ML19296A121

Amend 3 to Fuel Test Element SAR Which Covers Effects of Extended Cycle 2 Core Operation on Fuel Test Elements

ML19296A121
Person / Time
Site:	Fort Saint Vrain
Issue date:	01/26/1979
From:	PUBLIC SERVICE CO. OF COLORADO
To:
Shared Package
ML19296A120	List: ML19296A119 ML19296A121
References
NUDOCS 7902010160
Download: ML19296A121 (62)

Text

GLP-5494 Amendment 1 CONTENTS ABSTRACT . . . . . .. . . . . . . . . . . . . . . . . . . . . . . . v GLOSSARY . . . . .. . . . . . . . . . . . . . . . . . . . . . . . . vii

1. INTRODUCTION . . . . . . . . . . . . . . . . . . . . . . . . . . 1-1

2.

SUMMARY

. . .. . . . . . . . . . . . . . . . . . . . . . . . . 2-1

3. TEST OBJECTIVES AND DESIGN CRITERIA . . . . . . . . . . . . . . 3-1

4. TECHNICAL DESCRIPTION OF TEST FUEL . . . . . . . . . . . . . . . 4-1 4.1. Fuel Element Assembly and Location in the Core . . . . . 4-1 4.2. Fuel Particles . . . . . . . . . . . . . . . . . . . . . 4-3 4.2.1. Driver Fuel . . . . . . . . . . . . . . . . . . 4-3 4.2.2. Test Array Fuel . . . . . . . . . . . . . . . . 4-3 4.3. Fuel Rods . . . . . . . . . . . . . . . . . . . . . . . . 4-4 4.4. Fuel Element Graphite . . . . . . . . . . . . . . . . . . 4-5 4.5. Cure-In-Place Process . . . . . . . . . . . . . . . . . . 4-6 4.6. Fluence, Burnup, and Temperature Monitors . . . . . . . . 4-7

5. PERFORMANCE ANALYSIS - NORMAL OPERATION . . . . . . . . . . . . 5-1 5.1. Nuclear Analysis . . . . . . . . . . . . . . . . . . . . 5-1 5.1.1. Fuel Loadings and Burnable Poisons . . . . . . . 5-1 5.1.2. Power Perturbati: as . . . . . . . . . . . . . . 5-2 5.1.3. Fluence Perturbations . . . . . . . . . . . . . 5-3 5.1.4. Control Rod Worths and Reactivity Effects . . . 5-3 5.1.5. Fuel Handling . . . . . . . . . . . . . . . . . 5-3 5.2. Thermal Annlysis . . . . . . . . . . . . . . . . . . . . 5-4 5.2.1. Analysis Procedure . . . . . . . . . . . . . . . 5-4 j 5.2.2. Analysis Results . . . . . . . . . . . . . . . . 5-6 5.3. Fission Product Release Analysis . . . . . . . . . . . . 5-8 5.3.1. Gaseous Fission Product Release . . . . . . . . 5-8 5.3.2. Metallic Fission Product Release . . . . . . . . 5-10 5.3.3. Conclusions . . . . . . . . . . . . . . . . . . 5-12 5.4. Graphite Structural Analysis . . . . . . . . . . . . . . 5-13 5.5. Graphite Dimensional Change . . . . . . . . . . . . . . . 5-16 i

7 9 0 2 010 //ad '

CLP-5494 Amendment 3

6. SAFETY ANALYSIS . . . . . . . . . . . . . . . . . . . . . . . . 6-1 6.1. Introduction and Summary . . . . . . . . . . . . . . . . 6-1 6.2. Reactivity Events . . . . . . . . . . . . . . . . . . . 6-2 6.2.1. Rod Withdrawal Accidents . . . . . . . . . . . 6-2 6.2.2. Loss of Normal Shutdown Cooling . . . . . . . . 6-4 6.2.3. Moisture Ingress . . . . . . . . . . . . . . . 6-5 6.2.4. Design Basis Accident No. 1, " Permanent Loss of Forced Cooling (LOFC)" . . . . . . . . . . . . 6-9 6.2.5. Primary Coolant System Ruptures: Design Basis accident No. 2, " Rapid Depressurization/

Blowdown" . . . . . . . . . . . . . . . . . . . 6-11 6.3. Fuel Handling Accidents . . . . . . . . . . . . . . . . 6-13 6.3.1. Loading Errors . . . . . . . . . . . . . . . . 6-14

7. RESEARCH AND DEVELOPMENT . . . . . . . . . . . . . . . . . . . 7-1 7.1. Fuel Test Element Experience . . . . . . . . . . . . . . 7-1 7.1.1. Peach Bottom FTEs . . . . . . . . . . . . . . . 7-1

!.2. Graphite Development Program . . . . . . . . . . . . . . 7-2 7.3. Fuel Development . . . . . . . . . . . . . . . . . . . . 7-2 7.3.1. Introduction . . . . . . . . . . . . . . . . . 7-2 7.3.2. Fuel Particles . . . . . . . . . . . . . . . . 7-3 7.3.3. Fuel Rods . . . . . . . . . . . . . . . . . . . 7-7

8. REFERENCES . . . . . . . . . . . . . . . . . . . . . . . . . . 8-1 APPENDIX A: PROPOSED POST-IRRADIATION EXAMINATION (PIE) . . . . . A-1 2 IABLES 2-1. Description of Fort St. Vrain fuel test elements FTE-1 through FTE-8 . . . . . . . . . . . . . . . . . . . . . . . 2-4 2-2. As-built quantity of different fuel types in FSV FTE-1 through FTE-6 . . . . . . . . . . . . . . . . . . . . . . . 2-5 l3 2-3. Summary of high-temperature gas-cooled reactors utilizing coated-particle fuel . . . . . . . . . . . . . . . . . . . . 2-6 2-4. Data on coated particle fuels for high-temperature reactors. 2-7 4-1. Comparison of reference FSV, proposed FSV, and large HTGR test fuels . . . . . . . . . . . . . . . . . . . . . . . . 4-9 4-2. Test element as-built heavy metal loadings . . . . . . . . . 4-10 3 4-3. Test matrix for the six fuel hole test arrays in FTE-2,

-4, and -6 . . . . . . . . . . . . . . . . . . . . . . . . . 4-11 4-4. Fertile fuel particle characteristics comparison . . . . . . 4-12 11

GLP-5494 Amendment 3 ABSTRACT This report describes the eight fuel test elements to be loaded with Se gmen t 7 (first reload) of the Fort St. Vrain nuclear reactor. It also presents the results of the analysis of the effects of the test elenents on plant normal operation and plant safety. These analysee were done using as-built test element fuel loadings. The analyses also show the effects of 3 operation of Cycles 1 and 2, each up to 200 EFPD.

The analysis confirms that the test elements have a very small effect on the operation of the core under all conditions. The eight test elements rep resent a very small percentage of the core (0.54%). Only six test elements contain new fuel types, which represent only 0.40% of the total fuel.

The test elements will be manufactured f rom near-isotropic H-451 graphite in place of the needle-coke H-327 graphite used in the reference elements. This results in structurally stronger fuel elements with heat transfer and dimensional change characteristics superior to those of the re ference elements.

The fuel to be used represents an evolutionary improvement over the initial core fuel. Re fe rence fresh and recycle fuel for large commercial HTGRs as well as improved fuel contemplated for use in future Fort St.

Vrain reload segments is included in the test elements. Fuel improvements in this test series include Th0 f rtil ke rnels , UC nd weak acid resin 2 2 (WAR) fissile kernels, and the cure-in-place fuel rod process. A very small number of fertile particles with BISO coatings are included as a test of all aspects of fuel for the large HTGR.

In addition to * .e above-mentioned fuel types , a small number of fuel 1

rods containing medium-enriched uranium (MEU) fuel will be included. This v

GLP-5494 Amendment 1 fuel will be 19.5% enriched as compared with 93% enrichment of the standard O

high-enriched uranium (HEU) fuel and represents a proliferation-resistant 1 fuel cycle.

Extensive in-pile and out-of-pile test programs have demonstrated the good performance of the different HEU fuel types. Low-enriched uranium fuel has been tested extensively in Europe, and a parallel test program in 1

accelerated capsules will provide demonstration of satisfactory performance to full exposure before the test elements have received 1 year of exposure.

The test elements will further increase the confidence in HTGR fuel per-formance under realistic power reactor operating conditions.

O O

vi

GLP-5494 Amendment 3 GLOSSARY AVS agglomeration voie seche (dry route agglomeration process)

BISO two-coating fuel particle BOC beginning of cycle CEA Commissariat a l'Energie Atomique CIB cured-in-bed (green fuel rods cured in an alumina bed)

CIP cured-in-place (green fuel rods cured in the fuel element)

CIT cured-in-tube (then pushed out to form loose rods)

DNAA delayed neutron activation analysis DOE Department of Energy 3

EFPD ef fective full-power days

• EOC end of cycle EPRI Electric Power Research Ins titute FEVER a multigroup one-dimensional depletion program for analyzing a f uel region and subregion in the axial direction FIMA fissions per initial heavy metal atom FSAR final safety analysis report FSV Fort St. Vrain FTE fuel test element GA General Atomic Company CAUGE a two-dimensional few-group diffusion-depletion program for analyzing a fuel region layer HEU high-enriched uranium (93% U-235 enrichment) lI HTCR high-temperature gas-cooled reactor KFA/HOBEG Kernforschungsanlage JUlich GMBH/Hoch Tempetatur Brennelement GMBH, Hanau MEU medium-enriched uranium (19.5% U-235 enrichment) l1 LOFC loss of forced cooling LWR light water reactor MLC mid-length-center ORNL Oak Ridge National Laboratory PSC Public Service Company of Colorado vii

GLP-5494 Amendment 3 RTE recycle test elemnt O

RWA rod withdrawal accident SG sol-gel process TRICA acronym: training, research, and i_sotope production 3 reactor manufactured by General Atomic TRISO four-coating fuel particle UC 0 WAR fissile particle (see Table 2-1) xy UKAEA United Kingdom Atomic Energy Authority WAR weak acid resin (a bead-loading process)

Driver fuel UC TRISO and Th0 TRISO 3 2 O

O vtLL

GLP-3494 Amendment 3

1. INTRODUCTION This document presents the results of safety and performance analyses that have been carried out to support the planned insertion of eight fuel test elements (FTEs) of improved fuel into the Fort S t. Vrain (FSV) reactor during its first refueling. The report describes the proposed FIEs, their operational behavior, their ef fect on postulated accidents described in the FSAR, and irradia'cion test results on the fuel element materials. The evaluations presented are intended to form the technical basis for Nuclear Regulatory Commission approval for loading the FIEs into the FSV reactor.

The test elements will be fabricated and assembled by General Atomic Company (GA) under specifications and quality inspection requirements comparable to those used for manufacture of the initial core and reload segment fuel. All of the high-enrichment and fertile fuel materials have undergone full irradiation exposure in test capsules under DOE, ORNL, GA, l3 and foreign fuel qualification programs. The medium-enrichment fuel will be included in accelerated capsules which will provide demonstration of I satisfactory performance to full exposure before the fuel has received one year of exposure in FSV.

Inclusion of the test elements in an early reload segment is highly desirable to:

1. Obtain early irradiation experience on potential future reload materials under actual commercial reactor operating conditions, which cover a range of neutron fluence, burnup, and temperature.

2. Perform integral system demonstrat!on of the acceptability of fuel materials and processes planned for use in the large commercial HTGR and future FSV reload segments. l1 1-1

CLP-5494 Amendment 3 O

This report is divided into four major sections: (1) a technical description of the fuel elements and fuel, (2) performance analysis under normal steady-state conditions, (3) safety analysis of the effect of the test elements on accidents considered in the FSAR, and (4) status of research and development programs and experience on the proposed test ele-ment materials. The analyses preser.ted in this report were done using as-built FrE fuel loadings. The analyses also show the effects of 3 operation of Cycles 1 and 2, each up to 200 EFPD.

For the purposes of this document, "the reference core" is defined as the standard design FSV core operating with Cycle 2 loading distribution. l3 O

O 1-2

GLP-5494 Amendment 3

2. SUE!ARY The proposed FSV test elements, PIE-1 through FIE-8, are described in Table 2-1.* These elements are all standard fuel blocks without control rod channels. They are designed to operate within the limits of peak fuel tem-perature, neutron fluence, and burnup specified for the initial core and reload fuel elements. Instrumentation is included in the test elements to measure each of these parameters. Locations of the elements within the core have been selected to yield test results over a range of exposure con-ditions. One test element will be located in each of Segments 2 through 6, and three will be placed in Segment 7. The basic fuel materials (graphite, highly enriched uranium, and thorium) and the fuel block reactivity worth of the test elements are unchanged from the initial core fuel elements which are replaced. Thus, the neutronic characteristics of the core are essentially identical to those presented in the FSAR. Furthermore, in order to introduce an additional safety margin, the fuel loadings for al.

the test elements with a residence time greater than one year and smaller than 5.7 years are less than the loadings for the elements which they l3 replace. The maximum power perturbation to the elements surrounding the test elements is limited to !2% of their original power.

The physical properties of the test element materials are improved over the initial FSV core materials. For instance, the strength and dimen sional stability of H-451 near-isotropic graphite specified for the test elements under HTGR operating conditions are both substantially better thar those of the reference fuel H-327 needle-coke (anisotropic) graphite; and the thermal stability of the UC /Th0 2

and 2

WAR /Th0 2fuel Particle systcas used in some of the test elements is improved, over the range of critical HTGR operating conditions, relative to the reference (Th,U)C /ThC2 particle 2

system. Three of the eight test elements contain small amounts of BISO Tables and figures appear at the end of each section.

2-1

CLP-5494 Amendment 3 coated (buffer + PyC) Th0 2 fertile fuel. Although BIGO coatings are less retentive of the volatile metallic fission products at peak temperatures than TRISO coatings, the Th0 2 kerncis provide very high retention of stron-tium relative to ThC2kernels (Ref. 1, pages 4-1 to 4-8). However, even under very conservative assumptions, the contribution of the fuel test elements to the 30-year design inventory of metallic fission products is very small. Also included in these three test elements are 264 rods con-taining MEU (Th,U)0 2 fissile particles. Again, for the purpose of these test elements, very conservative assumptions are being made in the analysis for metallic fission product release. Tabic 2-2 summarizes the amounts of the different fuel types in test elements FTE-1 through FTE-6. The quantity of BISO fertile fuel introduced into the core by these test elements repre-sents only 0.02% of the total thorium present. Only 0.01% of the total core fissions will occur in the BISO particles over their lifetimes.

As shown in Section 4 of this report, there are 88 MEU fuel rods in each o f FTE-2, -4, and -6. Therefore, approximately 1.4% of the test ele-ment fuel rods contain MEU fuel. The amount of U-238 which the MEU fuel rods introduce is about 0.18 kg, as compared with approximately 45 kg already contained in the core.

Detailed performance analysis of the test elements was conducted to establish the power distribution, temperature history, fission product retention, graphite element stresses, and dimensional stability of each test element for verification of design margins. These results are described in Section 5. The steady-state performance results were incor-porated into the accident analysis to establish that:

1. There is no increase in the probability of occurrence or the consequences of an accident or malfunction previously evaluated in the FSAR.

2. Accidents or malfunctions of a different type caused by the insertion of the test elements other than any previously 2-2

GLP-5494 evaluated in the FSAR have been considered and found to be of negligible probability.

3. The margin of safety as defined by the technical specifications has not been reduced.

Irradiation experier.ce for BISO and TRISO HTGR fuel particles is presented in Appendix A of the FSV FSAR and is summarized in Section 7 of this report. Experience obtained on over 800 Peach Bottom Core 2 fuel and test element assemblies, over 200 BISO coated particle samples in irradi-ation test capsules, and over 100 fuel rods in test bodies gives a high degree of confidence in the performance predictions for these materials under the most severe combination of temperature, neutron fluence, and burnup expected for the FSV reactor. Experience at ORNL and European research and development groups has also been applied to the design of test elements. Tables 2-3 and 2-4 present a summary of operating and proposed HTGR plants which employ HTGR fuel technology.

Based on the above and specific results presented in Sections.5 and 6, it is concluded that the proposed insertion of eight FSV test elements during reload 1 operations will involve no increased hazard to the health and safety of either the plant personnel or the public.

2-3

TABLE 2-1 DESCRIPTION OF FORT ST. VRAIN FUEL TEST ELEMENTS FTE-1 THROUGH FTE-8 FTE-1 FTE-2 FTE-3 FTE-4 FTE-5 FTE-6 FTE-7 FTE-S Segment 2 3 4 5 6 7 7 7 Graphite type H-451 H-451 H-451 H-451 11-451 H-451 H-451 H-451 Fissile fuel type UC2 UC2 UC2 UC2 UC2 UC2 (Th,U)C2 (Th,U)C2 TRISO TRISO TRISO TRISO TRISO TRISO TRISO TRISO plus test plus test plus test fuel (a) fuel (a) fuel (a)

Fertile fuel type Th02 Th02 Th02 Th02 Th02 Th02 ThC2 ThC2 TRISO TRISO TRISO TRISO TRISO TRISO TRISO TRISO plus BISO plus BISO plus BISO Method of fuel rod curing CIP CIP CIP CIP CIP CIP CIB CIB Residence time, years 0.7 1.7 2.7 3.7 4.7 5.7 5.7 5. 7 Peak fluence x 1025 n/m2 0.73 2.13 3.06 4.03 4.45 6.97 6.98 6.76

{ (E > 0.18 MeV)

HEU MEU HEU MEU  !!EU MEU 3 Peak fuel temperature, *C 1208 1030 991 974 1070 1064 1062 1246 1254 1094 1092 Burnup, FIMA(C)

Fissile, % 32.6 47.0 6.4 59.1 66.3 11.5 71.3 74.4 16.1 19.6 19.6 Fertile, % 0.1 0.8 0.8 1.4 2.6 2.7 3.7 6.3 6.1 6.4 6.4

(*) Test fuel includes: (Th,U)C2 TRISO, 88 rods per element CIB UCxy0

• TRISO/Th02 BISO, 350 rods per element CIP HEU 1 UC2 TRISO/Th02 TRISO, 176 rods per element CIP (Th,U)o2 TRISO/Th02 TRISO 88 rods per element CIT} MEU g O

CIP = cure-in-place fuel rod carbonization; CIB = cure in alumina bed - reference FSV process; CIT = cure-in-tube, graphtto crucibles, simulating conditions as experienced in cure-in-place. @

(C FIMA = fissions per initial heavy metal atom.

• x and y represent the mean quantities of carbon and oxygen and do not signify a specific compound. These values will be explicit in the final fuel specification. All kernels of this type are derived from WAR beads.

G 9 9

GLP-5494 itmendment 3 TABLE 2-2 AS-BUILT QUANTITY OF DIFFERENT FUEL TYPES IN FSV FTE-1 THROUGH FTE-6 Uranium Loading Thorium Loading U Total Th-232 Number of Heavy Metal Heavy Metal Fuel Rod Type Fuel Rods (kg) (kg)

(Th,U)Cy TRISO 264 0.057 0.918 UC2 TRISO 17,097 3.323 61.425 Th02 TRISO UCx 0 TRISO 1,050 0.234 3.683 3 Tho,yBISO 1

(Th,U)02 TRISO 264 0.218 0.600 Th02 TRTSO Total 18,675 3.832 66.626 2-5

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10 4-3A

GLP-5494 Amendment 3 O

The experimental fuel in the test arrays has been subject to rigorous quality control inspection to verify coating properties and integrity, low contamination levels, and heavy metal loadings. Test fuel manufactured by 3

sources other than GA has been required to meet specifications equivalent to the GA-produced fuel and has been subjected to quality inspection by GA prior to fuel element assembly.

4.3. FUEL RODS The test element fuel particles and shim material have been bonded 3 together with a matrix consisting of an organic binder and graphite filler which was heat treated to outgas and carbonize the binder to yield a struc- l3 turally sound, bonded fuel rod. Table 4-1 includes a comparison of fuel rod materials for the test elements and reference FSV fuel. The FTE fuel rod is very similar to that used in the FSV initial core with the exception of certain process and product improvements as described below.

O The shim material consists of isotropic graphite granules about the same size as the coated fertile particle. The shim particles permit adjustment of fuel loading within the fixed volume of the fuel rod. The binder consists of a petroleum pitch into which is mixed a graphite flour.

The binder-graphite mix makes up the matrix, which is melted and injected into molds containing the appropriate amount of fissile, fertile, and shim particles.

Bonding the fuel particles in this manner contains the particles in a 3 well-defined free-standing body, and has the ef fect of increasing the ther-mal conductivity of the particle bed and decreasing the thermal gradients in the fuel particles by providing a relatively uniform heat flow path at the particle surface. The matrix formulation has been chosen to provide a rod with adequate strength to maintain integrity during irradiation without detrimental irradiation-induced mechanical interaction between the fuel rod matrix and fuel particles (Ref. 2, page 13).

O 4-4

GLP-5494 Amendment 3 heat treatment of most of the fuel rods in FTE-1 through FTE-6 has been l3 performed in-situ by the cure-in-place (CIP) process developed at GA and described further in Section 4.5. The fuel rods in FTE-7 and FTE-8 and those in selected locations of other test elements (Fig. 4-2) have been 3 cured in alumina beds. The alumina bed curing process is identical tc that used for the FSV initial core. The CIP process has been developed for large HTGR fuel as a process improvement. The product specifications and characteristics of fuel rods produced by the two carbonization techniques are identical. Certain rods are of the CIP type, as shown in Figs. 4-2 3

and 4-3, but have been cured and fired in graphite tubes prior to loading in the FTE in order to allow dimensional measurements af ter firing and before and after irradiation.

The initial fuel element and fuel rod stack height dimensions provide sufficient axial void space to accommodate the relative axial dimensional changes of the graphite block and fuel rod stack expected during operation.

Volatile spacers are used between pairs of fuel rods during the CIP proc-ess. The spacers are vaporized without leaviag a residue during heat treatment. One ef fect of these spacers is to lower the linear heat rate locally over that of loose stacked rods. This is accounted for in the thermal analysis. The thermal analysis also considers the effect of shim particles and BISO fuel particles on the dimensional stability of the fuel rods.

4.4. FUEL ELEMENT GRAPHITE The test element graphite blocks were manufactured by the Great Lakes Carbon Corporation from near-isotropic graphite, grade H-451, development lot 426. Grade H-451 is an extruded near-isotropic petroleum-coke graphite produced in 17-in.-diameter by 34-in.-long logs (Ref. 3, page 8). Use of near-isotropic petroleum-coke filler material produces a stronger, more isotropic graphite than the grade H-327 graphite currently used in the FSV reactor core. The essential advantage of H-451 graphite over H-327 is 4-5

GLP-5494 O

attributable to differences between the properties of the near-isotropic petroleum-coke fil'er used in the manuf acture of H-451 and those of the needle-coke filler used in H-327.

Unitradiated properties of H-451 (lot 426) graphite are given in Table 4-5 and properties for H-327 graphite are given in the FSAR, Table 3.4-1. Typical values are also presented in Table 4-5 for H-451 graphite 25 j properties after irradiation at 900*C to 6 x 10 2 (E > 0.18 MeV)HTGR (Ref. 4, page 78). Grade H-451 is more isotropic than H-327, especially in properties such as strength, elastic modulus, and thermal expansivity.

In addition, the absolute strength of H-451 is higher than that of H-327 in both the axial and radial directions.

The irradiation-induced changes in thermal properties are less pronounced in H-451 than in H-327. The thermal conductivity of irradiated H-451 is higher than that of H-327 (Section 5.4). Ultimate tensile strength and elastic modulus increase with irradiation in both graphites (Ref. 5, Tables 7-2 and 7-3). Irradiata n-induced dimensional changes for lll H-451 are more isotropic and show no net expansion over the temperature and fluence ranges of FSV in comparison with H-327 (see Section 5). As a result, fuel elements made from H-451 will be more dimensionally stable under normal reactor conditions.

Chemical impurities and oxidation reaction rates are approximately the same for H-451 and H-327 (Ref. 6, page 35).

4.5. CURE-IN-PLACE PROCESS The cure-in-place (CIP) process is a method of curing the green fuel rods to outgas and carbonize the binder while the rods are assembled within the fuel element. This is performed in a specially designed furnace which provides good temperature and atmosphere control during heat treating.

O 4-6

GLP-5494 Amendment 3 Axial gaps between fuel rods are controlled by the insertion of plastic spacers which volatilize during heat treatment. These spacers are used primarily to ease the manuf acturing process.

Specified limits on fired fuel rods are essentially the same as for 3 the initial FSV core.

All of the test elements except FTE-7 and FTE-8 have undergone CIP l3 processing of the fuel rods. Fuel rods in these two elements, and selected fuel stacks in the remaining test elements, havebeencuredinpackedaluminal3 beds, which was the reference process used for the FSV core. The cured rods have been loaded into the fuel blocks, without spacers, and sealed with a l3 special end cap at the top of the fuel stack.

4.6. FLUE: ICE, BUR:iUP , A'iD TEMPERATURE !!ONITORS Each of the eight test elements contains monitors for measuring the l3 operating fluence, burnup, and temperature of the experimental fuel and overall fuel element. The monitor design selected includes sic temperature monitors, neutron dosimetry wires, and fuel particles for burnup analysis ,

which are encased in an 11-451 graphite crucible. Figure 4-6 shows a sketch of the monitors. 3 A sufficient number of monitors have been located across the block to allow measurement of in-block temperature and fluence tilts and conditions 3

at each experimental fuel array in FTE-2, -4, and -6. Table 4-6 indicates the number of monitors for each of the test elements. Each test array contains one column with four monitors, one at each end and the other two 3 equally spaced along the fuel stack, plus six monitors located at each face of the element and two near the center. Additional monitors are required for the buffer element in order to measure interface effects between the two fuel loadings used (Fig. 4-3).

4-7

GLP-5494 Amendment 3 O

Because of the high temperatures required during curing, the fluence, burnup, and temperature monitors have been assembled after the curing process in test columns along with rods fired outside the FTE. These rods have been cured in place in graphite crucibles, then separated and measured for dimen- 3 sions after firing. Fluence, burnup, and temperature monitors have also been placed at the top of selected fuel columns that have been cured in place.

For FTE-1 through FTE-6, several fuel rods were submitted to delayed neutron activation analysis (DNAA) for as-built uranium content measurements.

These rods have been placed into the test elements at corner locations and in the vicinity of the fluence and burnup monitors. These rods can be used 3

far gamma spectrometric burnup measurements after irradiation. Some of these rods have also been subjected to fission gas release measurements in the TRIGA reactor facility at GA for fuel performance measurements prior to i r rad ia t ion.

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CLP-5494 Amendment 1 TABLE 4-1 COMPARISON OF RFFERENCE FSV, PROPOSED FSV, AND LARGE HTGE TEST FUELS FSV Fue!

Large HTGR Feature Reference Te e. t Element (d) Commercial Fuel MEU Fuel (a)

Fissile particles Kernel type (Th,U)C, UC 3 , UC xy0 (b) UC,, UCxv0 (b) (Th,U)0

' ' 7 (WAR) ~ (WAR)

Coating type TRISO TRISO TRISO TRISO Kernel diameter, um 140 and 225 l195, 305, and 360 195, 305, and 360 350 Total coating 160 and 170 220, 170, and 175 220, 170, and 175 210 thickness, um Fertile part icles i

Fernel type ThC Th0 Th0 3 I Tho.,

2 s f a Coating type TRISO BIl0 lTRISO and BISO lTRISO Kernel diameter, Lm 355 and 455 l450 and 500 j450 l500 Tota l ( oat ing ,145 and 155 175 l175 thickness, Lm l175 l  !

I  !

Fue1 rod '

Diameter, in. l 0. ',91 l0.4895 and 0.491 f0.619 !O.4895 I i i .

I.en g t h , in.

l1.94 1.94 i2.476 l1.94 liin de r l Coa l tar pitch Petroleum pitch I Petroleun pitch i l

l Petroleum pitch 1 Filler Natural flake ! Synthetic flake l Synthetic flake jSynthetic flake l r,raph it e igraphite  ! graphite jgraphite Shim particle f Wne Near-isotropic INear-isotropic lNear-isotropic

! rgraphite igraphite i graphit e

[ l Heat treatment jrure in Al,0 - 3 l Cure in fuel (Cure in fuel lCureingraphite l powder l block and alter- block jtube i lnativt l

l i Craphite fuel block j i l l t

Graphit e type Needle-ceke l Near-isot ropic Near-isotropic Near-isotropic I

Hole geometry 210/108 l210/10(c) 132/72 210/108(c)

(number of fuel  !

holes / coolant l' holes)

Coolant hole 0.625 0.625 0.826 0.625 diameter, in.

Fuel hole diameter. 0.500 0.498 and 0.500 0.624 0.498 in.

(a)See Table 2-1 for test matrix for FTE-1 through FTE-8. FTE-7 and FTE-8 contain reference fuel. FTE-2, -4, and -6 contain STU f uel.

(b)UCxy 0 is the reference recycle fissile kernel (360 um) and is a development goal for the large HTGR (305 km). It is derived from ion exchange resin beads that are loaded with uranium and then heat treated to obtain a controlled amount of conversion to carbide.

(c) Arrangement of each test fuel combinatien is six fuel holes surrounding one coolant channel.

4-9

TABLE 4-2 TEST ELEMENT AS-BUILT HEAVY METAL LOADINGS Uranium Loading Thorium Loading U-Total Th-Total lleavy Metal Heavy Metal Core Element Location Segment g/ rod kg Total g/ rod kg Total FTE-1 25.07.F.06 2 0.177 0.551 3.519 10.966 FTE-2 22.06.F.06 3 0.177( } 0.596 3.514(b) 10.832 FTE-3 30.04.F.06 4 0.177 0.551 3.519 10.966 FTE-4 27.02.F.06 5 0.177 f) 0. 59;, 3.514(b) 10.832 FTE-5 24.03.F.06 6 0.177 0.372 3.519 7.411 Buffer 0.135 0.136 4.648 4.681 p Total FTE-5 0.508 12.092 FTE-6 10.07 F.06 7 0.309( 1.030 3. 561 (b) 10.938 1 3 FTE-7 05.05.F.06 7 0.300 0.935 3.564 11.105 FTE-8 05.06.F.06 7 0.300 0.935 3.564 11.105 Total FTE-1 through FTE-8 5.702 88.836 Segment 7 reload 220.0 2,400.0 Initial core 773.3 15,905.0 k u

"?

P "

(a)There are 88 rods each of low-enriched uranium in FTE-2, -4, and -6. The "$

U-total loadings for these rods are 0.686, 0.686, and 1.105 g/ rod, respectively.

The Th-total loadings for the low-enriched rods in FTE-2, ~4, and -6 are 2.411, 2.411, and 1.996 g/ rod, respectively.

NOTE: The fuel loadings of the MEU fuel rod variety were adjusted to obtain peak and average irradiation temperatures similar to those for the llEU fuel varieties.

9 O O

GLP-5494 Amendment 3

5. PERFORMANCE ANALYSIS - NORMAL OPERATION 5.1. NUCLEAR ANALYSIS The ef fects of the test elements on the nuclear characteristics of the FSV core were evaluated using axial FEVER code and radial GAUGE code (Ref.

8, pages S2-92) diffusion-depletion calculations. These methods are the sane as those used to calculate the core physics parameters presented in the FSAR. The calculations were performed to establish the axial and radial power correction factors for FIE-1 through FIE-5, which replace partially depleted elements, and to estimate possible ef fects of the test elements on control rod worth, excess reactivity, and other nuclear-related parame te rs .

5.1.1 Fuel Loadings and Burnable Poisons As-built uranium and thorium loadings for the test elements are given 3 in Table 4-2. FTE-2 through FTE-5 contain uranium loadings which are lower than the uranium loadings for the elements which they replace. They con-tain no lumped burnable poison. FTE-6, -7, and -8 use the same fuel load-Ing as the Segment 7 elements they replace and will contain the standard Segment 7 lumped burnable poison loadings. FTE-1 contains a higher uranium loading than that for the element it replaces; it will contain no lumped burnable poison.

In order for the MEU fuel rods to experience peak and time average temperatures comparable to those in other fuel rods containing HEU, the 1

loadings were chosen so that the power perturbation factors for MEU fuel are comparable to those for the other fuel in each FIE.

5-1

GLP-5494 Amendment 1 5.1.2. Power Perturbations 0

The radial and axial power perturbations due to the test elements were computed with the FEVER and GAUGE computer programs and were combined to obtain a total power correction factor for the test elements. It is assumed that the radial power perturbation changes at approximately the same rate as the axial perturbation and will gradually approach unity during the segment life cycle. These combined correction factors are given in Fig. 5-1. This correction is defined as:

pg, power in test element at time (t) p power in element replaced at time (t)

Tine-dependent axial power correction factors were obtained using FEVER results for each of the test elements as shown in Fig. 5-2.

The FEVER calculations overestimate the power perturbation effect since the boundary conditions imply an infinite number of adjacent patches with test elements added.

Table 5-1 shows the beginning-of-cycle power perturbations in both the fuel column containing the test element and in the corresponding fuel region. The small power perturbations shown will be reduced further with burnup. There is no change for FIE-6, -7, and -8, which are located in Segment 7. These results emphasize that the test elements will not cause the bases for the limits specified in LCO 4.1.3 of the FSV Technical Specificationn to be exceeded.

Axial power perturbations for the MEU fuel are plotted in Figs. 5-2a, 5-2b, and 5-2c, which show that the perturbation factors for the FEU fuel are comparable to the rest of the FIEs. Radial power perturbations are 1 expected to be the same as those for the rest of the FIEs, since the MEU fuel occupies a snull volume of the block and will not affect the radial flux distribution significantly.

5-2

GLP-5494 Amendment 1 analysis takes into account these differences in graphite thermal conductivity, thermal expansion, and irradiation shrinkage.

The presence of shim particles in many of the test fuel rods results in higher fuel rod thermal conductivity (Ref. 11, page 1). This has been conservatively neglected, and a fuel rod thermal conductivity of 4 Btu /hr-ft *F has been used in the analysis, as was done for the analyses in the FSAR.

The test fuel rods containing BISO fuel particles experience higher irradiation contraction in the radial direction than fuel rods containing all IRISO fuel. This has been accounted for in the analysis.

A fuel performance analysis was also done for normal operating conditions for the MEU fuel rods in each test element to obtain projections of fuel temperature and performance. The analysis was based on the thermal analysis discussed above, using the power perturbation factors shown in Figs. 5-2a, 5-2b, and 5-2c.

Kernel nigration models developed for the VSM UC TRISO particle were 2

also applied for the WAR UC 0 TRISO particles. The models will give a xy censervative estimate of TRISO WAR fuel performance because the WAR kernels 1

are porous and less subject to kernel migration than dense VSM kernels. It has been shown that unirradiated WAR kernels do not migrate in a thermal gradient (Ref. 12, page 69). It was conservatively assumed, however, that they migrate at the same rate as VSM UC kernels. Kernel migration for the 2

Th0 TRISO particle was taken to be the same as that of the Th0 BISO par-2 2 ticle, based on similar kernel diameters and densities (Ref. 12, page 69).

Kernel migration rates for the (Th/U)O MEU fissile particles rare obtained from data contained in Ref. 12.

As discussed in Section 3.6.4 of the FSAR, uncertainties in engineer-ing tolerances and physical properties result in uncertainties in calcu-lated fuel centerline temperatures. The hot spot evaluation presented in the FSAR may be applied to the temperatures calculated for the test elements.

5-5

GLP-5494 Amendment 3 O

5.2.2. Analysis Results The fuel performance of test elements FIE-1 through FTE-8 has been compared individually with the corresponding FSV reference elements as a function of fast neutron fluence, burnup, and temperature under normal operating; conditions.

The performance of all the test element fuel versus the FSV reference fuel is summarized in Table 5-2, where the most severe environments of fuel centerline temperature, fast neutron fluence, and fuel particle burnup are conpared. In addition, the resultant end-of-life kernel migration dis-tances, calculated by the methods described in FSAR Amendment 16, Attach-ment A (page 3.3-12 and Fig. 3.3-5), are given. The test arrays, 1 to 7, and driver fuel in test elements FTE-2, -4, and -6 (refer to Table 2-1) are formed into four groups as shown in Table 5-2, and are expected to have similar particle burnup, fuel rod dimensional change characteristics, and kernel nigration rates. I FTE-1 will operate at slightly higher power levels than the element it replaces, as shown in Fig. 5-1. The resultant increase in peak fuel tem-perature over the reference fuel element temperature is about 180 C. This 3 increase in temperature results in a higher kernel migration rate, but the kernel migration distance at the end of 200 EFPD is only a small fraction of the buffer coating thickness, as seen in Table 5-2. The end-of-life fluence, burnup, and expected fuel particle coating failure fraction are small in FTE-1.

PIE-2 through FTE-5 will operate at lower power levels than the FSV reference elements they replace, as shown in Fig. 5-1. The test elements experience lower fuel temperatures, as seen in Table 5-2; they also expe-rience lower fuel burnup and fast neutron exposure, as their core residence time is 200 EFPD shorter than that of the elements they replace. The ker- l3 nel migration rates are dif ferent for the different kernel types, but the 5-6

GLP-5494 Amendment 1 end-of-life kernel migration distances are very small relative to the buffer coating thicknesses, and hence particle coating integrity will be maintained.

Tes t elements FIE-6, -7, and -8 will operate at the same power as the corresponding FSV reference elements apart f rom the IEU fuel in FTE-6. As l1 shown in Table 5-2, the temperatures experienced by these test elements are slightly higher (<5*C) than those of the FSV reference elements they replace. This is due to the combination of the differences in thermal properties of graphite and the fuel rod dimensional change characteristics.

ETE-7 and FIE-8 have the same type of fuel as the FSV reference elements, whereas FIE-6 contains test fuel. The test elements experience the same fluence and burnup environment and are expected to have about the sacu level of fuel performance as the FSV reference elements.

Table 5-2 also presents the results of the analyses for the MEU fuel in test array 2 o f FTE-2, -4, and ~6. A comparison of the results of the IEU analyses with those of other arrays shows that, with the exception of FT E-6 , the peak MEU fuel temperature will be less than that of the FSV re f e rence fuel it replaces. The peak FTE-6 temperature will be slightly higher (%13 C) than that of the reference fuel, o . still less than the 1260 C (2300*F) maximum temperature limit in the FSAR. The higher MEU fuel temperature in FIE-6 results from the power perturbation curve in Fig.

5-2c. 1 Kernel migration distances have been calculated at upper 95% confi-dence levels. Fissile kernel migration distances in FIE-6 are seen to be larger than those expected for other test fuel types. Nevertheless, the calculated migration distance is still less than the minimum combined thickness of the buffer and inner pyrolytic carbon (Table 5-3) , assuming a 15% standard deviation in the thickness of each coating, which is con-sistent with the standard deviation assumed in the MEU particle manu-facturing specification. Thus, it is expected that the coating integrity 5-7

GLP-5494 Amendment 3 O

of the IEU TRISO particles will be maintained during normal operation.

Expected (50% confidence) migration distances for the !EU fissile kernels 1 were also calculated and are 1cuer than those shown in Table 5-2 by about a factor of five.

5.3. FISSION PRODUCT RELEASE ANALYSIS

5. 3.1. Gaseous Fission Product Release Both pyrolytic carbon (PyC) and sic coatings are effective barriers to the release of gaseous fission products (including iodine). Consequently, gaseous fission products released into the coolant are from fuel particles with failed coatings and from uranium and thorium contamination of the as-manuf actured fuel. For this analysis, contamination is defined as the heavy metal not contained by intact fuel particle coatings at the beginning of irradiation. In the manufacture of test element fuel for the FSV reac-tor, fuel manufacturing specifications were set and rigorous quality con-3 trol procedures followed to assure that heavy metal contamination levels for the test fuel are equal to or less than those for the regular FSV fuel.

llence, the fission product release due to heavy metal contamination in the fuel will be equal to or less than that from the corresponding FSV fuel.

The contribution of fuel particles with f ailed coatings to gaseous product release is determined by the fuel temperature, particle failure fraction, and fission gas release rate of each f ailed fuel particle.

The fission gas release rate from fuel increases with increasing temperature, assuming a constant fuel particle coating failure fraction.

Table 5-2 indicates that the test elements will operate near or below the temperatures of the reference FSV fuel they replace. Consequently, consid-eration of only the operating temperatures of the test elements would result in a prediction that the total release of fission gases would be lower in the test fuel than in the fuel being replaced.

5-8

GLP-5494 Amendment 3 As indicated in Table 5-2, maximum kernel migration distances for particles in the test elements will be within reasonable limits. The maximum calculated migration distance in a BISO test particle is about 17 l3 micrometers, and the maximum calculated migration distance in a TRISO MEU test particle is about 75 micrometers, considerably less than the combined l3 buf fer and inner pyrolytic carbon coating thickness of the MEU test fuel.

This re s ult , along with lower operating temperatures (Table 5-2) and the irradiation testing experience already obtained for coated particles of the test fuel type (Section 7.3), leads to the conclusion that the coating performance of the test fuel particles will be entirely satisfactory.

It has been experimentally shown, however, that fuel particles with UAR fuel kernels release fission gases at a rate about six times greater than the reference FSV VSM fuel particle kernel type upon failure of the fuel particle coating. This is interpreted as being due to the lower material density of the WAR kernels (Ref. 13, page 7).

A hypothetical example is given here to show the safety margin in circulating activity on a more quantitative basis. The test elements represent less than 1% of the total fuel in the core, and the fuel rods containing WAR kernels rep resen t less than 0.03% of the total fuel in the core at the time they are loaded.

For the purpose of illustration, assume, arbit rarily , that the par-ticle failure rate of all the test fuel is 100% greater than that of the re fe rence fuel. Assume also that 0.03% of the total fuel in the core is WAR fuel, 0.37% is other test fuel, and the remaining 99.60% is reference FSV fuel. Assume further that the gaseous fission product release rate of failed WAR fuel and other oxide test fuel is 10 times that of the reference fuel. Under these highly improbable conditions, the amount of circulating activity in the circuit contributed by particle failure would be increased by about 4%. Since, as shown in Table 3.7-1 of the FSAR, the margin 5-9

GLP-5494 O

between expected and design circulating a,ctivity is about a factor of 10, a 4% increase in circulating activity clearly poses no rist- to the health and safety of the public.

In fact, no significant increase in circulating activity is expected.

Circulating activity contributed by particle failure is only part of the total circulating activity, the remainder being caused by as-manufactured heavy metal contamination. The amount nel in the core will decrease as the test elements are remsvu alug successive refuelings (Table 2-1). The gaseous fission product release rate of failed WAR particles is only a factor of 6 times that for the reference fuel, and the coating performance of the test fuel is expected to be similar to that of the re ference fuel.

5.3.2. Metallic Fission Product Release The presence of BISO coated fuel particles in the test fuel makes possible an increase in the release of metallic fission products. Unlike TRISO coated fuel particles, which contain an sic coating that is imper-meable to metallic fission products, BISO coated fuel particles have only PyC coatings. Irradiation tests have shown that some metallic fission products migrate through BISO fuel particle coatings when the fuel operates at a high temperature for a long period of time. Calculations have been performed to determine the ef fect of placing BISO fuel particles in FSV FTEs using methods discussed in Section 3.7 and Appendix A of the FSAR.

Calculations were done for cesium and strontium, since their migration properties are representative of a large group of chemical elements and since Cs-137 and Sr-90 are two of the more important radionuclides.

The results of these analyses indicate that the releases of cesium and strontium from ETE-6 are approximately 2 and 5 times greater than those from the FSV reference element replaced by the test element. It was assumed that BISO particle failure is 10% throughout the test element life, that the total metallic fission product inventory of each BISO particle is 5-10

GLP-5494 Amendment 1 released to the fuel rod matrix upon failure, and that the BISO fuel oper-ates at its peak tenperature throughout its life. These assumptions are very conservative. In particular, data presented in Ref. 1 (pages 4-1 to 4-8) indicate that Th0 kernels are strongly retentive of strontium. Thus, strontium release from failed BISO particles is limited to those strontium fission fragments which escape the kernel via recoil -- about 5% of the total.

The fractional contribution of rulease from the test elements to the 30-year cesium and strontium design inventories is accordingly very small, since only three elements out of 1482 will contain BISO fuel particles, and the BISO particles will be contained in only 12% of the fuel in each of those elenents. The release of cesium and strontium from FIE-2 and FIE-4, the only other test elements containing BISO fuel particles, is at least an order of magnitude less than the release from FrE-6 due to their substan-taally lower operating temperatures (Table 5-2). Thus, the increase in metallic fission product release due to the presence of a small quantity of BISO coated fuel particles will be imperceptible.

Although it is expected that MEU particle coating integrity will be maintained during rormal operation, the analyses discussed in Section 5.2 indicate that MEU oxide kernel migration distances are larger than those experienced by other FSV fuel types. Thus, the margin available for kernel migration during transients is less. Accordingly, although less than 1%

FEU particle failure is expected, the ef fects on fission product release discussed below are conservatively based on 10% "ailure of the MEU fuel particles in the FTEs.

As indicated in Table 4-2, MEU fuel particles will be loaded into a I

total of 264 fuel rods. The total initial uranium and thorium contents of the MEU particles in those rods will be 220 g and 215 g, respectively. In contrast, the initial FSV total core thorium-uranium loading was 1.6744 x 10 g. Thus, 10% failure of the MEU particles in the FIEs would result in a fractional thorium-uranium exposure of 2.6 x 10" . Since the " design" 5-11

GLP-5494 Amendment 3 fission product activity levels given in Tables 3.7-1 and 3.7-2 of the FSAR 9

~ ' '

are obtained assuming a 10% failure fraction at the end of irradiation for each segment, which is equivalent to a core average failure fraction of 0.05, and, in addition, fuel manufacturing specifications for FSV fuel

~

allow a fractional heavy metal exposure of 1 x 10 ' due to as-manufactured heavy metal contamination, an incremental thorium-uranium exposure of 2.6 x

-6 10 over that expected for the total core will have en insignificant ef fect on fission product activity levels.

With regard to plutonium buildup in the MEU rods, depletion analyses discussed in Section 5.1 indicate that the sum of the peak plutonium levels 1

expected in each of the three FTEs is approximately 0. 75 g. In contrast, the total core plutonium content in the equilibrium core is approximately 2300 g. Thus , 10% f ailure of the MEU particles in the FIEs would result in a fractional plutonium exposure (not release) of less than 3.3 x 10

-5 , g shown by comparison with the " design" core average failure fraction of 0.05

-0 or the as-manufactured exposure fraction of 1 x 10 plutonium exposure of 3.3 x 10

, an incremental wil'. clearly have an insignificant impact g

on plutonium-related fission product release.

In summary, tha presence of MEU fuel rods in FTE-2, -4, and -6 will have no significant effect on FSV fission product activity levels.

5.3.3. Conclusions Results from the thermal analysis (Table 5-2) indicate that FTE in-pile operating temperatures are essentially equal to, or are substantially less than, the reference fuel elements being replaced, with the exception of FIE-1, which operates at a higher temperature than the fuel element it replaces, but only for a 200 EFPD irradiation period. Fuel particle coating l3 performance is expected to be similar to that in the fuel elements which are replaced. Even if failure rates in the test fuel are higher, analyses show that gaseous fission product release (i.e. , krypton, xenon, iodine) from the test elements will be only slightly greater than that from the 5-12

GLP-5494 Amendment 3 elements they replace. The effect on metallic fission product release (i.e. , cesium, strontium) will be imperceptibly small.

5.4 GRAPHITE STRUCTURAL ANALYSIS The most significant structural difference between the reference fuel elements and the test elements is the use of near-isotropic H-451 graphite in place of the needle-coke H-327 used in the reference fuel elements. The tensile strengths of these two types of graphite are given in Table 5-4 for H-451 graphite and in the FSAR, Table 3.4-1, for H-327 graphite. It can be seen that the strength of H-451 graphite in both the radial and axial directions is higher than that of H-327 graphite.

The magnitude of fuel element stresses and deformation is determined by the following mechanical properties, which are compared for the two types of graphite:

1. Elastic modulus. The axial modulus of H-451 graphite is approx-imately 30% to 40% lower than that of H-327 graphite at all 1 fluence levels, and therefore, for a given strain, the axial stresses in H-451 graphite elements are significantly lower than those in H-327 graphite elements.

2. Creep properties. The steady-state creep behavior of both graphites is similar, with H-451 having slightly lower values.

3. Irradiation-induced dimensional change. Operating stresses are produced within the graphite elements by strains due to differ-ential irradiation-induced dimensional changes across the ele-ment. Figure 5-3 shows the irradiation-induced dimensional changes of H-451 graphite in both the axial and radial direc-tions. The irradiation-induced dimensional changes for H-327 graphite are shown in the FSAR, Fig. A.1.15-1. A comparison of l3 these two graphites shows that at peak temperatures and fluences, 5-13

GLP-5494 Amendment 3 the irradiation-induced dimensional change of 11-451 graphite in 9

the axial direction is about 50% lower than that of 11-327 graph-ite. Therefore, the stresses in fuel elements made from H-451 graphite are lower than those in the 11-327 graphit lements they replace.

4. Thermal strains. While the irradiation-induced strains make the najor contributions to the operating stresses within the graphite elements, the thermal strains contribute strongly to the shutdown stresses. The thermal expansion of 11-451 graphite is about 30%

and 100% higher than that of II-327 graphite in the radial and axial directions for all temperatures of interest.

5. Thermal conductivity. This important graphite thermal property determines the temperature distribution in fuel elements. The thermal conductivity of II-451 graphite is greater than that of 11-327 graphite. This tends to reduce operating and shutdown stresses in the II-451 graphite because of the smaller temperature g

differential than that for H-327 graphite.

The computer programs FESIC and SAFE GRAFITE were used to calculate the operating and shutdown strain and stress distributions, both axially and radially, throughout the lifetime of the fuel elements. The graphite web temperature dif ferences were calculated with conservatively chosen irradiated graphite thermal conductivity values. These low thermal conductivity values maximize the graphite temperature differences and thereby maximize the predicted thermal and contraction stresses.

The initial thermal stresses oduce a compressive stress near the hotter fuel channel and a tensile stress near the cooler coolant channel.

The maximum calculated FTE initial thermal tensile stresses were 972 kPa I

(141 psi) in the axial direction and 841 kPa (122 psi) for the radial direc-tion. This compares with maximum initial core 11-327 stresses of 1034 kPa (150 psi) and 1241 kPa (180 psi) for the axial and radial orientations, respectively, as shown in the FSAR, page 3.4-6.

g 5-14

GLP-5494 Amendment 3 Under irradiation by fast neutrons, the hotter graphite shrinks faster than the cooler graphite. After a period of operation varying from a few months to about a year, the irradiation-induced dimensional changes over-come the thermal strains in magnitude, and the cooler portion of the graph-ite which was originally in tension goes into compression while the hotter portion goes into tenoion, the FSV plant is expected to operate continuously between refueling operations. However, shutdowns can occur at any time. These shutdowns cause large changes in temperature distributions and hence elastic stresses in the fuel elements. In the absence of creep and irradiation strain, a shutdown to a state of uniform temperature would simply remove the oper-ating thermal stress and reduce the stress to zero. The influence of creep and irradiation strain, however, causes maximum shutdown stresses higher than maximum normal operating stresses.

Shutdown stresses were calculated for the test elements at several time points during their residence in the core by superimposing a shutdown on the normal continuous operating history. The maximum shutdown stresses were calculated for test element FTE-6. FIE-6 has the highest power density and operating temperature of the test elements irradiated to significant fluence and therefore has the highest stresses. The maximum calculated shutdown tensile stresses were 1268 kPa (184 psi) and 1475 kPa (214 psi) 3 for the axial and radial orientations. This compares with maximum initial H-327 core stresses of 3103 kPa (450 psi) and 1379 kPa (200 psi) in the j axial and radial orientations (FSAR, page 3.4-7). Although the H-451 fuel element has a slightly higher stress than the H-327 element described in the FSAR, it has a greater design margin (the difference between strength and stress) due to its greater strength.

A summary of operating and shutdown stresses is given in Table 5-5.

5-15

GLP-5494 Amendment 3 Based on the above analysis it is concluded that the inclusion of the test elements in the FSV reactor will improve the stress margins relative to those of the reference core.

5.5. GRAPilITE DIMENSIONAL CilANGE During core operation, the graphite test elements will be exposed to fast neutron irradiation which will induce dimensional changes. Because of the differences in dimensional change behavior between H-451 and H-327 graphite, an analysis was performed to calculate the expected axial and radial dimensional changes of new test elements. The maximum axial shrink-ages of all the 11-451 test elements are less than those of the correspond-ing H-327 elements that they replace. However, for the radial direction, the maximum calculated end-of-life shrinkage of FTE-6 is 1.1% compared l3 with the 0.8% listed in the FSV FSAR (page 3.4-9) for initial core H-327 fuel elements.

The maximum coolent hole offset that could be caused by differential thermal and irradiation strain across the interface of the FTEs and the remaining H-327 fuel elements in the column will result in an increase in centerline fuel temperature of less than 8'C (15'F). Because the FTEs are all standard blocks, their shrinkage will not af fect control channel offset.

The small change in radial gaps caused by the replacement of H-327 elements by the FTEs will have a negligible effect on coolant flow through inter-element gaps. This is due to the dominating effects of the five H-327 fuel elements remaining in the fuel element column.

Owing to the small difference in H-451 and H-327 dimensional change and the small number of H-451 elements in the core, the test elements will have no effect on the overall core seismic motion. The individual H-451 elements will have a larger design margin in a seismic event than the H-327 elements they r' place due to their higher strength. 3 5-16

GLP-5494 Amendment 3 Fuel element bowing is primarily caused by the large radial fast neutron flux gradient across an element. This situation occurs in the 1 buffer locations, and, as expected, FTE-5 was found to have the largest '

3 amount of bowing.

The maximum calculated end-of-life bowing for FTE-5 is about 0.05 in. 3 (maximum difference between bowed and original axial centerlines) as com-pared with the 0.09-in, maximum bowing predicted for the initial H-327 ele-ments (FSAR, page 3.4-8). Therefore, column *.ilting and the resulting element interface coolant flow will be reduced.

5-17

GLP-5494 O

TABLE 5-1 FUEL COLUMN AND REGION POWER PERTURBATIONS AT BOL FTE-1 FTE-2 FTE-3 FTE-4 FTE-5 Fuel column +8% -2% -5% -4% -2%

Refueling region +3% NO -2% -1% NO t

O O

5-18

TABLE 5-2 MOST SEVERE ENVIROD1ENT EXPERIENCED BY VARIOUS FUEL TEST ELDtENTS Max. Maximum Kernel Fuel ( Maximum FIMA Migration (a) (um)

Test Temp. Max. nvt Element Test Array ( C) (x 10 25 n/m2 ) Fissile Fertile Fissile Fertile I

FTE-1 Driver 1208 0.73 0.326 0.001 2 5 Ref. fuel 1029 1,49 0.081 0.004 <1 <1 FIE-2 1 1016 2.13 0.073 0.003 <1 <1 2 991 2.13 0.064 0.008 5 <1 3 to 6 1030 2.13 0.470 0.008 <1 <1 7, driver 1009 2.13 0.470 0.008 <1 ,

<1 Ref. fuel 1090 2.99 0.113 0.009 <1  ! <1 FTE-3 Driver 9'4 3.06 0.591 0.014 <1 <1 Ref. fuel 1066 3.96 0.156 0.020 1 <1 u FTE-4 1 1055 4.03 0.115 0.026 <1 2 1 2 1064 4.03 0.115 0.027 12 <1

• 3 to 6 1070 4.03 0.663 0.026 <1 2 1 3 7, driver 1045 i 4.03 0.663 0.026 <1 1 Ref. fuel 1092 5.09 0.164 0.034 1 1 FTE-5 Driver 1062 4.45 0.713 0.037 <1 4 Ref. fuel 1130 5.01 0.181 0.049 4 4 FfE-6 1 1243 6.97 0.193 0.063 10 16 2 1254 6.97 0.161 0.061 75 16 3 to 6 1246 6.97 0.744 0.063 6 17 7 driver 1236 6.97 0.744 0.063 5 15 Ref. fuel 1241 6.97 0.193 0.063 10 16 ka FT E-7 Driver su ' 6.98 0.196 0.064 2 1 S. $

Ref. fuel (b) 1089 6. ' i 0.196 0.064 2 1 @y

s u FTE-8 Driver 1092 6. 7t 0.196 0.064 Ref. fuel 1087 6. 7 , 0.196 0.064 2

2 1

1 ue (a) Calculated at upper 95% confidence level.

(b) Reference u,. i . environment which replaced FSV fuel would have experienced.

TABLE 5-3 NOMINAL FUEL PARTICLE DESIGNS IN FSV FTE-1 THROUGH FTE-8 Nominal Dimensions (pm)

Particle Kernel Coating Kernel Buffer IPyC sic OPyC Type Material Type Diameter Thickness Thickness Thickness Thickness Fissile UC TRISO 195 110 35 35 40 2

Fissile WAR U*Cy *0 ("} TRISO 305 and 360 60 and 65 35 35 40 7

Fissile (Th,U)C TRISO 140 and 225 80 25 25 30 and 40 2

Fissile (Th,U)02 (MEU) TRISO 350 95 35 35 45 1 5 Fertile ThC TRISO 355 and 455 55 30 25

• 35 and 45 2

Fertile Th0 BISO 500 95 - --

80 2

Fertile Th0 TRISO 450 60 35 35 45 2

(" Nominal value of X = 4.0 and Y = 1.0, N

aa

&G O O O

GLP-5494 TABLE 5-4 MEAN ULTIMATE TENSILE STRENGTHS OF UNIRRADIATED H-451 GRAPHITE H-451 Strength (#}

Log Position [kPa (psi)]

Axial Direction End center 15,270 (2215)

End edge 18,030 (2615)

Mid-length center 13,580 (1970)

Mid-length edge 19,170 (2780)

Radial Direction End center 13,930 (2020)

End edge 15,170 (2200)

Mid-length center 10,760 (1560)

Mid-length edge 13,960 (2025)

(" Ref, 3, page 13.

5-2J

GLP-5494 Amendment 3 O

TABLE 5-5 COMPARISON OF MAXIMUM INITIAL OPERATING AND SHUTDOWN TENSILE STRESSES IN TEST ELEMENT FTE-6 AND H-327 ELEMENTS OF THE REFERENCE FSV CORE Tensile Stress

[kPa (psi)]

Reference Core H-327 Elements FTE-6 Axini Startup thermal stress 1034 (150) 972 (141)

R2x. shutdown stresses 3103 (450) 1268 (184)

Radial 1 3 Startup thermal stress 1241 (180) 841 (122)

Max. shutdown stresses 1379 (200) 1475 (214)

O O

5-22

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GLP-5494 Amendment 3 0

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5-26

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GLP-5494 Amendment 3 valid: 0.012 Ak for Event a and 0.047 Ak for Event b. Furthermore, the FIEs will not affect the shutdown margin, excess reactivity, reactivity lifetime, temperature coefficients, and xenon worth of the reference core.

Therefore, the analysis presented in the FSAR regarding overall power and temperature transients of the RWAs will not be affected significantly by the FIEs.

When the transient is terminated by the first line of protection afforded for the respective RWA transients, the FSAR shows that the increase of fuel temperature in the hottest regions does not result in any fuel failure (even for Event a with an assumed rod worth of 0.025 Ak).

Since the perforuance of the FIE fuel is expected to be equivalent to that of the reference fuel, no fuel failure would occur in the FIEs when the RWA is terminated by the scram at 140% power.

For the limiting case of an uncontrolled withdrawal of the most reactive rod pair (0.012 Ak) at 100% initial power in an end-of-cycle equilibrium core, the FSAR discusses consequences in the event that the first line of protection fails; i.e., the scram at 140% power does not occur and the scram occurs when the reheat steam temperature reaches 1075*F. In the FSAR case for this transient with all reference fuel, the fraction of fuel failed as a result of this accident does not exceed the fraction assumed failed as the basis for the " design" circulating activity levels and is considered acceptable. Since all FTEs except FTE-1 run near or below the temperatures of the referense fuel they replace, FTE-2 through FTE-f would not affect the FSAR conclusion regarding fuel failure.

Although FTE-1 will oi- cate at abour 421 (B0C) to 15% (EOC) higher power 3 over the length of the cycle tht, reference fue3 at the same location, its residence time in the core is about one-half that of the replaced reference 3

fuel. It would thus produce only aoout 60% of the fission product inven-tory of the replaced reference fuel element.

6-3

GLP-5494 O

However, even if it is conservatively assumed that all the fuel in FTE-1 fails during the transient, it constitutes a fuel fraction of only 0.07% of the core. Even with this su:all increment added, the total fuel failure for the event as considered in the FSAR would still be less than fuel f ailure assumed for calculating the " design" circulating activity (see FS AR Section 3.7.4.1.2) .

In addition to fuel failure, the FSAR considers the ef fect of an RWA on the core outlet gas temperature and potential hot streaks. As stated in Section 5.1.2, the power perturbations induced by FTEs will not result in region peaking factors in excess of those given in LCO 4.1.3 of the FSV Technical Specifications. Thus, the FSAR values for maximum outlet gas temperatures for the RWA event remain valid, and the integrity of the primary coolant system boundary would be maintained.

Therefore, it is concluded that the inclusion of FTEs in the FSV core does not alter the findings of the RWA analyses described in the FSAR for the equilibrium case and poses no undue risk to the health and safety of the public.

6.2.2. Loss of Normal Shutdown Cooling This accident is defined in Section 14.4 of the FSAR as the unavail-ability 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 cooldovn modes are analyzed for this accident. The most severe case (Sectien 14.4.2.1, Case B2) was determined to be cooling with one circulator driven by the fire water system.

O 6-4

GLP-5494

\nendment 1 As shown in Section 5, all of the FTEs except FTE-1 operate at rela-tive powers and temperatures less than or approximately equal to those of l3 the reference fuel they replace. Over the entire cycle, the maximum fuel temperature of FTE-1 remains less than the core maximum fuel temperature.

Because of the latter, the thermal transients with FTE-1 in place can be no more severe than for the FSAR cases.

6.2.3. Moisture Ingress The analysis in Section 14.5.2 of the FSAR considers inleakage into the primary coolant system from an economizer-evaporator-superheater subheader or tube or from the helium circulator bearing water supply. All other water-containing systems in the proximity of the primary coolant system are at a pressure less than the primary coolant pressure during operation at power and cannot leak into the primary coolant. Design of the steam generator limits the steam and water inleakage flow rate to 15.4 kg/sec (34 lbm/sec) initially, decreasing to 0 sg/sec (22 lbm/sec) in about 5 sec.

The FSAR treats several moisture ingress cases. Of these, Case 5, a steam generator subheader rupture compounded by concurrent failure of the moisture monitor system and a dumping of the wrong (non-leaking) steam loop, has the dreatest potential for graphite oxidation and fuel hydrolysis in the shortest tiac following the accident. As shown in the FSAR, Cases 5 and 6 actually show comparable results. To evaluate the potential effect of the FTEs on the analysis of Case 5, the following phenomena were investigated:

1. Steam-graphite reaction.

2. Oxidation-hydrolysis of failed fuel.

3. Potential change in fission product release due to 1 and 2.

6.2.3.1. S team-Graphite Reactions. As described in Section 14.5 and Appendix A.11,1 of the FSAR, the steam-graphite reaction is dependent on five principal variables:

6-5

GLP-5494 Amendment 3 0

1. Fractional burnof f of graphite.

2. Steam partial pressure.

3. Ilydrogen partial pressure.

4. Catalyst concentration.

5. Graphite temperature.

These variables apply both for the H-327 non-isotropic needle-coke graphite in the FSV reference elements and for the H-451 near-isotropic graphite in the test elements. The steam-graphite reaction rates at various temperatures for these two graphite types are equal within experimental uncertainty over the temperature range of interest, 700* to 1300*C (1292* to 2372*F) (Ref. 6, Figs. 6-1 and 6-2). The reaction dependence on fractional graphite burnoff (item 1) is empirical and not particularly sensitive to the change in graphite between the FIEs and reference fuel.

With these t imilarities in reaction rate versus temperature and surnof f, there will be no significant change in the steam or hydrogen partial pressures (items 2 and 3) during this accident due to the test elements. Eor these reasons, only the effects of variations in items 4 and 5 are subjects of comparison in this report.

The ef fect of the barium catalyst concentration was conservatively evaluat ed. Whereas the barium concentration employed in the FSAR for the reference fuel is 0.001 mg/g carb' n in the graphite, it is expected to be dif ferent for the test elements which employ some BISO particle coatings.

Although the BISO fuel rods constitute only about 12% of the rods in FIE-2,

-4, and -6 and will not all be irradiated for the full 5.7 years, a 3 conservative estimate of the barium concentration was made assuming the entire f uel loading of FIE-2, -4, and -6 to be a TRISO/BISO mixture and that these test elements reside in the core for the full exposure. As a conservative estimate, it was assumed that the barium concentration throughout FIE-2, -4, and -6 is 0.04 mg/g carbon.

O 6-6

GLP-5494 Amendment 3 Based on this concentration, the reaction rate for these FTEs would be on the order of 10 times greater than that of the reference fuel (see FSAR, page 5-3) . It may be shown analytically from the relationship for steam diffusion and reaction within the fuel blocks that the amount of graphite reacted in a given time is proportional to the square root of the reaction rate. Hence, FIE-2, -4, and -6 would, at most, exhibit about three times as much graphite reacted as the reference fuel at otherwise equal condi-tions. Hence, for the three PIES containing TRISO/BISO fuel, the net increase in the amount of graphite reacted in any moisture ingress accident would be about 0.4%, relative to the FSAR results.

The effect of graphite temperature is not of general concern in the FTEs. The FTE power correction factors shown in Fig. 5-1 illustrate that all test elements except FTE-1 will operate at power levels lower than, or approximately equal to, the levels in the standard elements which they 1 1

replace. Since fuel and graphite temperatures are proportional to the power generation, the cooler FTEs would tend to produce a lower amount of graphite oxidation than the respective replaced reference elements. FTE-1, however, is shown in Fig. 5-1 to run at higher relative power than the replaced fuel:

about 42% at BOC and 15% at EOC. Analyses indicate that the maximum surface temperature in FTE-1 during its residence in the core is 2013 F (1373*K).

As a conservative evaluation of the effect of relative temperatures on the steam-graphite reaction rate of FTE-1 relative to the replaced reference fuel, it is assumed that the replaced reference fuel operates with the core average graphite surface temperature, 1600*F (FSAR Section 14.5.2.2), and that graphite surface temperature in FTE-1 is 2013 F throughout. Employing l3 this assumption in the expression for graphite oxidation reaction rate in FSAR Section 14.5, it may be seen that the ratio of the reaction rate of FTE-1 to that of the reference fuel is approximately 11. As stated above, the amount of graphite reacted varies as the square root of the reaction rate; the amount reacted in FTE-1 relative to core average reference fuel would be greater by a factor of about three. This corresponds to an 3 increase in the amount of core graphite reacted of about 0.2% relative to the value cited in Table 14.5-1 of the FSAR. Hence, at BOC when the BISO-bearing FTEs would not yet exhibit a significant effect of barium catalysis, 6-7

GLP-5494 Amendment 3 e

the net incr( 'se in graphite reacted would be 0.27. at most. At EOC the 3 net increase due to PIE-1 plus FTE-2, -4, and -6 with barium catalysis would be about 0.67. relative to the FSAR values. Therefore, even with 3 conservative evaluation the presence of FTEs does not significantly alter the amount of reaction products and therefore does not alter the transients described in the FSAR with respect to PCRV pressurei and margin to relief system setpoints. Furthermore, since the FTEs are located in the upper half of the core, they cannot have any significant effect on the FSAR conclusions regarding the core support structure and bottom reflector.

Since the temperature of fuel in FTE-1 is lower than that of the 3 allowed core maximum at all times during its residence in the core, the amount of graphite reacted in FIE-1 during a moisture ingress event would be less than that in the hottest element reference fuel.

Y 6.2.3.2. Oxidation-Hydrolysis of Failed Fuel. Oxidation-hydrolysis reac-tion with fuel particles occurs only when steam diffuses through the h graphite block and reacts with fuel kernels whose coatings have failed prior to the moisture ingress event. Since FIE-7 and FIE-8 have original Segment 7 fuel and would not exhibit a change in fuel failure, these elements will not affect the FSAR results. Although FIE-1 through FIE-6 contain experimental fuel, the performance eva.uation discussed in Section 5 indicates that these are equivalent to the reference fuel. Therefore, none of the FIEs will significantly affect the amount of previously failed fuel or increase the total oxidation-hydrolysis in the core relative to the FSAR analysis because of failed fuel.

During a moisture ingress event, the rate of hydrolysis and noble gas fission product release is dependent upon local fuel temperatures. Since the reactor would be automatically shut down in the first few seconds af ter the initiation of this event, the local temperature peaks would flatten out and approach the coolant temperature. The operating temperature differences between the test elements and the reference elements, shown in Table 5-2, would rapidly diminish and minimize any effect on the fuel hydrolysis rates.

6-8

GLP-5494 Amendment 3

1. The gasborne activity assumes a " design" fuel failure fraction 3

which increases to 10% over the 6-year life of a fuel element and assumes an equilibrium core at the time of the accident (FSAR 3.7.4.1.2).

2. The pla teout activity is assumed to be a 30-year accumulation.

The fission gas release rate from failed WAR particles employed in four test arrays of three FTEs is greater than that of reference fuel.

Ilowever, as discussed in Section 5.3.3, the overall gaseous fission product release from the FTE driver fuel and test arrays is expected to be only 3 slightly greater than from the replaced reference fuel. Also, the contri-bution from contaminants outside the particle coatings will be less than or equal to that of the reference fuel. Therefore, these contributors to circu-lating activity are only slightly affected by the FIEs, and the " design" level gasborne activity in the FSAR remains conservative.

I Although the FIEs contain some BISO coated fertile particles, Section 5.3.2 states that increased release of metallic fission products from these particles will be imperceptible relative to that from the entire core.

Furthermore, the use of the 30-year plateout activity is very conservative.

It is therefere concluded that the presence of the FIEs will not increase the circulating or plateout activities significantly beyond the conservative values which were employed in the DBA-2 evaluation in the FS AR. Thus, the resulting offsite doses for a hypothetical DBA-2 are no'.

affected by the FTEs and remain at least an order of magnit'ade below 10CFR100 guidelines as presented in Table 14.11-4 of the FSAR.

6. 3. FUEL IIANDLING ACCIDENTS When loading FIE-1 through FIE-8 into the core at tTe different locations, an increase in the amount of fuel element handling will be necessary. Elements have to be removed and replaced to allow the FIEs to 6-13

GLP-5494 9

he placed in the core. As a result of this, an additional handling acci-dent can be hypothesized: replacing an irradiated element in the wrong location.

6.3.1. Loading Errors The placement of fuel during the FTE insertion will be controlled by the computerized fuel handling system. This will be checked and verified to ensure that errors in loading do not occur, as part of the standard operating procedures for fuel handling.

o 6-14