core-reflector interface. The fitat seven FSV fuel segments used a five-row buffer fuel zone (" thin buffer") in standard fuel elements at the core-reflector interface.- (See FSAR, Fig. 3.5-3.) This buffer fuel was not axially zoned. In Segment 8, the width of the buffer 3-2

l i

I l

i zone in the standard fuel element will be increased to include the entire l

element, and the buffer fuel will be axially zoned to assure a proper axial pewer and temperature distribution. In the control fuel elements of the five-column regions at the core-reflector interface, the entire element will continue to be loaded with buffer fuel, as was the case in the first seven segments. However, this buffer fuel will also new be axially zoned. Adop-tion of this design change results in simplified fuel rod loading procedures during the manufacture of buffer fuel elements, increased assurance of pro-per fuel rod loading in these elements, and slight reductions in fuel ten-peratures and calculated fuel element stresses near the core-reflector interface.

There is one change to the fuel rod manufacturing process relative to the process described in the FSAR. The specified fuel rod final heat treat-ment temperature will be lowered from 1800* to 1650*C. Irradiation experi-ence with fuel rods containing petroleum pitch matrix binder and shim particles and heat treated over a range of temperatures indicate that fuel rod performance will not be adversely affected by this process change.

Manufacturing experience has shown that final heat treatment at the lower temperature minimizes the level of various impurities in the fired fuel rods, while maintaining all fired rod product attributes within specification.

All three of the fuel design changes discussed above and the l manufacturing process change are changes relative to the fuel design and process described in the FSAR. None of the changes require any changes to the Technical Specifications, including Design Features. As described in 1

subsequent sections of this report, evaluations of these changes have shown that they have no adverse impact on core performance or plant safety.

Accordingly, they present no unreviewed safety questions as defined in .

10CFR50.59.

t In addition to the design and manufacturing process changes described' above, a process change was introduced during the manufacture of Segment 8 to control the microporosity of the outer pyrolytic carbon coating of the 3-3

TRISO-coated fuel particles. Microporosity and its potential effects on core performance under both normal and transient operating conditions are described in the following Public Service Company of Colorado submittal to the NRC:

" Fort St. Vrain Fuel Particle Coating Failure," P-79157, July 24, 1979:

This submittal includes information on the different microporosity

requirements for particles in fuel rods with coal tar pitch binder (referred to as "R-1" matrix in the submittal) and petroleum pitch binder (referred to as "R-2" matrix). The process changes introduced to control microporosity are also described. These changes do not require any changes to the Tech-nical Specifications, nor do they constitute a change to the fuel design as -

described in the FSAR or an unresolved safety issue. (T. P. Speis, NRC, letter to J. K. Fuller, PSC, February 7, 1980.)

3-4

. y

= . . . ,,

B 87 7 3 . D N . G 13 5 37 21 17 2 10 16 11 5

30 28 RELOAD NO.1 RELOAD NO.4 (SEG 7) (SEG 10) 36 8 y 7 9 1 23 32 15 4 25 14 27 RELOAD NO.2 RELOAD NO. 5 (SEG 8) (SEG 11) l 18 19 33 22 6

3 31 24 13 12

(

RELOAD NO.3 RELOAD NO.6 (SEG 9) (SEG 12)

Fig. 3-1. Core regions refueled in reloads 1 through 6 3-5 l

l N ,

f f

-y ~$

'i  ;.

36 37 20 -

35 6 1 a '

34 18 19 3 21 2 4 1 5 6

j.

wp s 22 7N' 33 u 7 2 m 33 4 6 2 3 2 4 V

.,jl;;..

$h 23 " .- -,_

6 1 3 10

.. 32 16 4 5 3 1 2 4 6 2 N

, .,m 4 11 24 ,_p 31 15 5

,' 5 6 5 3 1 1

-x.

n g

"" 30 14 13 12 25 4 1 5 yd 3 2

' +14 .. N!

,,l .

,,l:is l 28 27 26 .,

~

. , ti; 23 4' 6 2 3 gy

$ st a "4 "

jiflQ W _ Q _ f6 W FUEL REGION REFUELING IDENTIFICATION REGION AGE NUMBER Fig. 3-2. Refueling region age distribution for the equilibrium cycle (before refueling) 3-6

v

4. F1JEL SYSTEM DESIGN 4.1. FUEL DESIGN The fuel elements in this second reload are of the same basic design as those in the initial core and Segment 7. TRISO-coated (Th/U)C2 and ThC2 particles are bonded by a carbonaceous matrix into fuel rods which are cured and loaded into H-327 graphite fuel blocks. The materials and processes used in manufacture are essentially the same as those for the initial core fuel elements, with the exception of the changes described in Section 3. Fuel and burnable poison loadings have been adjusted to accommodate the reactivity requirements of the cycle, as discussed in Section 5.

As described in Section 3, the fuel rods for Segment 8 feature two design changes relative to the fuel design described in the FSAR: (1) use of a new fuel rod matrix- formula and (2) use of near-isotropic graphite shim particles. The integrity of fuel rods containing petroleum pitch matrix binder and near-isotropic shim particles has been demonstrated to be excel-lent for fast neutron fluences up to 12.4 x 1021 n/cm2 (E > 0.18 MeV)HTGR and temperatures up to 1350*C (Refs. I through 8). These same experiments have shown little or no detrimental interaction between fuel particle coat-ings and the fuel rod matrix attributable to the matrix composition. Table i 4-1 summarizes the irradiation data base for fuel rods containing petroleta

! pitch matrix binder and shim particles. Many of these fuel rods have been exposed to fluences and temperatures which exceed the peak exposure conditions for FSV fuel as given in the FSAR. Of the 172 rods irradiated in the tests described in Table 4-1, 28% received fluences in excess of the FSV design peak l

fluence of 8.0 x 1021 n/cm2 (E > 0.18 MeV)HTGR, 22% experienced peak i

centerline temperatures in excess of the FSV design peak temperature of 1260*C, and 6% experienced both fluences and temperatures in excess of the FSV l

design peak values.

4-1 l

l

l 1

i In addition to these fuel rod design changes, there is one change to the fuel rod manufacturing process relative to the process described in the FSAR. In FSAR Appendix A.1.1.6.2, the fuel rod final heat-treat temperature is described as approximately 1800*C. In Segment 8 fuel production, the final treat temperature is approximately 1650*C. The irradiation experience with fuel rods containing petroleum pitch matrix binder and shim particles shown in Table 4-1 includes fuel rods heat treated over a range of tempera-tures. Table 4-2 summarizes data on those fuel rods fired at temperatures less than 1800*C. These data indicate that the irradiation performance of these rods is acceptable and comparable to that expected for other fuel rods in the FSV core. The range of exposure conditions for these fuel rods encompasses the peak design conditions for FSV fuel. Accordingly, it is not necessary to expose fuel rods containing the new matrix formula and shim particles to the higher heat-treating temperature.

The impact of these fuel design changes on core performance is discussed in subsequent sections. These design changes do not require any changes to the FSV Technical Specifications, and as shown in the following sections, do not involve any unreviewed safety questions, as defined in 10CFR50.59.

4.2. MECHANICAL DESIGN The fuel elements for Segment 8 of the FSV core are fabricated from H-327 graphite. Both the structural material and the configuration of these fuel elements are identical to those used in Segments 1 through 7. Table 4-3 provides a summary of the Segment 8 fuel element stress, strain, and bowing analyses described in this section.

Stress analysis was performed for the FSV Segment 8 fuel elements using the methods discussed in the FSAR. Operating and shutdown strain and stress 4-2

4 distributions were calculated for the axial and radial orientations through-out the lifetime of the fuel elements. All operating and shutdown stresses, including those in the thick buffer elements, were less than those predicted for the initial core (FSAR Section 3.4.2.1.1).

During core operation, the Segment 8 fuel elements will be exposed to fast neutron irradiation, which will induce dimensional changes in the H-327 graphite. An analysis was performed to calculate the expected dimensional changes of the Segment 8 fuel elements. The maximum axial dimensional changes of all Segne.nt 8 fuel columns were less than those of the initial core fuel columns shown in FSAR Section 3.4.2.1.2. The maximum calculated EOL bowing for the thick buffer elements is 0.09-in., which was the maximum bowing predicted for the initial and equilibrium core H-327 thin buffer elements (FSAR, Section 3.4.2.1.2).

4.3. THERMAL DESIGN The selection of fuel rod and burnable poison loadings and of the control rod program for cycle 3 (see Section 5) is made so that the cycle 3 power distribution falls within the limits described in the FSAR. No changes are planned for the operation of the core during cycle 3 (i.e.,

helium temperature at the core inlet and average outlet temperature will be enveloped by the FSAR reported values). Accordingly, the temperature limits presented in the FSAR will not be exceeded during cycle 3, and use of the thick buffer design will have no adverse impact upon fuel temperatures.

These conclusions are supported by analyses using the COPE code, which is discussed in the FSAR. The results of these analyses are also shown in I rable 4-3.

l l Introduction of petroleum pitch matrix binder and near-isotropic graphite shin particles in the fuel rod design for Segment 8 results in

increased fuel rod thermal conductivity. Measured thermal conductivities for fuel rods with this matrix composition and shin particles (Ref. 9) are well above the value of 4.0 Btu /hr-f t-*F assumed in the FSAR. Nevertheless, the conservative FSAR value of thermal conductivity was still assumed in the thermal analyses for Segment 8.

I 3

l

i l

i 1

l 4.4. FISSION PRODUCT RELEASE

\

During Cycle 3, the FSV core is expected to be operated within the limits presented in the FSAR and contained in the Technical Specifications.

Accordingly, the fission product release characteristics of the fuel are expected to be within design limits and the design radionuclide inventories presented in Section 3.7 of the FSAR will not be exceeded. These conclusions are consistent with operating experience gained during cycles 1 and 2.

Use of petroleum pitch matrix binder and near-isotropic graphite shim particles in the Segment 8 fuel rods is expected to result in lower fission product release as compared with fuel rods like those in Segments 1 through

7. As noted above, Segment 8 fuel rods will have higher thermal conductiv-ity, resulting in lower fuel temperatures. Since both gaseous and metallic fission product release are temperature dependent, lower temperatures will lead to lower fission product release. In addition, with regard to metallic fission product release, data have been obtained for sorption of metallic fission products (particularly cesium and strontium) on both coal tar pitch and petroleum pitch based matrix materials (Refs. 10 and 11). These data indicate that the sorptivity of metallic fission products on matrix mate-rials is the same or greater due to the replacement of coal tar by petroleum pitch. The increased matrix sorptivity results in lower release of metallic fission products from the fuel rods to the graphite webs, and hence, to the helium coolant. Thus, Tables 3.7-1 and 3.7-2 of the FSAR remain appropriate as circulating and plateout activity source terms for accident analyses.

4-4

TAPLE 4-1 IRRADIATION DATA RASE FOR TUEL RODS CorffAlHING t.T.An-ISOTROFIC CRArtilTE SHIM PARTICl.ES AND PETROLEUM-FITCH BINDER MATRIK(a) irradiation Conditions Structural thamber Volume Integrity of Peak Center!!ne Average Fluence Rode Temperature Tempe ra t ea re 1021 n/ce2 Fuel Shin Capsule Test Shin Material Tested ( *C)( b) (*C) (E > 0.18 MeV)HTCR(C) Rod Fatticles References Base Frogram Tests itRB-4, lite-5, IIRB-6 CI4C 1099(d) 12 1250 1050-1100 2.3-10.5 (e) (f) 1 HT-24 HT-25 UCC/TS-1240X 8 850-1020 950 2.7-8.6 (e) (f) 2 CLCC 1099(d) 6 3.6-9.3 Fl3Q Cl4C 1099(d) 18 915-1175 835-!!60 4.3-9.5 (e) (f) 3 F13R, F13S CLCC 1099(d) 36 995-1350 910-1295 3.6-12.4 (s) (f) 4 F13T CLCC 1099(d) 24 1100-1300(8) (h) 3.0-8.0 (e) (f) 5 UCC/TS-12401 6 3.5-6.4 2-

.$ OF-1 CI4C 1089(d) 21 770-1190 (h) 3.9-9.7 (e) (f) 6 Foreign Program Tests SSL-1 (loop) CLCC 1099(d) 27 a 12WI5) '(h) 2.7-4.6 (e) (f) 7 CF-1. CF-2, CF-3 CLCC 1099(d) 14 1050-1250(8) (h) 5.5-9.0 (e) (f) 8 (a)Matriz binder is petroleum pitch.

(b)FSV design peak temperature = 1260*C.

(c)FSV design peak, fluence = 8.0 x 1021,fe.2 (E > 0.18 Mev)nTCF.*

(d) Dest'gnation for near isotropte n-451 graphite particles.

(*)Resed on visual and metallographic observations defining fuel rod surface appearance, a minimal amount of surf ace

. cracking or metria spalling, and minimal particle debonding.

(f) Based on metallographic observations, no cracking or fragmentation in shin particles.

(8)posinal design values.

(h)not determined.

TABLE 4-2

SUMMARY

OF IRRADIATION DATA ON FUEL RODS FIRED AT TEMPERATURES (1800*C Irradiation Exposure Conditions lleat Visual Treatment Volume Appearance /

Irradiation Fuel Rod Firing Temp Av Temp Fluence *321 n/cm2 Structural Capsule Designation (*C) (*C) (E > 0.lb MeV)llTCR Integrity (a)

G3-2A 1500 855 4.3 Excellent P13Q G3-2B 880 5.7 Excellent P13R 2A 975 11.8 Excellent SC 1180 8.5 Good I 1045 9.3 Excellent ID

-P13S SC 1285 8.4 Good ID 965 9.1 Excellent 6E U 965 3.5 Good (a) Excellent implies no matrix spalling, cracking or particle debonding. Good implies a minimal degree of one of these visual defect categories.

4 0 .

TABLE 4-3 SEGMENT 8 CALCULATED PEAK OPERATING CONDITIONS VERSUS TSAR INITIAL CORE PEAK VALUES Parameter FSAR Peak Value Segment 8 Peak Value(a)

Axial operating stress (psi) 450 318 Radial operating stress (psi) 200 122 Initial axial thermal stress (psi) 150 65 Initial radial thermal stress (psi) 180 51 Column axial strain (%) 2.1 2.0 Buffer element bowing (in.) 0.09 0.09 Fuel temperature (*F) 2300 2141(b)

(* Values calculated using FSAR methods.

Peak fuel temperature in core during cycle 3.

l i

4-7

t

5. NUCLEAR DESIGN 5.1. SEGMENT 8 FUEL LOADING In the initial core design, the fuel was zoned both radially and axially to achieve the desired power distribution and to mock up the equilibrium cycle. The core was divided into four radial zones and two axial zones. The radial zones consisted of the central refueling region, the six refueling regions adjacent to the central regions, the 12 refueling regions adjacent to these six regions, and the outer 18 refueling regions adjacent to the side reflector. In addition, the five. rows of fuel rods in the outer 18 regions that were immediately adjacent to the side reflector interface contained a buffer fuel loading with low uranium and high thorium loading to reduce the power peaking at the reflector edge. There were two axial zones consisting of the top and bottom three layers of fuel elements.

The buffer zone was not axially zoned. These various fuel zones, combined with a partial mock-up of the equilibrium core fuel distribution, made it necessary to fabricate 13 different fuel blends or compositions for the initial core fuel elements.

For the design of the reload segments, these different fuel zones are essentially maintained, except that the second and third radial zones are combined and the width of the buffer zone is increased. Since the central refueling region is not reloaded in this reload segment, only six new fuel compositions or blends (numbered 19 through 24) are used for the Segment 8 fuel. Refueling regions 4, 8, and 15 require a top and bottom fuel loading, and refueling regions 25, 32, and 36 require a top and bottom fuel loading in addition to the two loadings for the buffer fuel at the core-side reflector interface, which is now axially zoned.

5-1

In addition to the six fuel blends for Segment 8 discussed above, a relatively small number of surplus Segment 7 production fuel rods will be loaded into the Segment 8 feel elements. Five different fuel blends (numbered 14 through 18) were used in Segment 7 fuel. The surplus Segment 7 fuel rods constitute approximately 3% of the total nunber of fuel rods in Segment 8. Their location in the core has been chosen to ensure that the core power distribution will be within the limits of the Technical Specifications.

A summary of the 11 different fuel blend uranium cod thorium loadings used for the second reload segment is given in Table 5-1. Fuel blends 14, 15, 18, 19, and 20 are used in the inner refueling regions, and blends 16, 17, 21, 22, 23, and 24 are used in the outer refueling regions. Blends 23 and 24 are used in the buffer zone.

It is anticipated that the reactor will have operated for up to 374 EFFD, generating a total of up to 312,000 MW(d) of energy at the time of the second refueling. The projected heavy metal loadings in the remaining core segments at the EOC 2 (374 EFPD) are given in Table 5-2. The maximum burnup in fissile particles is about 10.4% FIMA and in fertile particles about 0.9% FIMA. These burnups are substantially lower than the limiting values given in the FSAR, Appendix A, Table A.1.1-2. The maximum fast finx (E > 0.18 MeV) exposure in the discharged segment is about 1.6 x 10 21 nyt.

5.2. BURNABLE POISON LOADING Six holes are provided in each standard fuel element, one at each corner, for insertion of burnable poison rods; four holes cre provided in each control fuel element. For Segment 8 fuel, there will be no burnable poison in the control fuel elements; all six poison holes in the standard elements may be used depending on burnable poison loading requirements.

In the initial core and Segment 7, different poison rod types, varying in their boron loading, were used to provide reactivity control and to adjust power distributions in fresh fuel regions. The poison rods were 28.5 i

5-2  :

l 1

in. long. In the case of Segment 7, it was necessary to have available several different poison rod types, only two of which were ultimately used.

The poison rods to be used were chosen based upon the reactivity requirements of the core at the EOC 1 and were inserted into the reload segment just prior to refueling.

To avoid this unnecessarily complex and costly procedure, the design of the poison rods will be slightly changed for Segment 8 and later segments.

Instead of using single rods 28.5 in. long, shorter rods (1.98 in. long) will be loaded into the poison holes in combination with graphite dummy spacer rods (also 1.98 in. long). Each poison hole will either contain 14 poison rods or spacers or will be lef t empty, and the boron loading in each element will be controlled by varying the combination of poison rods and spacers. The new poison rods will be the same in diameter (0.45 in.) and made of the same graphite as the 28.5-in single rods previously used. The spacers will be made from HLM braphite and will also be 0.45 in. in diameter. This change does not require any changes to the Technical Specifications, nor does it constitute a change to the core design as described in the FSAR.

For Segment 8, 28.5-in. poison rods left over from Segment 7 will be cut into 1.98 in. nominal length rods and loaded into the standard fuel elements in combination with graphite spacers. The poison rods left over from Segment 7 which will be used in Segment 8 are known as " type 4" and

" type 11" rods. They differ only in their boron loadings. The standard i fuel elements in the upper core half will be loaded with 14 poison rods in each of the sia poison holes. In the lower core half, each of the six poison holes in the standard fuel elements will be loaded with five poison j rods, four spacer rods, and five more poison rods, in that sequence. The poison holes in the control fuel elements will be left empty. The use of l

burnable poison rods in Segment 8 is sunmaarized in Table 5-3, and the loading sequence for the type 4 and type 11 rods is shown in Fig. 5-1. The

~

burnable poison loading for Segment 8 described above was used for the cycle 3 depletion analyses referred to in the following sections. This loading will be used in Segment 8 unless the reactivity characteristics of the core or the length of cycle 2 deviate significantly from current projections.

l 5-3 1

5.3. CONTROL ROD SEQUENCE Technical Specification LCO 4.1.3 states that a control rod sequence will be specified for each fuel cycle and that the sequence will always be followed, except for rod insertion resulting from a scram or rod runback or during low power physics testing. The control rod sequence for use during cycle 3 is given in Table 5-4. The cycle 2 rod sequence is included for comparison. The identification of the control rod groups is shown in Fig.

5-2.

The regulating rod is located in the central refueling region (rod group 1). This group is partially withdrawn before criticality is achieved and then maintained in its most reactive control rod position for the remainder of the operation. In this manner, minor reactivity adjustments can be made most rapidly with the minimum amount of control rod motion.

This is consistent with the method of operation utilized for the control rod sequence of cycles 1 and 2.

A summary of the calculated power peaking factors obtained using the control rod sequence for all control rod configurations is given in Table 5-5. This includes all of the control rod configurations in which the control rod groups tce either fully inserted or withdrawn, including those suberitical configurations during the withdrawal of the first few control rod groups. Any configuration with a partially inserted control rod group will have peaking factors lying between those calculated when that group is fully inserted and fully withdrawn. The control rod worths and power peaking factors in Table 5-5 were calculated at the beginning of cycle (BOC) with equilibrium xenon. Rod worths at middle of cycle (MOC) and EOC will be essentially unchanged. Power peaking factors during cycle 3 depletion are discussed in Section 5.4.

The control rod configurations shown in Table 5-5 may be separated into .

four categories of reactor power operation: (1) full power, (2) 20% to 100%

power, (3) 0% to 20% power, and (4) suberitical. Full power operation may 5-4

i l

i

\

l l

l 1

be achieved with configurations ranging from one control rod group (three  !

rod pairs) fully or partially inserted at the EOC 3, when the excess reactivity is relatively low, to three control rod groups (nine rod pairs) fully or partially inserted at the BOC, when fission product poisons such as xenon are not present. (In accordance with the requirements of the Technical Specifications, only one control rod group in addition to group 1, the regulating rod, will be partially inserted at any time.)

At less than full power, in the range of exit gas temperatures between 1460* and 950*F, when the xenon level and the core temperature are lover, configurations ranging from four control rod groups (12 rod pairs) to five control rod groups (15 rod pairs) fully or partially inserted can be expected, depending on the temperature, power level, and the rate at which power was increased from the previous level.

The third category covers the rise-to power phase from the cold critical condition to about 20% power (gas outlet temperature 1950*F).

Different limiting operating conditions are applied to this phase of reactor operation by the Technical Specifications. The initial cold criticality, following the refueling operation, was calculated to be achieved with seven

rod groups (21 rod pairc) fully or partially inserted (control rod group 2A 45 in, withdrawn). ,

The last category covers the suberitical crentrol rod configurations where region peaking factors (RPF) or intraret; ion tilts are not meaningful, and consequently, are not given in Table 5-5.

From the data given in Table 5-5, it can be seen that the calculated l

l power peaking factors for the various power levels do not exceed those given 1

in Technical Specification LCO 4.1.3. This is true for both the radial region power peaking factors and the intraregion peaking (column tilt) factors. The radial region peaking is below 1.83 for all configurations l involving less than 19 control rod pairs. For low power operation (gas outlet temperature 1 950*F), when more control rod pairs are inserted, the region peaking in below 3.0 until the suberitical configurations are reached. In the same manner, the intraregion peaking factors are also 1

1 5-5 l

I l

i 1

i I

i acceptable. It is clear, tht.refore, that use of the thick buffer design at the core-reflector interface, use of the new matrix formula and graphite  ;

shim particles in the fuel rods, and the new burnable poison design have no adverse impact on power peaking factors.

5.4. PROJECTED CYCLE 3 OPERATION This section presents the results of cycle 3 depletion analyses using design methods discussed in Section 3.5 of the FSAR. Fuel and burnable poison loadings discussed previously were used as input (see Sections 5.1 and 5.2).

Figures 5-3a through 5-3d present envelopes encompassing projected RPFs and column tilts during cycle 3 depletion. The results indicate that RPFs and tilts during cycle 3 will be well within the allowable limits set by LCO 4.1.3 and that the use of the thick buffer design at the core-reflector .

interface has no adverra impact on power distributions.

Axial zoning of the Segment 8 fuel and burnable poison is provided (1) to produce a power distribution which tends to reduce axial fuel temperature peaking and (2) to maintain the desired axial power distribution with deple-tion. The calculated axial power distributions in Segment 8 fuel during cycle 3 are shown in Figs. 5-4a and 5-4b. The calculations were carried out with the FEVER code by modeling the Segment 8 fuel by itself. Use of this modeling approach is conservative relative to full-core models. It can be seen that the LCO 4.1.3 limit on peaking factor in the lower fuel layer (axial power factor <0.90) is not exceeded and that the new burnable poison design does not result in unacceptable axial power peaking.

5.5. MAXIMUM CONTROL ROD WORTH The rod withdrawal accident for the full power operating condition described in the FSAR (Section 14.2, per Amendment 24) assumes withdrawal of a control rod worth of 0.012 Ak at EOC and an equilibrium EOC temperature 5-6

coefficient. This 0.012 ak control rod worth, because of lower reactivity feedback resulting from the less negative temperature coefficient at EOC, is equivalent to a rod worth of about 0.016 Ak at BOC. Since the temperature coefficient of reactivity is more negative for cycle 3 than for the equilibrium cycle (see Section 5.7), a maximum single rod pair worth for cycle 3 BOC of less than 0.016 ak is acceptable. In addition, the calculated worth of any rod pair in any operating critical control rod configuration must be less than 0.047 ak, per LCO 4.1.3.

The =aximum worth control rod pair which occurs during the specified control rod sequence is given in Table 5-5. Since the temperature defect is more negative during cycle 3 (see Section 5.7) than for the equilibrium core, the consequences of the rod withdrawal accident (RWA) at rated power of a rod worth between 0.012 and 0.016 ak are no more severe than the withdrawal of 0.012 Ak at rated power discussed in the FSAR. A conservative estimate of the maximum worth of a rod. pair which is equivalent (in cycle 3 RWA consequences) to the FSAR results is shown in Fig. 5-5.

At full power, the maximum calculated rod worth at the BOC 3 (no xenon) is 0.015 Ak, which is less then the i3AR limit at BOC of 0.016 ak. During the cycle, the maximum rod worth at full power is less than 0.012 ak, which is less than the conservative estimate of the allowable rod worth at the EOC 3 shown in Fig. 5-5. For other categories of reactor operation discussed in l

Section 5.3 (i.e., at less than full power), the maximum worth rod remains below the allowable limit shown in Fig. 5-5, and consequently, meets the l requirement of LCO 4.1.3 of the Technical Specifications.

5.6. CORE SHUTDOWN MARGIN The Technical Specifications require that the reactar be capable of being shut down (0.01 ak suberitical) with any one control rod pair withdrawn at room temperature and with any two control rod pairs withdrawn at refueling temperature. This is consistent with the policy that, during the rod withdrawal accident postulated in Section 14.2 of the FSAR, the 5-7 l

i

(

I control rod pair being withdrawn is continuously withdrawn until fully removed during the accidens and is not capable of being reinserted. The requirement for being shut down with any two control rod pairs withdrawn is established to allow for operation with one inoperable control rod pair.

The net negative reactivity insertions following a scram during cycle 3 ara given in Table 5-6. Since the excess reactivity at full power with equilibrated xenon and protactinium is about 0.020 Ak, the instantaneous minimum shutdown margin (SDM) with the two maximum worth rod pairs inoper-able is about 0.120 Ak. This shutdown margin is reduced due to core cool-down (in a matter of hours), due to xenon decay (in a matter of a few days),

and due to transformation of Pa-233 into U-233 (in a matter of weeks). To show that the shutdown margin is satisfactory for these cases with one or more control rod pairs assumed inoperable, the summary in Table 5-7 gives calculated core shutdown margin at the BOC immediately after refueling and at the MOC and EOC. It can be seen that for all cases of interest, the core shutdown margin is larger than the 0.01 Ak as required in the Technical Specifications. Accordingly, the use of the thick buffer design at the core-reflector interface and use of petroleum pitch matrix and graphite shim particles in the fuel rods have no adverse impact on shutdown margins.

For the case with only the maximum worth rod pair inoperable, the minimum shutdown margin at room temperature with xenon and Pa-233 decayed (the most reactive case) is 0.045 Ak. At refueling temperature (220*F),

this value would be 0.050 Ak for the same condition. However, since the second maximum worth rod out can add as much reactivity as 0.035 Ak, it may not le yossible to maintain a shutdown margin of 0.01 Ak with these two rod pairs out if the Pa-233 is allowed to fully decay to U-233. This situation l

does not conflict with the LCO, however, since the requirement for the two rod pairs out at refueling temperature is established to allow for operation with one inoperable control rod drive assembly. An adequate shutdown margin can be maintained for a period of at least two weeks in this core condition.

If the inoperable unit cannot be repaired in this time period, the reserve shutdown system can be used to maintain adequate reactivity control.

5-8

5.7. KINETICS PARAMETERS The kinetics parameters for the initial cycle and the equilibrium cycle are given in Table 5-8 (taken from the FSAR, Amendment 14). The equilibrium cycle kinetics parameters represent a conservative estimate of cycle 3 kin-

' etics. The temperature coefficients for the initial cycle and the equilib-rium cycle are shown in Figs. 5-6 and 5-7, respectively. The temperature coefficient for cycle 3 lies between the values for the initial and equilib-rium cores as illustrated by the temperature defects sbown in Fig. 5-8. For conservatism, the temperature defect at the EOC 3 was used to determine the allowable reactivity worth of irdividual rod pairs in the specified control rod sequence (see Section 5.3). Since the maximum rod worth for the core at rated power does not exceed the value calculated for the equilibrium core, a RWA during cycle 3 would have less severe consequences than a comparable case at equilibrium. Thus, the requirements of LCO 4.1.3 are met for cycle 3 operation. Use of the thick buffer design and the changes in fuel rod matrix composition have no adverse impact on kinetics parameters.

5.8. ANALYTICAL ItTUT Nuclear analyses were carried out u, sing the same methods applied to the analyses presented in the FSAR and the Segment 7 SAR. The design of Segment 8 introduces no new aspects to high-temperature gas-cooled reactor (ETGR) core design techniques; consequently, there was no need to develop or adapt l

any new methods or procedures for the nuclear design..

i The depletion analyses described in this chapter were performed by l

l simulating the actual core power history for cycle 1 (174 EFPD) and the first

( 16 EFPD of cycle 2. Continuous operation at 70% power for the balance of cycle 2 and at 100% power for cycle 3 was assumed. Cycle 2 was assumed to l

' continue to a total length of 200 EFPD. Cycle 3 and subsequent cycles were assumed to be 300 EFPD in length.

I i

l l

5-9 a l i l ]

l L

0 5.9. CORE OPERATING PROCEDURES Core operating procedures will be the same as those for cycles 1 and 2 and those planned for equilibrium cycles. The ocly difference will be the control rod withdrawal sequence discussed in Section 5.3.

5-10

TABLE 5-1 SEGMENT 8 DESIGN URANIUM AND THORIUM LOADINGS Total Uranium Total Thorium Fuel Blend (kg) (kg) 14(a) 1.5 17.7 15(a) 1.7 24.7 16(a) 2.0 19.6 17(a) 2.2 26.9 18(a) 0.7 15.9 19 57.5 610.3 20 34.5 488.5 21 32.4 283.4 22 18.1 219.5 23 24.3 320.4 24 14.5 248.9 Total 189.4 2275.8 (a) Fuel blend originally manufactured for Segment 7.

l l

e 5-11

TABLE 5-2 PROJECTED CORE LOADINGS AT THE END OF CYCLE 2(a)

Nuclide Weight (kg)/ Segment Nuclide 2(b) 3 4 5 6 7 Th-232 2730.6 2572.2 2365.5 2841.2 2363.3 2254.1 Pu-233 + U-233 34.7 33.1 31.1 37.2 31.1 17.4 U-234 2.4 2.5 2.5 2.9 2.5 1.7 U-235 47.1 63.5 74.5 82.7 74.5 124.8 U-236 7.5 10.1 11.8 13.1 11.8 8.6 U-238 5.1 6.9 8.1 9.0 8.1 10.3 Np-239 + Pu-239 0.11 0.15 0.17 0.19 0.17 0.17 Pu-240 0.04 0.05 0.06 0.07 0.06 0.04 Pu-241 0.03 0.04 0.04 0.05 0.04 0.02 l (a)374 EFPD.

(b)This segment is discharged. -

1 .

l l

l i

l l

{

l

! 5-12 I

IABLE S-3 ,

USE OF BURNABLE POISON IN SECMENT 8 Top Axial Zone (a) Botton Axial Zone (b)

Type 4 Type 11 Type 4 Type 11 g Bnat/em3 0.0148 0.0553 0.0148 0.0553 Nominal rod diameter (in.) 0.45 0.45 0.45 0.45 Nominal rod length (in.) 1.98 1.98 1.98 1.98 Rods / standard fuel element 42 42 30 30 Rods / control fuel element 0 0 0 0 LBP inserted into fuel elements containing:

(a) Fuel blends 14, 18, 19, 21, and 23.

(b) Fuel blends 20, 22, and 24.

5-13

TABLE 5-4 CONTROL ROD SEQUENCE FOR CYCLE 3 Cycle 3 Cycle 2 Coctrol Rod Group Group Sequence Withdrawn Withdrawn Regions 1

1 2B(a) 2B(*) (3, 5, 7) 2 4F(a) 4A(a) (20, 26, 32) 3 4C 4C (22, 28, 34)

, 3A 1 (half out) 1 (half out) (1) 4 4D 4E (24, 30, 36) 5 4A 4F (25, 31, 37) 6 2A 2A (2, 4, 6) 7 4B 4D (23, 29, 35) 8 4E 3A (8, 12, 16)-

9 35 3C (10, 14, 18) 10 3D 45 (21, 27, 33) 11 3C 3D (11, 15, 19) 12 3A 3B (9, 13, 17) 13 1 (last half) 1 (last half) (1)

(a) Rod groups used for rod runback.

l l

l l

l l

5-14

r TABLE 5-5 CALCULATED CONTROL ROD CROUP WORTH AND POWER PEAKING FACTORS WITH PROPOSED ROD SEQUENCE 4 . (BOC EQ. XE)

Maximum Worth Rod Control Rod Group Cumulative Max Max Configuration- Worth Worth Max Tilt (b) Tilt (b) Worth Rods Inserted (AK) (AK) RPF(a) Rodded Unrodded (&K) Reg No rods in 0.0000 0.0000 1.28 --

1.32 -- --

Rod 1 (half in)' O.0026 0.0026 1.28 --

1.31 0.0026 1 4 rods (+3B)(c) 0.0174 0.0200 1.33 1.19 1.28 0.0086 17

() 0.0191 0.0391 1.54 1.19 1.23 0.0119 13 7 rods (+3D) 10 rode (+48) l 0.0113 0.0504 1.42 1.30 1.31 0.0145(e) 11 13 rods (+3C) 0.0175 0.0679 1.60 1.36 1.30 0.0140 11 2 1.29 0.0163 16 rods (+3A) 0.0161 0.0840 1.49 1.36 14 u, 19 rods (+4D) 0.0087 0.0927 1.75 1.49 1.29 0.0150 19 3 1.48 1.32 0.0172

d. 22 rods (+2A) 0.0184 0.1111 2.35 19 25 rode (+4F)

O.0225 0.1336 2.11 1.38 1.17 -- --

28 rods (+4E) 0.0103 0.1439 -- -- -- -- --

28 rode (1 fully in)L 4 0.0043 0.1482 -- -- -- -- --

31 rods (+4C) i 0.0190 0.1672 -- -- -- -- --

34 rods (+4A) 0.0063 0.1735 -- -- -- -- --

37 rode (+28) , 0.0385 0.2120 -- -- -- -- --

Initial criticality at 0 days was calculated with group 2A withdrawn 45 in.

(a)RPF = region peaking factor = region power / core average region power.

(b) Tilt = column peaking factor /RPF.

(c) Refers to rod group sequence (see Table 5-3) (4 rods = rod 1 + group 38).

( Power range defined in Section 5.3.

• }BOC maximum rod worth.

TABLE 5-6 COWrROL ROD BANK WORTH Cycle 3 BOC(a) MOC EOC Total bank worth, AK 0.222 0.213 0.216 (37 rod pairs inserted)

Total bank worth, AK, less 0.177 0.167 0.170 maximum worth rod pair Total bank worth, AK, less 0.139 0.134 0.136 two maximum worth rod pairs (a) Xenon assumed at equilibrium.

TABLE 5-7 SHUTDOWN MARGINS - CYCLE 3 Shutdown Margin,. AK Number of Inoperable Rods BOC HOC EOC O(a) 0.087 0.093 0.103 1(a) 0.050 0.045 0.055 l 2(b) 0.021 0.039 0.044 l

(a)With no or one rod pair inoperable, the core is assumed to be at room temperature with complete Pa-233 f decay.

(b)With two rod " pairs inoperable, the core is assumed l to be at refueling temperature with a two-week Ps-233 l decay. l l

i 1

5-16 i

, - -, --n _ _

TABLE 5-8 KINETICS PARAMETERS Cycle 1 Equilibrium Cycle BOC, with Xe 150 EFPD BOC, with Xe EOC Fractional productions From U-233 0.0 0.19 0.38 0.48 From U-235 1.0 0.81 0.62 0.52 Prompt neutron lifetime, see Hot 2.69 x 10-4 3.17 x 10-4 2.85 x 10-4 3.41 x 10-4 Cold 2.43 x 10-4 2.81 x 10-4 2.64 x 10-4 3.09 x 10-4 0.00650 0.00577 0.00505 0.00451 y Effective delayed neutron fraction U Delayed neutron decay constant, 1, sec-1 Precursor 1 0.01243 0.01249 0.01250 0.01251 2 0.03050 0.03088 0.03126 0.03164 3 0.114 0.1136 0.1170 0.1199 4 0.3013 0.3025 0.3047 0.3063 5 1.136 1.136 1.135 1.135 6 3.013 2.981 2.913 2.859 Delayed neutron fraction,.8 Precursor 1 0.000214 0.000219 0.000220 0.000222 2 0.001424 '01 0.001186 0.001099 3 0.001274 0.001 0.001046 0.000961 4 0.002568 0.0022 0.001874 0.001619 5 0.000748 0.0005 0.000516 0.000430 6 0.000273 0.0002 0.000204 0.000179

P s

l*]lSON R00 STACK IN:

TOP AXIAL BOTTOM AXf AL g ZONE ZONE 3 I POISON R00 TYPE MEAN BORON CONC.(G/CM )

1.98 IN. 11 11 0.0148 4

U 4 0.0553 k 11 l

11 J 4 4 G = GRAPHITESPACER R00 4 11

Y w -

• 4 G 4 G 4 G 4 G 4 11 11 4 11 4 11 11 11 11 l

Fig. 5-1. Segment 8 poison rod using surplus Segment 7 LBP rods 4

I O

l N

a

=

w. ,.

.j,

'! 'f 35 3 37 N 40 4A H

'S a 34 is ai 3C 30 3A 48 4C -

3 ss ,, s  ; t e n 38 28 2 38 4C 48 .

s:,:.

1. 4D g,

--- -- 32 is 1 in 23 ._ , ._

2A 1 28 ,, s 4A 3A 3C

<;,;"p. 4 11 24 31 15 5 ,

2A 3 4E

.., 4F 3 2B "i n n.

b 14 13 12 25 30 3 4F E.' ' ~

- 4E 3C 8 3 4A '" l 4$ 40 2: 2 rs

' 8 ,, 4 4 .-:. ';

1,, t/

,;g n 'N.id. , !y;,,

I 4;

.m - $

s FUEL REGION IDENTIFICATION NUMBER Fig. 5-2. Identification of control rod groups

?-19

I

=

=

4 3

1 s

n o

i g

e r

d e

d N d O o I

2 r T I n A 1 u

C I ) .

F a I

C: S (

ET PI  :

SM LI 3 L3 4 A e C 3.18. 13 I T L

l c

N 1 I y H 4: T: T c COF ECP L TLRT I

r o

f e

p o

l e

I 1

v 1

n e

t l

i T

3 5

g

_ i F

- - - 0 1

_ 5 0 5 0 1 1 0

&a Y$

!  : , : ; i i ii ' :

2 -

1 s

n o

i g

e r

d e

d d

o N r O

I y T l A l C

I a

F i I t C: S r E T a PI p SM LI 0 )

A L3 b C38.4 I 1 1 (

N 1 T H 4 : T: 1 L 3

COF IL ECP 1

I T

TLRT e l

_ c y

c r

o f

e p

o l

e v

n e

t l

i T

3 5

g i

F

- 0

- 1 0 5 0 1 0 tc Ia 4 e1 1

l i.0 h 0.5 -

m TECHNICAL SPECIFICATION h

" LCO 4.1.3 LIMITS:

RPF: 1.83 TILT: 1.46 0 I 1.0 1,1 j,2 TILT Fig. 5-3. Tilt envelope for cycle 3: (c) fully rodded regions

l' ,' i \

3 1

s n

o i

g e

r l

l l

2 a 1

)

d

(

3 e

_ l

_ T c L y

. I c T

- r o

f e

_ p o

l e

v t

1 n

1 e t

_ l i

T 3

5 g

i F

- 0

- - 1 5 0 5 O

_ 1 1 0 Eco

,a vghw

_ i! ;I 3li :.1 iii. 1 : i, .! !1 i:i4 I i! !4

• l i a .

2.0 BOC - CYCLE 3 I --- EOC - CYCLE 3 i I 1.5 -

e /\ f o J 13

/ ,/

2 /

I%

$ l l e

u I

I

,I ~~\

1.0 -

/

5 I / "\

w N \ / \

h m  % s

• - 2: N

& \

$ ~\

W

! 0.5 -

\\

I I I I I I I

0 1 2 3 4 5 6 LAYER NUMBER Fig. 5-4. Axial power distribution during cycle 3: (a) unrodded Segment 8, outer radial zone i

9

e n

o z

l a

i d

a r

r 3 3 e E E I n

L L n C C i Y Y ,

C C 6 8 C C s t n

O 0 B E .\

I e m

- s g e

N S

- \ 5

' d

- \ e d

\ d

\ I o o r n

N u

)

,N _ _ 4 b

R (

g

-s I E

B  :

f M 3

/. U

' 3 N

R l e

E c Y y A c I

L g n

i r

u 2 d

/

/ n g o I

i i t

/ u

/ b i

/ r

/ 1 t

/ s Il

\ i d

I r

e w

o p

l a

- - - i x

0 5 0 5 0 A

'2 1 1 0

a 4 a 4w g!m _4R $s5e -

5 g

i F

v'

l 1

0.018

/

0.016 -

END OF CYCLE 3 /

7 (EQ.Xe) ESTIMATED /

2 /

= /

E /

y 0.014 -

/

o /.

E /

m / END OF EQUILlBRIUM CYCLE (EQ. Xe) d j 0.012 --

s A

E 2 0.010 -

E 2

0.008 -

l 1 I I I I 1500 1400 1300 1200 1100 1000 900 1

1

AVERAGE GAS OUTLET TEMPERATURE (DF) 1 Fig. 5-5. Maximum allowable rod pair worth versus core average gas outlet temperatura l 5-26 f

1

9

- 4 1

m

- S d

, 0 n T I 0 a N 0 3 5 EI

' C I

F F

3 1

e E

T-N E

I O

C L

X m

C A u I

F F

E

- M R

i b

r E

O H i C T l L O i _

E U

F ' SI

)

h u

q e

- K

(

i t .

E 0 R w I 0 0

U e -

2 T A r R o

/

E P

c _

M l _

E T

a i .

E t _

R i _

O C

n E

G i

e p A R

h t _

E .

V n

/ A i ,

t n

e i

/ I 0

0 0

1 f i

c -

/ f e

o

/ c

/ e .

r .

u

/ t a

/ r e

/ p m

e T

/ _

/ .

6 5

- - O .

g 6 8 0 2 i 0 2

- 4 - - 1 1

- F _

2fEEQ :54chna Y0:

0 1

i i FUEL COEFFICIENT _

4 i

r ISOTHERMAL COEFFICIENT g

.a' o

=

H

$4 e

~

2 8

u E

, yi g -6 -

M &

2 i E F

I 4 _

l l

-10 2000 3000 IM AVERf 3E CORE TEMPERATURE (K)

Fig. 5-7. Temperature coefficient at end of equilibrium cycle

l o

l l

l l

l 0.10 0.09 - BOC, INITIAL CORE f t

(FSAR TRANSIENT ANALYSIS)/

0.08 -

, 7

/

0.07 -

/

? /

[ 0.06 -

/ BOC, CYCLE 3 (GAUGE)

S /

$ 0 .05 -

/

0.04 -

/

/ / ,- pf k / / ,/ / '

• 0.0s -

/ / / ,

0.02 -

/ EOC, CYCLE 3 (GAUGE)

EOC, EQUILIBRIUM COR

/ (FSAR TRANSIENT ANALYSIS) 0.01 -

/

I I I I I 200 600 1000 1400 1800 2200 AVERAGE CORE TEMPERATURE (0F) l Fig. 5-8. Temperature defect versus average core temperature 5-29 l

l - ..

6. THERMAL-HYDRAULIC DESIGN As noted in Section 4.3, the thermal design of the Segment 8 fuel ele-ments is essentially the same as that of the initial core. No changes have been made in f wl element geometry. The power distributions expected during the third cycle are within the envelopes defined by the Technical Specifi-cations (LCO 4.1.3), as discussed in Section 5.4 and shown in Figs. 5-3 and 5-4. Hence, except for the opening of cross-flow gaps, as discussed and accounted for in Section 3.6.2.2 of the FSAR, core coolant flow character-istics are also unaffected. Accordingly, the thermal-hydraulic design limits of the Segment 8 fuel elements are not changed from those of the initial core or the equilibrium core.

I l

6-1

~

1 l

7. SAFETY ANALYSIS 7.1. INTRODUCTION AND

SUMMARY

In this section, events and accidents previously analyzed in Chapter XIV of the FSV FSAR are reviewed to determine if the substitution of petroleum pitch binder for coal tar pitch binder, the use of graphite shim particles in the fuel rods, and the use of a thick buffer fuel element in the FSV core could alter the likelihood or consequences of postulated accidents. The purpose of such a review is to assure that the worst case conditions previously defined for accident analysis, and found to be acceptable during the FSAR review, are not exceeded as a consequence of introducing Segment 8 and these changes into the core, and that no new safety issues are presented.

Chapter XIV of the FSAR has been examined to identify those events which might be affected by the changes cited above. The results of this examination are given in Table 7-1. Those events which required a more detailed examination are:

1. Rod withdrawal accidents (RWAs).

2. Incidents involving fuel elements.

3. Loss of normal shutdown cooling (limiting case: cooldown on one firevater driven circulator).

l 4. Moisture inleakage.

5. Permanent loss of forced circulation (DBA #1).

6. Rapid depressurization/ blowdown (DRA #2).

As indicated in Table 7-1, RWAs are discussed in Section 5.5 of this document, in which it is concluded that the neutronic consequences of a RWA will be unchanged _and the RWA consequences are no more severe than those of the postulated RWA described in the FSAR.

7-1 1

I

In the sections which follow, it is demonstrated for the remaining events (items 2 through 6) that existing FSAR results of accident analyses conservatively bound any perturbations resulting from the introduction of petroleum pitch binder and shim particles to the fuel rods and of a thick buffer element.

7.2. INCIDENTS Three incidents are cited in Section 14.3.1 of the FSAR that could be affected by the use of the thick buffer element. In all three cases the change to thick buffer will improve or leave unchanged safety margins that are already satisfactory. For ' Column Deflection and Alignment," FSAR Section 3.3.1.2, the thick buffer elements (Section 4.2) will undergo less shrinkage differential than assumed in the FSAR and hence less deflection.

For " Fuel Element Malfunction," FSAR Section 3.4.2.1.1, the thick buffer blocks will experience lower stress buildup than the thin buffer blocks.

Under " Misplaced Fuel Element," FSAR Section 3.5.4.5, the uniformly loaded thick buffer elements cannot be loaded in a wrong orientation, an incident that was possible with the asymmetrically loaded thin buffer elements.

7.3. LOSS OF NORMAL SHUTDOWN COOLING This accident is defined in Section 14.4 of the FSAR as the unavailability of the normal number of helium circulators, the loss of normal driving power for the helium circulators, or the unavailability of the economizer-evaporator-superheater sections of one or both steam generators. Loss of the reheater sections of one or both steam generators is the same as loss of normal helium circulator driving power, since the circulators are normally driven with reheat steam. In the FSAR, several cases of cooldown modes are analyzed for this accident. The most severe case (Section 14.4.2.1, Case B2) was determined to be cooling with one circulator driven by the fire water syseen. As shown in Section 5, the introduction of petroleum based pitch binder and graphite shim particles to the fuel rods and the replacement of " thin" buffer elements with thick 7-2

buffer elements introduces no adverse changes in core neutronics values used in FSAR analysts of this accident. Specifically, no adverse impact on power peaking factors is created (Section 5.3), there is no adverse impact on power distribution (Section 5.5), core shutdown margins are not adversely affected (Section 5.6), and kinetic parameters are not adversely affected (Section 5.7). In addition, there are no adverse effects on core thermal analyses (Section 4.3).

Therefore, as there are no changes required for either the neutronic or thermal conditions employed in the FSAR analysis of this accident, the FSAR conclusions regarding this accident are not invalidated by the substitution of petroleum pitch binder for coal tar pitch binder, the addition of graphite shin particles, and use of thick buffer elements.

7.4. MOISTURE INLEAKAGE The analysis in Section 14.5.2 of the FSAR considers inleakage to the primary coolant system from an economizer-evaporator-superheater subheader I or tube or from the helium circulator bearing water supply. Of the moisture l

ingress cases treated in the FSAR, Case 5, a steam generator subheader i .

l rupture compounded by concurrent failure of the moisture monitor system and

! dumping of the wrong (nonleaking) steam loop, has the greatest potential for graphite oxidation and fuel hydrolysis in the shortest time following the accident. As shown in the FSAR, Cases 5 and 6 actually show comparable results. To evaluate the potential effect of the substitution of petroleum pitch binder, graphite shia particles, and buffer elements on the analysis of Case 5, the following phenomena were investigated:

l j

1. Steam graphite reaction.

l 2. Hydrolysis of failed fuel.

3. Fission product release from oxidized graphite.

l 1

7-3 l

l

• rwe -

w ~ q p ym r--v- - m e - ~ - -m -

7.4.1. Steam-Graphite Reaction j The core power distribution (Section 5) and thermal analysis (Maction 4.3) are not adversely affected by the change in binder, shim particles or buffer element loading, and therefore, the related initial conditions of the accident are unchanged or less severe than postulated in the FSAR.

Barium and strontium are known to catalyze the steam-graphite reaction.

As set forth in Section 4.4, the retention of metallic fission products in the fuel particles is expected to increase due to lower fuel temperature, which results from improved thermal conductivity at the fuel rod. In addition, the sorption of metallic fission products into the rod matrix is improved with the petroleum pitch binder over that with a coal tar pitch binder. These factors will tend to decrease the inventory of catalyst in the graphite of the fuel element and thus decrease the overall amount of the steam graphite reaction, since most graphite reaction occurs around the coolant channels. The rate of production of CO and H2 during moisture ingress events will therefore be reduced as the catalytic effect on the graphite is reduced. The peak pressure in the primary coolant system therefore is expected not to exceed present values cited in the FSAR.

7.4.2. Hydrolysis of Failed Fuel Hydrolysis of failed fuel can result in the release of a fraction of the noble gas fission product inventory from the failed fuel to the primary coolant system. However, since the primary coolant system remains essen-tially leak tight after a steam leak accident, this release represents no hazard to the general public.

The hydrolysis reaction with feel particles occurs only when steam diffuses through the graphite blo.k and reacts with fuel kernels whose coat-ings have failed prior to the moisture ingress event. The rate of hydrol-ysis and associated release of noble gas fission products is dependent upon local fuel temperatures and the steam concentration. .

7-4

The local fuel temperature is expected to oecrease (see Section 7.4.1, above). Concentrations of barium and strontium in the graphite will there-fore also decrease. Lower barium and strontium concentrations in the graph-ite, by reducing the graphite reactivity with steam, will allow higher steam concentrations in the fuel early in the accident whc: temperatures are high.

At longer times, when the core cools after trip, the steam concentration will be unaffected. Hydrolysis will continue after steaa-graphite reactions stop, and the large majority of noble gas release will occur during this long-tern hydrolysis. Since the eventual number of fissile particles hydro-lyzed will be little affected by early higher steam concentrations, the overall fission gas release due to hydrolysis is expected to be unaffected.

7.4.3. Fission Product Release from Oxidized Graphite The amount of activity released to the primary coolant system from oxidized graphite is proportional to the amount of graphite reacted and the concentration of fission products within the graphite. Both of these values will be reduced with the fuel rods containing petroleum pitch binder and graphite shias and with thick buffer elements. The reduction in graphite reacted is discussed in Section 7.4.1. A reduction of fission product con-tent in the graphite is expected owing to lower fuel particle temperatures and increased metallic fission product sorption on the fuel rod matrix (Section 4.4). Hence, the fission product release from the graphite as a result of moisture ingress will not exceed the FSAR value, in that less l graphite will be reacted and it will contain a lower concentration of fission products.

7.5. PERMANENT LOSS OF FORCED COOLING (DESIGN BASIS ACCIDEffr NO. 1)

\

This hypothetical event assumes permanent loss of forced circulation of primary coolant helium.

7-5

The FSAR considers consequences of this event in four categories:

1. Thermal results wherein metallic components might fail.

2. Structural results which might affect the core, reflector, barrel, or core support.

3. Nuclear consequences which affect shutdown capability.

4. Fission product release and offsite doses.

The substitution of the fuel rod matrix and shim particles and thick buffer elements will not alter thermal effects significantly, as the core afterheat relationship is not affected because fission product inventory and distribution are essentially unchanged. Therefore, the conclusions in the FSAR will remain unchanged. 9 The structural results mainly concern thermal stresses which might affect the structural integrity of the core support and lateral restraint.

Such structural effects, external to the core, are not icfluenced by the nature of the fuel rods or by the buffer fuel loading.

The nuclear consequences concern reactivity effects whereby the potential reduction of shutdown margin in the overheated core could result from compaction and melting of control rods et spatial redistribution of control poison, fission product poisons, or uranium and thorium. The FSAR shows that these effects do not result in an increase in reactivity at any time during the accidant. The petroleum pitch binder matrix and the thick buffer do not affect control rod melting and compaction or the distribution of control rod poisons. Core neutronics are also not affected (Section 5).

Thus, no change is expected in the prior FSAR analysis of fuel or fission product redistribution during core heatup, and shutdown capability will not be altered.

7-6

, o.

The offsite doses for DBA-1 reported in Table 14.10-1 of the FSAR are small fractions of the 10CFR100 guidelines even with the conservative assumption of TID-14844 release fractions. These offsite doses will not be adversely affected by the substitution of the rod matrix or the thick buf-fer. The new rod matrix and the thick buffer are expected to give fission product releases during core heatup similar to or lower than those predicted for the FSAR elements using the old reference fuel rod constituents.

7.6. RAPID DEPRESSURIZATION/ BLOWDOWN (DESIGN BASIS ACCIDENT No. 2)

The FSAR considers three classes of events for loss of primary coolant:

Primary Coolant Leakage (Section 14.7), Maximum Credible Accident (Section 14.8), and the hypothetical Design Basis Accident No. 2, " Rapid Depressurization/ Blowdown" (Section 14.11). Since the third class envelops i any potential consequences of the other two, it alone was considered for possible impact by the fuel rod and buffer changes.

The aspects of a Design Basis Accident No. 2 (DBA-2) which are

, discussed in Section 14.11 of the FSAR are as follows:

l

1. Integrity of reactor.

2. Continuation of adequate primary coolant circulation.

l 3. Ingress of air.

4. Effects of operator actions.

l S. Vertical thrust on the PCRV.

6. Effect on building pressure.

7. Radiological consequences.

Item 1 concerns the pressure differentials that could develop across ,

reactor components during the depressurisation. Item 2 concerns the ability of the helium circulators to continue operation and provide adequate core cooling in the event of a PCRV penetration failure. Item 3 concerns the possible pathways for air ingress into the PCRV following a DBA-2. These l items are obviously not affected by the presence of petroleum pitch binder 7-7 l . - - . . _ - _ _ _ -_

.o ,

and graphite shim particles in the fuel rods and of thick buffer elements.

In addition, as shown in Sections 5 and 4.3, the use of petroleum pitch binder and graphite shim particles in the fuel rods and of thick buffer elements does not introduce any neutronic or thermal changes.

Regarding item 4, the FSAR describes how automatic reactor scram and alarm would alert the operator to the fact that a leak has developed in the PCRV and to be ready to perform necessary monitoring and system adjustments.

The operator's responsibilities are neither increased nor diminished by the presence of petroleum pitch binder and graphite shin particles in the fuel rods and thick buffer elements in the core. Items 5 and 6 concern mechan-ical processes external to the PCRV and core. Therefore, only item 7 needs to be examined in further detail.

As discussed in the FSAR, the radiological release from a postulated DBA-2 consists of essentially all the activity which exists in the circu-lating primary coolant prior to the depressurisation event plus a small fraction of the activity plated out on the surfaces of the primary circuit.

As discussed in Section 4.4, use of petroleum pitch binder and graphite shim particles in the fuel rods and of thick buffer elements will not result in any increase in the circulating and plateout activity. Bence, the radiolog-ical consequences of the accident would be essentially unchanged or reduced from those reported in the FSAR.

7.7. CONCLUSION A review of Chapter XIV of the FSAR identified six postulated accident conditions which required more detailed examination for potential impact from the substitution of petroleum pitch binder and graphite shim particles in the fuel rods and thick b~ffer elements in the core. These accident con-ditions were discussed in this section and no requirement for additional analysis has been identified. It is concluded that the worst case condi- .

tions previously defined for accident analyses and found to be acceptable during FSAR review are not exceeded as a result of introducing petroleum .

7-8

pitch binder and graphite shin particles in the fuel rods and thick buffer elements in the core. In addition, since the change in burnable poison configuration, as discussed in Sections 5.2 and 5.4, does not result in power peaking above previously assumed values, the new burnable poison raises no new safety issues. Accordingly, these design changes present no unreviewed safety questions as defined in 10CFR50.59.

7-9

TABLE 7-1 POTENTIAL EFFECTS ON FSV ACCIDENT ANALYSIS OF PETROLEUM-DERIVED FITCH BINDER AND CRAPHITE SHIM PARTICLES IN FUEL RODS AND THE USE OF WOLE BthCK BUFFER ELEMENTS Petroleue Graphite Thick FSAR Chapter XIV Event Fitch Binder Shie Particles Buffer 14.1 Environmental Disturbances Earthquake None - any reactivity effects would be bounded by rod withdrawal events (FSAR)

Wind Effects l Flood j Fire The core is not affected by these events Landslides

{

Snow and Ice 14.2 mesetivity Accidents and Transient y Response

$ Summary of Reactivity Sources Excesolve removal of control poison Loss of fission product poisons Rearrangement of core components Reactivity insertions in these evente are less than rod withdravel events (FSAR)

Introduction of steme into the core Sudden decrease in reactor temperature Rod withdrawal accidents Evaluation required, see Section 5.5 14.3 Incidente Incidents Involving the Reactor Core Coluen deflection and misalignment No change from Section 3.3.1.2 of FSAR Discussed, se:u Sections 7.2 and 4.2 Fuel element entfunctione No change from Section 3.4.2.1.2 of FSAR Discussed, see Sections 7.2 and 4.2 Nisplaced fuel element No change from Section 3.5.4.5 of FSAR Discussed, see Section 7.2 O

TABLE 7-1 (Continued)

Petroleue Crephite Thick FSAR Chapter XIV Event Fitch Binder Shie Particles Buffer Blocking of coolant channel No change from Section 3.6.5.2 of TSAA Control rod malfunctions No change froe Section 3.8 of FSAR Orifice selfunctions No chacge from Section 3.6.5.1 of FSAR Core support floor loss of cooling No chsage from Section 3.3.2.2 of FSAR Incidente lavolvint the primary coolant None system Incidente favolving the control and instrumentation system Incidente involving the PCRT y facidente involving the secondary y coolant and power convereine system i

Incidente involving the electrical system Malfunctions of the helium purification systee Melfunctions of the helium storese eyetee Malfunctions of the nitrogen system 14.4 lose of Normel Shutdown Cooling Evaluation required, see Section 7.3 14.5 Secondary coolant System Leekage Steae leaks outside the primary coolant None systee Leaks inside the primary coolant.eystee Evaluation required, see Section 7.4 (moisture inleakage)

i TABl.E 7-1 (Continued)

Petroleum Craphite Thick FSAR Chapter XIV Event Fitch Binder Shia Particles Buffer 14.6 Auxiliary Systen 1.eakage Failures involving the helium purification system 14es of both purification trains 1 Failure of regeneration line Fossible effect would be bounded by Design Basis Accident No. 2. FSAR w/ simultaneous valve failure and Section 14.11 operational error s Accidents involving the gas weste No change froe Section 14.6.2 of FSAR systes y Fuel handling and storage accidents Fuel headling accidents No change f rom Section 14.6.3 of FSAR Fuel storage accidents i 14.7 Priesty Coolant Leakage Fossible effects would be bounded by Design Basie Accident No. 2. FSAR

• • • I**

• 14.8 Henteum Credible Accident 14.9 Nazimus Eypothetical Accident Some se FSAR. Section 14.11 14.10 Destga Basis Accident No. 1 Evaluation required; see Section 7.5 2

"Fernanent Loss of Forced Circulation (IDFC)"

14.11 Desiga Basis Accident No. 2

• Rapid Evaluation required; see Section 7.6 Depressurisation/ Blowdown" o

.i i

. o.

1

8. PROPOSED MODIFICATIONS TO TECIDlICAL SPECIFICATIONS I

No changes to the plant Technical Specifications are necessitated by the insertion of Segment 8 into the reactor core.

i l

.l 1

?

i j

1 i

i l

l l

1 I

i 8-1 t

~

9. STARTUP TESTS Following refueling, a stepwise approach to full power will be performed. During these power steps, the following tests will be performed:

1. Differential control rod calibration measurements. These tests are required by Technical Specification SR 5.1.5. The method used for these measurements is the same as that in SUT B-9.

2. Temperature coefficient (temperature reactivity defect) measure-ments. This test is required by Technical Specification SR 5.1.3.

The method used for these measurements is the same as that in SUT B-8.

3. Reactivity status surveillance check. This test is performed at each startup and once per week as required by Technical Specification SR 5.1.4.

9-1

(

10. REFERENCES

1. Scott, C. B., and D. P. Harmon, "Postirradiation Examination of Capsules HRB-4, HRB-5, and HRB-6," ERDA Report GA-A13267, General Atomic Company, November 28, 1975.

2. Johnson, W. R., et al., "Postirradiation Examination of Capsules HT-24 and HT-25," ERDA Report GA-A13486, General Atomic Company, September 15, 1976.

3. Young, C. A., and C. B. Scott, "Postirradiation Examination of Capsule P13Q," ERDA Report GA-A14174, General Atomic Company, September 1977.

4. Scott, C. B. , D. P. Harmon, and J. F. Holzgraf, " PIE Examination of Capsules P13R and P13S," ERDA Report GA-A13827, General Atomic Company, October 8, 1976.

5. "HTGR Fuels and Core Development Program, Quarterly Progress Report for the Period Ending February 28, 1977," ERDA Report GA-A14298, General Atomic Company, March 1977.

6. Tiegs, T. N. , et al. , "0RR Irradiation Experiment OF-1: Accelerated Testing of HTGR Fuel," Oak Ridge National Laboratory Report ORNL-5234, August 1977.

7. Ballagny, A., et al., "SSL-1 Spitfire Loop Experiment Postirradiation Examination and Analysis (Doesier No. 6)," General Atomic Report GA-A14628, July 1378.

8. Blanchard, R., and M. L. Pointud, " Rapport de Synthese des Irradiations GF.1, GF.2, GF.3," DMG DR 34/78, Commissariat a L'Energie Atomique, November 16, 1978.

9. Johnson, W. R., " Thermal Conductivity of Large HTGR Fuel Rods," USAEC Report GA-A12910 (GA-LTR-9), General Atomic Company, March 15, 1974.

10. "HTGR Base Program Quarterly Progress Report for the Period Ending February 28, 1974," USAEC Report GA-A12916, General Atomic Company, March 29, 1974.

11. "HTGR Fuels and Core Development Program, Quarterly Progress Report for the Period Ending May 31, 1975," ERDA Report GA-A13444, General Atomic Company, June 30, 1975.

10-1