ML19269E301
| ML19269E301 | |
| Person / Time | |
|---|---|
| Site: | Washington State University |
| Issue date: | 05/31/1979 |
| From: | Wilson W WASHINGTON STATE UNIV., PULLMAN, WA |
| To: | |
| Shared Package | |
| ML19269E300 | List: |
| References | |
| NUDOCS 7906270184 | |
| Download: ML19269E301 (205) | |
Text
SAFETY ANALYSIS REPORT for the WASHINGTON STATE UNIVERSITY MODIFIED TRIGA NUCLEAR REACTOR Submitted To U.S. NUCLEAR REGULATORY COMMISSION FOR RENEWAL OF FACILITY LICENSE R-76 Prepared by W. E. Wilson WASHINGTON STATE UNIVERSITY NUCLEAR RADIATION CENTER Pullman, Washington 99164 May 1979
}h 2132 052
TABLE OF CONTENTS
1.0 INTRODUCTION
AND GENERAL INFORMATION 1.1 Introduction and Summary 1.2 Principal Design Criteria 1.3 Design Highlights 1.4 Conclusion 2.0 SITE CHARACTERISTICS 2.1 Location 2.2 Population Density 2.3 Climatolony 2.4 Meteorology 2.5 Geology 2.6 Seismology 3.0 FACILITY STRUCTURE 3.1 General Description 3.2 Heatino and Air Conditioning System 3.3 Pon1 3.4 Liquid Waste Collection System 4.0 REACTOR DESCRIPTION 4.1 General Descrintion 4.2 Bridge Structure 4.3 Grid Box 4.4 Fuel 4.5 Control Blades 4.6 Transient Control Rod 4.7 Control Elements Drives 4.8 Control and Instrumentation 4.9 Cooling System 4.10 Pool Makeup and Deaineralizer System 4.11 Experimental Facilities 4.12 N-16 Diffuser 5.0 0PERATIONAL PERFORMANCE 5.1 General Reactor Data 5.2 Steady State Operation 5.3 Pulsing Characteristics 2132 053
6.0 SAFETY ANALYSIS 6.1 General Considerations 6.2 Design Bases 6.3 Design Limits 6.4 Accident Analysis
7.0 REFERENCES
8.0 APPENDICES A. Safety Analysis for Conversion to FLIP Fuel B. Environmental Impact Appraisal C. Emergency Plan D. Facility License E. Technical Specif', o Lions 2132 054
LIST OF TABLES Chapter 2 TABLE NUMBER 2.2-1 Population Distribution Around Site 2.4-1 Total Number of Hours of Wind by Direction and Velocity 2.4-2 Total Tine for Winds of all Velocities 2.4-3 Monthly Average Precipitation, Daily Mean Temperature, and Mean Daily Minimum-to-Maximum Temperature Difference 2.6-1 Histori' Earthquakes, 1872-1979 within 200 Miles of Site Chapter 3 TABLE NUMBER 3.1-3 Nuclear Radiation Center Room Listing Chapter 4 TABLE NUMBER -
4.4-1 Standard and FLIP Fuel Parameters 4.8-1 Minimum Reactor Safety Channels Chapter 5 TABLE NUMBER 5.1-1 Control Element Worths in Core 30-A Chegter 6 TABLE NUMBER 6.3.5-1 Typical 110-Rod TRIGA Core Pulsing Response 2132 055
LIST OF FIGURES Chapter 2 Figure 2.1-1 State of Washington Map 2.1 - 2 Whitman County Map in Pullman Aru 2.1 -3 Campus Topography Map 2.1 -4 Site Area Topography Mao 2.1 - 6 Site Aerial Photograph 2.1 - 6 Site Photograph 2.2-1 Population Distribution Map 2.4-1 Frequency Distribution of Winds 2.4-2 Frequency of Occurrence of Winds Less than 3 Kilometers /Hr 2.5-3 Geologic Cross Section of Site 2.6-1 Location of Major Faults Within 100 Miles of Pullman "
Chapter 3 Figure 3.1 -1 Nuclear Radiation Center Photograph 3.1 - 2 Nuclear Radiation Center Floor Plans (a) Ground Floor (b) First Floor (c) Second Floor (d) Penthouse 3.2-1 Reactor Ventilation System 3.2-2 Continuous Air Monitoring System 3.2-3 Gaseous Effluent Monitoring System 3.3-1 Pool Structure .
3.4-1 Retention Tank System 2132 056
Chapter 4 Figure 4.2-1 Bridge Structure Photograoh 4.3-1 Core Grid Box 4.3-2 Reflector Element 4.4-1 Four-Rnd Cluster Drawing 4.4-2 Four-Rod Cluster Photograph 4.4-3 TRIGA Fuel Rod Construction 4.4-4 Instrumented Fuel Rod 4.4-5 Three-Rod Cluster and Transient Rod Guide Tube
- 4. 5-1 Safety Blade 4.5-2 Regulating Blade 4.5-3 Shroud Assembly 4.6-1 Transient Control Rod 4.7-1 Blade Drive Mechanisms 4.7-2 Blade Support Shaft 4.7-3 Transient Rod Drive 4.7-4 Schematic of Transient Rod Drive 4.8-1 Control Console Photograph 4.8-2 Safety and Linear Indication Channels 4.8-3 Pulse Mode Instrumentation 4.8-4 Scram Circuitry 4.8-5 Wide Ranae Channel 4.9-1 Pool Cooling System Schematic 4.9-2 Pool Cooling System Horizontal Layout 4.9-3 Pool Cooling System Vertical Layout 2132 057
Chapter 4 (Cont.)
4.10-1 Pool Water Treatment System 4.11-1 Experimental Facilities 4.11-2 Pneumatic Transfer System Chapter 5 Figure 5.1-1 Aonroach to Critical for Core 30A 5.1-2 Core 30A Layout 5.1-3 Blade No.1 Integral Worth Curve 5.1-4 Blade No. 2 Integral Worth Curve 5.1-5 Blade No. 3 Integral Worth Curve 5.1-6 Blade No. 4 Integral Worth Curve 5.1-7 Blade No. 5 Integral Worth Curve 5.2-1 Fuel Temperature vs. Power Level 5.2-2 Excess Reactivity vs. Power Level 5.2-3 Excess Reactivity vs. Megawatt-Days Burnup 5.2-4 Excess Reactivity vs. Indicated and Average Core Temperature 5.2-5 Average Temnerature Coefficient vs. Average Core Temperature 5.3-1 Peak Power and Energy Release vs. Reactivity Insertion 5.3-2 Peak Indicated Temperature vs. Reactivity Insertion Chapter 6 Figure 6.4-1 Argon-41 Concentration Levels about Site 2132 058
1-1
1.0 INTRODUCTION
AND GENERAL INFORMATION 1.1 Introduction and Summary This report describes the Washington State University modified TRIGA reactor and the operation of the reactor with a core of mixed Standard and FLIP
- fuel . This report supersedes and replaces all previous SAFETY ANALY-SIS REPORTS and descriptions of the Washington State University reactor.
The Washington State University reactor has been in operation since March 1961. From 1901 to 1967 the reactor was fueled with MTR type fuel elements and operated at a maximum power level of 100 kilowatts. In 1967 the reactor was shut down and the core and control systems were modified so that the reactor could operate with TRIGA type fuel . The orginal core grid box was retained and the MTR fuel elements were replaced with a special 4-rod cluster of TRIGA fuel rods de::igned to replace a MTR fuel el ement. From July 1967 to date the reactor has operated as a modified TRIGA reactor with a maximum steady state power level of 1 MW. In February of 1976 the co e was loaded with a mixture of Standard and FLIP fuel.
1.2 Principal Design Criteria The existing reactor system operates with a mixture of Standard and and FLIP types of TRIGA fuel in the steady-state or pulsed modes. The maximum continuous steady state power level is 1 MW and the average maxi-mum pulsed power level is 2000 MW. Standard TRIGA fuel contains uranium-zirconium hydride enriched in U-235 to 20%. FLIP TRIGA fuel contains uranium-zirconium hydride enriched in U-235 to 70% and 1.5 wt% erbium as a burnable poison to offset the added U-235. The reactivity worths of
- FLIP (Fuel Life Improvement Program) is a new type long-lifed fuel .
developed by Gulf Energy and Environmental Systems for TRIGA reactors.
2132'059
1-2 both types of fuel are about equal. The increased U-235 content of FLIP fuel along with the burnable poison yields a fuel that has a significantly larger core lifetime potential than standard TRIGA fuel.
The safety of the modified system, as with all TRIGA reactors, comes from the large prompt negative temperature coefficient that is inherent in a water-moderated, U-ZrH fueled reactor. The overall operating charac-teristics for the Washington State University Modified TRIGA reactor fueled with various combinations of fuels is discussed in the report attached to Appendix A. The data in this report were calculated using the EXTERMINATOR-2(1) code and multigroup cross-section data (2) obtained from Gulf Energy and Environmental Systems.
1.3 Design Highlights The Washington State University Modified TRIGA reactor is located in the Nuclear Radiation Center which is situated in the northeast corner of the campus at Pullman, Washington. The core is located in a 242,000 liter above-ground pool which is, in turn, cooled and purified by external cooling and purification systems. Reactor experimental facilities include incore irradiation positions, a thermal column, and numerous beam tubes.
1.4 Conclusions Past experience with the Washington State University TRIGA reactor and other TRIGA reactors clearly indicates that a properly designed reactor system fueled with TRIGA type fuel can be safely operated at steady-state power levels of 1 MW and pulsed to a power level of 2000 MW. This history of safe and conservative reactor design has permitted TRIGA type reactors to be sited in urban areas without the need for specially designed contain-ment structures. Furthermore, the W.S.U. reactor fueled with a mixture of 2132 060
1-3 Standard and FLIP fuels has operated for over three years without a single fuel related probic.n.
The information presented in this safety analysis report indicates that the continued operation of the Washington State University Modified TRIGA reactor will pose no health or safety hazards to the public. Further-more, the system is safe when operated in a normal manner or even if a highly abnormal condition occurs. The three major accidents considered are: (1) accidental fuel addition, (2) pulsing of the reactor (transient rod ejec-tion) while operating at full power, and (3) loss of coolant accident. In all of these postulated accidents, no loss of fuel cladding integrity would occur. Also the Design Base Accident, which is the loss of the inteority of the cladding on one TRIGA rod in air, is shown to not present a signifi-cant hazard to the general public.
o 2132051-
O 2132 062
2-1 2.0 SITE CHARACTERISTICS 2.1 Location Washington State University is located in the southeastern corner of the State of Washington in the town of Pullman as shown in Figure 2.1-1. The town of Pullman, Washington has a population of 23,500 and is located in Whitman County about eleven kilometers from the Washington-Idaho border as shown in Figure 2.1-2. In addition to the town of Pullman, the town of Moscow, Idaho is located approximately 13 kilom-eters east of the site just across the Washington-Idaho border. Moscow has a population of 17,700 and is the location of the University of Idaho. The Palouse region surrounding the towns of Pullman and Moscow is a rural agricultural area devoted to dry land farming.
The actual reactor site is 3.2 kilometers east of the center of the town of Pullman and 1.6 kilometers east of the main portion of campus as shown on Figure 2.1-3. The site is surrounded by university-owned prop-erty for at least .4 kilometers in all directions used for the grazing of livestock as shown in the site photograph of Figure 2.1-6. The Moscow-Pullman airport is located 3 kilometers east of the site and the closest occupied dwelling is 411 meters west of the site.
2.2 Population Density The oopulation distribution about the site in 500 meter increments out to 3 kilometers in eight directional segments is shown in Figure 2.2-1 and tabulated in Table 2.2-1. The population distribution was calculated for a typical day with the university students, faculty, and staff present on the campus. A circle with a radius of 400 meters about the site has no permanently occupied dwellings.
2132 063
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2132 072
2-9 TABLE 2.2-1 POPULATION DISTRIBUTION AROUND REACTOR SITE Number of Residents per Octant s " N E SE S SW W NW NE e er 0 - 500 35 500 - 1000 0 0 4 2 800 1,032 66 92 1000 - 1500 0 0 0 0 27 1,574 3,393 233 1500 - 2000 0 0 0 15 18 5,454 5.280 0 2000 - 2500 0 0 0 0 0 700 2,800 76 2500-3000 10 2 4 2 30 588 3,200 280 3500 4 0 4 9 4 340 233 8 TOTALS 14 2 12 28 879 9, 68 8 15,007 689 2132 073
2-10 2.3 Climatology Pullman is situated at latitude 47 north of the equator and con-sequently is about midway between the equator and the North Pole. From May to August, when the sun remains above the horizon from 14 to 16 hours1.851852e-4 days <br />0.00444 hours <br />2.645503e-5 weeks <br />6.088e-6 months <br /> a day, Pullman receives more solar radiation than does the equator.
In Decenber, the sun rises only about 20 above the southern horizon at noon and is in the sky only about 8 hours9.259259e-5 days <br />0.00222 hours <br />1.322751e-5 weeks <br />3.044e-6 months <br />. Therefore, the daily accumu-lation of solar radiation in winter is less for two reasons: 1) the days are shorter, and 2) the sun's rays, striking the earth at an angle, are spread over a larger area. Because of this great variation in energy intake, Pullman experiences pronounced differences in temperature and other weather conditions from summer to winter.
The latitude of Pullman is only one factor influencing the climate pattern at the site. Other factors are its location with respect to land and water areas, mcuntain barriers, and prevailing winds. Pullman is approximately 480 kilometers inland from the Pacific Ocean; and the Cascade Mountains, which average more than 2 kilometers in height, separate Pullman from the coast. The combined effect of the distance from the ocean and the existence of the mountain barrier creates a climate with a continental characteristic. However, because the prevailina winds blow inland from the Pacific Ocean, winters are considerably warmer than other-wise micht be expected 480 kilometers inland at a latitude of 47 north.
Winters in Pullman are characterized by cloudy skies and frequent snow-storms. On the averace, the sun shines in Pullman only about 30% of the time durino the winter months.
During the summer months, the westerly winds weaken, and contin-ental climatic conditions prevail. Rainfall, cloud cover, and relative 2132 074
2-11 humidity are thus at their minimum; the daily mean temperature and daily temperature variation are at their maximum. Sumers in Pullman are char-acterized by warm clear days and cool nights. On the average, the sun shines in Pullman about 80% of the time during the sumer months.
2.4 Meteorology 2.4.1 General Washington State University is located in eastern Washington in a dry-land agricultural area, known as the Palouse region. The cli-mate of this region is moderate, being a transitional region between the Columbia Basin and the mountains of Idaho. Precipitation and temperature data at the Washington State University campus have been accumulated since 1893 by the Dc9artment of Agronomy at the school.
This extensive backlog of data was utilized to prepare the tempera-ture and precipitation tables given in this section. The wind data were obtained from a detailed analysis of wind velocity and direction data for the year 1953.
2.4.2 kind Velocity and Direction In order to obtain wind data that are relevant to the site, charts taken from a wind recorder located on top of Wilson Hall duririg 1953 were analyzed in detail. Wilson Hall is approximately 1.6 kilometers WSW of the site. The monitoring station was =at an elevation of 824 meters and the reactor site at 808 kilometers.
All more recent wind data are taken at the Moscow-Pullman airport which is about 3.2 kilometers ENE, at an elevation considerably below the site, and thus not valid.
A wind rose indicating the frequency of occurrence of winds at the site is given in Fiaure 2.4-1. It is to be noted that 'the pre-vailing winds are from a westerly direction and blow over Pullman 2132 075
2-12 N
7 NE NW -
Io A
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Frequency Distribution and Mean Wind Speed at Site from Direction Shown (Hatched Area is Percent Occurrence -
Solid Line is Mean Speed in MPH)
- 1 M.P.H. = 1.61 Kilometer /Hr -
FIGURE 2.4-1 7j{ Q]b
2-13 and the campus toward tne site. The major population density is upwind from the site about 57% of the time and downwind only about 21% of the time. Furthermore, about 79% of the time the wind blows in a direction in which there are no inhabitants for about .8 kilo-meters around the site.
The total number of hours of wind by direction and velocity is given in Table 2.4 and total time for winds of all velocities for each month is given in Table 2.4. These tables indicate that the average annual wind velocity is 16 kilometers /hr. Winds in January, the high month, average 21 kilometers /hr. Furthermore, the wind ve-locity was greater that 5 kilometers /hr 94% of the time and greater than 8 kilometers /hr 76% of the time. In general, one may conclude that there is almost always a light breeze blowing over the site.
2.4.3 Precipitation and Temperature The monthly average precipitation, monthly mean temperature, and monthly mean daily variation from minimum to maximum temperature at the site are tabulated in Table 2.4. The seasonal variations depicted in this table are a graphic representation of the climatic conditions that prevail at the site as previously described.
2.4.4 Temperature Inversions s Quagtitative data on temnerature inversions in the vicinity of the site are nonexistent. The closest points for which inversion data are available are at Spokane and Richland. However, the meteor-ological conditions at these two cities are significantly different from those at Pullman, making these data not applicable. The fre-quency distribution of winds of less thun 3 kilometers /hr is depicted in Figure 2.4-2. These low velocity winds blow only about 6% of the 2132 077
TABLE 2.4 -l Total Number of Hours of Wind by Direction and Velocity, 1953 Velocity DIRECTION Kilometers ,
Per Hour ,
N NE E SE S SW W NW 0-3 31.6 33.2 84.4 63.8 25.8 57.2 134.4 112.5 4-6 135.3 109.0 131.9 141.7 90.8 123.5 514.5 273.5 7-10 34.6 71.6 185.8 191.0 97.o 118.8 612.1 229.2 11-13 15.5 39.9 201.3 260.7 114.6 136.9 548.0 104.2 14-16 13.6 17.7 227.4 161.6 63.6 76.6 490.1 46.3 17-21 0.2 o.5 296.3 130.5 75.8 109.5 454.o 8.6 22-24 103.7 76.8 ,25.5 72.6 164.o 0.5 ha 25-32 1.1 230.1 78.6 19.6 72.2 256.9 2.0 gj 33 19.2 6.8 8.5 17.2 49.4 c3 $
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2-15 TABLE 2.4-2 Total Time for Winds of all Velocities Month N NE E SE S SW W NW 1953 hours houre hours hours hours hours hours hours January 2.3 3.4 43.3 123 7 93.7 135.3 243.4 16.o February 18.0 5.7 52.4 87.o 63.2 56.9 278.7 90.o March 14.2 10.0 146.5 123 5 31.5 81.8 308.3 19.4 April 13.9 24.0 152.2 45.5 31.6 73.2 310.4 69.o May 23.'9 29.3 157.4 51.1' 32.2 56.8 2 96 .6 88.2 June 27.0 38.1 64.5 25.5 18.1 68.8 347.0 106.5 July 58.0 57.3 81.3 63.8 17.3 31.0 257.5 124.1 August 24.4 39.3 96.3 92.2 54.1 48.3 267.1 110.6 September 28.5 34.4 141.0 81.5 46.7 75.1 265.5 38.4 october 15.0 37.0 208.5 106.0 34.6 54.8 225.0 62.2 November 5.5 3.9 257.2 134.9 33.5 36.3 158.7 37.5 December 1.5 o .~1 112.4 157.6 71.1 78.4 284.9 36.3 Total 232.2 282.51 1,513.0 1092.3 527.6 7 96.7 3243 1 798.2 Percent 2.7 3.3 17.8 12.9 6.2 9.4 38.2 9.4 Av. Dura- .94 1.26 1.46 .85 .75 .66 1.47 .73 tion (hrs.)
2132 079
2-16 YABLE 2.4-3 Monthly Average Precipitation, Daily Mean Temperature, And Mean Daily Minimum To Maximum Temprature Difference, 1893-1970 Daily Mean Precipitation Temperature Mean Daily Month in Centimeters in Degrees C. Minimum to Maximum January 6.78 -2.6 5.7 February 5.33 0.2 6.7 March 5.38 3.8 8.4 April 3.78 8.5 10.8 May 3.71 12.7 11.8 June 3.91 15.4 12.7 July 0.99 19.9 15.7 August 1.32 19.1 15.3 September 2.74 14.3 ,
13.0 October 4.85 10.0 10.4 November 6.27 3.2 6.7 December 6.96 0.1 5.7 Annual Total Precipitation - 49.50 centimeters Annual Average Temperature - 8.7 C Annual Average Difference between Minimum and Maximum Temperatures - 10.2 2132 080
2-17 time whereas winds in the 5 to 7 kilometer /hr range blow about 18%
of the time.
If the assumption is made that a temperature inversion can only be maintained with winds of below 3 kilometers /hr, then inversions could occur only about 6% of the time. The distribution of the low velocity winds further indicates that the population center west of the site would be downwind only about 22% of the time during which inversions could possibly occur.
2132 08i
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2-19 2.5 Geology 2.5.1 Regional Geologic History Pullman is situated in Eastern Washington near the eastern margin of the Columbia Rive Plateau. In early Miocene times the area was mountain-ous with a relief of over 1400 meters. These mountains, composed mostly of pre-Cambrian sedimentary and metamorphic rocks and Cretaceous granite, formed the basement rock across which the Columbia River Basalts would flow dt. ring the Miocene epoch. These basal t flows were numerous as well as extensive and advanced from the west and the south into the region.
The basalts of the Columbia plateau are somewhat unique in that a large thickness of volcanic material accumulated in a relatively short period on the geologic time scale. The lava flows extended over a 160,000 square kilometer area in the short span of about 3 million years about 16 to 13 million years ago. The total thickness of the basalt varies from 1000 meters in the Pullman area to a muimum of over 3 kilometers in the Pasco Basin. Individual flows were enormous and involved of the order of 300 cubic kiloneters of lava. The source of the immense amount of heat needed to create the lava flows is postulated to be a " Hot Spot" in the magma below the region. Some geologists believe that the " Hot Spot" re-mained stationary as the Pacific Plate moved west. This theory accounts for the young basalts in southern Idaho and the geothermal activity in Yellowstone National Park. The " Hot Spot" is thus postulated to presently reside under the Yellowstone Park region.
The basalt that flowed into the pre-i' low terrain of the region progres-sively submerged the basement features and darrrned up the well-established drainage systems. Numerous lakes were created along the margin of the grow-ing basalt plateau. Weathering of the exposed basement uplands prodyced.
2132 083
2 *20 detritus materials which rapidly filled in the temporary Miocene lakes estab-lished by the advancing basalt. Such lacustrine deposits were subsequently buried by flows from renewed basaltic eruptions triggering a repetition of the accumulation cycle. The solidified lava flows were nearly horizontal, however, the lava evidently eruptino from many different locations at ,iffer-ent times so that individual flows are not continuous across the plateau.
The orginal upper surface of the basalts were nrobably quite rouch but very low in relief.
At the end of the outpourings of the lavas of the Columbia River Basalt in early Pliocene times, mild folding of the basalt began. The folding con-tinued through middle and late Pliocene and intn Pleistocene time. Defor-mations in this age include the Cascadian orogeny which greatly affected the climatic conditions of the region. The main tetonic events during this period include; the uplift of the Cascade Range, Oregon Coast Range, Olym-pic Mountains, and Blue Mountains; the downwaroing of the Lewiston Grade, the Snake River Region, and the Walla Walla Plateau; a slightly 5:esterly increasing subsidence of the Columbia Plateau; the isostatic depression of the Pasco Basin; and block faulting of the Great Basin and Payette section.
Volcanic eruptions accompanied tN deformations particularly in the Middle Cascades giving rise to the volcanic peaks of the Cascade Range.
Following the cessation of the major igneous activity in early Pliocene times, the basalts and lacustrine deposits became subjected to moderate erosion as the drainage patterns began to develop. This initiated the dis-section of the plateau surface. During the Pleistocene enoch the modified surface was canned with the lcess of the Palouse formation and produced the rolling-hill topogranhy of the region. The most significant aeoloaical event during the nast million years is the Sookane flood at the end of the 2132 084
2-21 Ice Age. The advancing ice sheet dammed the Columbia, Spokane, and Clark Fork Rivers. The water that was impounded behind the dams filled the tri-butary valleys for many miles.
The lake created by the damming of the Clark Fork contained an esti-mated 1000 cubic kilometers of water or about half the volume of present day Lake Michigan. When the ice dam at the mouth of the Clark Fork failed, the lake drained at an estimated flow rate of 15 cubic kilometers per hour.
The incredible force of the massive flood scoured the Rathdrum Prairie and Spokane Valley creating the " Channeled Scablands" in the Sprague-Cheney area. Similar euents during the ice age created tta present features of the Columbia Plateau including Grand Coulee.
The region has seen a very unique sequence of geological events begin-ning with a vast series of lava flows. The lava flows were followed by a regional tiltino of the land and by the deposition of a 30-60 meter layer of wind blown silt. The great glacial lake formed by the damming of the Clark Fork and the destructive flood created by the sudden release was the final event that brought this area to its present character.
2.5.2 Site Geology The Columbia River formation in the Puliman area is approximately 1000 meters thick and consists of alternating layers of basalt and the silts and clays of the Latah formation. A geoloaic cross section of the Pullman area is shown in figure 2.5-1. The Palouse Formation Soil at the site is 35 to 55 meters thick. Structurally the layers of basalt in the Pullman area have not been disturbed since their deposition. The major movements in this section of the Columbia Plateau have been the Lewiston downwarp and the westerly subsidence.
2132 085
2-22 2.5.2 Geologic Haztrds No knJwn geologic hazards such as Karst terrain, cavernous conditions, tectonic depressions, surface or subsurface subsidence or uplif ts, or active volcanoes, are present at the site or in the immediate vicinity of Pullman. Also, there are no conditions present which could produce rock-falls, avalanches, floods, tsumani, mudflows, or permafrost at the site.
2.5.4 Site Groundwater Conditions The main aquifers in the Pullman area are associated with the Latah Formation interbeds between basalt flows as shown in Fiqure 2.5-1. Hori-zontal migration within an aquifer may also occur in the vesticular or porous top of the basalt layers. The cities of both Pullman and Moscow obtain their water from deep aquifers over 200 meters below the surface.
Carbon-14 dating of the water from the deep aquifers indicate that no measurable recharae has occurred in recent times. Accordingly it is be-lieved that a layer of imnervious basalt of about 100 meters below the sur-face prevents the downward migration of surface waters.
Recharge of the shallow aquifers is believed to occur at the eastern end of the Moscow-Pullman basin where the basalts contact the Pre-Tertiary Moscow Mountain Formation (see Figure 2.5-1). Additional recharge also occurs by infiltration from streams and precipitation water 2. However, surface waters percolate slowly downward due to the high water retention capacity of the Palouse Formation as well as the thickness of such soils.
Accordingly liquids discharged at the reactor site in the event of an accident will not enter the local aquifer. In addition, there are no rivers or streams within one kilometer of the site.
2132 086
2-23 -
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, 2132 087 e
5 2-24 2.6 Seismoloay Realistic predictions regarding earthquakes, or earth shocks, as well as their frequency and severity, can only be based upon seismic history of the area. Significant geological features, such as known slip-planes or faults, play an important role in the seismic history of any region. Thus these features as well as past shocks must be taken into consideration in depicting the seismology of the site.
The overall geolooical features of the Pullman area are described in the section on geology and will not be repeated here. In this section we are concerned with the geological features of this area that could possibly produce earthquakes. The significant faults within a circle of a 100 mile radius of the site are shown in Figure 2.6-1. From this drawina it is evi-dent that there are no known significant faults in the immediate vicinity of Pullman. The closest active fault is the Walla Walls fault, some 70 miles from Pullman. The closest inactive faults are the Vista and Wilma faults associated with the Lewiston Downwarp.
Historically, the seismic activity within 200 miles of the site is low, with infrequent earthquakes of low intensity and magnitude. The occurrences of earthquakes within 200 miles of Pullman are listed in Table 2.6-1. It is noteworthy that only two shocks have occurred at Pullman in recorded history, both of them of low intensity.
Based on the geology of the Pullman area and the past seismic activity, the probability of the occurrence of significant earthquakes in the future can be said to be very small.
2132 088
LOCAL N OF FAULTS i 5(LAKE COEUR D'ALENE SYMBOLS WITHIN A 100 MILE ,
RADIUS OF PULLMAN, FAULT WASHINGTON i ST. MARIES e g .f- ,
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2-26 TABLE 2.6-1 HISTORIC EARTHQUAKES 1872-1979 WITHIN 200 MILES OF SITE
- Anoroxima te Intensity location of Distance from at Year Date Epicenter Pullman Epicen'ter**
1872- Dec. 16- Walla Walla 70 miles Unknown 73 Jan. 1 1874 Unknown Ya kima 155 Unknown 1875 May 6 Yakima 155 Unknown 1875 May 7 Yakira 155 Severe 1887 April 29 Wall *a Walla 70 Fel t 1898 Feb. 22 Ellensburg 160 Fel t 1906 Jan. 2 NE Washington ---
Felt Over 200 Square Miles 1906 Nov. 2 Colville 130 V 1909 May 24 47. 6N , 120. 0W 140 Fel t 1911 July 5 Ellensburg 160 V 1915 Dec. 10 Sookane 65 III-IV 1918 Nov. 1 46. 7N , 119. 5W 105 V-VI 1920 Nov. 28- Spokane 65 Felt 29 1921 Sept.14 Walla Walla 70 V-VI 1922 Jan. 31 Republic 150 Fel t
- Data abstracted from " Washington State Earthquakes 1840 through 1965", " Seismic Trends in Washinaton State", and " Washington State Earthauakes Jan.1969 - June 1979", all by N. H. Rasmussen, Seismologist, Geology Department, University of Washin9 ton, Seattle, Washington.
- Modified Mercalli Intensity Scale of 1931.
2132 090
2-27 Approxira te Intensity _
Location of Distance from at Year Date Epicenter Pullman Epicenter **
1922 June 1 Spokane 65 IV 1922 Oc t. 16 Hermiston, Ore. 120 III 1924 Jan. 6 Walla Walla 70 IV 1924 May 27 Walla Walla 70 IV 1926 April 11 Walla Waila 70 III 1926 April 23 Walla Walla 70 IV 1930 Sept. 3 47.3N, lii.8W 70 V 1935 Oct. 24 Ellensburg 160 Fel t 1936 July 16 46.0N, ll8.3W 70 VII 1936 July 18- Walla Walla 70 Fel t 20 1936 July 30 Freewater, Ore. 80 VI 1936 July 30 Walla Walla 70 .III-VI 1936 Aug. 4 45.8N, ll8.6W 115 V 1936 Aug. 28 Walla Walla 70 IV 1936 Nov. 17 Walla Walla 70 II' 1937 Feb. 8 Walla Walla 70 III 1937 Feb. 9 Walla Walla 70 IV 1937 June 4 Walla Walla 70 IV 1937 June 17 Walla Wal1a 70 Felt 1937 Aug. 11 Spokane 65 Fel t 1937 Sept. 20 Walla Walla 70 Fel t 1938 May 9 Walla Walla 70 Felt 1938 May 24 Walla Walla 70 Fel t 1938 Aug. 11 Mil ton, Ore. 80 VI 1938 Oct. 27 Mil ton, Ore. 80 VI 2132 091
2-28 Anproxima te Intensity Location of Distance from at Year Da te Epicenter Pullman Epiceiiter**
1939 Feb. 6 Ellensburg 160 Fel t 1940 Jan. 6 Ephra ta 120 Fel t 1940 No v. 14 47. 7N, 121. 5W 165 III 1941 Jan. 3 Pullman 0 Fel t 1941 April 7 Republic 150 VI 1941 July 29 Spokane 65 Felt 1942 Nov. 1 48.0N, ll6.7W 85 VI 1943 April 24 47. 3N, 120. 6W 110 VI 1944 Sept. 2 Walla Walla 70 IV 1945 April 29 47.4N , 121. 7W 150 VII 1945 April 30 47.4N, 121. 7W 150 VI 1945 May 1 47.4N , 121. 7W 150 V 1945 Sept. 23 Walla Walla 70 IV 1949 Feb. 6 Wapato 155 III 1949 April 14 Pullman 0 Fel t 1950 June 25 Cheney 55 IV 1952 Mar. 4 Spokane 65 V 1952 July 27 47.8N, 121.9W 155 IV 1952 July 29 47.8N, 121. 9W 155 Fel t 1952 Nov. 10 47. 6N, 121. 5W 165 Fel t 1955 Feb. 6 Grand Coulee Dam 120 IV 1955 July 15 Soao Lake 120 IV 1955 Nov. 3 48. l N, 121. 7W 170 V 1956 Feb. 24 Electric City 120 V 1956 Nov. 18 48. l N, 121.8W 165 Fel t 2132 092
2-29 Approximate Intensity Locanon of Distance from at Year Da te E 61 center Pullman Epicenter **
19E7 Feb. 11 47. 5N , 121. 7W 150 VI 1957 Nov. 1 47. 0N, 121 W 185 V 1958 Apr. 12 48N,120W 150 VI 1958 Apr. 12 Electric City 120 IV 1959 Jan. 21 Walla Walla 70 IV 1959 Aug. 6 47.8N , 120. 0W 145 VI 1959 Nov. 23 46. 7N , 121. 7W 140 V 1961 May 22 47. 6N, 120. 2W 145 IV 1961 June 28 Rocky Reach Dam 145 IV 1961 Oct. 31 48.4N, 120W 170 V 1961 Nov. 7 Spokane 65 Fel t 1962 Ja n. 15 47. 8N , 120. 2W 155 VI 1963 Jan. 25 La Grande, Ore. 105 III 1963 Dec. 22 48. 3N , 119. 3W 130 V 1964 Oc t. 18 47.9N, 121.9W 155 IV 1966 Det. 24 47.9N, 121. 3W 155 III 1967 June 6 48. 2h , 119. l W 125 IV 1969 Oct. 9 46.8N , 121. 7W 140 VI 1969 Nov. 1 47. 9N , 121. 9W 165 V 1969 Nov. 10 48. 5N, 121.4W 190 V 1971 Oct. 25 46. 7N, 119. 6W 105 IV 2132 093
2-30 Approximate Intensi';
Location of Distance from at Year Date Enicenter Pullman Eoicenter**
1974 July 14 47. 6N, 120. 7W 177 IV 1975 June 28 46. 2N, 119. 7W 125 III 1975 Sept. 18 47.8N, ll8.2W 89 III 1975 Dec. 3 45. 6?l, 118.9W 113 III 1976 Apr.13 45. 24N, 120. 2W 174 IV 1976 May 15 47. 71 N, 120. 03W 1 51 III 1976 June 15 46.45N, ll7.68W 31 III 1976 June 15 47.63N,120.3W 160 III 1976 July 23 46.08N , 118.75W 87 III 1976 Aug. 30 47. 62N, 120.18W 1 54 III 1976 Dec. 13 47. 64N, 120.13W 153 III 1977 Jan. 27 46.94N,119.59W 115 III 1977 Ma r. 10 45.89N, 119. 68W 132 III 1977 Apr. 21 49.12N,117.67W 168 IV 1977 July 13 47. 06N , 120. 95W 179 IV 1978 June 27 46. 94N, 121.14W 188 III 1979 Ja n. 19 47.92N, 119.69W 144 IV 1979 April 8 46.0N, ll8.42W 78 IV 2132 094
.s*
2132055
3-1 3.0 FACILITY STRUCTURE 3.1 General Description The W.S.U. TRIGA reactor is located within the Nuclear Radiation Center which is pictured in Figure 3.1-1. The Center is a 1200 square me'.er lab-oratory devoted to nuclear related research and educational acti fities. The laboratory building is a concrete structure located east of the main portion of campus. All the utilities for the laboratory including heating, cooling, ventilation, and power distribution are contained within the structure. A set of floor plans for the Center is shown in Figures 3.1-2(a) to 3.1-2(d).
The reactor core is located in the reactor pool, the top of which is accessible via Pool Room 201 depicted in Figure 3.1-2(c). The pool room, control room, and reactor shop have a combined volume of 1 x 109 cm and 3
only two external exits on the east side. The reactor console is situated adjacent to the pool room in Room 201B and the reactor shop is located in Room 201A. A radiochemistry laboratory is located in Room 101 just belovi the control room and the beam room is located in Room 2 on the ground level below the radiochemistry laboratory. A list of rooms in the Center is given ir Table 3.1-1.
- 3. .? Heating ano Air Conditioning System The Center contains two independent heating and ventilation systems.
One system serves the laboratory areas exclusive of the reactor and the other serves the reactor. A third air handling system serves the fume hoods in the various laboratories. The hoods in all radiochemistry laboratories have ,
absolute filters.
The reactor ventilation system as shown in Figure 3.2-1 has provisions for operation in the nornal, isolation, and dilution modes. In the normal mode 7.08 x 10 5 cm3 /sec of externally treated air is discharged into the pool room. This loop is shut off in the isolation and dilution modes.
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TABLE 3.1-1 Nuclear Radiation Center Room Listing Rppm_ Description R.o_om De_ scrip _tj o n_ 002 Neutron Beam Room ll7A Faculty Office Source and Material Storage 119 Standards Laboratory 002A 003 Rest Room - Women 121 Sample Preparation Laboratory 004 Lounge - Women 122 Storage - BL Iding Maintenance 005 Rest Room - Men 123 Bay 006 Electrical Panel, Utilities 123A Gas Cylinder Storage 010 Boiler Room 124 Rest Room - Men B10 Transformer Vault 150 Counting Room B10A Transformer Vaul t 1 51 Conference Room 020 Storage 021 Radioactive Waste Storage 201 Reactor Pool Room B21 Neutron Genera tor 201AA Reactor Supervisor 050 Secretary, Receptionist 201A Reactor Shop 050A Director 20lB Reactor Control Console 050B Associa te Director 201C Hea t Excha. ,:r Dump Room 050V Entryway 210 Sample Preparation Laboratory 210A Effluent Sampler 101 Radiochemistry Laboratory 21 2 Geologic Materials Laboratory 101A Pool Trea tment Pump 214 Organic Chemistry Laboratory 105 Janitor Storage 21 5 Atomic Absorption Laboratory 106 Pool Water Trea tment 21 7 Facul ty Of fice 110 Radiochemistry Laboratory 217A Facul ty Of fice 112 Faculty Office 21 8 Electronics Shop 112A Passageway 218A Electronics Office 1128 Photo Darb room 220 Storage
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3-9 5 During normal operation, 9.44 x 10 cm3 /sec of air from the pool room is mixed with the beam room exhaust and discharges out the monitored exhaust. In the isolation mode, dampers on the pool room supply and exhaust lines serve to prevent all air flow into and out of the pool room. In the dilution mode,1.42 x 10 5cm 3/sec of pool room air is passed througii an absolute filter, mixed with 8.02 x 105 cm3 /sec of outside air, and dis-charges up the exhaust stack. The main control panel for the system is located in the reactor control room and a set of emergency controls is located in the main office. The continuous air monitor and gaseous efflu-ent monitors are shown in Figure 3.2-2 and 3.2-3. 3.3 Pool The reactor pool is a reinforced, above ground, unlined concrete pool with a volume of 247,000 liters. The pool is penetrated by a thermal col-umn and a nunber of beam ports as described in Section 4.11. A cross sec-tion of the pool is shown in Figure 3.3-1. 3.4 Liquid Wa' te Collection System The Nuclear Radiation Center has two separate waste systems. The sanitary waste system handles all the normal non-radioactive liquias and the " hot drain" system handles all the radioactive liquids. The sanitary waste system connects all the washroom fixtures and cold laboratory drains to the campus sewer system. The " hot drain" system connects all the drains from the radiochemistry laboratories and reactor areas to a reten-tion tank system. Radioactive effluents from the Center are collected in the retention tank system shown in Figure 3.4-1. Prior to discharge into the sanitary sewer, the contents of the retention tank are pumped to a sampling tank. The sampling tank is sampled, evaluated for activity con-tent, and diluted as necessary during discharge. 2132 108 s
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4-1 4.0 REACTOR DESCRIPTION 4.1 General Descrintion The WSU modified TRIGA reactor is a one megawatt pool-type research reactor using light-water as the moderator, coolant, and shield and TRIGA type solid fuel rods. The reactor core is inmersed in a large concrete water-filled open-topped pool . The pool is snanned by a manually-operated bridge structure from which the core sunnort structure is suspended. The core is situated in a grid box into which 4-rod clusters of TRIr-A fuel are positioned. Control over the reactor is exerted by insertina or withdrawing neutron absorbing control elements susoended from control drives mounted on the bridge. Heat generated by the fission process is transferred from the fuel to the pool water by natural convection cooling. The heat from the pool is dissipated to the atmosphere by means of a coolina tower-heat exchanger arrangement. A mixed bed demineralizer system maintains the purity of the pool water. 4.2 Bridge Structure The WSU reactor is suspended in the pool from a movable bridge which is mounted on rails. The bridge and entire reactor structure may be moved laterally. The bridge and core suspension framework are shown in figure 4.2-1. The all-aluminum framework is suspended from the bridge and supports the grid box into which the fuel is inserted. The hollow t.orner posts of the suspension framework serve as guide tubes for the nuclear instrumenta-tion detectors. The control element drives are connected to and supported by the bridge structure. Deck plates mounted on the top side of the bridge structure form a ficor area around the control drives. The floor area provides a work 2132 114
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M I:n SECTIONAL VIEWS OF TRIGA STANDARD 4-ROD CLUSTER S FIGURE 4.4-1
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.y FIGURE 4,4_.2 Photograph of Four-Rod Fuel Element 2132 120
4-8 STAINLESS STEEL TOP END FITTING A LT e a - a - - 3.45 IN.
- GRAPHITE bl '
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STAINLESS STEEL T,UBE n CLADDING ZlRCONIUM THICKNESS HYORIDE-0.02 IN. 8.5 WT-% ~30 IN. 15 IN. 1.37 IN. , _ _
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y BOTTOM END FITTING FIGURE 4.4-3 2132 12!
4-9 TABLE 4.4-1 Standard and FLIP Fuel Parameters Fuel Element Type FLIP STANDARD Fuel-moderator material U-ZrH l.6 U-ZrH l.7 Uranium content 8.5 wt% 8.5 wt% U-235 enrichment 70% 20% U-235 content (avg) Der element 123 9 35 g Burnable poison natural erbium none Erbium content 1.5 wt% -- Shape cylindrical cylindrical Length of fuel meat 15 in. 15 in. Diameter of fuel meat 1. 371 in. 1.371 in. Cladding material Type 304 SS Type 304 SS Cladding thickness 0.020 in. 0.020 in. length of 15 inches. A 3.45 inch graohite reflector plug is positioned in each end of the fuel rod and top and bottom end fittings are welded onto the claddina. I" addition to standard fuel rods, one or more instrumented fuel rods as shown in fioure 4.4-4 are used in the core. This type of fuel rod is fitted with three thermocouples used to measure the fuel temperature. A special 3-rod cluster with a transient rod quide tube as shown in figure 4.4-5 is positioned in the center of the reactor grid. The transient control rod is positioned inside the guide tube as described in section 4.6. 4.5 Control Blades The safety and regulatina control elements of the WSU reactor are blade type elements as shown in figures 4.5-1 and 4.5-2. The poison section of the safety blades is a boral sheet 40.5 inches long and 10.5 inches wide. The boral sheet is 3/8 inches thick and is clad with 1/8 2132 122
4-10
=. .
1.3 SOFT SOLDER [- ,. STAINLESS STEEL FILLER PLUG STAINLESS STEEL , i LEAD-0VT TUBE ~ ~ - ; ; (3/4 IN.DIA,) c d j- , - D $ SPACER - 5
-END PLUG GRAPHITE END REFLECTOR Z lRCONIUM ROD ;
FUEL-MODERATOR _ ' MATERIAL I.0 '! kM b f 26-7/8 IN. THERMOCOUPLES ( 3 ) -- 42-M-)a- .. 7 -- (CHROMEL-ALUMEL) .- 1 IN. 15 il -{-- CL IN. k ' bl.' 1 IN. d; W4 L. 0.02 IN. STAINLESS STEEL CLADDING g]y E-9 e b GRAPHITE END REFLECTOR fh Rs1 i Instrumented Fuel Rod FIGURE 4.4-4 iL_ _ 1.41 IN. . . 2132 1-23
4-11 h
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i f i 2132 124 Y FIGllRE 4.4-5 Transient Rod Guide Tube with 3-Rod Cluster
l 9 :
. . x
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i 3/8" BORAL SHEET WITH 1/8" ALUMINUM CLADDING g o 2: TOP OF ACTIVE n FUEL ELEMENTS j l 4 _s_ N C REACTOR CORE N {
+
--* s . LN N . i m BOTTOM OF ACTIVE FUEL ELDtENTS ( s Y , e 10.5"' NOM : a
.L Safety Blade FIGURE 4.5-1
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.I v 1 l .l Regulating Blade >
FIGURE 4.5-2 2132 126
4-14 inch aluminum. The regulating blade is a stainless steel sheet about 11 inches wide and 40 inches long. Each blade is guided through its travel by a shroud, as shown in figure 4.5-3. The shroud consists of two thin aluminum plates 38 inches high seoarated by aluminum spaces to provide a 3/4 inch control blade sloc. Small flow holes are drilled at the bottom of the shroud to reduce the effects of viscous damping on the blade fall time. 4.6 Transient Control Rod The transient control rod is a solid borated graphite cylinder con-tained in a 1-1/4 inch diameter stainless steel or aluminum tube as shown in figure 4.6-1. The poison section of the transient rod is 15 inches in length. The transient rod is connected to the transient rod drive via an end fitting welded on the top end of the rod. The rod is held in position laterally by the guide tube inserted into a 3-rod cluster. A hold-down tube extends from the top of the guide tube up to the bottom of the tran-ient rod drive, as shown in figure 4.7-3. 4.7 Control Element Drive The drive mechanism for the blade type control elements are shown in figure 4.7-1 and are' activated by reversible electric motors with an inte-gral worm-gear drive mechanism. The worm-gear assembly serves to reduce the drive speed and to minimize over-travel of the drive after power is removed from the drive motor. A mechanical slip clutch on the output shaft limits the force on the blade to approximately 75 pounds. A ball-bearing screw and nut system is used to raise and lower the control element. Each safety blade is coupled to its associated drive mechanism by means of an electromagnet and steel armature disk, as shown in figure 4.7-2. De-energizino the electromagnet allows the safety blade to fall into the 2132 127
4-15 L e -- 27 4'- - 4
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Photonpapy 0f yrunsient Controp pgg FIGUgg 4 g-I
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- # % i
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--MOUNTING FL ANCE Cll $( A WACNET LEADS D "
drive TUBC GUICK DISCONNECT PIN # SERVO CONTROL AND SCR A:A WACNET iq i 7 j START-UP COUNTER DRIVES ,,
~
CONTROL DRIVE DRIVE MECHANISMS
4-18 h CONTROL SHAFT [ 1 Armature Disc p ' Armature Rod k % - s. f ... UPPER GUIDE ASSEMBLY
*h L 0
Bushing 1, 1: Jam Nut I 'l Connecting Shaft
. Piston f
F 4 g?v;! k Lower Tube d - UPPER SUPPORT ASSEMBLY Y' GUIDE TUBE ASSEMBLY
/
-h ),
Dashpot Upper Bearing l Set Screw i l Guide Tube 4, (' ! i e y- ; e i [0 Oc , Lower Bearing ; Packing Nut Ry . ;
- \ ,
LOWER SUPPORT ASSEMBL [4 's.s k i e ' n y CONTROL BLADE
;C. #
-Lifting Hole 0,
- ll s
.y' i
CONTROL SHAFT, DASH POT,8 BEARING DETAILS FIGURE 4.7-2 2132 131
4-19 core by the action of gravity within 700 milliseconds. A shaft connects the armature disk to the blade and is fitted with polyethylene sleeve-bearings which control the lateral position of the blade drive shaft. A dashpot is positioned at the end of the shaft travel to decelerate the last 5 inches of fall . The blades are recovered after a scram by running the driv mechanism down and re-energizing the electromagnets. The transient control rod drive employs a combination pneumatic-electromechanical drive assembly shown in figures 4.7-3 and 4. The mech-anism is designed to allow the rod to be used both as a control rod and a transient rod. The pneumatic portion of the pneumatic-electromechanical drive, re-ferred to herein as the " transient" rod drive, is basically a single-acting pneumatic cylinder. A piston within the cylinder is attached to the transient rod by means of a connecting rod. The pistan rod passes through an air seal at the lower end of the cylinder. Compressed air is admitted at the lower end of the cylinder to drive the piston upward. As the piston rises, the air beina compressed above the piston is forced out through vents at the upper end of the cylinder. At the end of its stroke, the piston strikes the anvil of a shock absorber. The piston is thus decelerated at a controlled rate during its final inch of travel. This action minimizes rod vibration when the piston reaches its upper-limit stop. An accumulator tank mounted on the movable bridge stores the com-pressed air that operates the pneumatic portion of the trar.sient rod drive. A three-way solenoid valve, located in the piping between the accumulator tank and the cylinder, controls the air suoplied to the pneumatic cylinder. De-energizing the solenoid valve interrupts the 2132 132
4-20 8 SHOCK ABSORBER RETAINING NUT
, D s
. '" [ SHOCK ABSORBER
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SUPPORT s SHOCK AESORBER AIR VENT r SHOCK ABSORBER l l ANVIL 1 PISTON DR M TOR ' EAR G l i BALL NUT Gx p CYLINDER SCREW h i NBALL NUT y ,-- 7 1; h 8= , DRIVE GEAR kd' POTENTIOMETER DRIVE WORM
)
.L
! BALL NUT 's POTENTIOMETER DRIVE WORM DRIVE GEAR CYLINDER SCREW
; END CAP "O" RING SEAL
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AIR SUPPLY l ROD GUIDE N' 4 i i SUPPORT DN PISTON ROD ROD CONNECTOR
, DAMPER CONNECTING ROD
= Pneumatic-Electromechanical Transient-Rod Drive FIGURE 4.7-3 j}} }}}
- 4-21 SHOCK ABSORBER b
i VENT HOLES - EXTERNALLY THREADED CYLINDER PISTON SUPPORT r1 r1 . _ , L J L / BE ARINGS- '
/
[ h = WORM WORM GEAR
~
-= HOUSINO 3 CEi~
PISTON ROD = VENT
- SUPPLY AIR SUPPLY HOSE SOLENOID VALVE b
BOTTOM LIMIT = CONTROL ROD Schematic Drawing of Transient Rod Drive FIGURE 4.7-4 - 2132 134
4-22 air supply and relieves the pressure in the cylinder so that the oiston drons to its lower limit by gravity. With this operating feature, the transient rod is inserted in the core except when air is supplied to the cylinder. The electromechanical portion of the transient rod drive consists of an electric motor, a ball-nut drive assembly, and the externally threaded air cylinder. During electromechanical operation of the tran-sient rod, the threaded section of the air cylinder acts as a screw in the ball-nut drive assembly. These threads engage a series of balls contained in a ball-nut assembly in the drive housing. The ball-nut assembly is in turn connected through a worm-gear drive to an electric motor. The cylinder may be raised or lowered independently of the piston and control rod by means of the electric drive. Adjustment of the posi-tion of the cylinder controls the upper limit of piston travel, and hence controls the amount of reactivity inserted for a pulse. A system of limit switches is used to indicate the position of the air cylinder and the transient rod. Two of these switches, the Drive Up and Drive Down switches, are actuated by a small bar attached to the botton of the air cylinder. A third limit switch, the Rod Down switch, is actuated when the piston reaches its lower limit of travel. Durina steady state operation the transient rod may be withdrawn and used as a control rod by means of the ball-nut drive. 4.8 Control and Instrumentation 4.8.1 Control Console The Control System for the WSU TRIGA reactor consists of a Con-trol Console and associated instrumentation. The Control Console is pictured in figure 4.8-1 and was desianed and constructed by the 2132 135
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4-24 Nuclear Radiation Center Staff. Considerable thought and exnerience went into the human engineerino aspects of the console layout. All indicating devices are placed for ontimum readability and accessi-bility. All the alarm and interlock functions are in a single block of dual color back-liahted switches. The controls for the control elements are grouped together and located for ease of use. The con-sole is positioned to allow the operator to not only watch the con-sole but also to view the activities on the reactor bridge. A closed circuit TV system allows the operator to view the activity in the radiochemistry laboratory and beam rooms. The electronic systems use solid-state circuitry and high reli-ability components wherever possible. All essential relays and termi-nal boards are mounted on slide-out trays for easy maintenance. A permanent record of the important narameters is provided by strip-chart recorders mounted in a rack on the left side of the console. Ready access to comunications facilities are nrovided for operator convenience. The control system instrumentation contains electronic subsys-tems to perform the functions listed in table 4.8-1. A block dia-gram of the Linear Power indication ano safety subsystems is shown in figure 4.8-2. A block diagram of the pulse mode instrumentation is shown in figure 4.8-3. A block diagram of the scram circuitry and wide range channel are shown in figures 4.8-4 and 4.8-5. 4.9 Cooling System The heat aenerated within the fuel during operation of the reactor is transferred to the pool water by natural convection heat transfer. The heated cool water is pumned through the tube side of a conventional g 2132 137
4-25 TABLE 4.8-1 MINIMUM REACTOR SAFETY CHANNELS
^' Number Operable in Specified Mode Safety Channel Function S.S. Pul se Fuel Temperature Scram if fuel temperature 1 1 exceeds 500 C Power Level Scram if power level exceeds 1 125% of full licensed power Manual Scram 1 1 Wide Range a. prevent initiation of a l pulse above 2 kw
- b. prevent control element I withdrawal when neutron count is less than 2 cps High Voltage Monitor Scram on loss of high voltage to 1 1 power channels Pulse Mode Switch Prevent withdrawal of standard 1 control and regulating elements in pulse mode Preset Timer Transient rod scram 15 seconds or 1 less after pulse Pool Level Alarm if pool level falls below 1 1 16 feet over the core Transient Rod Control Prevent application of air unless 1 fully inserted 2132 l38
~
a
) R R O .
RR E T D A AE R I EW y C 0 l NO v O e C N I P E U n L N n R N a 1 h - C l
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n o , i !i ! C t I _ I a !
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[ i l j( d n I ll r a e E n . G i P HA L M GT O I L C l l O V 2 I 8 4 E R U G I F
!l I T
N E M I v 1l ' E j!; L E S T i ' fv RR S NRD OE T EI T E I AD P TII v i! li I AF N v S N RI I G NDC TRL CP A M AOf RRn v T R T AM WMI A Ill i OAU LRC SCR Y v SI T C E1 l e n n a F# A S O h J' C y t e f a I! NR S t OE E I B N G i l SM
- HA SA GGT !IlI' il i .'
I l OI L ,I [ FC LHO Il V N uN .ge
A) FUEL TEMPERATURE CHANNEL Instrumented s T.C. Temp. s Temperature Fuel Rod ' 4 ' XMTR Trip Readout
) Recorder _) So oid s, +
Logic Slow s Element Scram s Blade TAA ' Magnets
/
B) PULSE POWER CilANNEL Gamma s g , High Speed Chamber ' ' Recorder C) PULSE MODE CONTROL CllANNEL l Mode Interlocks
) Transient Position Rod Fire Readout SS Pulse ) ,
'#"""I #
External . Timed Motor Drive N Control u Blades y , s, Transient' Rod N Magnets ( Magnet Air , ) Air Drive ,h 3 CD Scram Air Solenoid 4 s Pulsing Mode Instrumentation e FIGURE 4.8-3
SCRAM CIRCUITRY 4-28 Scram Signal from upAmm #1 125% v min. , Scram Signal Scram Signal from C.I.C b q from Period Amp. H.V. Failure LOGIC ELEMENT T<5 sec. Power from Slow Master Sw. Scram Relay
- : : F
___V__V V__V VV
#1 #2 #3 Trip Trip Trip Actuator Actuator Actuator Ampli fier Amplifier Ampli fier V V y
#1 #2 #3 Transient Blade Blade Blade > Rod Air Magnet Magnet Magnet Solenoid s
(Safety , Cnannel (Period Open Open I) Open Amp) ~' Power from Open on on Open on Open Open Master Sw. on C.I.C. Log-N on High on on Seismic H.V. H.V. High Fuel Short Bldg. Detection Failure Failure Flux Temp. Period Evac. Manual
. l .
I 125% T<5 sec. min. q
, NOTE: Period Trip installed for training use only and a.ot requinad at all times .
J 2132 141 FIGURE 4.8-4 '=
s t
. HIGli VOLTAGE It.V. FAILURE I l SUPPLY MONITOR g 1
, I .
I v I v b I
, I LOG LW COUNT I I
POWER RATE NRIOD v I I SLOW SCRAM l I d CIRCUIT I I LOW COUNT PULSE PERIOD l 1 INHIBIT ,g gjT ALARM , t - v I PERIOD LOGIC
! SCRAM } ELEMENT A
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- I I PERIOD g. I I I_ _ _ ___ __ -- --- I FISSION RECORDER Wide Range Channel CllWBER a d> FIGURE 4.8-5
- 4-30 shell-and-tube heat exchanaer constructed with stainless-steel tubes. The primary pump has a stainless steel pump housing and rotor, and the primary piping is aluminum. A secondary side pump takes water from the sump of an induced draft type cooling tower and passes it through the shell side of the heat exchanger. The heat generated is dissipated to the atmosphere through the latent heat of vanoration of water in the cooling tower. A schematic diagram of the cooling system is shown in figure 4.9-1.
The physical layout of the cooling system is shown in figures 4.9-2 and 4.9-3. The system includes appropriate siphon breaks in the pool to prevent drainina of the pool in the event of a piping rupture. A thermo-statically controlled heater is installed in the cooling tower sump to prevent freezing during cold weather. Instrumentation for the cooling system in the control room allows the operator to control the system pumps as well as to monitor the conductivity, tenperature, and pressure at vari-ous points in the system. 4.10 Pool Make-up and Demineralizer System The pool water makeup and r.ool water deniineralizer system is shown in figure 4.10 l . Water is circulated from the surface of the pool, through a mixed bed demineralizer, and then discharged at the botton of the pool. This loop maintains the purity of the pool water and collects the majority of the radionuclides created in the pool water by activation. Make-up feed water to compensate for pool evaporation is suoplied by a Culligan demineralizer system. A float switch in the cool controls a solenoid valve in the make-up water line which automatically maintains the pool level . 4.11 Exoerimental Facilities - Exnerimental facilities are provided in the WSU TRIqA reactor. for 2132 143
NORMAL CORE POSITION w
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U PUM P N y S 0 Cooling System Schematic A FIGURE 4.9-1
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s NUC LE A R REACTOR LABORATORY M {s\'-O FIGURE 4.9-2
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POOL WATER PURIFICATION & t%KE-UP SYSTEM FIGURE 4.10-1
4-35 research purposes. These include a oneumatic transfer system, a thermal column, and numerous beam ports. The layout and dimension of the thermal column and beam norts is shown in fiaure 4.11-1. A schematic of the pneu-matic transfer system is shown in figure 4.11-2. The pncumatic transfer system discharges into the monitored facility ventilation system. The ventilation system also maintains a negative pressure on the thermal column cavity to insure that air flows into the thermal column cavity and up the monitored exhaust. 4.12 N-16 Diffuser A fraction of the oxygen present in the o001 water is activated to nitrogen-16 by a fast neutron (n.p) reaction as the water passes upward throuah the core. To reduce the dose rate on the bridge caused by the upward migration of N-16, a diffuser system consisting of a centrifugal pump and discharge nozzle was installed. This system pumps water from the surface of the pool and discharges it downward and across the top of the core. The net result is an increase in the transport time of N-16 to the surface and a significant reduction of the N-16 related dose rate at the bridae level. 2132 148
c. BEAM PORT SCHEDULE l2,@Mb T3 H-4 y LENGTH, INQ4ES { HEIGHT, INOES
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j . FRCN FR3A F RO*4
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BEAM IO" La id' to 8" 8" La CORE POC4. BEAN RN
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- PORT SECTION REDUCER TCTION EXTDSON { FLDOR FLOOR H-3
- H-l 31.50 15.5 0 21.75 5 5.13 -0.75 60.50 , 33.50 l '-
33.g \ 33 , H-2 31.50 15.5 0 21,75 56.00 8.25 6930', 42.50 H-3 31.50 t5.50 21.7 5 - - - - 8.25 6930- 42.50
\ g H4 31,50 5.50 21.7 5 54.63 -0.75 FROM 6050 FROM 33.50 FRCM
\ o (- BEAM 4" 1.D. 4" lo 2" 2" 1.D. CORE POOL BEAM ."*8 E-3 ;, ., , ,,
.,,.p ,. . PORT mrTION REDUCER SECTIOrJEXTENS!ON [ FLOOR F1.OCH l . . , , , . , . i e" ., E-1 3&40 4.50 37.00 55.38 - 173 55.50 28.25
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E-2 39.19 4.50 37.00 69.75 -1025 5 f .00 23.63
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- W E-3 39.19 4.50 37.0 0 70.50 -IO25 51.00 24.13 a.',- '
E-4 3840 4.50 37.00 54.50 - 5.75 55.50 2 8.13 T-1 38.40 4.50 3700 -2125 36.00 9DO .-- rr'( To watt. -* .
' T-2 3840 4.50 37.00 -2025 41.00 13.75
\ T-3 38.40 4.50 37.00 -I.525 36.00 925
... , ,. ] I T-4 3840 450 37.00 -2025 4fDO I3.87
' NOTES:
s coat mica AT 'o*2" 1. THE H-l EXTENSION IS 550" 1.D.. O.25" WALL THICKNESS.
, 2. THE H-2 EXTENSION IS 2 067"I.D,2.575"O.D.
~ ' 3. THE H-4 EXTENSION IS TAPERED WITH AN INNER DIMENS!CN TFERMAL COLUMN l
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'l CF 5.00" SO. AT THE CORE END 8 3.625" SO. AT 53.00" FROM THE EyD. THE SrDES ARE.O.50" PLATE AND THE
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END IS O.25 PLATE.
' '. 4. THE E-SERIES AND T-SER:ES EXTENSIONS ARE 2.067" I.0,,
2.375" O. D. S THE L OF H-18.H-4 IS 9.50" FROM
- THE THERMAL COLUMN
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POOL FACE. .
' [ 6. THE PONT WHERE THE th OF E-2 8 E-3 CROSS THE
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- 7. THE POINT WHEF:E THE CS OF H-2 8 'H-3 CROSS THE
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POOL L IS 17.00" FROM THE TFERMAL COLWN FACE. Y s-reesd , g . N,/ '.:'.'/..
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WASHINGTON STATE UNNERSITY
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FIGURE 4.11-1
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8; i i I e . o**m _M "A "JL lL. 2132 159
5-9 TABLE 5.1-2 Four-Rod Fuel Cluster Reactivity Worths Core Position Cluster No. Reactivity in Dollars C-6 F43 (FLIP) $ 2.80 E-6 F48 (FLIP) 2.13 E-3 S30 (Standard) 1.79 5.2 Steady State Operation Escalation to full power on mixed core 30-A was commenced on February 20,1976 using 100 kw steps. Routine full power operations on this core began on March 12, 1976. As of May 1,1979, a total of 2700 Mw-hrs of operation without a single significar.t problem has been accumulated on the W.S.U. TRIGA reactor fueled with a mixture of FLIP and Standard type fuels. The fuel temperature as measured by the instrumented FLIP fuel rod in core position F42 versus reactor power level is shown in Figure 5.2-1. The measured excess reactivity loss due to heating of the fuel versus po:ser level is shown in Figure 5.2-2. The measured excess reactivity of the core as a function of megawatt hours of operation is shown in Figure 5.2-3. The indicated fuel temperature and average core temperature versus excess reactivity is shown in Figure 5.2-4. The average core fuel temper-ature versus average temperature coefficient of reactivity is shown in Figure 5.2-5. No unusual or significant problems have been encountered during the three years that the W.S.U. TRIGA reactor has been operated in the steady state mode at a maximum power level of one megawatt. Future plans call for the replacement of one of the FLIP instrumented fuel rod.s due to the 2132 160
400'- -- - - - - FUEL TEMPERATURE VS. POWER LEVEL CORE 30-A - 350 - - - 300 - - - - - - - - - - - - 250 - - - - - - o .. . _ . . ,. ggg , W i a FIGURE. 5.2-1:.. ._ _ .. _ _ D *
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B - M a 0 8 6 1.0 - N ,a LeJ N 0 ' ' ' ' ' ' ' I i 8 200 400 600 800 1000 Ch N Power Level, Kw Y FIGURE 5.2-2 0
I 7.9 ' EXCESS RE ACTIVITY vs MEG AWATT- D AYS . CORE 30-A X Burnup = 5.011/ MWD 7.7 - t 7,5 - 7.3 - X 7.1 - N . 6.9 u N X
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5-13 8-Excess Reactivity vs. Indicated and Average Core Temperature FIGURE 5.2-4 7 6 - Indicated Fuel Temperature G <- T
- B M
e U a 5 - Average Core - Temperature) 4 - 2132 1 4 1 i i 1 100 200 300 375 Temperature (*C)
5-14 _ Average Temperature Coefficient vs. Average Core Temperature FIGURE 5.2-5 1.6 - 1.4 - m 'o 1.2 - x 7 1.0 - G o U U 3 .8 - 0 o i a E .6 -
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.2 2132 165 I I i 1 50 100 150 200 250 300 350 Average Fuel Temperature, C
5-15 failure of two of the thermocouples. This type of failure is common in TRIGA reactors and the only solution is replacement. An increase in the number of FLIP fuel rods in the core is also planned to lengthen the core life. The performance and safety considerations for mixed cores with more FLIP fuel than that of the current core was previously considered in the criainal S. A.R. for conversion to FLIP fuel. The information is contained in Apnendix A of this revised S.A.t. 5.3 Puls1ng Characteristics Pulsing tests for core 30-A began on February 23, 1976, starting with a $1.25 insertion and increasina un to the license limit of $2.50 in $.25 increments. The results of this initial series of nulses is shown in Table 5.3-1 below. The measured neak oower and enerov released in a TABLE 5.3-1 Core 30-A Pulse Testing Data p, $ No. of Pulses Avg. Peak Fuel Temp., C Avg. Peak Power, Mw 1.10 2 235.5 7.11 1.15 1 248 16.59 1.25 3 271. 7 60.0 1.50 6 309.3 155.3 1.75 4 333.8 304.8 2.00 3 367.3 622.3 2.25 3 403 1197.3 2.50 11 442.4 1850.5 pulse versus the reactivity insertion for core 30-A is shown in Figure 5.3-1. The peak indicated fuel temperature in core 30-A versus the reactivity insertion is shown in Figure 5.3-2. The measured pulsing , 2132 166
5-16 Peak Power and Energy Release vs. Reactivity Insertion 2000 - 1 1750 1500 T m 1250 - - 25 Q E o M m 3 1000 - - 20 - E % J. E'
- 750 - -
15 500 - o 10 250 - 5 O I I I t i 1.25 1.50 1.75 2.00 2.25 2.50. Reactivity Insertion ($) FIGURE 5.3-1
5-17 Peak Indicated Fuel Temp vs. p 600 - 550 500 - 450 - 4 400 "o d' 350 - 300 o 250 - C'
~
I I I i 200 1.25 1.50 1.75 2.00 2.25 2.50 Reactivity Insertion ($) }j}} )fg FIGURE 5.3-2
5-18 characteristics cf the W.S.U. TRIGA reactor fueled with a mixed core of Standard and FLIP fuels were as expected. No significant nulsino perfor-mance changes have been obserud to date. In September of 1976 a fue' p-hkm developed in the FLIP fuel region of a mixed core fueled TRIGA reactor similar to the W.S.U. reactor. The reactor involved was being pulsed with insertions up to $2.70. At the request of the Commission, the FLIP fuel in the central portion of the W.S.U. reactor was inspected but no damage was discovered. At the time of the inspection the W.S.U. reactor had been pulsed 55 times at a maximum insertion of $2.50. The current number of pulses on core 30-A is 141. A detailed cnmparison of the two reactors revealed the fact that the water gap between the transient rod guide tube and the adjacent fuel rods varied from .050 to .0204 in the reactor with the damaged fuel whereas it is a constant .0825 inches in the W.S.U. reactor. Thus the fuel damage problem is due in part to water-gap peaking effects not present in the W.S.U. reactor. Additional information is contained in the report of December 6, 1976 f'.ied with the Commission on this subject. 2132 169
6-1 6.0 SAFETY ANALYSIS 6.1 General Considerations The safety related aspects of the operation of the W.S.U. Modified TRIGA reactor will be considered in this section. First we will examine the fundamental characteristics and operational parameters of a TRIGA-type reactor as they relate to safety. Then we will examine the consequences of a number of postulated acci-dents. 6.2 Design Bases The primary design criteria for the Washington State University TRIGA reactor are determined by the maximum safe operational capa-bilitles of the solid-fuel moderator elements used to fuel the W.S.U. modified TRIGA reactor and the core configuration described in this report. The combination of fuel and core configuration must be selected to provide a high degree of operational safety independent of mechanical, electrical, or human errors. To attain this, the following characteristics and properties must be inherent in the reactor system.
- a. Large prompt negative temperature coefficient of sufficient magnitude to control the effects of a sudden large insertion of positive reactivity.
- b. Metallurgical properties of the fuel-moderator alloy that would insure integrity of the fuel-moderator alloy during either a sudden increase in temperature or prolonged periods of operation dt high temperatures.
2132 170
2132 171 6-2
- c. A suitable cladding that would contain the fuel-moderator material and associated fission products under all operating conditions. Cladding integrity must be maintained under the expected thermal and mechanical stresses and strains resulting from sudden increases in temperature and prolonged periods of operation at high temperature.
- d. A core configuration that is slightly undermoderated to pro-vide a negative void coefficient and to insure safety in case of a loss-of-water accident.
6.3 Design Limits A considerable amount of theoretical analysis has been performed and a large amount of operational experience has been accumulated on TRIGA-type reactors over the past decade. This accumulation of know-ledge and experience has led to the establishment of certain design limits for TRIGA-type reactors. These limits may be categorized as (1) shutdown margin limit, (2) reactivity addition rate limit, (3) fuel operating temperature limit, (4) operatina power limit, (5) reactivity addition limit during pulsing, and (6) Fuel dimensional variation limit.
- 6. 3.1 Shutdown Margin The aggregate worth of the control elements of a reactor must be set so that a safe shutdown margin is obtained with the highest worth control element fully withdrawn from the core. This requirement insures that the reactor remains sub-critical during core change with one control element with-drawn. The total control element worth necessary is obviously determined by the excess reactivity available and the worth of 2132 172
6-3 the individual control elements. The safe shutdown margin for TRIGA reactors has been set at 0.2% AK/K or about $.25. The highest worth control element in most TRIGA reactors is the pulse rod which generally has a maximum worth limit of
$4.00 as determined by pulsing considerations. Accordingly in normal operation a TRIGA reactor is shut down by $4.25 with all the control elements inserted.
6.3.2 Reactivity Addition Rate The reactivity addition rate to a reactor during normal operation is a function of the worth of the control elements, the speed of element withdrawal, and the number of elements being withdrawn at one time. In a TRIGA reactor the control system is generally designed to allow the withdrawal of only one element at a time. Thus the maximum reactivity addition rate is equal to the product of the maximum differential worth of the most reactive element in dollars per inch times the element speed in inches per second. In practice the maximum reactivity insertion rate is set at a level where an operator can retain control of power changes during steady-state operation. No limit is set on the transient rod during pulsing as the transient rod must be re-moved in a very short time to prevent clipping of the power transient resulting from the pulse. The normal reactivity insertion rate limit that has been set for TRIGA reactors is about 0.2% AK/K per second or $.25 per second. 2132 173
6-4 6.3.3 Fuel Operating Temperature The th mal limit for the fuel used in the W.S.U. TRIGA reactor is based on the combined characteristics of the fuel-moderator alloy and the associated cladding naterial. The limit depends upon the metallurgical properties of the fuel-moderator alloy, the pressure of the gases in the cladding gap, and the yield stress of the cladding (see page 63 of AppendixA). For TRIGA reactors using high hydride fuel it is a well established and documented fact that the limiting factor is the pressure buildup from out-gassing of hydrogen from the uranium-zirconium hydride fuel-moderator alloy. TRIGA fuel with a hydrogen-to-zirconium ratio of 1.6 as used in the W.S.U. TRIGA reactor is single phase for temperatures in excess of 1150 C. The fuel-moderator alloy actually melts at about 1800 C. Furthermore, the higher hydride fuels do not undergo any significant thermal diffusion of hydrogen. These two facts and the intensive testing of the fuel-moderator alloy by the manufacturer plus extensive in-core experience clearly demonstrate that the fuel-moderator alloy characteristics would allow safe operating temperatures up to at least 1150 C. The currently accepted limiting fuel temperatures for high hydride type TRIGA fuels are 1150 C for FLIP fuel and 1100 C for Standard fuel. It is customary to employ a 200 C safety margin to yield a limiting condition of operat_ ion of 950 C in FLIP fuel and 800 C in Standard fuel. 2132 174
6-5 The cladding material for the W.S.U. TRIGA reactor fuel is type 304 stainless steel with a thickness of 20 mils. It is a well known and documented fact (see Page 57 of Appendix A) that the tensile and yield strength of this cladding ma-terial is not significar.tly reduced up to a temperature of 850 C. The analysis contained on Page 55 to 65 of Appendix A of this S. A.R. establishes the fact that the ultimate strength of the 304 stainless steel cladding for TRIGA (H-Zr 1.6) fuel is 940 C. This cladding temperature is the limit-ing condition for a Loss of Coolant accident.
- 6. 3. 4 Operating Power The limitation on the maximum steady-state power level of a TRIGA-type reactor is determined by the ability of the cooling system to remove heat at a rate to assure that the fuel cladding tenperature is held well below the safety limit during normal steady state operation. A limitation on the maximum power level is also imposed by the decay heat of fission products if an accident occurs in which all or part of the cooling water is lost. Sufficient cooling must be pro-vided under this circumstance to insure integrity of the clad-ding. These considerations are analyzed in Appendix A of this section.
It is a well established fact that a TRIGA reactor can safely operate at a steady state power level of one megawatt with natural convection cooling if the pool coolinn system is designed to remove the heat oroduced and to limit the 2132 175
6-6 core cooling water outlet temperature to below 100 C. At power levels significantly above one megawatt, forced cool-ing systems are needed during steady state operation and an emergency spray cooling system for cooling in case of a loss of core cooling water. The W.S.U. TRIGA reactor utilizes natural convection cooling and is thus limited to a steady state power level of one megawatt. 6.3.5 Reactivity Addition During Pulsing During a transient, the limiting factor, as with steady state operation, is the fuel-moderator alloy temperature and the corresponding hydrogen pressure beyond which a cladding rupture may occur. Thus the limiting temperatures are 1150 C for FLIP fuel and 1000 C for Standard fuel for pulsing. Applying the 200 C safety margin, the limiting conditions for pulsing operation become 950 C in the FLIP fuel and 800 C in the Standard fuel. The peak fuel temperature in the core during a pulse is given by: TP(max) = (T(ave) + ATP) x P/A (6.3.5-1) where TP(max) = Peak Core Temperature,- C T(ave) = Average Core Temperature before pulsing, C ATP = Average Core Temperature increase during pulse, C P/A .a Peak to Average Temperature
,atio.
The average core temperature increase may be calculated using the Fuchs-Nordeim model equation given in Se'ction. 4.4 of Appendix A. Equation 6.3.5-1 and the Fuchs-Nordeim model 2132 176
6-7 equation contain a number of parameters that are very core speci fic. Thus a universal pulsing limit for all TRIGA reactors can not be set. The pulsing response of a typical 110 rod TRIGA reactor is tabulated in Table 6.3.5-1 for the specified parameters. Detailed calculations for the W.S.U. TRIGA reactor as fueled with a mixture of Standard and FLIP fuels are given in Appendix A. This analysis established a conservative pulsing limit of $2.50 for the W.S.U. modified TRIGA reactor. In addition, the S. A.R. for the conversion of the Texas A&M TRIGA reactor to TRIGA fuel established the fact that in a mixed core during pulsing for a given reactivity .nsertion, the average increase in core temperature decreases as the fraction of FLIP in the core , creases. TABLE 6.3.5-1 Typical 110-Rod TRIGA Core Pulsing Response Reacti vity Temp. Coeff. Li fetime
$ -$/ C Microseconds AT, C 1.5 .012 20 81 2.0 .012 20 160 2.5 .012 20 236 1.5 .014 34 70 2.0 .014 34 138 2.5 .014 34 203 2132 177
6-8 6.3.6 TRIGA Fuel Rod Inspection The rapid increase in power and the resultant increase in fuel temperature in a TRIGA reactor during pulsing subjects the fuel rod cladding to stress and to thermal cycling effects. In order to insure that the fuel rod cladding integrity has not been significantly deteriorated, it is customary to inspect the fuel rods periodically. This inspection at some specified in-terval of time involves checking the transverse bending and elongation of TRIGA fuel rods. In order to inspect the fuel rods, they must be removed from the reactor core and placed in a ji; or fixture in the reactor pool. This operation involves a considerable amount of manipulation of the fuel rods using underwater handling tools. During the manipulation there is a possibility that nhysical damaae to the fuel rod may result from mishandling. A few fuel rods have even been dropped during such operations at some facilities. While it is important' that adequate in-spection frecuency be maintained to quard aaainst nossible nulsing induced damage, it is also important to minimize the number of inspections in order to reduce the possibility of physical damage to the fuel rods. The strain produced during the pulsing transient results from the stress of internal pressure in the heated rod and the differential expansion of the fuel-moderator rod and cladding. The increased pressure results from the increased temperature of the air in the cladding gap, the fission products released from the fuel, and the hydrogen released from the partial dis-associat ion of 'he zirconium hydride. Actual measurements 2132 178
6-9 made by General Atomic on specially instrumented fuel rods during pulsing reveal equilibrium pulsing pressure increase to be only about 20 psia (22). A .25-inch gap is provided in a TRIGA fuel rod between the lateral end of the graphite reflector and the end piece welded onto the cladding. This gap reduces to about half this value (23) at a fuel rod temperature of 1200 C and a cladding temperature of 200 C. Because of this gap, differ-ential expansion during pulsing will not produce a signifi-cant amount of lateral strain on the fuel rod cladding. The predominant and most significant effect that pulsing has on the cladding is that of radial differential expansion in new, unpulsed rods. Near the middle region of the fuel rod the uranium-zirconium hydride is in close contact with the stainless steel cladding. The effects of differential expansion between the fuel rod and the cladding is greatest in the middle region. Assuming a tenperature of 1200 C for the fuel-moderator rods and 200 C for the cladding, the amount of strain is equal to the fractional increase in the cladding circumference due to the fuel rod expansion and is calculated as follows:
- 1. Change in circumference of cladding due to increase in temperature above nominal 25 C ACc = Circumference x Linear Coefficient of Exgansion x Temperature Change = 1.41 x n x 17 x 10- (200 - 25)
= .0132 in.
2132 179
6-10
- 2. Increased area of fuel rod due to increase in temperature above nominal 25 C At = Area (1 + 2 x Linear Coefficient of Expansion x Temperature Change)
At " ( '2 ) x n[(1 + x 10h x (1200 - 25)] At = 1.452672 x 1.01645 = 1.476569 in.2
- 3. Increase in circumference of fuel-moderator rod A
ACr " "(D r -D)" c
"(I r -I)*
c "I ) - 2 AC =2 (.68577 .6800) = .0350 in. r
- 4. Difference between circumference of heated cladding and heated fuel rod = strain AC - C = .0350 r c
.013? = .0218 in.
- 5. Radial strain on cladding due to differential radial thermal expansion Circum-Strain = Fractional Deformation = Increase in Cla_otng ference Cladding Ci. imference Strain = .0218/1.41n = .0049 or about .5% strain.
The above amount of strain would cause some per-manent deformation in the claddii.g but is well within the safety limit for the expansion of type 304 stainless steel. The fact that some permanent deformation is produced is substantiated by the fact that the heat transfer between the fuel-moderator rod and the cladding of a TRIGA fuel rod decreases in new fuel after pulsing. In other words, the time required for new instrumented TRIGA fuel rods to cool down from a given temperature to ambient increases 2132 180
6-11 after pulsing. This fa.: tor is indicative of an increase e in the space between tha fuel rod and the cladding due to permanent stretching of the cladding by pulsing. The amount of strain produced during extended puls-ing will in actual fact be significantly lower than that calculated above. The permanent deformation of the clad-ding will obviously reduce the value of the strain. The actual fuel rod temperature will be below that assumed in the calculation and will not exceed 750 C. Furthermore, during pulsing, film boiling causes an increase in the fuel cladding tempera'uh above the steady-state case. All of these factors vill decrease the magnitude of the strain so that the salue calculated above is very conser-vative. Studies made for N. A.S. A. (24) on low-cycle fatigue indicate that the cladding could receive over 7,000 cycles of the postulated strain previous to the appear-ance of cracks that could allow fission product gases to escape. This value was obtained by the following con-sideration N = (K/S p
)2 = (.69/.0049)2 = 79,307 where N = number of cycles previous to fracture Sp = the plastic strain (less than one-half of the total strain on the cladding)
K = a measured constant for a given material which is related to the fracture directly (for type 304 stainless steel, K has bee,n measured to be .69). 2132 181
6-12 Cracks may start to appear at about 10% of the cycles at which fracture occurs. Applying the 10% factor to the above calculation, a conservative estimate of 7,000 cycles of pulsing will occur before the onset of possible crack-ing that could cause a fission product leak. The pulsing limit for the W.S.U. TRIGA reactor fueled with a mixed core is $2.50. If the core was pulsed 7,000 times with the maximum allowable pulse, this limit would amount to a total of $17,500 worth of pulses. As a pru-dent rule one could set an inspection frequency at 20% of the expected pulsing life or a total of $3,500 worth of
$2.50 pulses. This inspection frequency limit is estab-lished for the W.S.U. TRIGA reactor.
In actual operation a variety of sizes of pulses up to the maximum allowable value are shot. Only those pulses above $1.50 will produce a fuel temperature higher than thai, attained during normal steady-state operation. Thus in making the tabulation for fuel rod inspection, only those pulses over $1.50 should be summed for inspec-tion purposes. 6.4 Accident Analysis 6.4.1 Mixed Core Operation A detailed analysis of the safety aspects of the W.S.U. Modified TRIGA reactor fueled with a mixture of Standard and FLIP fuels is given in Appendix A. This document establishes the following limiting conditions of operation: - 2132 182
6-13 (1) Maximum FLIP fuel temperature of 950 C (2) Maximum Standard fuel temperature of 800 C (3) Maximum reactivity insertion of $2.50 (4) LSSS of 500 C for fuel temperature scram (5) Maximum allowable power density of 23.5 Kw/ rod for FLIP and 22.3 Kw/ rod for Standard fuels. The analysis also demonstrates that no realistic hazard to the general public would result from the Design Base Accident, a Loss of Coolant Accident, the accidental addition of one 4-rod cluster, or the accidental ejection of the transient rod at full power. 6.4.2 Argon-41 Releases Section 6.5 of the S.A.R. for the conversion of the W.S.U. TRIGA reactor to FLIP fuel as given in Appendix A of this report substantiates a 3 x 10-4 dilution factor for Argon-41 release due to the atmospheric wake effect in the lee of the building. A more thorough analysis of the distribution of Ar-41 in the atmosphere about the site may be obtained by the use of equation F-1 of Appen-dix F of Regulatory Guide 1.109, " Calculation Reactor Effluents for the Purposes of Compliance with 10 CFR Part 50, Appendix I." The annual release of Ar-41 from the W.S.U. facility for the past five years has averaged 10 Ci/ year. Thus the daily release is .027 Ci/ day which is equivalent to a release rate of 3.17 x 10 -7 Ci/sec. However, for the purposes of our calculations we shall assume a' 100 Ci/ year 2132 183
6-14
-6 Ci/sec release rate. Using a release and a 3.17 x 10 100 Ci/ year total release, the wind distribution data of Figure 2.4-2, and Equation (F-1), the Ar-41 concen-tration in the atmosphere about the site may be calcu-lated. The results are shown in Figure 6.4-1 in terms
-8 3 Ci/cm ,
of the % of the 10 CFR 20 limit of 4 x 10 The ground level Ar-41 concentration levels about the site for a 1000% normal release rate are significantly below the 10 CFR 20 limit as well as the ALARA criteria of 2% of the 10 CFR 20 limit. The closest occupied location to the site is 411 meters west and thus would be exposed to a 2.2 x 10
-II pCi/cm3 annual average Ar-41 concentration for the postulated release. Accordingly the Ar-41 released to the atmosphere by the operation of the W.S.U. TRIGA reactor does not endanger the health and safety of the general public.
2132 I84
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7.0 REFERENCES
- 1) E. G. Holzman, " Safeguard Renort for Open Pool Reactor for State College of Washington," General Electric Company Report GEAP-3100, February 1959.
- 2) H. Stern and H. W. Dodgen, " Safety Analysis for the Washinaton State University Reactor Core Conversion and Power Increase," Washington State University, October 1966.
- 3) General Atomic Staff, "TRIGA MARK III Reactor Hazards Analysis," General Atomic Report GA-3886, February 1965.
- 4) Gulf General Atomic Staff, " Safety Analysis Report for the Torrey Pines TRIGA Mark III Reactor," Gulf General Atomic Report GA-9064, January 1970.
- 5) J. M. Batch, " Standard TRIGA 4-Rod Cluster Conversion Reactor Maintenance and Operating Manual for Washington State University," General Atomic Report GA-7677, February 1967.
- 6) R. J. Cashwell, " Safety Analysis Report for the University of Wisconsin Nuclear Reactor," University of Wisconsin, April 1973.
- 7) D. A. Swanson, " Geologic Map of the Columbia River Basalt in the Pullman and Walla Walla Quadrangles of Southwest Washington," U.S.G.S. Open File Report 77-100, 1977.
- 8) U.S.G.S. Staff, "The Channeled Scablands of Eastern Washington," U.S.G.S.
Information booklet IF-72-2(R-1),1972.
- 9) Bates McKee, Cascadia: The Geologic Evolution of the Pacific Northwest, McGraw-Hill Book Company, New York,1972.
- 10) N. Rasmussen, " Washington State Earthquakes 1840 through 1965," Bull .
Seism. Soc. Am., 5_77 (1967), pp. 463-476.
- 11) N. Rasmussen, Unpublished Additions to Washington State Earthquake List, June 1965 to 1979.
- 12) R. C. Newcomb, " Tectonic Structure of the Main Part of Basalt of the Columbia River Group, Washington, Oregon and Idaho," Miscellaneous Inves-tigations Map I-587,1:500,000, U.S. Geologi: 1 Survey,1970.
- 13) J. W. Crosly and R. M. Chatters, " Water Dating Techniques as Applied to the Pullman-Moscow Ground-Water Basin," W.S.U. Bulletin 296, 1965.
- 14) W. A. Goodwin and M. E. Wyman, "The Measurement of Radial Power Distri-butions in TRIGA Fuel Elements During Reactor Power Excursions," Nuc.
App. & Tech. , V.18, March 1970.
- 15) W. E. Wilson and T. A. Lovas, " Washington State University Conversion to Mixed Core and Test Program," Report to N.R.C., April 1976.
2132 187
7.0 REFERENCES
(Cont.)
- 16) V. Ichimura, " Uranium Concentrations in Ground Waters of the Pullman-Moscow Basins," W.S.U. Dept. of Geology M.S. Thesis,1979.
- 17) D. E. Feltz, " Amendment II to the Safety Analysis Report for the Texas A & M TRIGA Reactor," Texas A & M, Nov. 1972.
- 18) J. D. Randell," Status Report on Damage to FLIP fuel During Oneration of the NSCR at Texas A & M University," Report to the N.R.C., Nov.1976.
- 19) W. E. Wilson, "Results of Inspection of W.S.U. TRIGA FLIP Fuel in Light of Texas A & M Fuel Damage," Report to N.R.C, Dec.1976.
- 20) Gulf Atomic Staff, " Summary of TRIGA Fuel Fission Product Release Experi-ments," Gul f Atomic Report Gul f-EES-A10801, Sept.1971.
- 21) W. E. Wilson, " Amendment I to Safety Analysis Report of October 1966 for the W.S.U. TRIGA Reactor," Report to the N.R.C. , May 1974.
- 22) General Atomic Staff, " Safety Analysis Report for the Romania Annular Core Pulsing Reactor," General Atomic Report E-117-323, Vol. III.
- 23) G. Beck, " Safety Analysis Report for the Illinois Advanced TRIGA Reactor,"
University of Illinois, August 1967.
- 24) R. W. Smith, " Fatigue Behavior of Materials Under Strain Cycling in Low and Intermediate Life Range," NASA TN 0-1574, April 1963.
2132 188
O b 2132 189
WASHINGTON STATE UNIVERSITY NUCLEAR RADIATION CENTER Pullman, Washington 99164 SAFETY ANALYSIS FOR CONVERSION TO FLIP FUEL May 1979 2132 190
TABLE OF CONTENTS Page
1.0 INTRODUCTION
. . . . . . . . . . . . . . . . . . . . . . . . . . I 2.0 CONTENTS OF REPORT . . . . ................... 1 3.0 FUEL DESCRIPTION AND SAFETY LIMITS . . . . . . . . . . . . . . 2 4.0 CALCULATIONAL METHODS ..................... 4 4.1 Core Calculations . . . . . . . . . . . . . . . . . . . . . 4 4.2 Fuel Element Temperature ................. 5 4.3 Prompt Negative Temperature Coefficient . ......... 9 4.4 Pulsing Calculations ................... 11 5.0 GENERAL SAFETY ANALYSIS .................... 13 5.1 Fuel Rod Temperature and Power Density .......... 13 5.2 Operational Mixed Cores . . . . . . . . . . . . . . . . . . 20 5.3 Pulsing Characteristics of Operational Core . . . . . . . . 20 5.4 Fu21 Temperature Scram .................. 40 6.0 DESIGN BASIS ACCIDENT ..................... 42 6.1 Whole Body Dose in Pool Room ............... 45 6.2 Lung Dose in Pool Room .................. 45 6.3 Thyroid Dose i n Pool Room . . . . . . . . . . . . . . . . . 46 6.4 Discharge of the Fission Products into the Environment . 48 6.5 Dilution of Discharge in the Lee of the Building ..... 51 6.6 Whole Body Dose Outside Facility ............. 52 6.7 Thyroid Dose Outside the Facility ............ 52 6.8 Summary of Results of D.B.A. ............... 53 7.0 REACTIVITY EFFECTS OF ACCIDENTAL FUEL ADDITION . . . . . . . . . 54 2132 191
TABLE OF CONTENTS (continued) Page 8.0 REACTIVITY EFFECT OF TRANSIENT R0D EJECTION . . . . . . . . . . 54 9.0 LOSS OF COOLANT ACCIDENT ................... 55 10.0 FLIP FUEL LOADING AND MIXED CORE PERFORMANCE TESTS ...... 59 APPENDIX A - Fuel Rod Temperature Calculations ........ 60 APPENDIX B - Temperature Coefficient Weighting Factor for a Mixed Core . . . . . . . . . . . . . . . . . . 61 APPENDIX C - Calculation of Temperature Coefficient of Core 30E at Full Power ............... 62 APPENDIX D - Fuel Temperature Limitation for TRIGA FLIP Fuel. . 63 References .......................... 66 2132 192
1.0 INTRODUCTION
This report considers the safety aspects of the operation of the W.S.U. modified TRIGA reactor with cores containing a mixture of standard and FLIP fuel s. This document was orginally submitted in May of 1974 to the Commis-sion as " Amendment I to the Safety Analysis Report of October 1966." The S.A.R. of 1966 is now replaced and superseded by the S.A.R. of May 1979 sub-mitted to the Commission in application for the renewal of facility license R-76. This report now becomes " Appendix A" to the S. A.R. of May 1979 and constitutes the Safety Analysis for the use of FLIP fuel in the W.S.U. reactor. 2.0 CONTENT OF REPORT This report only considers FLIP fuel related safety aspects of the opera-tion of the W.S.U. TRIGA reactor. All other items are covered in the S.A.R. to which this report is an appendix. The specific items considered are:
- 1. Fuel Description and Safety Limits
- 2. Calculation Methods
- 3. General Safety Analysis
- 4. Design Base Accident
- 5. Reactivity Effect of Accidental Fuel Addition -
- 6. Reactivity Effects of Transient Rod Ejection (Pulsing Limits)
- 7. Loss of Pool Water Accident Each of these items will be considered in detail in the following sections of this report.
3.0 FUEL DESCRIPTION AND SAFETY LIMITS The fuel in the mixed core will be both standard TRIGA and FLIP fuel-el ements. The two types of elements are identical in construction and differs 2132 193
2 only in U-235 enrichment, burnable poison content, and hydronen-to-zirconium ratio. The dimensional and physical description data on the fuel is contained in the revised SAR of May 1979. It is possible to visually distinguish the element types, however, by the markings on the upper tip of the FLIP fuel. Table I lists the principal design parameters of both FLIP and standard TRIGA elements. The safety limitations on the fuel are those limiting values imposed to preclude a loss of fuel element integrity. During a reactivity excursion the limiting condition is fuel temperature and the corresponding hydrogen overpressures at which clad rupture may occur. Studies show that in FLIP fuel the hydroqen pressure which would result from a transient for which the peak fuel temperature is 1150 C* would not produce a stress in the clad in excess of the ultimate strength.(l) TRIGA fuel with a hydrogen-to-zirconium ratio of at least 1.65 has been pulsed to temperatures of about 1150 without any damage to the clad.(2) As a safety limit, the peak adiabatic fuel tem-perature to be allowed during transient conditions is set at 1150 C for FLIP fuel. Since standard TRIGA fuel nominally contains more hydrogen than FLIP fuel, its corresponding safety limit is reduced to 1000 C. For steady state operation (non-adiabatic case) fuel temperatures are dependent upon the heat transfer characteristics of the element and coolant; thus, an experimental limit on power density is selected to insure fuel integrity. This limit is well below the maximum allowable power density which corresponds to a heat flux value at which there is a departure from nucleat boilii.g. The maximum stady-state power density generated in the Torrey Pines TRIGA fiARK III is 32 kw per element. (I' ) Since the WSU TRIGA pool depth is somewhat deeper than the Mark III, improved cooling characteristics are ex-pected. Thus, for steady state operation at power densities of up to 32 kw
- See Appendix D 2132 194
3 TABLE I STANDARD AND FLIP FUEL PARAMETERS Fuel Element Type FLIP STANDARD Fuel-moderator material U-ZrH l.6 U- rH l.7 Uraniuni content 8.5 wt% 8.5 wt% U-235 enrichment 70% 20% U-235 content (avg) per element 123 9 35 g Burnable poison natural erbium none Erbium content 1.5 wt% -- Shane cylindrical cylindrical Length of ruel meat 15 in. 15 in. Diameter of fuel meat 1.371 in. 1.371 in. Cladding material Type 304 SS Type 304 SS Cladding thickness 0.020 in. 0.020 in. I 2132 195
4 per element, no cooling problems are expected. As will be seen later on, the loss of coolant accident imposes a more restrictive limitation on the maximum safe allowable power density per element. 4.0 ALCULATIONAL METHODS 4.1 Core Calculations The calculations of the characteristics of the WSU TRIGA reactor core with standard fuel, and mixtures of FLIP and standard fuel were performed using the EXTERMINATOR-2 codeI4) and temperature dependent cross section data obtained from Gulf General Atomic. The fast cross section data were generated with the GGC-4 code (5) and the thennal cross section data using the SUMMIT and THERMIDOR codes.(6,7) Seven energy groups were used in the core calculations as well as group dependent buckling for the FLIP fuel as listed below. This group structure and energy dependent buckling is identical to that conventionally used by Gulf General Atomic in their TRIGA s ector calculations including the Puerto Rico FLIP fueled reactor calculations. TABLE II Energy Groups and Group Dependent Buckling Use in WSU TRIGA Core Calculations Standard Fuel Group Buckling FLIP Fuel Buckling 1 15.0 .694 MeV 0.0041 0.00546 2 639 - 9.12 kev " 0.00435 3 9.12 - 0.001125 kev " 0.00347 4 1.125 - 0.414 eV " 0.000324 , 5 .414 .14 eV 0.00469 6 .14 .05'eV -0.0146 7 .05 .0002 eV -0.0614 2132 196
5 The results of calculations made with the EXTERMINATOR-2 on the existing core of the WSU TRIGA reactor compare favorably with measured values. Furthertiore, calculations on the Puerto Rico FLIP core at WSU with the code yield results comparable to those obtained by Gulf General Atomic. Thus the calculations performed on all standard and all FLIP fueled cores are known to be accurate and reliable. Conse-quencly, the results obtained on mixed cores can be expected to be reasonal ,y accurate and reliable. Furthermore, Texas A&M has had good success in using this code for calculating the characteristics of mixed cores. 4.2 Fuel Element Temperature The direct theoretical calculation of accurate fuel element tem-peratures for steady-state operation in a TRIGA core using natural convection cooling is very difficult. This is due to the fact that the contact coefficient between the fuel and the ciudding is not known accurately, especisily in fuel rods that have been pulsed. In other words, the fuel-cladding thermal contact coefficient is a function of the fuel temperature and the pulsing history and age of the fuel. In addition, the film coe'ficient of heat transfer between the cladding and the coolant under conditions of natural convection is uncertain. The film coefficient is a function of the coolant temperature, coolant velocity, and effective hydraulic diameter of the coolant channel. In order to circumvent these uncertainties, we have chosen to utilize experimentally measured fuel temperature data obtained from the instrumented fuel rod coupled with calculated power distribution 2132 197
6 data obtained with the EXTERMINATOR-2 code. A graph of the maximum fuel temperature of the instrumented fuel rod as a function of reactor power in WSU TRIGA core No. 28A during steady-state operation is shown in Figure 1. Calculations with the EXTERMINATOR-2 code indicate that the axial average power density in the instrumented fuel rod is .283 watts per kw of reactor power per cm of core height. Combining this result with the experimental power-temperature data yields a relationship be-tween axial average fuel rod power density and fuel rod temperature. This relationship as detennined by least squares fitting of the data to a polynominal is given in Appendix A. The fuel temperatures calculated using the derived equation are shown in terms of power density in Figure 2. The maximum observed power density used in deriving the fuel temperature equation was 12 kw/ rod and thus a linear extrapolation is used above this value. A comparison of the curve in Figure 2 with experimental measurements at Texas A&M and PRNC which are also displayed on the graph substantiate the validity of the relationship. That is, fuel rod temperatures calculated by the WSU equation are essentially identical or more conservative than measured values. Due to the large number of calculations involved and the numoer of core configuraticns studied, a special program, PLDEQC , was written to calculate and display the neutron flux, power, and temperature dis-tributions. This program utilized the output group flux and power generation matrixes from EXTERMINATOR-2 to make these calculations. The temperature and power density data given in Section 5.0 of this report were generated by this method. 2132 198
7 Figure 1 1000_ 800 - 600 _ 3 e - L B U Z 400 , i 200 , 2132 199 0 , , , , , . 0 100 200 300 Instrumented Fuel Element Temperature, C INSTRUMENTED FUEL R0D TEMPERATURE AS A FUNCTION OF REACTOR POWER
8 Figure 2 700_ 600 _ a
/
/
Linear Extrapolation- / 500 _ ,'
/
E / a - % / 5 ' + R 400 _ s' E x E o ) 300 _ ,$ +WSU Calculations 5 x 200 _ o Texas A & M Measurement
+ PRNC Measurement 100 _
2132 200 0 e i i t i i 0 5 10 15 20 25 30 Power Density, KW/ Rod STEADY STATE FUEL RCD TEMPERATURE AS A FUNCTION OF POWER GENERATION
9 4.3 Prompt Negative Temperature Coefficient Calculations of the prompt temperature coefficient of the WSU TRIGA reactor for all standard fueled cores were performed using the EXTERMINATOR-2 code and temperature dependent cross sections. The data thus obtained is plotted on the graph in Figure 3. The results thus obtained locally are essentially identical to those reported by Gulf General Atomic for similar type fuels. Experimental measurements of the overall temperature coefficient of WSU TRIGA core No. 28A at full power with all standard fuel yield a result of -$.014 per degree centigrade. The average fuel temperature for this core at 1 mw is 194*C which would give a theoretical value of -$.0147/*C for the temperature coefficient. Since it is a known fact that a TRIGA reactor has a slightly positive bath coefficient, the calculated and measured values agree as well as can be extected. The direct calculation of a meaningful temperature coefficient for a mixed standard-FLIP fueled core with the EXTERMINATOR-2 code is not possible with the limited temperature dependent cross section data that was readily available. An alternate technique was adopted involving the suming of the appropriately weighted temperature coefficients of the standard and FLIP fueled regions of a mixed core (see Appendix B). Accordingly, the mixed core prompt negative temperature coefficient is given by: Tcm = Tcf(Tf) x h + Tcs(Ts) x h where Tcm = Temperature coefficient of mixed core at mean core temperature Tcf(Tf) = Temperature coefficient of all FLIP core at temperature Tf , 2132 201
10 Figure 3
.026
.024 -
.022 -
u .020 - o 3 _ J
.018 .f c % WSU TRIGA O .016 -
All Std g 3320 MWib B 2 E .014 -
+PRNC 5 WSU TRIGA cold clean all FLIP
$ cold clean-
'5 .012 _
+PRNC (3000 MWD)
E g .010 - 1
.008 -
m
.006 _
2132 202
.004 e i e i r i 0 100 200 300 400 500 600 -700 Average Fuel Temperature in Core, C CALCULATED PROMPT NEGATIVE TEMPERATURE C0 EFFICIENT
11 Tcs(Ts) = Temperature coefficient of all-FLIP core at temperature Ts Tf = Mean temperature of FLIP fuel region Ts = Mean temperature of standard fuel region FF = Fissions in FLIP fuel region FS = Fissions in standard fuel region FT = Fissions in total core. 4.4. Pulsing Calculations The peak power of a TRIGA reactor during a transient can be accurately described using a Fuchs-Nordheim model with variable heat capacity. Additional analysis of transient fuel rod temperatures by the group at the University of Illinois has shown that the values cal-culated by the model are conservative (high).(10) According to this model the peak reactor power during the transient, P(max), is given by P(max) = Po + C(p - 1)2 1 + 3o at 6a where C = (Co + y To)N = Heat capacity of core at initiation of pulse N = Number of rods in core o = aC/ N(p - 1) Co = Heat capacity of TRIGA fuel at 25*C = 769 wat-sec/*C/ rod To = Average initial temperature of the core above 25*C t = Prompt neutron lifetime of core a = Prompt negative temperature coefficient y = Rate of change in heat capacity of TRIGA fuel = 1.47 wat-sec/*C/ rod 2}32 2 p = Reactivity inserted.
12 In addition, the average peak core temperature, T , is given by T= = b - I ) f (a - 1) + f /(, _ j)2 ,160 + (To + 25). 3-The results of a series of calculations using the above two equations for a 100 rod and a 125 rod core and different values of the prompt negative temperature coefficient are given in Section 5.3 of this report. 2132 204
13 5.0 GENERAL SAFETY ANALYSIS 5.1 Fuel Rod Temperature and Power Density The limiting parameters which insure the safe operation of a TRIGA reactor are the maximum allowable fuel rod temperature and the maximum allowable fuel rod power density. These two parameters are interdepen-dent and their limiting values were previously considered in Sections 3.0 and 4.2 of this report. The maximum allowable fuel rod temperature for FLIP fuel was found to be ll50*C and that for standard fuel found to be 1000*C. In order to provide a reasonable margin of sa%ty during steady state operation, a safety margin of 200'C is established for the WSU TRIGA reactor. Thus, the limiting safety system settings would correspond to maximum temperatures of 950*C in the FLIP fuel and 800*C in the standard fuel. The limiting value of power density will be con-sidered in Section 9.0. A number of mixed cores were studied at WSU in order to establish their expected performance and characteristics. The current layout of the WSU TRIGA reactor core, core 29A, fueled with all standard fuel is - shown in Figure 4. The maximum centerline fuel temperatures and rod power densities at a power level of one megawatt are shown in Figure 5 and Figure 6 respectively. A thermal neutron flux profile along the D row of this core is shown in Figure 7. Starting with core 29A, six mixed cores of standard and FLIP fuel were studied by incremental replacement of standard fuel with FLIP fuel. The FLIP fuel was located in a contiguous block in the center region of the core. The maximum centerline fuel rod temperatures, rod power den-sities, and flux profiles along the indicated rows are shown in Figures 2132 205
~
CORE ARRANGEMENT CORE NO. 29A : DATE AUGUST 13 ~ ,197 3 k* I 2 s. 4 s s T s s . 4 y y y 1 y v v v v LEGEND
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2 3 4 5 6 7 0 0 0 0 202 227 0 0 237 175 0 0 A 0 0 0 0 210 234 C 0 223 183 1 0 188 197 215 229 240 252 256 262 24) 2 16 193 137 2 214 237 259 275 286 291 273 283 279 2a3 234 225 240 264 286 332 318 326 331 323 31. '02 272 252 C 233 256 278 293 337 323 336 221 313 0 272 241 234 260 281 295 303 325 0 323 315 301 277 253 D 234 262 0 294 3J2 217 327 213 3]o 2 d9 263 253 231 256 276 289 306 315 317 211 211 2 32 25' ?' 7 E 237 260 281 296 314 323 325 313 3:M 2o7 263 249 g 211 233 255 27] 2 i0 235 235 281 272 257 234 120 N r 184 192 210 225 234 239 239 235 ? 2 f2 211 191 1 32 N CD " P 2 A K / t.VI = 1.27 '2M
- = 254.6 Figure 5
0 ) 7 2 7 5 s 5 4 6 0 6 3 3 3 5 6 6 5 6 S 6 5 3 7 0 3 9 ) 1 1 0 4 1 4 7 3 3 0 3 6 3 S 2 7 7 7 S 3 3 6 9 5 3 0 7 6 0 5 0 6 3 3 4 7 0 0 0 9 9 9 7 4 6 1 1 4 3 ) 7 3 7 3 2 4 7 2 3 4 5 S 3 1 1 1 1 3 3 3 5 1 1 1 1 1 1 0 0 4 5 5 3 3 7 6 1 9 7 0 0 7 9 2 2 2 1 1 2 5 5 5 1 1 1 1 1 1 0 0 7 0 0 3 0 7 0 6 3 0 6 0 0
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/ 3 1 1 1 1 1 W
K 0 0 8 2 3 6 9 0 5 9 9 6 N . . . . . . . . . . . . I 0 0 4 7 9 0 3 0 8 d 6 4 Y T I 0 0 1 8 5 0 3 4 0 3 6 9 S . . . . . . . . . . . . N 0 0 4 5 7 7 7 7 7 7 5 3 A 2 7D 2 R 0 0 8 7 J 6 7 7 5 9 6 6 E: . . . . . . . . . . . . C . 0 0 3 4 6 5 5 5 5 5 4 3 O3 CF h S C D 5 F
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18 8 to 25. The FLIP region of each of these cores is outlined on each figure. Table III summarizes the results of the studies on these mixed Cores. TABLE III Max Power No. Max Fuel Peak / Ave Mean Fuel Density, Designation FLIP Rods Temp, *C Fuel Temp,*C Temp,*C kw/ rod 30A 11 681 2.48 275 21.7 30B 23 501 1.86 269 18.8 30C 31 623 2.30 271 20.9 300 31 559 2.07 271 19.9 30E 34 442 1.67 265 17.4 30F 45 403 1.54 262 16.3 The data on the six mixed cores studied indicates that these con-figurations, which all have a contiguous block of FLIP fuel surrounded by standard fuel, are safe to operate under steady state conditions. However, the temperature and power peaking effects in core 30A with only 11 FLIP rods and in cores 30C and 300 which have 4-rod water holes are rather severe. Furthermore, the large peaking factors in these cores would greatly limit the allowable reactivity addition under pulsing con-ditions. Thus, to provide a wider margin of safety and to permit a reasonable pulsing capability, the arrangement of a mixed core should be limited to configurations with moderate peaking effects. The following configuration specification is established for the WSU TRIGA core. The FLIP-fueled region in a mixed core shall contain at least 22 FLIP fuel rods in a contiguous block of fuel in the central region of the reactor core. Water holes within the FLIP regio'n shall be limited to single rod holes. 2132 210
Cl; E 33' FUFL 6 03 M aXI *.UM C E NT 5 0.L ! r!! Ti"?T Rl 'U? S , 2EGR5:5 C: 2 3 4 5 6 7 0 0 0 0 191 211 0 3 197 163 0 0 A C 0 0 0 2 )J 226 0 3 21+ 177 0 3 174 183 232 218 231 247 262 23d 23d 211 133 191 8 198 222 247 266 279 236 237 233 275 261 239 219 225 250 275 293 238 292 639 356 247 299 270 243 C , 217 243 267 285 3 10 291 629 ;70 204 3 27} 265 i l 219 248 271 287 297 296 : 0 562 2 3 ', 299 263 25) D ' 219 249 0 236 295 237 !631 496 277 237 251 249 216 244 266 292 299 232 4 19 414 263 274 255 244 E , 222 247 270 239 305 290 537 273 235 263 245 523{ 196 219 243 261 274 230 2 10 277 2?) 2 35 231 215 F 17C 180 199 215 227 234 236 233 224 2 '8 137 179 N (3a oll,K / AVE = 2.48 12 2 P! = 274.6 PO . Figure 8 U M
20 5.2 Operational Mixed Cores The actual arrangement of the FLIP fuel that will be initially loaded into the WSU reactor is that of core 30E containing 34 FLIP rods. However, due to unforeseen needs and conditions that may arise in the future the mixed core configuration will not be limited to this arrange-ment. Any configuration that satisfies the requirement established in the previous section would be permissible. On the other hand, the arrangement of core 30E is not expected to be changed significantly and additional FLIP fuel will not in all probability be added to the core for a number of years. The standard fuel removed from the central por-tion of the core will be cycled into the standard region to maximize the fuel burnup per rod in that region. It is anticipated that the operational mixed core will have a prompt negative temperature coefficient for pulsing of about -$ 014/"C at the beginning of core life reducing to about -$.012 by the end of core life. Some of the other characteristics of this core are shown in Figures 20, 21 and 22. 5.3 Pulsing Characteristics of Operational Core The exact pulsing characteristics of a mixed core are not as easily predicted as those for a core with a single type of fuel. The problem stems from the uncertainty in the exact magnitude of the prompt negative temperature coefficient and the temperature dependence of this parameter. The measured pulsing characteristics of the Texas A&M mixed core are shown in Figure 26. (11) This core contained 35 FLIP fuel rods and 64 standard fuel rods for a total of 98 fuel rods. The characteristics of a 100-rod TRIGA core with a temperature coefficient of -$.0145/*C.and 2132 212
k' 0 0 5 0 5 3 6 5 2 3 9 4 C 0 3 5 6 6 6 6 6 6 4 3 7 3 0 7 3 0 0 3 3 4 3 5 7 0 0 3 5 8 3 7 7 6 7 5 3 1 4 6 3 5 0 5 4 8 2 9 5 3 3 4 7 0 0 0 9 3 9 6 4 6 1 1 2 4 7 3 5 5 3 4 2 1 2 1 5
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0 3 1 1 9 1 0 7 6 2 5 6 0 C 7 9 9 3 3 3 6 ) 3 5 5 1 1 2 1 1 l' 0 C 4 4 7 0 0 7 3 9 8 e 9 0 0 7 9 0 1 0 1 7 9 8 5 e 2 2 2 1 1 r i' u g i 6 3 4 3 9 8 2 4 9 7 8 7 F 4 5 6 9 9 9 0 9 8 9 8 5 4 1 3 2 6 7 3 6 3 2 5 1 3 4 3 4 5 S 1 0 0 0 0 1 3 5 1 1 1 1 1 1 0 0 0 9 7 9 2 4 4 9 5 3 8 0 . . . . . R 0 0 4 7 9 9 9 9 8 9 7 4
/ 3 W
K 0 0 3 4 4 8 0 0 7 0 2 1 N . . . . . . . . . . I 0 0 4 6 8 7 8 0 7 8 6 4 Y T I 0 0 6 . 6 2 4 6 2 4 0 5 S . . . . . . . . . . . N 0 0 3 5 6 6 6 6 6 6 5 3 AE 2 0O 3 F 0 0 3 1 2 9 0 0 3 1 0 2 FE . . . . . . . . . . . . F A 0 0 3 4 5 4 5 5 4 5 4 3 O2 CP A 0 C 0 E F A. N UN ra t
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CC0c 3]B FUEL F0D MAXIMUM CENTS LINE "PE0A'UD.E, OEG7E~5 C: 2 3 4 5 5 7 0 0 0 0 189 223 0 0 192 160 1 3 0 0 0 0 199 223 0 3 277 167 0 3 - 174 184 204 219 232 245 258 253 22) 203 170 L71 B 200 224 250 266 2 76 232 233 279 27) 253 226 207 227 253 275 270 432 4d6 'e l l 5]l 201 291 2sl 233 C 219 245 268 250 332 357 335 433 27' 3 261 '35 221 250 27L 259 394 2'J4 0 499 271 291 250 230 D 221 252 0 258 318 3s5 3S7 44U 27) 27:3 252 239 218 246 266 254 369 232 335 2635 252 2 /1 246 233 E N 224 249 271 264 , 448 335 3E3 467 271 277 251 235 t __ . a q ,a 196 220 245 261 270 275 276 2?2 253 2 ',7 221 205 r\) F
~
171 180 199 215 227 233 233 222 217 203 17 3 163 N LM 3'AK/
- AVE = 1.36 3. AN = 259.2 to w
Figure 11
C JRE 303 F C .4 E D DENSI7Y IN KW/RG7: 2 3 4 5 d 7 0.0 0. 0 0.0 0.0 3.9 4.5 0.0 2.3 3.9 2.9 3.3 3. 0 A O.0 0.0 0. 0 0.0 4.1 5.2 0.0 0.0 4.4 3.1 0.1 1.0 3.3 3.6 4.3 5.0 5.6 6.3 7.1 1. '3 5.5 4.3 3.4 3.2 B 4.2 5.2 6.6 7.7 8.5 9.0 9.1 3.7 7.9 6.7 5.3 4.5 5.3 6.3 S.4 8.0 13.4 16.4 L6.5 13.3 3.9 9.J 7.3 S.0 C 5.0 6.3 /.8 7.1 15.5 14.5 15.6 17.1 3.a 3.3 7.3 5.7 5.1 6.6 8.0 7.2 16. 0 16.3 0.0 13.3 3.o 9.7 7.2 6.0 0 5.1 6.7 0.0 7.1 15.7 14.8 16.1 17.6 J.3 8.7 6.7 5.9 4.9 6.3 7.7 6.8 15.0 13.3 13.3 15.6 7.4 S.1 6.3 5.6 E ' 5.2 6.5 8.0 7.6 ! 17.6 15.6 15.0 13.1 3.1 U.5 5.6 5.7 IN} 4.1 5.0 6.3 7.3 3.0 3.4 3.4 3.1 7.5 6.4 5.3 4.4 F u pg) 3.2 3.5 4.2 4.3 5.3 5.6 5.6 5.4 4.9 4.2 3.4 3.! Figure 12 %
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COR E 2]C FUEL RCD M AXI MUM CE *l7 F R L IIE T i."PE 7.AT*F.E . 02G3.225 C: 2 3 4 5 6 7 0 0 0 0 185 232 C 3 132 147 0 0 3 0 0 0 0 195 218 0 0 195 157 0 0 - 174 183 202 216 228 240 251 244 217 189 16S 153 8 199 224 249 264 273 277 277 27] 262 245 213 194 228 253 275 269 b h68 371 38? 623 3 J' 273 223 C 1 221 246 268 258 379 352 372 447 0 0 256 224
' l 223 252 272 260 395 371 0 - 01 311 + 16 i 220 235 D -
223 254 0 260 , 393 367 3 )3 253 317 322 210 2 39 221 248 263 256 1377 336 338 323 312 317 21' 235 E 227 252 274 268 472 400 433 2 34 353 355 237 241 200 224 249 265 274 283 231 277 263 252 230 216 F 175 184 204 221 232 238 239 235 226 213 133 179 rs) PEAK / AVE = 2.30 *!E AN
= 270.9 trJ N 9t Figure 14 00
U 0 0 0 0 2 3 3 4 5 4 4 5
'.5 0 5
3 5 1 6 3 4 3 4 7 0 0 0 7 2 9 4 0 9 3 5 3 0 J 3 4 8 o 5 5 4 5 5 3
. t 6 8 8 3 0 9 7 4 1 '. 7 6 2 2 3 6 - 0 0 6 2 2 4 6 4 6 1 1 1 1 5 3 7 - 3 0 3
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. ~. . . . . . . .
C 0 6 3 0 7 6 4 2 5 8 5 5 2 1 1 1 1 1 5 2 0 7 5 8 1 0 9 4 3 _ 9 0 1 0 0 6 3 5 5 0 5 3 6_ 3 6 e 1 1 1 1 1 r u g i 3 9 3 6 8 2 9 9 3 2 8 4 F 4 4 6 c 5 4 5 4 3 6 8 5 4 1 1 1 1 1 1 7 0 4 2 1 4 0 9 3 2 3 6 3 4 5 3 3 5 6 5 5 3 8 5 1 1 1 1 1
,'1 0 0 0 9 5 9 1 2 2 0 8 6 0 0 . . . . . . . . . .
P, 0 0 4 7 7 7 7 7 7 7 7 5
/ 3 N
K 0 0 3 5 4 8 1 0 9 3 5 3 N . . . . . . . . . . . I 0 0 4 6 8 7 d 0 7 8 6 4 Y T I 0 0 6 2 8 4 7 8 5 7 2 6 S . . . . . . . . . . . . N 0 0 3 5 6 6 6 6 6 6 5 3 CE 2 0D 3 R 0 0 3 2 4 0 1 2 0 3 2 3 EE . . . . . . . . . . . . RW 0 0 3 4 5 5 5 5 5 5 4 3 CO CP A 8 C D E F g p uN N"o
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C2FS 300 FUSL 000 VAXIVT4 CENTF7LI'if T2?'P"_3* W_ ,
, .? i G 3 E I ", C:
2 3 4 5 6 7 0 0 0 0 137 237 C 0 194 164 3 1 A 0 0 0 0 197 222 0 0 209 173 0 0 - 171 131 201 216 230 244 259 255 232 207 134 17' B 196 221 246 263 274 23) 233 2 79 257 253 237 215 223 249 272 267 Go2 377 400 332 352 355 220 243 C 215 242 264 255 370 353 374 357 395 433 231 2T5 217 246 267 256 381 331 0 559 0 0 263 242 0 217 248 0 254 375 355 334 493 0 0 253 241 214 242 262 253 JSC 326 326 339 364 414 229 233 E 220 245 267 261 42a 374 377 353 3 ', J 343 224 235 192 216 241 257 267 273 274 270 261 245 222 F 200 y 167 177 196 212 224 231 232 223 213 202 101 172 'A ' N
- PEAK / AVE = 2.07 M: Afl = 27C.7 N
N e e" Figure 17
COP: 330 P0nER DENSITY IN K'4 / D.CO : 2 3 4 5 6 7 0.0 0.0 0.0 0.0 3.7 4.5 0.0 C.3 4.3 3.) 0.0 0.3 A 0.0 0.0 0.0 0.0 4.1 5.1 0.C C.3 4.5 3.3 0.0 0.3 - 3.2 3.5 4.2 4.3 5.5 6.2 7.1 0.9 5.6 4.5 3.6 3.4 B 4.0 5.0 6.4 7.5 8.3 S.3 9.1 9.7 7.) 6.8 5.5 4.3 5.1 6.5 8.2 7.7 17.9 16.1 16.2 13.5 14.2 14.3 5.4 6.0 C 4.3 6.1 7.5 6.9 15.1 14.1 15.2 14.7 16. 1 17.1 5.5 5.3 4.9 6.4 7.8 6.9 15.5 15.5 0.0 19.9 0.0 0.0 7.3 6.1 0 4.9 6.5 0.0 6.9 15.2 14.4 15.6 19.3 0.1 0.s 7.3 6.1 4.7 6.1 7.4 6.6 14.5 12.7 ! ?.8 13.5 15.5 16.3 5 . ', 5.' rs) t --. 5.0 6.3 7.3 7.3 17.1 15. 15.3 14.7 13.6 13.7 5.2 5.3 u ._. ___ _ _ _ _ _ _ _ _ __ rs) ' 3.9 4.8 6.1 7.1 7.8 8.2 :3 . 3 5.3 7.3 6.3 5.1 4.5 F I'3 3.1 3.4 4.0 4.7 5.2 _' . 5 5.6 5.4 4.9 4.3 3.5 3.2 N N w O Figure 18
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i noIx) o nos oNolu x01J NO LO3N liell$lt 2132 223
CORE 30E FUEL CCC
- A X I M L"i C E N T E R L I t c. T E" P i :17 'E Z , 3500: 25 ::
2 3 4 5 6 ' C C 0 0 134 234 0 3 '73
. 163 J J t
0 0 0 0 193 219 0 3 233 173 0 1 - 167 177 196 212 226 241 257 253 232 207 183 173 2 191 216 242 260 271 2/3 231 273 26) 254 232 213 218 245 268 263 442 333 396 393 378 . 9) 2 4 ', 245 C ' 210 237 260 251 352 346 371 344 3 4 '. 9 2 ', 4 2' ? I 212 242 263 252 371 370 0 374 535 361 ' 237 245 0 212 243 0 251 366 351 379 349 3t7 326 224 24; 209 237 2S9 247 351 323 327 321 334 3 16 217 233 E rs) 215 241 263 257 '12
, 367 3/3 363 3 ',1 345 22? 239
- _. -- -- a (#J 18a 212 237 254 2$4 271 273 269 261 24s 224 211 rs) p p) 163 172 192 209 221 223 230 226 217 222 '02 3 N 4 PEAK / AVE = 1.67 '1 :41 = 254.9 w N Figure 20
CCP5 30E POWER 0:NSITY IN KW/QOD: 2 3 4 5 6 7 0.0 0.0 0.0 0.0 3.6 4.3 0.0 0.3 3.9 3.2 0.0 0 . ') 0.0 0.0 0.C 0.0 3.9 5.0 0.C C.) +.5 3.3 J.0 0.0 - 3.1 3.4 4.1 4.7 5.3 6.1 7.0 6.8 3.a 4.5 3.6 3.4 9 3.9 4.3 6.1 7.2 3.1 E.6 S.9 .E . 7 7 . -) 6.3 5.6 4.0 4.9 6.3 7.8 7.5 '17.9 .5.S 16.0 15.5 15.4 17.4 6.2 6.3 C
.6 5.3 7.3 6.6 14 .7 12.9 15.1 13.3 13.4 9.] S.2 6.!
4.7 6.1 7.5 6.7 15.1 15.3 0.0 15.2 13.2 15.5 5.c 6.3 0 4.7 6.2 0.0 6.6 14.8 14.2 13.4 14.0 12.3 12.7 5.2 6.2 4.5 5.9 7.1 6.4 14.1 12.4 12.7 12.3 11.3 11.9 4.9 5.9 N E 4.5 6.0 7.5 7.1 ,16.6 14.9 15.2 14./ .3.6 13.9 5.3 6.) U i~ N 3.7 4.7 5.8 6.3 7.6 8.9 3.2 7.9 ?.3 6.3 5.2 4.6 N r N 2.9 3.2 3.9 4.5 5.1 5.4 5.5 5.3 4.9 4.3 3.5 3.? L.n w Figure 21 w
34 8 d 8
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00'0a1 00o$1 cod 9 00*b9 00*dh -00' 2 00' l tiOIX) 3 MOU ON010 Xnld NOU103N 10WU3HL 2132 226
CCoE 30: FUEL F OD M AXIMUM CE NTEEL INd T E M P E l'. T U R E , O'G4: 25 C: 2 3 4 5 6 ' O O O O 184 202 0 3 137 157 0 0 a 0 0 0 0 194 217 0 J 292 167 0 0 173 182 201 215 226 239 253 249 226 200 173 171 3 199 221 243 259 271 276 278 273 ?64 247 224 210 22' 223 335 332 357 377 381 355 353 4 03 237 237 C 216 210 305 299 313 338 359 333 331 0 237 234 213 218 319 304 316 365 0 355 323 353 230 23' C 218 222 0 305 315 3'3 360 337 309 315 216 236 215 214 312 299 308 318 320 313 331 307 2J9 230 E 221 219 328 324 344 350 302 349 329 331 N _ 221 232 {}) - 196 217 238 254 264 269 270 265 255 239 217 203 es> 170 178 197 211 222 227 227 222 212 196 175 166 N N N P6AK/ AVE = 1.54 9?AN = 261.9 ca m Figure 23
)
oO 0 3 2 6 0 7 9 3 5 6 3 0 0 0 3 4 5 5 5 5 5 5 4 3 7 0 0 4 2 8 9 5 9 6 3 0 3 0 0 3 5 5 5 5 4 4 5 4 3 3' 3 1 2 4 - 3 3 5 9 2 3 ) 0 2 3 4 6 6 3 4 1 1 3 6 4 6 1 1 1 1 1 7 3 3 3 5 3 5 4 7 9 9 7 3 4 5 7 't 3 2 l 3 2 6 4 l 1 1 1 1 1 3 0 5 3 0 1 5 4 7 1 6 1 0 0 6 S f. 3 4 3 1 4 7 5 5 I 1 1 1 1 1 4 0 0 8 6 5 5 0 9 3 7 0 3 2 0 0
. . . . . e 6 8 3 4 0 4 2 4 3 5 r 1 1 1 1 1 u g
i 9 F 2 0 5 3 5 0 7 1 6 9 3 4 4 6 8 5 2 4 3 2 4 7 5 4 1 1 1 1 1 1 6 9 3 0 4 7 1 9 3 8 5 1 3 3 5 3 4 1 2 1 1 3 7 5 1 1 1 1 1 1 0 0 0 8 2 1 5 0 0 4 5 8 6 J . . . . . . . . . . R C 0 4 7 3 0 1 1 0 2 6 4
/ 3 1 1 1 1 1 1 W
K 0 0 2 2 2 1 1 0 6 0 9 1 N . . . . . . . . . . . . I 0 0 4 6 3 1 2 0 1 2 5 4 1 1 1 1 1 Y , ' T
? J 0 6 1 1 6 9 1 7 0 9 4 5 . . . . . . . . . . . .
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)
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f Figure 26 800 _ Lt. WSU Calculations + g~700 - w + Texas A & M Reactor Performance l- _1 w 3 Lt. gQQ _ O w l-- a . . . - Z 500 - FUEL TEMPERATURE VS. REACTIVITY
>d
' INSERTION -
2 400 _
~~
N u - na ~ N EO ' f ' i " 2.00 2.% # ci 1.00 1.50 u, REACTI VITY INSERTION, CDLLARS
39 with a prompt neutron lifetime of 32 microseconds as calculated by the equations in Section 4.4 is also plotted on this figure. The theoreti-cal curve predicts a somewhat greater temperature than the measured values and assuming a 200*C core temperature before pulsing. The FLIP fuel content of the initial WSU mixed core 30E will be comparable to that contained in the Texas A&M mixed core. We can thus expect the two cores to have comparable negative temperature coeffi-cients and pulsing characteristics. If we assume a more conservative value of -$.014/ C for the temperature coefficient for the WSU mixed core along with a 32 microsecond prompt neutron lifetime, we obtain the results listed in Column A of Table IV. The expected performance at the end of core life is shown in Column B of Table IV. TABLE IV PULSING CHARACTERISTICS OF 125-R0D MIXED TRIGA REACTOR CORE 30E Average Peak Fuel Reactivity in S Peak Power in MW Temp Increase, *C (A) (B) (A) (B) 1.5 218 256 70 81 2.0 893 1048 139 1 61 2.5 2047 2411 204 237 3.0 3708 4380 269 312 (A) Tc = $.ci4/*C, (B) Tc = -$.012/ C In order to provide at least a 2C0 C margin of safety during pulsing of mixed core 30E, pulsing should be limited to $2.50 so that the average core temperature increase is 250 C or less. Combining this 250 C limit 2132 231
40 with the average core temperature of 265*C and a peaking factor of 1.67 yields a maximum possible fuel temperature of 861*C in this core. The peak fuel temperature obviously would be within the FLIP fuel region and the peak calculated value is below the established 950*C safety setting limit for FLIP fuel. In actual operation pulsing is not permitted above a power level of 1 kw and thus a wider safety margin actually exists than that calculated above. 5.4 Fuel Temperature Scram In order to provide a further degree of safety for the WSU TRIGA reactor when operated with a mixed core, a fuel temperature channel with a scram capability shall be added to the control system. This channel shall scram the reactor if the instrumented fuel rod temperature exceeds the fuel temperature safety system setting. The instrumented fuel ele-ment for this channel shall be located in the FLIP fuel region. The stud (es on mixed cores in Section 5.1 indicates that the core with the smallest allowable number of FLIP fuel rods has the highest FLIP region fuel temperatures. The cores with four (4)-rod water holes in the FLIP region are excluded as not being allowable configurations. Thus under the guidelines at the end of Section 5.1 core 308 is the core with the smallest allowable number of FLIP fuel rods. Figure 11 for this core indicates the following temperature characteristics for the FLIP region: maximum region temperature 501 C, minimum region tempera-ture 332*C, and maximum-to-minimum f"el temperature ratio in the region 1.51. Taking this core as the w;rst case and applying the limiting safety system setting of 800 C for standard fuel to the FLIP region for an added margin of safety we obtain a value of 800/1.51 = 529 C for the 2132 232
41 safety system setting. That is, if the instrumented fuel element is located in the FLIP fuel rod with the lowest power density in the FLIP region of core 30B and has an indicated temperature of 529 C, the tem-perature of the FLIP fuel rod with the highest power density would not exceed 800 C. An examination of Figure 21 for core 30E indicates that this core has the following characteristics for the FLIP region: maximum region temperature 442*C, minimum region temperature 309"C, and maximum-to-minimum fuel temperature ratio in the region of 1.43. In core 30E the maximum possible fuel temoerature for a 529"C indicated FLIP region temperature would be 758 C. Thus the 529 C safety system setting pro-vides a wider margin of safety in this core. A value of 500 C is established as the fuel temperature safety systw setting for the WSU TRIGA reactor operating with a mixed core. This limit will provide an ample margin of safety as demonstrated above and will permit operation of all allowable mixed core configurations. 2132 233
42 6.0 DESIGN BASIS ACCIDENT The Design Basis Accident for a TRIGA reactor is defined as the loss of the integrity of the fuel cladding of one fuel rod in air. The hazard associ-ated with this theoretical accident is thus the effects of the postulated fission product release within the facility and to the surrounding environ-ment. For the purposes of the D.B.A it is assumed that all the fission products are released directly into the pool room air and that none are retained in the pool water. The fission product release fraction for TRIGA-type reactor fuel has been measured experimentally (12) and documented before the AEC hearings on the Columbia reactor as being 1.5 x 10-5 at a fuel temperature of 300 C. The release fraction, FR, is, however, a function of the fuel temperature, T, in C given by the relationship (12): FR = 1.5 x 10-5 + 3.6 x 103 EXP - (1.34 x 104 /T + 273). Assuming a fuel temperature of 500 C, the release fraction is calculated to be 1.2 x 10 -4 by the above relationship. A release fraction of 1.2 x 10-4 will ' used in the calculations for the D.B. A. A power density of 30 kw per fuel rod and an infinite irradiation time will also be assumed for the D.B.A. Under these conditions the fission product inventory for one TRIGA fuel rod and the associated released fission products are tabulated in Tables V and VI. This tabulation was derived from the basic data of Perkins and King (13) along with the documented fact (12) that only the gaseous fission products escape when the cladding of a TRIGA fuel rod ruptures. A summation of the fission product release data tabulated in Tables V and VI yields a total fission product release for the D.B.A. of 1.72 curies of gamma emitters and 2.25 curies of beta emitters. If this activity is uniformly 2132 234
42a distributed in the pool room and control room which have a combined volune of 1 x 10 9cm ,3 the specific gamma activity would be 1.72 x 10 -3 Ci/cm 3 and the specific beta activity would be 2.25 x 10 -3 pCi/cm 3, 2132 235
43 TABLE V SOLUBLE GASEOUS FISSION PRODUCTS CONTAINED IN AND RELEASABLE FROM A SINGLE TRIGA FUEL ROD
- Saturated Released Isotope Inventory Activi ty Hal f-Li fe Ci mci Br-82 40 4.8 35.3 hr 83 137 16.4 2.3 hr 84 253 30.4 31.8 min 85 330 39.6 3.0 min 87 780 93.6 55 sec Total Br 1540 184.8 I-130m 260 31.2 9.2 min 131 734 88.1 8.1 days 132 1115 133.8 2.3 hr 133 1672 200.6 21 hr 134 2027 243.2 54 min 135 1546 185.5 6.8 hr 136 785 94.2 86 sec Tot-' I 8139 976.6 Total Released Soluble Gaseous Fission Products = 1.16 Ci Total Gamma Emitters = 1.03 Ci lotal Beta Emitters = 1.16 C'
- Power Density = 30 kw/ rod, fuel temperature = 500 C, release fraction = 1.2 x 10-4 2132 236
44 TABLE VI INSOLUBLE GASE0US FISSION PRODUCTS CONTAINED IN AND RELEASABLE FROM A SINGLE TRIGA FUEL ROD
- Saturated Released Isotopo Inventory Activity Hal f-Li fe Ci mci Kr-83m 137 16.4 1.9 hr 85m 330 39.6 4.4 hr 85 67 8.0 10.7 yr 87 634 76.1 78 min 88 912 109.4 2.8 hr 89 1115 133.8 3.2 min Total Kr 3195 383.3 Xe-131m 7 0.8 12 days 133m 40 4.8 2.3 days 133 1672 200.6 5.3 days 135m 457 54.8 15 min 135 1621 194.5 9 hr 137 1545 185.4 3.9 min 138 1166 139.9 17 min Total Xe 6508 780.8 Total Released Insoluble Gaseous Fission Products = 1.16 Ci 2132 237 Total Gamma Emitters r 0.69 Ci Total Beta Emitters = 1.08 Ci
- Power Density = 30 kw/ rod, fuel temperature = 500 C, release fraction = 1.2 x 10-4
45 6.1 Whole Body Dose in Pool Room In the calculation of the whole body dose from the gamma emitters,we may approximate the pool room by a hemisphere with a radius of 808 cm. The dose at the center of the hemisphere may. then be calculated from the equation SV (1 - e-IR) D = 2EC where D = dose rate, mr/ min 3 S y
= photons /sec cm of air I = attenuation coefficient for air R = outer radius of hemisphere (808 cm)
C = flux to dose rate conversion. In the case of fission products, the average energy of the gamma emit-ters is .7 MeV which gives a value 3.5 x 10-5 cm -I for I and 4.2 x 10 4 2 y/cm sec per mr/ min for the conversion factor C. The specific photon activity of the pool room air is 63.6 photons /sec cm3 . Substituting these values in the above equation yields a maximum dose rate of .6 mr/ min or 36 mr/hr. Using a quality factor of 1 for gamma rays, the maximum whole body exposure rate for operating personnel in the pool room as a _ result of the D.B. A. would be 36 mrem /hr. 6.2 Lung Dose in Pool Room The primary hazard to operating personnel in the pool room, however, would be from the inhalation of gaseous fission products rather than the whole body dose. In this case the critical organs are the lungs and the thyroid and the beta emitters are the source of the internal exposure. The " standard man" has a lung capacity of 3 x 10 3 2132 238
46 3 3 cm and breathes 1.25 m of air per hour. After 0 few breaths the lungs of a person in the pool room would contain 2.25 x 10-3 Ci/cm3 x 3 x 10 10 3 cm = 6.8 Ci of beta emitters. The lung dose from this in-haled beta activity may be calculated from ACR 0 fE il D = (1 - e~Ai t) rads (beta) , i 1 where A = activity in the lungs on leaving the reactor room, pCi 4 4 erg /MeV) C = conversion factor (3.7 x 10 s/sec 100 uti)(1.6 ergs x 10
/g-rad R =
retention factor (0.125) m = mass of lungs (1000 grams) f4= fraction of total activity Eg= energy of beta from nuclide i, MeV A; = radioactive decay constant + biological release rate t = time of exposure (assumed to be infinite). Re weighted average energy of the beta particles emitted by the gaseous fission products is .78 MeV. Substituting this and the other values in the above equation and assuming an infinite exposure time yields a lung exposure of 8 mrads. This is significantly below the limiting biologi-cal exposure of 50 rads for the lungs. 6.3 Thyroid Dose in Pool Room The thyroid dose to a person in the pool room was calculated using the conventional assumptions that the individual remains in the room for 10 min after the fission product release. All of the iodine is assumed to be absorbed into the bloodstream and carried immediately (6.7 x 10-8 sec ~l) 2132 239
47 tothethyroidwhichisassumedtohavearetentionfactoroT25%. The thyroid dose resulting from the inhaled activity may be Iculated from the equation D thyroid
=
Bt I (QjE j) where D thyroid
= thyroid dose 1.' rem 3
B = inhalation rate = 200 cm /sec t = time = 600 sec Q9 = concnntration of the i th isotope in Ci/cm3
= th isotope in rem /Ci inhaled.
E j effectivity of the i The results of the calculations are shown in Table VII. The 29.7 rem thyroid dose does not exceed the i early maximum permissible thyroid dose of 30 rem. It is to be notad that the thyroid dose would be re-duced by about a factor of 10 if the water remained in the pool as 90% of the iodine would remain in th? psol water. TABLE VII Ten-Minute Thyroid Dose to Persons in the Pool Room Following as a Result of the D.B.A. Oi E j D Curies 3 thyroid Isotope Released uc77.m rem /ucT inhaled rem I-131 .088 8.8 x 10-5 1.48 15.6 I-132 .134 13.4 x 10-5 .0535 .9 I-133 .201 20.1 x 10-5 .40 9.7 I-134 .243 24.3 x 10-6 .025 .7 I-135 .186 18.6 x 10-6 .125 2.8 TOTAL . . . . . . . 29.7 rem 2132 240
on 6.4 Discharge of the Fission Products into the Environment The rate at which the fission products from the D.B. A. are released into the environment is primarily dependent upon the effects of the pool room ventilation system. If the ventilation system is in operation, air is exhausted from the pool room at the rate of 2000 cfm 5 3 or 9.44 x 10 cm /sec. If the ventilation system is off, the release to environment would only be by leakage from a sealed tuilding wnich is 4 3 estimated to be of the order of 100 cfm or 4.81 x 10 cm /sec. The time integrated activity exhausted from the facility is given by the equrtion
=
Ag Aj g/(Aj + q/V), pCi. where Aj = the concentration of the i th isotope in the room at t = 0 in Ci/cm3 q = the building exhaust rate in cm3/sec V = the volume of room in cm A g
= the decay constant of the i th isotope in sec-I .
The activity discharged into the atmosphere as a result of the D.B.A. with the ventilation system on and off as well as the relevant MPC values ate tabulated in Table VIII. Column D is the released activiuy with the ventilation system on divided by the annual volume of the ventilation system. Column E is the released activity with the venti-lation system off divided by the annual volume of the ventilation sys-tem. Column G is the ratio of the annual average released specific activity with the ventilation system on (Column D) divided by the MPC for unrestricted areas for that isotope. An examination of Column D of Table VIII indicates that the only 2132 241
TABLE VIII FISSION PRODUCT RELEASE CONCENTRATIONS AS A RESULT OF THE D.B.A. A B C D E F G Isotope Activity Released in mci Release Concentration in pCi/cm3 MPC Ratio Ventilation System Averaged Over One Year
- Unrestricted D/F ON OFF ON OFF Areas Br-82 4.8 4.3 -10 -10 -8 -3 1.6x10 1.4x10 4x10 4x10
-10 83 15.1 5.9 5.0x10 2.0x10 -10 2x10
-6 2.5x10
-4 84 22.0 3.6 7.4x10 -10 1.2x10
-10 2x10
-6 3.7x10
-4 85 7.7 .49 2.6x10-10 1.6x10 -II -5 -6 3x10 8.7x10 87 6.5 .36 2.2x10-10 1.2x10
-II -5 -6 4x10 5.5x10 I-130m 4.6x10 -10 13.8 1.1 3.7x10 " ----- -----
131 88 88 3.0x10 -9 3.0x10
-9 lx10-10 30 132 123 49 4.1x10-9 1.6x10
-9 -9 3x10 1.4 133 -9 -9 -10 201 169 6.7x10 5.7x10 4x10 17 134 198 45 6.6x10 -9 1.6x10
-9 6x10
-9 1.1
-9 N
135 180 117 6. ~ r10 3.9x10 -9 lx10 -9 6 3.2x10 -10 -7 -3 136 9.4 .6 2.0x10 " lx10 3.2x10 N N
- Annual Volume of Ventilation System = 2.98 x 1013cm3 gj
TABLE VIII Continued: A B C D E F G Isotope Activity Released in mci Release Concentration in pCi/cm3 MPC Ratio Ventilation System Averaged Over One Year
- U astricted D/F ON OFF ON OFF Areas Kr-83m 14.8 5.2 5.0x10
-10 -10 -6 -4 1.7x10 2x10 2.5x10 85m 38 21 1.3x10 -9 7.0x10 -10 1x10
-7 1.3x10-2 85 8 8 2.7x10 -10 2.7x10-10 3x10
-7 9x10
-I
-9 87 66 18.6 2.2x10 6.2x10 -10 2x10
-8 1.1x10
-I 88 103 45 -9 -9 -0 3.5x10 1.5x10 2.x10 1.8x10
-I 89 -10 -8 4.7x10 -2 28 1.8 9.3x10 6.0x10 " 2x10 Xe-131m .8 2.7x10 " -7 6.8x10 -5
.8 2.7x10 " 4x10 133m 4.8 4.8 -10 -10 -7 1.6x10 1.6x10 3x10 5.3x10 -4 133 201 -9 -9 -7 -2 201 6.7x10 6.7x10 3x10 2.2x10 N 135m 3.3 -9 -10 -7 -3 31 1.0x10 1.1x10 2x10 5x10 U -9 -9 lx10 -7 135 190 135 6.4x10 -2 4.5x10 6.4x10
-10 N' 137 45 3.0 1.5x10 1.0x10 -10 3x10
-8 5x10
-2 u 138 81 9.2 2.7x10 -9 3.1x10 -10 4x10
-8 6.8x10 -4
- Annual Volume of Ventilation System = 2.98 x ,13 3 cm g
51 isotopes released that exceed the MPC values are Icdine 131, 132, 133, 134, and 135. Due to the long half-life of these iodine isotopes, shutting off of the ventilation system will have very little effect on the quantity of these isotopes that escape to the environmeJt under the postulated leakage rate of 100 cfm (see Column F). On the other hand, however, the ventilation system of the pool room is provided with a dilution mode of operation in which 1700 cfm of outside air is mixed with 300 cfm of air from the pool room and discharged from the facili ty. The 300 cfm of air from the pool room in this mode of opera-tion is passed through an absolute filter system which would remove sufficient iodine from the pool room air so that the quantity of iodine discharged when averaged over one year would be below the MPC values. 6.5 Dilution of Discharge in the Lee of the Building The gaseous radioactive material discharged from the facility ventilation system will be diluted by atmospheric air in the lee of the building due to turbulent wake effects. The dilution is proportional to the product of the cross sectional area of the building times the wind speed. That is, 3 4 = dilution factor = 1/cAu (sec/cm ) C = constant (2 to .5), select 1 (cm37,3) where A = building cross-sectional area in square meters U = wind speed in meters /sec. Thus for a nominal 2/ msec wind velocity (4.4 mph)* and a 56 x 28 ft building, the dilution factor is 4 = 3.4 x 10-3 , Average annual wind speed is in excess of 5 mph (14) 2132 244
52 6.6 Whole Body Dose Outside Facility A simple and conservative estimate of the maximum whole body dose to an individual outside the facility may be obtained by assuming that the individual is at the center of a large hemispherical cloud of fission products. The co..:entration of the fission products in the cloud is assumed to be equal to the concentration in the pool room divided by the turbulent wake dilution factor at a wind speed of 2 m/sec. The size of the cloud is assumed to be infinite and the expo-sure time to the individual is assumed to be equal to twice the time for one change of air in the pool room (20 min). On the basis of these assumptions and using the equation in Section 6.1 we find that Diluted release = 5.9 x 10-5 Ci/cm 3total gamma emitters 3 Sv = .22 photons /sec cm of air
=
D .075 mr/ min. For an exposure time of 40 minutes and a quality factor of 1, the maximum whole body dose from a cloud of fission products discharged from the facility as a result of the D.B.A. is 1.5 mrem. 6.7 Thyroid Dose Outside the Facili q The maximum thyroid dose to an individual outside the faci-lity may be calculated using the equation in Section 6.3 and the assump-tion of Section 6.5. On this basis the maximum possible thyroid dose to an individual outside the facility is found to be .26 rem. If the water remained in the pool this value would be reduced by a factor of 2132 245
53 10 and if the ventilation system were in the dilution mode this dose would be reduced by another factor of about 100. 6.8 Summary of Results of D.B.A. The preceding calculations on the consequences of the D.B.A. indicate that the only significant radiation exposure is the thyroid dose to a person in the pool room. The conditions necessary to produce this exposure are the failure of the cladding of one fuel rod along with a complete loss of pool water. The maximum possible radiation exposure to an individual outside the facility under the postulated conditions is minimal. Thus, no realistic hazard to the general pub-lic would result from the Design Base Accident. 2132 246 A 1" activated charcoal filter has a removal efficiency of 99.9% for elemental iodine
54 7.0 REACTIVITY EFFECTS OF ACCIDENTAL FUEL ADDITION The conditions for this postulated accident are the accidental addition of a FLIP four(4)-rod fuel bundle to the central region of the core with the reactor operating at full power. In actual practice the Technical Specifi-cations prohibit operation of the reactor with a vacant grid position. The maximum worth of a FLIP four(4)-rod cluster in the central region of the WSU TRIGA core is calculated to be less than $3.75. Thus the maximum possible reactivity insertion for an accidental fuel addition would be less than
$3.75.
A step increase of $3.75 in core 30E at the end of core life would pro-duce an average core temperature increase of 422 C. Adding this average core temperature increase to the average full power core temperature of 265*C and multiplying by a 1.67 peaking factor yields a maximum possible peak fuel temperature of 1142 C. This is below the 1150 C safety limit for FLIP fuel. The high temperature scram would in this accident limit the maximum fuel tem-perature to belaw the value calculated above and thus increase the margin of sa fet.r . Thus no additional hazard is caused by the addition of a four(4)-rod FLIP fuel cluster to the core during full power operation. 8.0 REACTIVITY EFFECT OF TRANSIENT R0D EJECTION The conditions for this postulated accident are the accidental ejection of the full worth of the transient rod with the reactor operating at full power. During normal operation, pulsing will be administratively limited to the maximum value established in Section 5.3. However, for the purposes of this accident it will be assumed that the operator deliberatel.y violates the established limit and also bypasses the interlock that inhibits pulsing 2132 247
55 above 1 kw. Thus the total maximum worth of the pulse rod of $3.75 would be added to the core at fs:ll power. The effects of a $3.75 reactivity addition at full power were shown not to exceed the safety limit in Section 7.0. Thus no additional hazard is caused by the ejection of the pulse rod at full power. 9.0 LOSS OF COOLANT ACCIDENT The conditions for this postulated accident are that the reactor has been operating at 1 W for essentially an infinite length of time and then a sudden complete loss of coolant (pool water) occurs. The loss of water will shut down the reactor; however, the decay heat from the fission products will continue to produce heat in the fuel elements. In order to insure the safety of the reactor in the event of this postulated accident, the fuel cladding temperature in air must be maintained below the point where a clad-ding failure could occur. The strength of the fuel element cladding is a function of its tempera-ture which is a function of the fuel element temperature. The conservative assumption is made for purposes of this analysis that the fuel and cladding are at the same temperature. The stress imposed on the cladding by hydrogen disassociation within the fuel is a function of the fuel temperature, the fuel burnup, and the free gas volume with the fuel rod. Figure 27 (1,3) shows the stress imposed upon the cladding by TRIGA reactor fuels with H-Zr ratios of 1.6 and 1.7 as a function of fuel temperature. This figure also shows the yield and ultimate strength of the cladding as a function of tem-perature. An examination of Figure 27 indicates that the maximum temperature that standard fuel (H-Zr 1.7) can tolerate in air without damage to the cladding 2132 248
56 and subsequent release of fission products is 900*C; that is, the point at which the hydrogen pressure equals the yield strength of the cladding. This point for FLIP fuel (H-Zr 1.6) is 940*C. Thus the operating conditions of the reactor must be limited such that in the event of a loss of water acci-dent, the fuel cladding temperature does not exceed the above limiting values. The maximum fuel cladding temperature after a loss of pool water as a function of fuel rod power density is shown in Figure 28. This curve was generates from data calculated with a two-dimensional transient transport computer code TAC developed by Gulf General Atomic. It was assumed that the fuel has been operated for 7700 Mw-days, and that the reactor is shut down 15 minutes prior to the loss of coolant. The 15-minute delay wct selected because it would take approximately 15 minutes for the pool to drain down from a point where a low level alann would occur to the point where the core was uncovered from the catastrophic failure of a 10-inch beam port. The results of the TAC code calculations sumarized in Figure 28 indi-cate that Standard and FLIP fuel rods operated at power densities up to 22.3 and 23.5 Kw/ rod respectively would not fail in the event of a loss of cool-ant accident. That is, the loss of coolant imediately after shutdown from 21 continuous years of operation at full power would not produce a fuel clad-ding failure and subsequent release of fission products. An eyamination of the power density data given in Section 5.0 of this report indicates that all the analyzed core configurations have fuel rod power densities below the above-mentioned limiting values. Thus a loss of coolant accident would not precipitate a fuel cladding failure and the fis-sion product decay heat would be removed primarily by natural convection of air. 2132 249
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tb 7 tti ,i t t- -" rh tT-10 0 400 500 600 700 800 900 1000 l100 TEMPERATURE, C FIGURE 27 STRENGTH AND APPLIED STRESS AS A FUNCTION OF TEMPERATURE FOR l.7 AND l.6 H-Zr TRIGA FUEL 2132 250
58 1000 - 900 - o J U a E 800 - 2 E C f 700 - 3 2 i B "a 600 - 2 5 = 500 - 1 1 10 15 20 25 30 Fuel Rod Power Density, Kw, Figure 28: Maximum Fuel Rod Temperature as a Function o. Fuel Rod Power Density for a Loss of Coolant 15 Minutes After Shut Down from 7700 MW-Days of Operation. 2132 2SI
59 10.0 FLIP FUEL LOADING AND MIXED CORE PERFORMANCE TESTS The addition of FLIP fuel into the Washington State University TRIGA reactor will be perfonned as a normal fuel loading in accordance with the standard procedure for such operations. The excess reactivity of the new core, control element worths, and shutdown margin will be measured at low power. The loading of the full operational core will be approached in a stepwise fashion to insure that the operating limits set forth in the technical specifications are not exceeded. The power level of the full operational core will be increased in a stepwise fashion with checks of power indication and fuel temperature at cach step. The first step will be at 100 Kw, the second at 200 Kw, and then in 200 Kw increments or less up to the licensed power limit. The power level of the reactor will then be measured by the calirometric method and the power level indication adjusted accordingly. The pulsing characteristics of the W.S.U. TRIGA reactor will be measured af ter completion of the power calibration. At least two pulses of each reactivity input up to the licensed limit will be performed to assure repeatibility of the data. The reactor will then be placed in routine operation. 2132 252
60 APPENDIX A Fuel Rod Temperature Calculations The experimental instrumented fuel rod temperature data shown in figure 1 was fit to a polynomial expression by the method of least squares and the following equation wb obtained: T = 31.77 + 0.6148 (Pw) - (5.396 x 10-4) (Pw)2 + (1.958 x 10-7) (Pw)3 where T = Instrumented fuel rod temperature in C Pw = Reactor power in KW. Calculations with the EXTERMINATOR-2 code for this core indicated that the axial averaged power density in the instrumented fuel element is 283 watts per centimeter of fuel rod heighth at 1 MW. If we assume that the power density in the instrumented fuel rod as well as all fuel rods is a linear function of reactor power and that all rods behave identically to the instrumented fuel rod, then we can calculate the temperature of any rod from the data on the instrumented fuel rod. That is, axial averaged fuel rod power densities calculated with the EXTERMINATOR-2 code may be used with the experimental data on the instrumented fuel rod to calculate the centerline fuel rod temperature o{ any rod. The power density, Pd, in the instrumented fuel rod at any power, Pw, is Pd = .283 x Pw, where Pd = axial averaged power density in watts per cm of core height and Pw is the reactor power in KW. Thus Pw = Pd/.283. This may be combined with the instrumented fuel rod temperature equation to yield a relationship between fuel rod power density and fuel rod temperature. This equation is: . T = 31.77 + 2.172 (Pd) - (6.738 x 10-3) (Pd)2 + (8.639 x 10-6) (Pd)3 , 2132 253
61 APPENDIX B e Temperature Coefficient Weighting Factor for a Mixed Core The multiplication factor, k, of a reactor according to modified one group reactor theory is given by: 2 k = npfce
-B , g -B r If(fuel) 2 1+B L' Ea+DB 2 = pe -B r . Fast Neutron Production.
2 Ia+DB 2
-B Thus Tc = dk = d pe . Fast Neutron Production dT dT 2 Ia+DB for a TRIGA r,eactor
-B2 r Tc : d_ pe . Fast Neutron Production dT 2 Ia+DB O
2132 254
62 APPENDIX C Calculation of Temperature Coefficient of Core 30E at Full Power im = 265 C is = 229 C ii = 360 C FS/FT = .5105 FF/FT = .4895 Tcs(229 C) = .0155 Tcf(360 C) = .0152 Tcm = .0155 x .5105 + .0152 x .4895 = $.0154/ C + 4 2132 255
63 APPENDIX 0 Fuel Temperature Limitation for TRIGA FLIP Fue'i The determining factor that sets the temperature safety limit for FLIP U-Zrt1 fuel is the disassociation pressure of the hydrogen in the fuel. 1.6 As the fuel temperature increases the hydrogen pressure increases and imposes a stress on the fuel cladding material. If the fuel temperature were to in-crease without limit, some point would be reaci.ed at which the internal pres-sure could cause the cladding to yield and eventually rupture. The purpose of the fuel temperature safety limit is to limit the hydrogen pressure to preclude a cladding failure. The hoop stress exerted on the cladding by the hydrogen pressure is given by the equation r 6=P " h 35.25 (tc = .02, r c= .705) h c where 6 is the hoop stress in psi, 'h is the hydrogen pressure in psi, re is the radius of the cladding, and tc is the thickness of the cladding. Under normal steady state operating conditions at full power the temperature of the type 304 stainless steel cladding will not exceed 140 C. At 140"C the yield strength of type 304 stainless steel is 38,000 psi and the ulti-
~
mate strength is 68,000 psi (16). Thus at a 1'40"C cladding temperature a 1078 psi hydrogen pressure would produce a .2% deformation in the cladding and the maximum allowable hydrogen pressure would be 1929 psi. The hydrogen-to-zirconium ratio in FLIP fuel has a nominal value of 1.6 and a r3ximum value of 1.65. AdetailedanalysisperfonnedbyG.A.(17) Stress to produce a .2% deformation 2132 256
64 indicates that tie equilibrium hydrogen pressure in a U-Zrii l.65 fuel rod which is constant temperature over the entire volume of the fuel is given by 9 4 Pe = 2.59 x 10 EXP[-1.997 x 10 /(T = 273)] where Pe is the equilibrium hydrogen pressure in psi and T is the fuel tem-perature in degrees centigrade. Solving this equation for the maximum allow-able hydrogen pressure of 1929 psi produces a maximum allowable unifonn fuel and temperature of ll42*C. The equilibrium condition pressure defined above never occurs, however, in the real case because a fuel rod is not at constant temperature over the whole volume of the rod. Consequently, the hydrogen pressure will be much lower than the equilibrium value calculated from the maximum fuel rod tem-perature. The axial power distribution along a typical TRIGA fuel rod varies from P max in the center to about .63 P max at the end of the fuel region (1). If we make the conservative assumptions that the fuel rod temperature does not vary in the radial direction and that the axial fuel rod temperature dis-tribution follows a the fuel rod power distribution, then the average fuel tem-perature equals the maximum fuel temperature divided by 1.2. Under these conservative conditions the maximum allowable fuel temperature under steady state conditions that does not exceed the yield strength of the cladding is 1370 C. In addition to the steady state case, the effects of the transient fuel temperature increase during pulsing must be considered. A detailed analysis performed by G.A. (17) has shown that the hydrogen pressure during pulsing increases to a maximum valt.e of 22% of the equilibrium value in about .3 seconds and then falls off. Under pulsing conditions some film boiling 2132 257
65 occurs and the peak cladding temperature is greater than under steady state full power operation. A conservative value of 500 C is selected as the maximum cladding temperature under conditions of film boiling. The ultimate strength of type 304 stainless steel at 500*C is 57,000 psi. At this temperature the maximum allowable hydrogen pressure is 1617 psi. A peak transient fuel temperature of 1290*C would be required to pro-duce this pressure during pulsing. As a safety limit the peak adiabatic fuel temperature occurring during the pulse mode of operation for U-Zr l.65 is selected to be 1150 C. The peak hydrogen pressure that would result at this temperature is approximately 460 psi. This would produce a stress of 16,0C0 psi which would not exceed the ultimate strength of the cladding at a temperature of 700 C. Actual measurements of the peak hydrogen pressure during pulsing have been made at G.A. Five special instrumented fuel rods were tested in the ATPR during these experiments that involved a total of 426 pulses. The maximum peak fuel temperature during these tests was 1175*C and the maximum observed peak transient hydrogen pressure was 41 psia. The peak pressure transient occurred during the first pulse and decreased to about 20 psia by the 220th pulse. These experiments clearly indicate that the actual peak hydrogen pressure during pulsing is at least a factor of ten below the cal-culated values. The 1150 C fuel temperature safety limit for TRIGA Reactor FLIP U-ZrH l.65 fuel is seen to be a conservative limit. This limit will preclude a fuel cladding failure due to the internal hydrogen pressure in both the pulsed and steady state modes of operations. Furthermore, TRIGA fuel with a hydrogen-to-zirconium ratio of 1.65 has been pulsed to a temperature above 1150*C without damage to the cladding. 2132 258
66 References
- 1. " Safety Analysis Report for the Torrey Pines TRIGA Mark IV Reactor,"
GA-9064, Gulf General Atomic, January 5, 1970.
- 2. " Annular Core Pulse Reactor," General Dynamics, General Atomic Division Report GACD 6977, Supplement 2, 9/30/66.
- 3. " Amendment II to the Safety Analysis Report - Texas A&M University Nuclear Science Center," November 1,1972.
- 4. " EXTERMINATOR-2: A FORTRAN IV Code for Solving Multigroup Neutron Diffusion Equations in Two Dimensions," ORNL-4P78, Oak Ridge National Laboratory, T. B. Fowler, M. L. Tobias, and D. R. Vondy, April 1967.
- 5. " Theory of Methods used in GGC-Y Multigroup Cross Section Code, Gulf General Atomic Report GA-9021, October 1968.
- 6. " SUMMIT: An IBM-7090 Program for the Calculation of Crystalline Scattering Kernels," General Atomic Report GA-2492, February 1, 1962.
- 7. "THERMIDOR: A FORTRAN II Code for Calculating the Nelkim Scattering Kernel for Bound Hydrogen," General Atomic Report GAMD-2622, November 10, 1961.
- 8. Private Communication with D. E. Feltz, Assistant Direcbr, Texas A&M University, Nuclear Science Center.
- 9. "PLDEOC: A FORTRAN Code for Calculating Power sensity, Fuel Temperature, and Plotting Neutron Flux Profiles for the W.S.U. TRIGA Core,"
unpublished report by T. R. Evans and W. E. Wilson, Nuclear Radiation Center, Washington State University,1973.
- 10. " Safety Analysis Report for the Illinois Advanced TRIGA," University of Illinois, August 1967.
2 62 259
$ 6/
- 11. "Results of Critical Test Program for Loading Number III, Nuclear Science Center, Texas Af41 University," November 1973.
- 12. " Summary of TRIGA Fuel Fission Product Release Experiments," Gulf-EES-A10801, September 1971.
- 13. " Energy Release from the Decay of Fission Products," Nuclear Science and Engineering, 3, 726 (1968), J. F. Pecking and R. W. King.
- 14. "Safeguardt Report, Washington State College Research Reactor," March 1955.
- 15. " Safeguards Sui. mary Report for the TRIGA-FLIP Paactor at the Puerto Rico Nuclear Center," PRNC-123, Revision C, November 11, 1969.
- 16. Reactor Handbook, Vol. 1, p. 569.
- 17. " Safety Analysis Report for the Romanian Annular Core Pulsing Reactor",
G.A. Report No. E-ll7-323, Vol. III. 2132 260 ,
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