ML19323F716
| ML19323F716 | |
| Person / Time | |
|---|---|
| Site: | 05000054 |
| Issue date: | 05/31/1980 |
| From: | UNION CARBIDE CORP. |
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| Shared Package | |
| ML19323F693 | List: |
| References | |
| NUDOCS 8005290371 | |
| Download: ML19323F716 (5) | |
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{{#Wiki_filter:.. __ _ ENVIRONMENTAL IMPACT ASSESSMENT UNION CARBIDE RESEARCH REACTOR (UCNR) INTRODUCTION The Union Carbide Research Reactor (UCNR) is a 5-megawatt swimming-pool-type reactor located on an industrial site in the sparsely-settled Sterling Forest development region of Orange County, New York State. The UCNR began operation in September 1961 (first criticality) and has been operating on varying but regular schedules ever since. Current operation is around the clock with a duty cycle of about 97%. The reactor is used principally as a source of neutrons for activating target materials used in producing radioisotopes for medical applications. This Assessment addresses the possible environmental impacts of continued operation of the reactor in the next 20-year period. The effects of normal operation are treated first, followed by abnormal operation, and concluding with alternatives to continued operation. ENVIRONMENTAL EFFECTS OF OPERATION
- 1. Direct Radiation Within the reactor restricted area, all zones in which nuclear radiation exists are posted and classified in accordance with 10CFR Part 20 regulations. These zones are surveyed regularly. The nearest reactor Radiation Zone (> 2 mR/hr) is about 110 meters from the unrestricted area (public highway). The external radiation dose rate from the reactor at this area is completely negligible.
- 2. Gaseous Effluents
- a. Wet steam:
A forced-draft cooling tower in a secondary (non-radioactive) cooling loop is used to dissipate the 5 megawatts of thermal energy generated by the reactor. Approximately 2000 gallons of
- water per hour is evaporated and discharged to the atmosphere.
No adverse effects from this discharge have ever been observed nor are they expected. Whenever fogging from this source has occurred, it has been confined to the cooling tower locality within the site boundary. 8006290371
f
- b. Airborne radioactivity:
The principal gaseous radioactive effluent from the reactor is argon-41, formed through neutron-activation of air. Minor amounts of other species may also be formed, but in amounts much less than for argon. The impact of argon-41 in the unrestricted area has been analyzed in the Safety Analysis Report (Reference a, Appendix 2, Section Cl). For 5-MW operation at 100% duty cycle, the yearly-averaged dose rate at the site boundary is estimated as 0.4 mR per year, an amount much less than natural background. An estimate of the 50-mile radius population dose from argon-41 (1200 curie:,/ year) utilizing EPA methodology (References b, c) gives a total of only 1.2 man-rem per year. The small magnitude of the environmental impact of airborne radionuclides is confirmed by the results of local environmental monitoring. These results are several orders of magnitude below Part 20 standards for unrestricted areas.
- 3. Liouid Effluents
- a. Radioactive liauids t
All aqueous radioactive liauids generated in the reactor are collected and treated by distillation. Included in liquid waste t is that from the occasional regeneration of ion exchange resins used for reactor primary water cleanup. The clean distillate (water) is sampled and tested for freedom from radioactive content before discharge to the plant drainage system. The concentrated bottoms containing the radioactive material are solidified and treated as solid waste (see below).
- b. Process water The principal waste water (non-radioact.ve) effluent is that from cooling-tower blow-down, to remove dissolved salts that would otherwise accumulate as the water is evaporated. This effluent consists of about 40 gallons of water per hour containing a small amount of an EPA-approved additive to retard precipitation. This is discharged into the plant storm sewer that flows into Indian Kill lake. No adverse impact has been observed for this small addition of waste water. This discharge is approved under a N.Y.
State D.epartment of Environmental Conservation permit.
- c. Sanitary: 1 The reactor staff (operations and administration) numbers only 30 people out of a total site population of 266. The contribution of reactor and related personnel to this effluent load is only about 10-20%.
i *
- 4. Solid Wastes
- a. Radioactive:
The principal sources of solid, or solidified, radioactive wastes are evaporator bottoms from liquid waste, including liquid waste from regeneration of resins (see above), and spent resins. Included also are other miscellaneous materials incident to the operation of the reactor, for example, sample holders, gloves, and paper. All such material is packaged- in accordance with federal regulations and shipped to a licensed burial ground. No disposal of solid waste is permitted on site.
- 5. Transportation
- a. Radioactive waste material:
Solid wastes from reactor operations require about two 55-gallon drums to be shipped to a burial ground each month. Such shipments are made in accordance with appropriate federal regulations,
- b. Spent fuel:
At a power level of 5 megawatts and 100% duty cycle, a total of about 36 fuel elements of all types would become spent each
. year. It can be expected that at intervals of about two years, a series of spent fuel shipments would be made to the Savannah River reprocessing facility. Three shipments, each comprising about 24 elements, would be made over a 2-3 week period. Such shipments are made in a specially-licensed fuel cask and in accordance with 00T and NRC regulations. The precautions taken are such that the environmental impact of such shipments is negligible,
- c. Reactor operating personnel:
The number of such people is a small fraction of the total plant staff, many of which are engaged in other activities. The impact ' of reactor personnel on transportation and access routes is minimal.
- 6. Abnormal Operations In Appendix 2, Section C2 of the Safety Analysis Report (SAR), a Design Basis Accident (DBA) is postulated. This accident is one with postulated _ consequences in excess of any considered credible. The results of the DBA are given in the SAR. The dose results at the boundary of the exclusion area and low population' zone are much less than the guidelines of 10CFR Part 100. Specifically, the 2-hour doses are only l~-3% of the Part 100 standards.
a-s l
- 7. Non-renewable Resources
- The principal non-renewable resource consumed through reactor operation is the uranium-235 contained in the fuel elements. This consumption rate is 2.2 kg per year.
- 8. Alternatives for Operation The most important and beneficial result of continued operation of the reactor is to provide irradiated targets from which are extracted the most frequently and widely-used radioisotopes in nuclear medicine. Examples of these are 99Mo/99mTc, 133X e, 1311 ,
125 I This reactor is the only commercial domestic supplier of the first three of these in the high specific activity grade required.
- a. Irradiation of targets might be dona in other reactors. There is no domestic commercial reactor available to do this. One government-owned reactor at Oak Ridge National Laboratory does have the capability of providing the required irradiation service. There would be potential environmental impacts involved in the shipment of enriched uranium targets to Oak Ridge and of highly-radioactive irradiated targets back to Sterling Forest for processing. Reliability of radioisotope supply to the medical community would be reduced somewhat, due to the additional transportation steps.
- b. ' Semi-processed radioisotopes might be obtainable from other suppliers and final purification done at Sterling Forest. There is no domestic supplier at present that could fulfil these needs. Unless one could be developed, which would probably involve a government-owned reactor and isotope-processing facility, the supply would have to be imported from foreign sources. It is by no means certain that a foreign supplier would be willing or able to do this.
- c. Finished radioisotopes might be obtained by present users (radiopharmaceutical manufacturers, nuclear pharmacies, hospitals, etc.) directly from suppliers of the purified radioisotopes. Again, no domestic supplier exists, but would have to be developed probably at a government-owned facility.
Otherwise, all supplies would have to come from foreign sources. Total dependence on the latter could have adverse environmental impacts in the domestic medical and health care field.
- d. Each of the above alternatives would also require the decommissioning of the reactor. The cost of decommissioning is estimated elsewhere in this renewal application. There are also potential environrental impacts related to the dismantling, shipment, and burial of the radioactive components and the decontamination of the immediate site.
, ..v -
- 9. References
- a. Safety Analysis Report, Union Carbide Research Reactor (UCNR),
May 1980.
- b. J. A. Martin, Jr., et al, "A Computer Program for Calculating Doses, Population Doses, and Ground Depositions due to Atmospheric Emissions of Radionuclices", U.S. Environmental i Protection Agency, EPA-520/1-74-004 (May 1974).
- c. Holzworth, George C., " Mixing Heights, Wind Speeds, and Potential for Urban Air Pollution throughout the Contiguous United States",
U.S. Environmental Protection Agency, Office of Air Programs, Report AP-101 (Jan. 1972). l
APPENDIX 1 FUEL OdSIGN
- 1. General ,
The type of fuel core that has seen most service in the UCNR is uranium-aluminum alloy, which has a long history of satisfactory opera-tion in research and test reactors using plate-type elements in which the fuel core is clad with aluminum. Two other types of fuel core that have been demonstrated to give equally satisfactory performance and which are viable alternatives to uranium-aluminum alloy are uranium oxide-aluminum powder compacts (or cermets) and uranium aluminide (UALx)-aluminum powder compacts. These two types are discussed below. Fuel plates made to the same dimensions and configuration with each of these three types of core material will not differ significantly in overall thermal or neutronic performance. In fact the better process control that can be exercised during manufacture of the oxide (U33 0 ) and aluminide (UAlx) compacts results in cores that are more homogeneous and less likely to exhibit clad bonding problems than alloy (U/A1) cores.
- 2. Uranium-Aluminum Alloy .
The composition of these alloys is limited by metal-working consider-ations to about 25 wt. % uranium, with 18-22 wt. % being the normal range. There have been many years of experience with U/Al-alloy fuel cores in low and high power research reactors. Typical of the latter are MTR, ORR, GETR, ETR, and ATR. Burnups in the range 50-60% are routinely attained for elements in the ORR, which corresponds to an average fission density of about 8 x 1020 fissions /cc. Fueled control elements can attain'up to % 100% burnup, or 1.5 x 10 21 fissions /cc. For UCNR fuel elements the geak fission density in fissions /cc is given approximately by 1 9 x 10 51 x 23)U Burnup (%). 3 Uranium Oxide-Aluminum Compacts (U308-A1) Cores of this type have been used in the form of plate-type elements in the high flux isotope reactor (HFIR) at Oak Ridge National Laboratory for the past 12 years with highly satisfactory results. In this period a total of over 80,000 plates has been used in the thermal and hydraulic environment peculiar to this reactor, an environment more rigorous than in most research reactors, including the UCNR. The manufacturing process for these cores is described in Ref.1, a copy of which is also enclosed. The reference also describes the experience with these cores, their thermal characteristics, the operating characteristics l l l L
9 of the HFIR, and also discusses the safety-related questions that arise in the use of this type of fuel. The latter include (a) the physical behavior under irradiation, (b) exothermic reactions, and (c) fission product release under accident conditions, it is concluded that plate-type research reactor fuel assembled using U3 08-Al cermet cores is equal to or super to U-Al alloy fuel and can be used without loss of safety margin. While #6061 Al is used as cladding for HFIR elements, #1100 Al can continue to serve as the cladding for UCNR elements in view of the much lower hydraulic forces resulting from the very low UCNR core pressure . drop (table below). Therefore UCNR elements using U3 08-Al cores can be identical physically and dimensionally with the current U-Al elements. For comparison purposes, typical operating characteristics of the HFIR and the UCNR relevant to the thermal and hydraulic performance of j U3 03 cermet fuel, are given in the following Table: Parameter HFIR UCNR Thermal Power (MW) 100 5 Peak Heat Flux (W/cm2) 630 40 Peak Fuel Temp. (oC) 330 95 _ Peak Fission Density (fiss/cm3) 1 9 x 1021 1 x 1921 Plate Thickness (mm) 1.27 1.27 Fuel Core Thickness (mm) 0.762 0.508 Core U3 0g Content (wt %) 41 22 Coolant Velocity (m/s) 12.8 1.2 Core Pressure Crop (kg/cm2) 4.8 0.05
~
- 4. Uranium Aluminide-Aluminum Compacts (UAlx-A1)
Reference 2 presents a thorough survey and analysis of the use of UAlx-Al cores (as well as U3 08-Al) in plate-type reserach and test reactor fuel elements. This survey will not be repeated in detail. The UAlx type has been used successfully in the MTR, ETR, ATR, MURR, and MITR reactors. In the ATR some 20,000 such plates have been irradiated under conditions (temperature, flow rate, and thermal flux) that are much more severe than in lower power research reactors (including the UCNR). In comparison with U-Al fuel it is found that UAl x fuel is superior in its retention of fission products, and that swelling during irradiation is less. Blister temperatures, which vary with irradiation, exceed 4000C which is well above the UCNR maximum plate temperatures of 95 C (sect. 2,above). As with the U 330 -Al cores discussed above, fuel elements made with UAix -Al cores can be identical physically and dimensionally (except for
the core) with current UCNR fuel . No reactor physics changes will occur as a result of the substi tution of the aluminide for the alloy. Thermal characteristics will also be identical, except for an insignificant I (fraction of loc) rise in internal plate temperature. Fission densities attainable with aluminide fuel exceed 2 x 10z1 fission /cm3 1 5 References (1) Binford, F. T. and Knight, R. W.: "The Use of U33 0 -Al Cermet Fuel in Research Reactors", Transactions of A.N.S. Vol. 27, 834-835 (1977). Copy of this reference is enclosed (by permission). (2) Safety Analysis: Utilization of Intermetallic' Uranium Aluminide (UA1, 3 VA14, UAl2) and Uranium Oxide (U30g, UO2) Cermet Fuel Cores in the Ford Nuclear Reactor, University of Michigan 4 (June 1977), License R-28, Docket 50-2. s I e l l 1 1 I l l l I
REFERENCE I Trs actions of American Nuclear Society, Vol. 27, pp. 834-835. (1977 Winter Meeting)
- 8. The Use of U 0s-Al 3 Cermet Fuel in Re-search Reactors, F. T. Binford, R. W. Knight .
MM . There are three areas of interest related to safety 2 Because there are at present no commercial manu- (a) the physical behavior of the UaO. cermet under facturers in the United States regularly producing urant- Irradiation conditions, (b) the possibility of exothermic um-aluminum alloy plate-type reactor fuel, it may be chemical reactions, and (c) the potential for fission necessary for many research reactor operators to obtain product release under accident conditions, fuel assembled using a different type of fuel core. Experimental results obtained by Martin et al." indi-Two viable alternatives to the use of alloy, which permit the fabrication of plate-type elements, are avail- este that dispersions of U 0. perform quite well to burnup able. levels of 1.5 x 10 to 2.4 x 10" fiss/cm' at irradiation temperatures of 60 to 120*C. They found that radiation-One of these employs the so-called UAlxcores which induced swelling was quite small and actually decreased have been used successfully at the advanced test reactor with increased U 0, loading. Postirradiation examination (ATR) at Idaho Fal'.s since 1967. - showed no indication of blisters, core-cladding separa-In the second alternative, which has been used to tion, matrix cracking, or other types of structural , produce fuel for the high flux isotope ructor (HRR) at defects. Operating conditions in the HFIR would imply an the Oak Ridge National Laboratory since 1965, the fuel average burnup of about 5.3 x 10 fiss/cm' with a peak of cores are assembled by powder metallurgy and are about 1.9 x 10". The temperatures involved range from 114 to *30*C. composed of a UaO.-Al cermet. It is the purpose of this paper to briefly describe the manu'acturing process, An exothermic reaction involving U30, and Al to summarize our experience with U 0.-Al fuel, and present produce UOn, U-Al solutions, and A1:03 is thermo-physical data which may be useful to those who wish to consider shifting to this material. dynamica!!y possible. This reaction has been studied by a number of investigators.* It was found that no signifi-The U30, currently used is a high-fired oxide which is cant reacton occurs at temperatures below about 950*C. calcined at 1350 to 1400*C. The material is supplied at The energy release may be as high as 800 J/g of U 0,;
-100 +325 mesh with 25% fines and is blended with however, high energy releases occur only when the U O.
aluminum powder to yleid the required 8 U content. The I ading is on the order of 60 to 75 wt%. At 30 wt%, the blended powder is compacted at ~22.5 tsi, vacuum de- self-heating was found to be negligible. gassed, and then inserted into an aluminum picture frame. The aluminum _ covers are welded on and the , assembly hot rolled To producOhe fuel plate. , The dimensions and configuration of the oxide plates normal Characteristics of U,0,.Al Ful Corn can be made identical to those of fuel plates made by the U,os-A2 Cermet Corn alloy process. Hence, there are no basic differences in e,,,,, the overall thermal performance relative to heat removal ver% .t% Desirva un eb at the surface. Because of the better process control that Uso, U,0 (s/cm') s u/ cia" (w/em c) p/s c)
. can be excercised during the manufacture of oxide fuel, .
,,,g.
these fuel cores are more homogeneous and are less . likely to exhibit significant areas of nonbcnifing at the 13.4s 31.s1 s.20s o.sss 1.72 0.?* fuel-cladding interface than are alloy cores. nos 43.2r s.424 a.254 1.as aso McElroy' has measured values of the thermal con. m-rirw odd. ductivity for several U,0.-Al cores, and these are 12.45 so.s2 . s.3:s- o.ses 1.tv o.ts presented in Table I. is.os 4 .as 3.5s0 8.2s4 1.se o.st - Through June 1977, 148 reactor cores have been "Me===rw essen.' consumed in the HRR. They contained a total of 25,308 ,beajc.tatw bas.4.. ..r nta sracu.na. plates containing 30.2 wt% U 0 and 54,612 plates con- u ,: eass s uf, ,= ,,,,,,,,,,,,, , ,,,,,,,,,g,,3,,,,,,,,,, ,,, ,, talning 40.1 wt% U2 0.. The performance of these plates state core toadw to b e is s u/piat.: 3.254 s u/ca' corresponds to in a thermal and hydraulle environment considerably ancus 23 ss u/otat,. more rigorous than that found in most research reactors - has been highly satisfactory. Since operation began in 1965 only one fission product leak has been detected. Relevant operating parameters for the HHR' are given l's Table !!. _4 O
r _ e t ., TABLE II Operating Characteristics of the HFIR Av. Max Powerlevel(M%1 100 --- Power density (MW/ litre) 1.92 4.26 Heat flux (W/cm') 243 630 , Fuel temperat re (*C) 114 330 t Fisston derv ' r1 (fiss/cm') 5.3 x 10'* 1.9 x 10 Plate thickr .s (mm) 1.27- --- Fuel core 5 ekness (mm) C.762 --- Coolant et .nel thickn 1.27 --- Inlet pressure (kg/cm,ess-g) (mm)45.7 ---
- 1. D. L. McELROY to R. W. KNIGHT, ORNL Internal Correspondence (Jan. 5,1971).
Core pressure drop (kg/cm') 4.8 --- Coolant velocity (m/s) 12.8 --- 2. F. T. BINFORD and E. N. CRAMER, Eds., "The Eigh Flux Isotope Reactor, X Functional Description,"
"Dased on 2300-mwd operating cycle. ORNL-2572 (1964).
"' ~~'
- 3. M. M. MARTIN, A. E. RICHT, and W. R. MARTIN,
" Irradiation Behaviour of Aluminum Based Fuel Dis-persions," ORNL-4856 (1973). .
In a later series of experiments performed in the 4. T. J. THOMPSON and J. G. BECKERLY,c Eds., The TREAT facility, which simulated actual accident condi- Technology of Nuclear Reactor Sa/ety, Vol. 2, MIT
' tions in the HFIR, the results indicated that even with Press Cambridge, Mass. (1973); L. BAKER, Jr. and 41 wt% HFIR fuel the U 0.-Al reaction is not an important R. C. LIMATAINEN, " Chemical Reactions," Chap.17.
energy source. Since most research reactor fuel that employs highly enriched uranium contains less U 0. than 5. R. O. IVINS and F. J. TESLA, " Studies with Aluminum-this, this t eaction does not constitute an additional safety UaO. Cermet Fuel in TREAT" in Argonne National problem. Laboratory, Chemical Engineering Division Semi An-
, Juh-Dec., N, M-M W M.
There have been very few studies of the release of fission products from oxide-aluminum fuels. Creek' 6. Ibid, G. W. PARKER and C. J. BARTON, " Fission melted trace-irradiated UOn-Al in air and obtained an Product Release," Chap.18. average release of 5.6% of the rare gases and 0.003% of the lodine. Experiments reported by Parker and Barton' for U-Al alloy samples held above the melting tempera-ture for 10 to 17 min in steam-air mixtures indicate a release of virtually all of the rare gases and from 27 to . 97% of the lodines depending on the maximum tempera-ture achieved. ' Because these two investigations were conducted under quite different conditions, it is not possible to definitely conclude that the fission product release from oxide cores will be less than that from U-Al alloy under accident conditions. However, the alloy melts at about 850*C, whereas the Ui O. melting point is 1250*C In either cast, the aluminum cladding melts at about 650*C. It would therefore appear that the fission product release from a gross cladding failure would be no greater in the case of the oxide than in the case of the alloy. . Based on the foregoing, it may be concluded that l plate-type research reactor fuel assembled using U O.-Al I cermet cores is equal or superior to U-Al aUoy fuel and can be used as an alternative without incurring any loss in safety margin. The same can probably be said for the U-Al xcores. Thus, the choice can be based on avail-ability and cost. l l 1
~1 D R l
l APPENDIX 2 SECTION A: Thermal-Hydraulic Safety Analyses SECTION 8: Reactivity Transients SECTION C: Radiological Safety Analyses (The material in this Appendix was formerly entitled SUPPLEMENT NO. 2 TO FINAL HAZARDS
SUMMARY
REPORT and was dated April 1977.)
? ?
TABLE OF CONTENTS Page SECT. A. THERMAL-HYDRAULIC SAFETY ANALYSIS . . . . ........... I 1.0 BURN 0UT HEAT FLUX FOR FORCED COOLING MODE . .......... 1 . 1.1 Burnout Heat Flux . .. . . . . . . . . . ........... 1-2 1.2 Ef fect of Coolant Bulk Temperature Change . .......... 3 13 Peak Heat Flux per MW . . . . . .. . . . ........... 3 1.4 Burnout Heat Power vs. Flow-rate . . . . ........... 4 1.5 Effect of Change in Pool Temperature . . ........... 4 1.6 Safety Limit on Power Transients . . . . ........... 5 1.7 Hot Channel Factors . . . . . . . . . . . ........... 5 2.0 BURNOUT HEAT FLUX FOR NATURAL CONVECTION MODE . ........ 6 2.1 Burnout Heat Flux . .. . . . . . . . . . ........... 6 3.0 INCIPIENT BOILING IN FORCED COOLING MODE ........... 13 3.1 Basic Relations . . . . . . . . . . . . . ........... 13-14 3.2 Evaluation of Heat Transfer Coefficient (h) . . . . . . . . . . 15 33 Temperature Margins for Non-Boiling . . . ........... 15-18 3.4 Comparison with Boiling Experiment . . . ........... 18-20 l 35 Fuel-Plate Surface Temp. , Normal Operating Conditions . .... 20 4.0 INCIPIENT BOILING IN NATURAL CONVECTION MODE ......... 21 _
5.0 REFERENCES
FOR SECTION A, THERMAL-HYDRAULIC SAFETY ANALYSIS . . 22 i SECT. B REACTIVITY TRANSIENTS . . . . . . . . .. ........... 23
- 1. Effective Delayed Neutron Fraction . . . ........... 23
- 2. Inhour Relation for UCNR . . . . . . . ............ 23-24 3 Step Reactivity insertion Transient . . . ........... 24-27
- 4. REFERENCES FOR SECTION B, REACTIVITY TRANSIENTS . ....... 28
' t I
d 4 TABLE OF CONTENTS -(cont'd.) i i 1 Page i SECT. C RADIOLOGICAL SAFETY ANALYSES .. . .............. 29 _
- l. Radiation Dose at Site Boundary . . .............. 29-30
- 2. Design Basis Accident . . . . . . . .............. 31-37 3 Fueled Experiments . . . . . . . . .............. 38-43
- 4. Effluer.ts in Unrestricted Areas . . .............. 44-54 5 REFERENCES FOR SECTION C, RADIOLOGICAL SAFETY ANALYSES .... 55 I
] 1 I i i i e I A 1 6 1 I. l t I t-I l. t ., __ . _ _ ._ , . __ _ _ . _ _ _ _ . ._ _ _ _ _ _ . _ - _ _ , . _ _ . _ . _ _ , _ . . _ _ _ .
4 9 1. A. THERMAL-HYDRAULIC SAFETY ANALYSIS 1.0 BURNOUT HEAT FLUX FOR FORCED COOLING MODE 1.1 Burnout Heat Flux Forced cooling in the UCNR results from gravity-induced flow of . primary cooling water from the open pool down through the core into the underground holdup tank. The superposition method of Gambill(l) for a forced sub-cooled water system is used to derive the burnout heat flux. His relationship, given below, is valid over a range of conditions (coolant velocity, pressure, subcooling, and hydraulic diameter) that include those values typical for the UCNR. Ypool = 0.1 A pv.5 (o 3p).25 g.5 [1 + (pg/py).75 CpATsub/9.8 A] This term represents the contribution from pool boiling and generally is the predominant one for UCNR conditions. It should be noted that Gambill's factor gc a (his equation 8 in ref.1) is equal to g 2 for normal gravity. ' Ynb = F (Tw - TB ) y.8/De .2, where F E 21.9 k2/3 Cpl /3 pg.8fy.467, with the units for V and De in ft/sec and in., respectively. The term Ynb is the contribution from forced convection in the absence of boiling. It is noted that, while Gambill's(1) Fig. 1 is correct, the numerical factor in his formula for F is incorrect. For the UCNR this contribution is about 15% of that from pool boiling. The burnout heat-flux is the sum of the above two contributions. Y BO " Yp ool + Ynb-For evaluating YBO, the following values (2) (3) are used, with coolant physical properties evaluated at Tsat: I
4 f 2. 1 1 Ty = fuel plate wall temp., F TB = coolant bulk temp., F Tsat = saturation temp. at core pressure (23.5 psia) = 2370F Tp = pool temp. , l sub " sat - B i D = coolant channel hydraulic diam. = .236 in. = .0197 it.
-T w B = AT + AT AT =T -T = 610F (Ref. 1, Fig. 2) sat w sat V=coolantspeed=1.626x10-36ft./sec.
pg = water density = 59.2 lb./ft. py = vapor density = .0582 lb./ft - A = latent heat = 954 BTU /lb. i o = surface tension = 3.8 x 10" lb./ft. Cp = specific heat = 1.012 BTU /lb. ap=pg-py= 59.1 lb./ft. g = acceleration of gravity = 4.17 x 10 ft./hr. k = thermal conductivity = .396 BTU /hr.ft. F 4 p = viscosity = .604 lb./hr.ft. 4 F = 267, calculated from above values - The foregoing values apply to a core at 20-fc. depth, containing 30 fuel elements (each control element counts as 1/2 element) and ) experiments equivalent in flow to 6 fuel elements. Each element
-3 2 1
contains 17 coolant channels, each 2.24 x 10 ft. . Total flow area is 36 x 17 x 2.24 x 10-3 = 1.37 ft 2, i 4
- , - - _ , m - -
. . _ - . .-- .- - - . - - ~.
I f 3. 1.2 Effect of Coolant Bulk Temperature Change: For a valid application of the expression in 1.1, above, it is necessary to estimate the temperature of the bulk coolant at the burnout This is taken in the hot channel, using the appropriate power (PBO). hot channel factor (Sect. 1.7). P = core thermal power, MW 1 Q = coolant flow rate, gpm F = hot channel factor = 1.66 Tp = pool temp., F AT = average temp. difference thru' core, F
~4
= P/ (1.456 x 10 Q)
TB=c lant bulk temp. in hot channel at hot spot Thus: 4 [ TB = Tp + F *AT = Tp + 1.14 x 10 P/Q. 1.3 Peak Heat Flux per MW: A conservative situation is considered in which a new fuel element
; is located next to a water-filled sample position in the core central region.
! Peak / average heat flux = Peak / avg. element power x peak / avg. axial flux , l = 2.55 x 1.30 = 3.31 4 For a 30- element core the total heat transfer area is: A = plate surface area x total no. of fuel plates s 2 2
= .93 ft. x 30 x 16 = 448 ft . ^l The peak heat flux /MW of reactor thermal power is therefore:
6 4 2 4 Yp = 3.31 x 3.414 x 10 /448 = 2.52 x 10 BTU /hr.ft. per MW { s-. m. , , _ , - , . , . . -
. _ .___=__ __.. _ . ._ _ ._ __ _ ._ ' t A
4. i 1.4 Burnout Heat Power vs. Flow-rate Combining the above expressions, with the physical properties given, yields the following relation: PBO = thermal power at burnout, MW 4 PBO " IYpool + Inb)/2.52 x 10 I
= 12.84 [1 + 1.95 x 10-2 (237-Tp - 1.14 x 104 PBO/ 3 + l 8.31 x 10-5 (g).8 (298 - Tp - 1.14 x 104 PBO/ )
i Evaluating PBO for various practical values of flow-rate (h) at a pool temp. (Tp) of 1200 F, gives the results:
- Primary . Burnout Safety Limit h/P gn Flow-rate (Q) Power (PBO) Power (PsL)
(qpm) (MW) (MW) (qpm/MW) 1800 17.2 14 131 j 2000 18.5 15 135
- 2200 19.6 16 140
.) 2400 20.8 17 144 1 The third column defines a " safety limit" power (P;g, as the value of PBO reducad by the factor 1.25 and rounded to the nearest whole MW. This factor allows for uncertainties in the Gambill burnout correlation employed in this analysis and establishes the safety margir.. For pool temperatures less than 1200F the values of burnout power are greater than given above, vir.., a change of approx. 1 MW for each 100F change. 1.5 Effect of Change in Pool Temperature (Tp) : As Tp increases, tha major effect will occur through decrease in AT sr o-This will decrease.Y pool leaving Ynb almost unchanged. For example an , increase in pool temperature from 120 to 1300F will decrease P BO by only about 1 MW. Adopting a high pool temperature, as in Sect. 1.4, is conservative.
' r 5. 1.6 Safety Limit on Power Transients: P ma e a e s for steady-state conditions. For a transient BO power " burst", the power below which fuel cladding failure would not occur could be considerably higher provided the total energy release is less than a certain maximum. Results of Sport IV tests (Ref. 10) show that this maximum energy is 12 MW-sec. 1.7 Hot Channel Factors: The hot channel, hot-spot factors used in this analysis are tabulated below. F = factor for bulk temp. rise to hot spot in hot channel = 1.66 F fa tor for peak heat flux at hot spot vs. core average = 3.3 2 F = factor for film temp. rise at hot spot = 3.4 3 1 2 3 Nuclear: Radial neutron flux distribution : 1.7 1.7 1.7
" " " 1.3 1.3 Axial :
Radial fuel distribution : 1.4 1.4 1.4 Bulk temp. rise to hot spot : 0.6 Product of nuclear factors : 1.4 3.1 3.1 Engineering: Channel gap : 1.07 1.01 Pressure distribution : 1.01 1.01 Power measuring error : 1.05 1.05 1.05 Fuel concentration : 1.02 1.02 1.03 Product of eng. factors : 1.16 1.07 1.10 Total Product : 1.66 3.3 3.4
' t
- 6. i 2.0 BURNOUT HEAT FLUX FOR NATURAL CONVECTION MODE i
l 2.1 Burnout Heat Flux The semitheoret.ical prediction (or homogeneous) method of Gambill I and Bundy(4) has been selected as rest applicable to the UCNR. This method has been used successfully to predict natural convection burnout . for the ORR and HFIR reactor fuel. The former is close in design to UCNR fuel and the method is valid for the physical and mechanical parameters of the UCNR, which are: Li = fuel inlet section length = 8.3 in. Lh = fuel heated length . = 23.5 in. Le = fuel outlet section length = 2.4 in. Lt = fuel total section length = 34.4 in. i De = coolant channel hydraulic diam. = .0197 ft. = .236 in. Ah = heated area = .425 ft
~
ATS = coolant channel area = 2.24 x 10 ft. i p = pressure at core depth = 23.5 psia . T = water saturation temp. = 2370F - sat Tp = pool temp. = 1200F ATsub = Tsat-Tp
= ll70F AH = entropy change (Tsat-Tp) = 117.4 BTU /lb.
sub Hgg
- latent heat (at Tsat) = 954 BTU /lb.
= water density at 1200F =
po 61.7 lb./ft. at osat = water density at Tsat
= 59.16 lb./ft.
1
= 17.18 ft. /lb.
Vy = vapor specific volume at Tsat
= .0169 ft. /lb.
Vg = liquid specific volume at Tsat p = water viscosity at T = .605 lb./hr. ft. sat
, , ,, - - - + , -
- ,- ,-- 9
7. J The iterative procedure of Ref. 4 (pp. 37-40) is now followed, with a recommended modification to allow for the non-uniform heat generation along the fuel element length. Instead of the uniform heat source, a fraction Fx is read from a plot of ATxvs. distance (x) along the element (Fig. 1). AH t = AHsub + Xe*Hfg = 117 + 954 x. 117 x 23.5
- LNB /Lt" (Li + AHsub'Lh/AHt )/Lt = (8. 3 + 117 + 954 x e Lx/Lt "
(Li + Fx Lh l/Lt = (8.3 + 23.5 Fx)/34 4 ATx = (AHsub + xHfg)/AH t = (117 + 954 x)/(117 + 954 x ) px = Vyx + Vg(1-x) = 17.18 x + .0169 (1-x)
#2
=V y x, + Vg (1-xe) = 17.18 xe + .0169 (1-xe)
Assumed x, Value AHt LNB/Et ATx 09 ! 0.5 594 .376 .197 (1 + 8.15 x) .116 0.2 308 .501 .38 (1 + 8.15 x) .58
' O.1 213 .617 .55 (1 + 8.15 x) .29 i ! For each value of xe , px and ATx are now calculated as a function of x, and FxLh is read from Fig. 1 for each ATx so calculated.
Results are i I e 4
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i m-FIG.1
.. . -.. . ~ . .
9. i x =.5 x =.2 xe = 1 x & Fy Lw Ly/L* g Fx.Lh Ly/L* & Fy.Lh Lv/Lt i A
.5 .116 1.0 23.5 .92 j
.4 .145 .84 20. .82
.2 .29 .52 11.8 .58 1.0 23.5 .92 i .15 .386 .84 18.7 .78
.58 .36 .49 .69 15.2 .68 1.0 23.5 .92
.1 8.7
.45 .53 12.0 .59 .77 16.8 .73
.05 1.14 .28 7.2
.21 .41 .41 9.7 .52 .60 13.2 .62 j .01 5.31 5.8 14.6 .39 9.4 .51 .56 12.5 .60
.003 5.5 .40 .38 9.2 .51 .55 12.3 .60 I, 10-3 29.4 .20 10-4 53.7 .40 .60
{ i i
. Lx/Lt is now plotted against px (Fig. 2) and the curves are <
j numerically integrated to give the mean density (pm) . Other quantities are as follows: 4
' f = friction factor (Fig. F-1 in Ref. 2) g = acceleration of gravity = 4.17 x 108 ft./hr.2 = 32.2 ft./sec.2 i ! G2 = 2 g (po-pm) Lt/(fL /D t ce m + 2/p2 .5/po) f iB0 = 2.85 x 105 c.85 o,.65/Lh*
9 " $ Bo*Ah ! gl = 3600 AH *G*ATS t 1 j For each value of x e , the above quantities are: i ( I_ ~ _ . . _ - . _.. ._ - - - _ ,, _ . . _ _ - _ . - . . -
i
- 10. I i
i l 4 4 eu a .,:
,a N f4 r0 C
u a .r , u l
- ?
d
<0 e*
s% rs
! E
- e e l i
)
( ' C X h x x 2- , os 4 :. .: . 4 J I I. o ' N x M. . ~.
.3 l
i ,f
. .4 i . = ax ,
t i I z . . ,x .w. a _ - .+ J
=*x
- J .S \
g . I l . f I r
'?
. e4 o o o O ao J .t u O g ., ,o, .o ,
~ -
1 -
- a u e FIG,2
~11.
- x. = .5 x,= 2 x. = 1 AH t 594 308 213 BTU /lb.
pm 24.7 31.5 36.9 lb./ft.3 (po-pm) Lt 106 86.6 71.1 lb./ft.2 . p2 .116 .29 .58 lb./ft.3 f .05 .042 .038 G 19.7 28.0 35.7 lb./see.ft.2 i 4B0 1.57 x 10 5 2.13 x 105 2.62 x 105 BTU /hr.ft2 q 6.69 x 104 9.05 x 104 1.11 x 105 BTU /hr.ft.2 gl 9.44 x 104 6.95 x 104 6.13 x 104 BTU /hr.ft.2 q/ql .709 1.30 1.81 In Fig. 3, q/q is l plotted against $BO. The intercept of the resulting curve at the peak / average flux value (1.30 from Fig. 1) yields the correct burnout heat flux (VBO) for the UCNR core in netur,al circulation. 5 i YBo = 2.1 x 10 BTU /hr.ft.2 . 5 Hence PBO = 2.1 x 10 /2.52 x 104 = 8.3 MW. For minimum design values of PBO, Gambill and Bundy(4) recommend a de-rating factor of 0.8 for ATsub between 80 and 1300F, which results in a final value of 6.7 MW. for the natural circulation burnout power, which can then be regarded as a " safety limit". f l _- . _, . ~
_.. . m e 12. Y -I'h EITIIN NNN -ICE-hi? I+NE' 'I -h'E hh i N Iri 'iF % i~b'O~+h"+~ I;iC
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---- m t .u. . -::=.~n.:.
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- =: =
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.+ -
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=
13. 3.0 INCIPIENT BOILING IN FORCED COOLING MODE 3.1 Basic Relations In this section, relationships are given between the physical parameters of the coolant system and the temperatures that occur in , the coolant and on the fuel-plate surface due to local deposition of all fission energy. As in the preceding section, a 30-element core with 6 experiment positions, at a 20-ft. depth, is considered as a limiting case. For estimating fuel plate temperatures necessary for incipient boiling the correlation of Bergles and Rohsenow(5) is used. This correlation is a particularly conservative one, so there is no need to allow a wide temperature margin for non-boiling. For calculating heat transfer coefficients the Seider-Tate (6) relation, which takes into account variations in coolant physical properties i when large temperature differences exist, is employed. This relation has been successfully used(7) for the thin rectangular channels of reactor fuel elements. Ac = c re flow area = 1.37 ft.2 (see 1.1, above) Tp = pool temperature, OF p = absolute pressure at core depth = 23.5 psia J
! (ATsatl i " (Twli-Tsat for incipient boiling, OF '
i F1 = hot channel factor = 1.66, (Sect. 1.7) i Tw = fuel plate surface tamperature, OF Tsat
= saturation temperature at core depth = 2370F fi
= heat flux for incipient boiling, BTU /hr.ft.2 AT = bulk coolant temp. rise, OF P = reactor thermal power, MW Q = primary coolant flow rate, gpm (TB)HC = bulk coolant temp. in hot channel at hot spot, OF
- - - .m_
14.
= .0197 ft. (see 1.1, above)
De h = heat transfer coeff., BTU /hr.ft. OF V = coolant velocity, ft./hr. Yp = peak heat flux in hot channel at hot spot,
= 2.52 x 10 4 BTU /hr.ft.2 per MW (see 1.4, above)
The following relationships apply:
* !P 4 Yi = 15.6 p * (ATsatl i (Ref. 5) 4
= 600 (ATsatl i So, (AT sat l i " (Yi /600)
O
= Tsat + (?i/600) * = 237 + (2 52 x 104 P/600)'
(Tw)f
= 237
- 5.75 P*
I P = 1.456 x 10-4 6 AT (TB)HC =Tp + F1 AT (Tw)HC (TB )HC + Y p /h, in the hot channel Thus, (Tg)HC =T p + F1 AT + Yp/h 4
=T p + 1.66 P/(1.456 x 10-4 6) + 2.52 x 10 P/h
=Tp + 1.14 x 10 4 P/6+2.52x104 P/h
=T p+(1.14/h+2.52/h) x 10 4 P and (TB)HC =T p + 1.14 x 10 4 P/h Nu = .027 Re
- P r (U/Pw)
So, h = .027 (D,vp/u)* (Cp u/k)* (u/pw) .14 k/D e
~
~*
= .027 (Vp) 8 (C p/u)
- k* ,eD uw ,
- where u, is evaluated at Tw, and other properties at (TB)HC' V = 2.23 x 10-3 ft.3/sec-gpmxhgpm/l.37ft.2=1.628x10-3hft/sec
=5.86hft/hr
15. 3.2 Evaluation of Heat Transfer Coefficient (h) : The determination of reactor thermal power that results in a desired wall temperature margin below incipient boiling requires an iterative 1 procedure that is complicated by the fact that the coolant physical properties are temperature-dependent. Considerable simplification results by pre-selecting the highest power at which automatic protective action .
; will take place. This selected power is 6.25 MW. The corresponding value ' of the predicted wall temperature for incipient boiling is thus:
(Tw)i = 237 + 5.75 (6.25) * = 250.60F Allowing a non-boiling margin of 50F, the value of uwcan then be specified as 0.58 lb/hr.ft. (i.e., at 245 F). The expression for h can 4 4 now be written as: i
* ~~
h = .027 (5.86 6p)* (Cp/p)* k* ( 0197) ( 58)
= 0 26 p (cp/u)* k h*
= R(T) 6' , where T is the bulk coolant temp.
where R(T) = 0.26 p* (Cp/u)* k* Values of R(T) have.been evaluated for the limited range of bulk temperatures of interest, as follows: T : 160 150 140 130 0F
! R(T): 3.89 3.79 3.68 3.56 a
~
i
- _ 3.3 Temperature Margins for Non-Boiling
The expressions given above allow temperature margins to be calculated 1 for many combinations of power (P), coolant flow rate (Q), and core inlet or pool temperature (Tp). It is sufficient for protective purposes to select a limited number of values of some of these parameters. For example, I- the minimum limit on flow rate is taken as 1800 gpm. Pool temperature is a slowly-varying parameter so typical values between 100 and 1250 F are selected. i t. j l i a
.i
- 16. f I
i l I 4 The temperature margin (M) is given by: I I i M= (Tw)f - (Tw) HC i
= 237 + 5.75 P' - Tp - 10 P (1.14/6 + 2.52/R(T) h*)
i 3.31 TemEera tu re Marg in s __ a t _1800_gemi l At the limiting flow-rate of 1800 gpm, the foregoing becomes: M = 237 + 5.75 P*4 - Tp - 6.33P - 62.7 P/R(T) A table of values for M (to nearest O F) at several elevated pool i temperatures (normal pool temperature is less than 100 F) and reactor
- j. powers is given below. The corresponding value of the bulk temperature at the hot spot, (T *# * * * # * * * * " "^# "
B HC' the proper value of R(T) to be selected. i (T B HC Tp + 7.33 P (at 1800 gpm) . ]; Pool Temp. (Tp) Power (P) Bulk Temp. (T er ,) . Margin (M) (DF) B HC (MW) (OF) (OF) 125 5.25 158 6 i i 120 5.25 153 10 1 120 5.45 154 6 115 5.45 150 10
> 115 5.65 151 6 110 5.65 146 9
) 110 5.8 147 7 i 105 5.8 142 10 105 6.0 143 6 100 6.0 138 10-5 100 6.25 140 5 t i - i a
- + e ,, .-- - -- - ,.-4 -. .r .- < u .~q . . , , * . . . . <p- + -
17. 3 32 Tegegraguge_gaggigg_a3_2199_HEsi The expressions for temperature margin and bulk temperature at thic flowrate, which is close to that normally used in operations, are: M = 237 + 5.75 P*4 - Tp - 5.43 P - 55.4 P/R(T) (T Tp + 5.43 P B HC Pool Temp. (Tp) Power (P) Bulk Temp. (TB'HC *E' '# 9 (OF) , (MW) (OF) (OF) 125 6.0 158 7
" 11 120 153 6.25 154 6 115 149 10 6.45 150 6 110 145 10 6.65 146 6 105 141 10 6.85 142 6 100 137 10 7.05 138 6 h
' 18.
i
- 3.33 Temperature, Margins,at_,2400,gpm
The relevant expressions are: 4 M = 237 + 5.75 P* - Tp - 4.75 P - 49.8 P/R(T) l ( (TB HC " E+ * ' - Pool Tamp. (Tp) Power (P) Bulk Temp. (T Temp. Margin (M) B HC CUF) (MW) (OF) (OF) 125 6.8 157 6 l 120 6.8 152 10 1 6 j 7.05 154 115 " 148 10 7.3 150 6 , 110 " 145 9 i
" 7.5 146 6 .
105 " 141 10 7.7 142 7
; 100. "
137 10 7.9 138 7 f, 1, 3.4 Comparison with Boiling Experiment: 1 i Boiling experiments (8) conducted at the ORR in Oak Ridge provide ( a direct check of the validity of the predictive method used in the - preceding sections. Both the fuel design and the ambient conditions (pressure, inlet temperature, etc.) used closely resemble those at , the UCNR and thus the results are comparable. t i l f _ . , ~.. . , _ , . . _ , - - . , - . _ , , . _ , - , ,_
19. ! Choosing the following example at the UCNR: i T = pool temperature = 1200F p
; 6 = primary flow rate = 1800 gpm V = coolant speed = 1.626 x 10~ x 1800 = 2.93 ft./sec. ,,
T = 2370F + sat For incipient boiling to occur: (see Sect. 3.3, above) l (T,g) 1
= # "O
- (Tg)HC '
1 Thus, O = 237 - 120 + 5.75 P' - 23.2 P Solving for the power P, by an iterative procedure, yields the l l following results: j (T B HC (T,g) g
= 2500F l P = 5.75 MN t
1 1 Incipient boiling is therefore predicted, by the method of 3.3, to occur at 5.75 MN. The ORR boiling curve ( } is entered at the same coolant velocity, . 2.93 f t./sec. , using the ORR conversion factor of 1.6 x 10' ft./sec.
~
! per gpm. The corresponding ORR flow rate is 2.93/1.6 x 10 = 1830 gpm,
; and the icwer at onset of boiling is given by:
PBoil " * *
- l 1
l I i i I l l
20. 4 Ref. 8 states that the peak / average heat flux in the ORR core { l was 2.02. The'UCNR value is 3.3 (Sect. 1.7). Therefore the ORR ^ boiling power has to be reduced by 2.02/3.3 = 0.61 in order to obtain a comparable peak heat flux. , 1 Thus, the experimental P 9.94 x .61 = 6.08 MW. boil The method used to predict boiling in the UCNR is clearly 2 conse rvative. 3.5 Fuel-Plate Surface Temperature, Normal Operating Conditions: It is of interest and significance to determine the fuel-plate surface temperature at the hot-spot, (Tw) HC, under normal or " design center" operating conditions. This will be the maximum surface temperature that a fuel plate will see for most of its lifetime. - Normal operating conditions are: P = 5 MW
= 2200 gpm ,
Tp = 98oF R(T) = 3.48 4 4 Thus: (Tw)HC = 98 + (1.14/2200 + 2.52/3.48 (2200) .8) x 5 x 10
= 154 0F.
o b I 1 l l i l l l
21. i 4.0 INCIPIENT BOILING IN NATURAL CONVECTION MODE 9 A pool temperature of 1200F is assumed, as the extreme case. For free convection between parallel plates, Elenbaas(9) has shown an empiri al relationship between Gr x Pr x and Nub, with all 1 coolant physical properties (except 8) evaluated at the plate - surface temp. (2500F), as follows: b = distance between plates = .01 ft L = length of plate = 2.05 ft Pr = 1.45 p = density of H 2O = 58.8 lb/ft 3 g = gravity = 4.17 x 108 ft/hr 2 S = expansion coeff. of bulk water = 2.5 x 10-4 0F (at 1200F) u = viscosity = .569 lb/ft hr k = thermal conductivity = .396 BTU /hr*ft0F AT = T wall
-T bulk
= 250-120 = 1300F Thus Gr = b32 p g 8 AT/u2 = 1.44 x 105 3
and Gr x Pr x b/L = 1.02 x 10 From Elenbaas' data this corresponds to Nu = 3.16 (Ref. 9, Fig. 7-10) Thus h = kNu/b = .396 x 3.16/.01 = 125 BTU /hr ft2 op Under these conditions the heat flux at incipient boiling is given by Yi = h x ATi = h[Tw-(Tb)aul
=
125 [250-(237 + 120)/2]
=
8.94 x 103 BTU /hr ft 2 This corresponds to a power level of 8.94 x 10 /2.52 3 x 104 = 0.35 MW. 1
+ .
22. i 5.0 PIFERENCES FOR SECTION A, THERMAL-HYDRAULIC SAFETY ANALYSIS. 4
- 1. W. R. Gambill, Design Curves for Burnout Heat Flux in Forced-
.! Convection Subcooled Light-Water Systems, ORNL-TM-2421, Nov. 14, 1968.
- 2. M. M. El-Wakil, Nuclear Heat Transport, International (1971) .
- 3. Handbook of Applied Engineering Science, Chem. Rubber Co. (1970).
- 4. W. R. Cambill and R. D. Bundy, Burnout Heat Fluxes for Low-Pressure Water in Natural Circulation, ORNL-3026, Dec. 20, 1960.
- 5. A. E. Bergles & W. M. Rohsenow, Determination of Forced-Convection i Surface-Boiling Heat Transfer, Journal of Heat Transfer, 365-372, Aug. 1964.
- 6. S. Glasstone & A. Sesonske, Nuclear Reactor Engineering, Van Nostrand (1963).
- 7. W. R. Gambill and R. D. Bundy, ORNL-3079 (1961) .
- 8. J. A. Cox, ORR Operations for Period April 1958 to April 1959, ORNL 59-8-39.
- 9. Frank Kreith, Principles of Heat Transfer (2nd Edition),
International (1965) . -
- 10. Reactor Power Excursion Tests in the SPERT IV Facility, J. G. Crocker and L. A. Stephan, (USAEC IDO-17000, Aug. 1964).
I t a 4
, ,. ~ -,r ,
e 23. B. REACTIVITY TRANSIENTS
. The transient resulting from the step insertion of 0.25%
SK positive reactivity into the cere while operating at a steady power of 7.5 MW is calculated below. l -
- 1. Effective Delayed Neutron Fraction (Seff) .
For calculation of kinetic effects, it is necessary to use the effective value I1' of the delayed neutron fraction (S) in order to allow for the differing effectiveness of prompt vs. delayed neutrons. The correction was derived in the following manner (2)t S = delayed neutron fraction = .0065 -' 2 Tp = Fermi lifetime for prompt neutrons in UCNR = 45 cm Td
= " " "
delayed " "
= 20 cm2 B2 c = Critical buckling in UCNR = .009 cm-2 S/Seff = (1-8) exp[- (Tp-Td )B2cl+ 0 Seff = 1.25 8 = .0081 This vtlue of Segg compares favorably with the value of .0080
^
experimentally determined (3) in the BSR reactor at Oak Ridge National tab.
- 2. Inhour Relation for UCNR: .
The standard relation is used, with effective delayed neutron fractions avaluated as in the preceding section. Delayed neutron parameters are from Ref. 2. p = reactivity Ai = decay constant of i-th neutron group (sec-1) Si = yield of i-th delayed group per fission neutron L = prompt neutron lifetime = 5.1 x 10-5 see T = reactor period (sec)
-r, e ,
24. Group No. Ai Si (81 ) eff 1 .0124 .000214 .000267 2 .0305 .001424 .001780 3 .111 .001276 .001592 4 .301 .002568 .003205 - 5 1.13 .000750 .000937 6 3.00 .000272 .000340 Sum Totals: .0065 .0081 p = 1/ (t + T) + T/ (t + T) I (Si )eff/II + A iT) A plot of this inhour relation for periods between 1 and 1000 sec. is shown in Fig. 4.
- 3. Step Reactivity Intertion Transient:
p = reactivity step (AK) S,gg = .0081 (above, B 1) T. = stable period, sec. (Fig. 4) P(0) = initial reactor power, MW P (t) = power at time to sec., MW t1 = control rod magnet release time = .05 sec. (max.) _ t2 = scram time = .75 sec. (max.) c = time required for step to occur = .1 sec. a = control rod acceleration, ft./sec. (avg.) S = total control red travel = 2 ft. Sp = rod insertion for -p reactivity c'hange, ft. tp = time for rod insertion Sp, sec. g = acceleration of gravity = 32.2 ft./sec.
25. Then: a = 2S/ (t2 -ti) 2 = 8.16 f t./sec.2 = 0.25 g tg = ty + (2 S p/a) * = .05 + .50 Sp The maximum reactor power is conservatively estimated by ignoring the negative step due to control rod insertion, and is then given by P (to) , where : P (tp) = P(O) xS gg (1-p)/(S gg-0) x exp(tp-c)/T
= P (O) x .0081 (1-p ) / ( . 0081-p ) x exp (tp .1)/T The following specific cases are considered, in each of which the rods are assumed conservatively to be fully withdrawn, which maximizes their travel needed to cancel the step.
- a. p = .25% AK = .0025 T = 17 sec. (Fig. 4)
P(O) = 5 MW Sp = 2 in. = .17 ft. (max.) tp = .05 + .50(.17)* = .26 P ( . 26) = 5 x .0081 (1 .0025)/(.0081 .0025) x exp (,26 .1)/17
= 7.3 MW Similarly, if P (0) = .25 MW, then P (.26) = .36 MW; l
l l l
26.
;. p = .5% AK = .005 T = 3.1 sec. (Fig. 4)
P (0) = 5 MW Sp = 5 in. = .42 ft. (max.) . to = .05 + . 5 ( .42) * = .37 sec. P(.37) = 5 x .0081(1 .005)/(.0081 .005) x exp (.37 .1)/3.1
= 14 ffd Similarly, if P(0) = .25 Ifd, then P (.37) = .71 MW 1
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28.
- 4. REFERENCES FOR SECTION B, REACTIVITY TPANSIENTS
- 1. G. Robert Keepin, Physics of Reactor Kinetics, Addison-Wesley 1 (1965).
- 2. Reactor Physics Constants, Argonne National Lab. ANL-5800 (1963).
l
- 3. J. D. Kington et al, Nucl. Sci. and Eng. 12, 505 (1962).
l l 4
29. C. RADIOLOGICAL SAFETY ANALYSES
- 1. Radiation Dose at Site Roundar/
The radiation dose at the site boundarf due to routine release of 41Ar, the only significant contributor to the rentine dose, is now considered. The experiment facilities (bea= tubes, ther .al column, and 4 pneumatic tubes) and the hold-up tank all vent into the main exhaust duct leading to the stack. The measured emission rate of 41Ar is 725 . Ci/yr for 5 MW operation with a 0.60 duty c/' cle. Ad, justed for T.5 MW cperation at a 1.0 duty cycle the average 41 Ar emission rate is there-fore 57uCi/sec. Wind data records in 22-1/2 sectors at the stack show that the most probable direction is 45 E of N, with a mean speed of h m/see towards the site boundarf distant h60 m. Elevation of the discharge above the site boundary is 91 m. The frequency of inversions leads to a yearly average of 14% F and c6% Fasquill C conditions. Doses at the boundar/ in the most probable direction are calculated using the methods, and the same symbols, of Ref. 1. Definitions: D = Annual 7 dose from 41Ar at site boundary. 7 41 Ar at site boundary. Dg = Annual S dose from 3 m1 o = Total absorption coeff. for 7 in air = T.2x10 3
= Energy abs rpti n coeff. for 7 in air = 3.hx10 m ~1 a
) k = (u-ua)/u a =l1 nergy of 41 Ar 7 photon = 1 3 Mev 7 - E 3
= Avg. 41Ar 8 energy = .4 Mev b = Activity release rate = 5.Tx10 5 Ci/sec iI = Mean wind velocity in most probable sector = h m/see f = Fraction of time wind blows into most probable sector = .21 6 = Angular width of sector (22.5*) = .39 radian R = Distance from release point to site bou.;dary = 460 m t = Exposure time (1 year) = 3.16x107 sec c = Vertical standard deviation of plume = 23.4 m n = Number of sectors considered = 16 h = Release height above grour.d level = 91 =
I1 = Integral from Ref.1 (p. 352) = 0.4 I2 = Integral from Ref.1 (p. 352) =05 Tm Avg. long term concentration of 41Ar
30. The annual ground-level y-dose at the site boundary, from Ref. 1 (p. 352), is: . D 7 =.2865pabEft(Il y + kI2)/uRG
= 6.4 x 10-4 m/fr .
The average long term 41 Ar concentration and the resulting annual B-dose at the site boundary are obtained using expressions from Ref. 1 (pp. 113, 328) as follows: i = (2/n) .5 g gf(qz [2(2nR/n)) x exp(-h2 /20 22)
= 5.65x10-10 x 5.2 x 10-4 = 2,9x10-13 Ci/m 3 Dg=.457Eg[t (infinite cloud)
= 1.7 x 10 6 rem /yr The total annual beta plus gamma dose at the site boundary in the most probable direction, for 7.5 FM operation at a duty cycle of 1.0, is then:
D =D +D = 0.6 mrem /yr Tto Y S l
. - . ., -. - .. -__._ - . ~ . . . _ _ _ , . . _ - . - .
31. i I
- 2. Design Basis Accident:
To evaluate the containment we concider some hypothetical reactor ] accident resulting in a melt-down of an amount of reactor fuel con-4 l taining 10% of the total accumulated fission products. A concurrent l total loss of water is assumed so no reduction of halogens due to absorption in the coolant is considered. i i The reactor is considered to have operated for a long time at a power level of 7.5 MW. As a result of such a melt-down, it is I i
' considered that 10% of the total core noble gases and 5% of the total i core halogens are released from the fuel. The solid fission products are assumed to remain in place.
The equilibrium concentration of any fission product may be found by 6 A = y x 0.1x1010) x (7.5x10 ) cur es 3.7 x 101g i Where: l I' y = fission yield of isotope considered J 3.1 x 1010 = fissions /second/ watt 7.5 x 106 = power level in watts 3.7 x 1010 = disintegrations /sec/Ci The fission products being considered are listed in Table I. 7 9 i The iodine released is considered to be reduced by a factor of 2 due to plate-out on the interior surf aces of the containment building. l It is further assumed that the effluent is released, after mixing with the building air, via the emergency exhaust system at a flow rate of 9.44x10-2 m 3 /sec (200 cfm). The iodines are further reduced by a- ' factor of 20 by the charcoal filters in this system, but the noble gases are unaffected. The containment building has a free air volume of i 7700 m 3 , resulting in the initial concentrations (Co) as given in the i second column of Table II. 4 i c
.~- .- . - . _ _
32. Except for the short-lived Xe and Xe, we assume no depletion in the inventory over the initial 2-hour period of release. Decay
- corrections are, however, applied to the inventory in calculating release rates for subsequent time periods. In the special case of 135 Xe, burnup and buildup are taken into account. The decay correction factors (DF), which are averaged over the time interval, and the ,
resulting average concentrations (C1 , C2) within the containment are given in Table II. TABLE I 4 Total Core Total Isotope y Inventory- Ci Release- Ci I .0277 1.74 x 10 8.7 x 10 12 I .0413 2.6 x 10 1.3 x 10 I .0676 4.25 x 10 2.12 x 10 I .0718 4.51 x 10 2.25 x 10 I .0639 4.02 x 10 2.01 x 10 "Kr .0133 8.36 x 10 8.36 x 10 Kr .0237 1.49 x 10 1.49 x 10 88 5 4
, . Kr .0364 2.29 x 10 2.29 x 10 4 ; Xe .0676 4.25 x 10 4.25 x 10 135 Xe .0105 6.6 x 10 6.6 x 10 4
Xe .0672 7.05 x 1C 7.05 x 10 137 Xe .059 3.71 x 10 3.71 x 10 Xe .062 3.89 x 10 3.89 x 10 e i I l-a I w . , - m - , - - . - , . , ,r,. , . . . ..e.. - w,.-.-- -- e r #., ,#,,
33. TABLE II O-2 Hours 2-8 Hours 8-2k Hours Isotore Co(Ci/m3 ) DF C, (Ci/m3 ) DF Cg(Ci/m3 ) . 1311 56 1 56 94 53 132I .84 .25 .21 .017 .014 133I 1 38 .85 1.17 59 .81 134 I 4 4 1.h6 .0h2 .058 1.3x10 2x10 135I 1 31 .60 79 .027 .035 esmKr 10 9 .h7 5.1 .10 1.1 87Kr 1.hx103 3 19 .13 .25 2.Tx10 88 Kr 3.0 32 96 .034 .10 1 Xe 55 97 53 92 51 1355Xe .86 0 0 13sXe 92 1.7 1.6 137Xe k.82 0 0 13eXe 5.05 0 0
- Note: DF's for 137Xe and 138Xe for 0-2 hrs. are .Oh6 and .170, respectively.
34. Following the guidelines of Ref. 3, doses at the site bcundary distance of 250 m are calculated for an elevated release at a height of 69 m. Fumigation conditions are assumed for the first half hour, result-ing in an average dispersion factor (X/k) for the first 2 hours of 3 1.8x10~4 sec/m. The X/k values for 2-8 hours and 8-24 hours are 4x10 s and 5x10~5, respectively. The appropriate concentrations (X) at the site boundar/' for the dose calculations are obtained from the concentra-tions (Co, C1 , or C 2) in the containment, the exhaust flow rate (.094h m3 /sec), and the filtration factor (1 for noble gases, 20 for iodines) as follows: X = Cx.0944/20xX/k (iodine) X = C x .0944 x X/b. (noble gases) The accumulated thyroid dose (De ), from Ref. 9, is given by: De = 5 92x10 E,f,5xt/mx, rem where: E = effective energy absorbed per disintegration (Mev) f, = fraction of inhaled iodine in thyroid = .23 5 = breathing rate = 3.4Tx10 ~4 m3 /sec in first 8 hours, and
= 1.75x10 4 m/seethereafter.
t = exposure time (sec) m = thyroid nass = 20 g 1 = effective decay Constant (see 1) X = iodine concentration at bou"Aary (as above)"(Ci/m3 ) Inserting these values and substituting for X rentlts in the following expressions for each time interval: -
- a. 0-2 hrs: (X/k = 1.8x10 4; 5=3.kTx10 4; C = Co; t = T200)
De = 1.44x105 CoE/Ae e rem
- b. 2-8 hrs: (X/k = 4x10 5; 5 = 3.hTx10~4; C=C; 1 t = 2.16xlO4)
(= 9.6 x 10 S C1eE e/l rem
- c. 8-24 hrs: (X/Q = 5x10~5; 5 = 1.75x10 4; C = C2 ; t = 5 76x104 )
(= 1.61x10 s C 2Ee/A e rem
35. The whole-body beta and ga=a doses g(D and D , respectively) in rem from a cloud of concentration X Ci/m3 are given by: Dg = .h57 EKt (infinite cloud) D 7
= .25 E7 Xt (semi-infinite cloud) where:
5 = average beta energy w 1/3E, (Mev) , E = maximum beta energy per disintegration (Mev) E = total ga=a energy per disintegration (Mev) t = axposure time (sec) Sutstituting for the value of X for noble gases (icdines centribute little to the whole-body dose and are neglected) gives the folicwing expressions for each time interval:
- a. 0-2 hrs. (X/b, = 1.8x10 4; C = Cc; t = 7200)
D = .0186 CoE rem g D = .0306 CoE 7 rem 7
- b. h8 hrs. (X/Q = Ex10 s; C=C; 1 t = 2.16x10 )
4 D = .0124 C 1E , S . D,
=
.020h C1 E, c . 8-24 hrs . (X/6; = 5x10 s ; C=C2; t = 5.76x104)
D = .Chlh CE, 2 s D,
= .C68 C2 5, Values of effective energy (E e
) and effective disi:.tegration constant (le) for the thyroid, maximum beta energyg(~ ), and total g9m energy for each isotope are taken from the data in Refs. 4 and 5, and are tabulated below.
36. TABLE III
~
e E E
, E m 7 e
Isotope (Mev) (sec~1) (Mev) (Mev) 131I .23 1.06x10 e 132I .65 8.27x10 ~5 13"I 54 9 22x10 e
~4 134I .82 2.23x10 5
135I 52 2.87x10
.63 .151 SS=:e 3,1 ,g@
67gy es@ 1.02 1.75 133xe 346 .032 135mxe 0 .h35 90 .26 135Xe
~
137Xe 39 31 138Xe 2.h ~ 2. 4
37. I Whole-body and thyroid doses are row calculated using the foregoing expres-sions and data for each isotope. As some parameters (e.g., breathing rates and dispersion coefficients) var / in accordance with the guidelines of Ref. 3, time intervals of o-2, 2-8, and 8-24 hours after release are used. The results are given in Table IV below: TABLE IV DOSES (Rem) i o - 2 hrs 2 - 8 hrs 8 - 24 hrs D D D D Isotore D D 6 7 6 1 6 7
.040 .015 .029 .oll SmKr .128 .050 S7 Kr .no .057 .010 .005 o o ea rr .161 .012 .034 .004 .012
.o57
.005 .023 .003 .073 .011 133Xe .035 4
o .oll o o o o 135mXe
.015 .007 .019 .009 .060 .028 i 135Xe
.016 .002 o o o o 137Xe 1
.o53 o o o o 138Xe .o39
.40 .10 .07 .17 .06 B, 7 Totals 35
' D D D 1.75 1.17 1.86 4 131I 4
.10 .02 0 132I 133I 1.17 .66 77 i
.08 o o 134I
.14 .01 issI 34 Thyroid Totals 3.h 2.0 2.6
! The acetnulated total doses are thus as follows: L L o.- 2 hrs: 3.4 rem thyroid; .75 rem whole-body. o - 24 hrs: 8.0 rem thyroid; 1.2 rem whole-body. l
. -- . - ~ .. . - -
,s 38.
- 3. Fueled Experiments:
An in-core fueled experiment is considered in which the total
- inventory of radioactive iodines and noble gases equals 500 curies of 131 I dose equivalent. For such an experinent submerged in the pool at core depth the significant fission products from the point of view of radiological consequences are those that are volatile. Exposures both on and off site due to the release of the inventcry are calculated.
The significant iodine and noble gas inventories appropriate to 1 500 curies of I dose equivalent (as determined using MPC's from 10CFR20, Appendix B) are tabulated below. Also given are initial containment air concentrations (Co ) assuming 100% release of the inventory, a decontamination factor of 10 for iodines in pool water, and mixing with 50% of the building air volume. Thyroid and whole-body doses are calculated inside the containment and at the site boundary. For the latter, dispersion factors (X/Q) are taken from Ref. 1 for an elevated release at 69 meters. 4 Containment building air volume is 7700 m3 . The containment air is released via the emergency exhaust system at a flow-rate of 200 cfm through absolute and charcoal filters. i I
39. Isotope Inventory Concentration (Cn) I 196 Ci 0.00483 Ci/m I 412 .0107 I 842. .0219 134 .0255 I 983 I 783 .0203
*Kr 166 .043 Kr 287 .074 Kr 433 .112 Xe 449 .117 "Xe 239 .0621 Xe 140 .0364 _
137 .0244 Xe 94 Xe 368 .0956 (1) Thyroid Doses: Total Iodine Inventory = 500 Ci I (dose equiv.) Pool Decontamination Factor = 10 Containment Volume = 7700 m Co E concentration in building = 500/10 x 2/7700
= 1.3 x 10-2 Ci/m 3 (131I equiv.)
Efficiency of charcoal filter for iodine removal = 90% l l f
. - , . -, - , ,- ,~.
40. (a) In containment building: t = exposure time = 30 sec. 3 5 = breathing rate = 3.47 x 10 ~4 m/sec. l = effective 131I decay const. = 1.06 x 10~8/sec (Ref. 5) e " E = effective energy 131 I in thyroid = .23 Mev e " m = thyroid mass = 20 g f = fraction in thyroid = .23 - Thyroid dose D = 5 92 x 10 2 Ef ea 0 e (Ref. 1)
= 1.k8 x 108BCot
= 200 rem. ,
For a 70 curie inventor / (131I equiv. ) the thyroid dose would be 200 x 70/500 = 28 rem. (b) Offsite: Containment exhaust rate =
.09hh m/sec. (200 cfh) 6, = iodine emission rate =
.09h4 co /10 = 1.23 x 104ci/sec no credit taken for decay of above iodine incentor/ in the contain- -
ment. Atmospheric dispersion factors are taket at 250 m distance from an elevated (69 m) release with fumigation for the first half hour (Ref.3 I . 4 3 (x/k = dispersion factor = 6 x 10 sec/m (first1/2-hour)
= 1.8 x 10~4 (average, 0-2 hrs)
= 4 x 10 s (2-8 hrs)
= 5 x 10 5 (8-24 hrs) 5 r breathing rate .
t = exposure time De = thyroid dose = 1.48 x 108 5 (X/k) Q t D Time Period 5 (X/6) k t 4 .0d2 0 - 2 hrs 3.ETx10 4 1.8x10~4 1.23x10 7200 2-8 " " 4x10~5 2.16x104 .055 4 5 " 4 8 - 2k " 1.75x10 5x10 5 76x10 .092 Total thyroid dose 0 - 2 hrs = .C8 rem 0 - 24 " = .23 rem
41. (2) '4 hole-Body Doses 100f; of noble gases are released with initial concentrations tabulated above. Iodines contribute little to the whole-body dose and are neglec ;ed. Beta dose ;: D g
= .457 EgXt = .152 E m Xt (infinite cloud)
Ga=ma dose ED 7
= .25 E Xt 7
(semi-infinite cloud) 3 X E concentration (Ci/m ) t = expos' ire time (sec) E 3
= total max. beta energy (Ref. 4 )
E g = total average beta energy = 1/3 "g E 7
= total gn-ma energy (Ref. 4 )
(a) In containment building: X = Co (given previously) t = 30 sec. Isotcpe X Em Ey De Dy 85 Kr .124
.043 .63 .151 .049 87 1.046 Kr .074 31 98 544 eeKr .112 1.02 1.75 521 1.47 123xe .117 346 .032 .185 .028 135Cxe .0621 0 .k35 0 .203 _
135xe .036k 90 .26 .149 .071 137xe .0244 3 94 305 .438 .018 138xe .0956 2.4 2. 1.046 1.434 TOIALS: 3 51 3.82 Total whole body = 7 3 rem.
62. (b) Offsite: Containnent exhaust rates and dispersion factors are the sere as in sub-section (1)(b) above. k = activity emission rate = .0944 co ci/see Co a initial concentration in contairment (see above) X a concentration at stack = (X/k) x k Ci/m3 E,and E7 are given in (a) above D g= .152 E,(X/d)k t (as before) D = 7 .25 E 7(X/Q)Q t (as before) (1) 0 - 2 hr. period: t = T200; X/Q = 1.8x104 Isotope k De Dy esnrr 4.06x103 5 0hx104 1 99x10 4 e7 3 Kr 6.99x10 k.27x103 2.22x10 3 1.06x102 3 eeKr 2.13x10 6.0lx103
'33 2
- Xe 1.10x10 7 50x10 4 1.14xlO4 135mXe 5.86x103 0 8.26x104 13sXe 3.4hx103 6.10x104 2 90x10 4 137Xe 3 3 2 30x10 1.79x10 2.27x104 3
13eXe 9 02x10 h.26x103 5.84x103 TOIALS: 1.hxlO' 1.6x102 - (ii) 2 - 8 hr. period: t = 6 hr = 2.16x104 see; y/k=hx10 3; inventory is decayed for 5 hours. 13sXe buildup is taken at 5 hours. Qs = inventory at 5 hours (curies in containment) Cs = concentration at 5 hours = Qs/3850 k = emission rate = .09h4 c s= 2.45x10 3 qs ci/sec l l
40. Isotope q j D D _7 _s _e 3 4 3 es=Kr 76 1.86x10 1.Shx10 6.0Tx10 4 5 87g7 1g g,gggyg 4 1 90x10 9.86x10
**Kr 125 3 06x10 3 h.10x104 1.16x103 2 4 133Xe hh9 1.10x10 5.00x10 7.60x105 -
3 4 4 13sXe 260 6.38x10 7.5hx10 3 58x10 TCfIALS: 2.0x10 3 1.8x103 (iii) 8 - 2h hr. perfod: t = 16 hr; X/k = 5x10 5; inventory is decayed to 16 hours 135Xe buildup at 16 hours Qte = inventory at 16 hours 5 Q = emission rate = 2.h5x10 Q1e De D7 Isotope b d. 4 esmKr 13 3 2x10 8.8x105 3 5x10 s 38 4 4 Kr 8 2.0x10 8.9x10 s 2 5x10 3 4 133Xe 410 1.0x10- 1 5x10 2 3x10 3 135Xe 2h6 6.0x103 2.hx103 1.1x10 TOTALS: h.1x10 3 1.6x103 (iv) Total Offsite whole-body doses: . 0 - 2 hrs: .03 rem 0 - 2h hrs: .0h rem
A 44.
- 4. EFFLUENTS IN UNRESTRICTED AREAS (1) INTRODUCTION This Supplement gives additional information to substantiate proposed tecnnical specifications on the concentrations of airborne radioactive emissions from the reactor stack. The methods prescribed in Regulatory
! Guides 1.109 (for dose methodology) and '1.111 (for meteorological methodology) are followed in order to derive release rates that are within . the annual dose limits or 10CFR20 for unrestricted areas. Meteorological input da.:a used are long-term averages obtained from the Final Hazards j Summary Report (FHSR), supplemented by observations of wind speed and , direction taken at the stack location. Three locations in the unrestricted area have been selected for detailed treatment - one is the most limiting location (designated " Hogback"), and the other two are the limiting location in each of the two nearest residential areas (designated " Laurel", , and "Clinton"). Of the latter, one is closest to the stack and the other is in the direction of greatest wind-direction frequency. Six radionuclides
~
are identified as possibly being limiting either from the point of view of average annual dose or because of predominance in the gaseous inventory. l These are two radiolodines (131I, 125I ) and four noble gases (41A, 88K r, 135xe, 133xe).
~
l (2) METHODOLOGY
- a. Meteorological:
The Gaussian straight-line airflow model, cf. Regulatory Guide
~
! 1.111 equation (3), is used to calculate long-term average dispersion. Yearly average conditions of 86% Pasquill Class C (slightly unstable) and 14% Class F (moderately stable) are taken as typical of this locality. X/h. = 2.03 f($)/ru - 0.86/o zc e c+o,jg/az f ** _ f($) = fractional time wind is in 22.50 sector at azimuth 4
#zc, #zf = vertical plume spread for C S F categories, resp. (m).
- h = elevation of release point (m).
u = mean wind speed (m/s) for azimuth 4 I r = distance from release point (m)'.
45
- b. Dose Estimation:
The methods and tables in Regulatory Guide 1.109 Sections 2 and 3, and in TID-24190 Sec. 7-5-2.5., are used to calculate child
- thyroid doses from inhalation of radiciodines or whole-body and skin doses from noble gases. '
(1) Iodines: D = 2700 x 31.65 x DFA
= 8.54 x 10 x DFA rem /sec.
Where: 2700 = child breathing rate (m3 /yr), Table A-2 X = concentration, (Ci/m ) b = thyroid dose rate (rem /sec) DFA = fodine dose factor, (mrem /pCi), Table C-3 ' For D not to exceed 1.5 rem /yr (4.75 x 10 -8 rem /sec), the maximum release rate , averaged over one year, is given by: km= 5.56 x 10-I3 /(x/6 x DFA) Ci/Sec. (2) Noble Gases - Ground Releases: T D = 1.11 x .7 x X x DFB x 31.65
= 24.59 x DFB b = 1.11 x .7 X
- DFY x 31.65 + X DFS x 31.65 3
= 31.65 x ( 777 DFY + DFS)
Where: i
.T D = total body dose rate (rem /sec) 6 = skin dose rate -(rem /sec)
X = concentration (Ci/m ) T 3 DFB, DF , DFS = dose constants, Table B-1 (mr m / pCi yr) 3 31.65 = conversion from mr m /pCi yr to r-m3/Ci-sec. 1.11, .7' : see Reg. Guide 1.109-12.
(-.. , 46. I i t i For 6 and b not to exceed 0.5 rem /yr and 3.0 rem /yr, resp.,thecorrespondingmaximumconcentrationskm are: < 1 .
-10 Ci/Sec.
I (Total body) Qm = 6.44 x 10 /(X/k x DFB) . (Skin) hm= 3.0 x 10~9 /(Xh)(.777DFY/ + DFS) Ci/Sec. l-(3) Noble Gases - Elevated Releases:* . T gT = 1.11 x .7 x .258 'q f(4) p, E f ( I j+KI 2) e "a /.390r
~T
= 0.51 d f(c) uayy E f 1(I +KI 2)e "a /Or ,
D =1.11x.7x.258df(4) p E f (I +KI ) y 2
/ 390r 1
+ 31.65 x DFS 1
- 0.51 6f(t) u,E f (I y+KI2) /r+31.656(X/h)DFS Where:
.T D = total body dose rate (rem /sec)
*S 0 =
skin dose rate (rem /sec)
- h. = emission rate (Ci/Sec) 4 u, = air energy absorption coeff. (m" )
-E = photon energy (Mev) f = no. of photons / disintegration K- = (p v )/D a a p = air linear absorption coeff (m)
I,I from TlD-24190 39 = 22.5 sector. width (radian) 9 e -- ,- t n- r.- -- -- .- ,-w urr - we-..r- - - - r- y q 1- m
47 3 X = 6. x (X/k) = Concentration (Ci/m ) r = distance (m) u = mean wind speed (m/s) .. DFS from Reg. Guide 1.109, Table B-1 31.65 = conversion from mRm /pCi yr to r-m /Ci-sec p = tissue energy absorption coeff. (cm /g) 2 t = tissue density = 5 g/cm For D'T and D 'S not to exceed 0.5 and 3 0 rem /yr, resp. , the Ci/sec, can corresponding values of the maximum emission rate, m
~
be derived from the above for each radionuclide and for the X/q value (sec/m 3
) at each location of concern. This is done below for the " Laurel" location, the only one of the three for which the release is elevated.
l (3.) HOGBACK SITE: (uninhabited) Site parameters:
~
l $ = 213 r = 314 m h =0 (ground release) f(t) = .0437 I2 = 1.38 m/s
= 22 m (C); 5.3 m (F) i "z
Resulting dispersion index is: X /d = 1.40 x 10 -5 sec/m 3 Maximum emission rates:
= 5.56 x 10
-I3 /1.40 x 10 -5 x DFA = 3 97 x 10 /DFA ~
(Iodines) dm
-6
! (I3II) dm = 3.97 x 10- /4.16 x 10~3 = 9.5 x 10 Ci/sec
-8 (1251 ) hm = 3 97 x 10 73,3 3 x 30. -3 = 1.3 x 10-5 Ci/sec
l
- 48. I l
(Noble Gases - Total Body) k = 6.44 x 10-10 /1.40 x 10 -5 DFB
= 4.6L x 10-5/DFB (NobleGases-Skin)k m = 3.0 x 10 -3/1.4 x 10-5 (.777DFY + DFS)
= 2.14 x 10'4 (.777DFY
/ + DFS) .,
Nuclide km (Ci/sec) Total Body Skin A 5.20 E-3 2.16 E-2 Kr 3 13 E-3 1.51E-2 Xe 1. 56 E-1 3.69E-1 135 6.38 E-2 Xe 2.54 E-2
- 4. CLINTON SITE: (Residential) i Site parameters: 4 = 65 .
r= 1370 m , b=0 (ground release). f($) = .0731 o = 80 m (C); 17 m (F) 7 Thus,dispersionindex,X/d=4.77x10-7 sr.cfm 3 Maximum eraission rates:
-7 x DFA = 1.17 x 10 /DFA~
('Iudines) d = 5.56 x 10-I3 /4.77 x 10
~ ~ ~4
( I) km= 1.17 x 10 /4.16 x 10 = 2.8 x 10 Ci/se
~
~
~3 Cl/sec
( I). k = 1.17 x 10 /3.13 x 10 = 3 7 x 10 (Noble Gas - Total Body) d = 6.44 x 10 -10 /4.77 10~ DFB
= 1.35 x 10-3 DFB/
(Noble Gas - Skin) km = 3.0 x 10-9 4.77
/ x 10-7(.777DFT + DFS)
[
= 6.29 x 10-3/(.777DFY + DFS)
I
3 49 Nuclide 6 , (Ci/sec) i Total Body Skin NI A 1.5E-1 6.3E-1 0 Kr 9.2E-2 4.4E-1 i 133 1.1El xe 4.6E0 135 Xe 7.4E-1 1.9E0 5 LAUREL SITE (residential) Site parameters: 4 = 109 r = 450 m l h = 54 m (elevated release) f(4) = .127 u = 2 9 m/s
#z= 32 m (C); 8 m (F)
Thus, dispersion index x/6 = 1.3 x 10-6 ,,cf,3 , i 4 a. Iodines: .
-6 m
=
5.56 x 10 -I3 /1.3 x 10 DFA = 4.28 x 10-7/DFA (I3I I) 6, = 4.28 x 10-7/4.16 x 10 -3 = 1.0 x 10 -0 Ci/sec (1251 ) y = 4.28 x 10-7/3.13 x 10
-3 = 1.4 x 10 -N CI/sec
- b. Noble Gases:
For these elevated releases the C & F conditions produce closely similar total-body dose results, and thus are lumped to-gether as 100% C in the following calculations. Because of the need to evaluate I + KI in which the nuclidic and the site 2 properties cannot be separated, each nuclide is treated in' turn. Certain site parameters (r,f(4),5) can, however, be inserted into the expressions in para. 2b(3) above to give the following: i n
4 50. 6 = 5.0 x 10-5d p,E f (I j+KI2 ) * "" 68 = 5.0 x 10-5 g -5 q DFS p,E f (I +KI j 2) + 4.1 x 10 (1) Argon-41: f E = 1.3 Mev i Y I f =1 Y
-1
= 3.4 x 10 -3 p m a
p = 7.2 x 10-3 ,-1 K 5 1.12 Ij i .8 ; paz = .23 .
- I2 5 .5 ; az /h = .59 pT = 2.86 x 10-2 cm 2 /g ; e pa't = .87 DFS = 2.69 x 10-3 mr-m3/pCI yr (Table B-1)
Substituting these values in the relation given above: . 6T = 5.0 x 10-5kx6xi'3 x .87 = 2.6 x 10-7 k rem /sec 1 1 DS = 5 x 10 -5 *q x 6 x 10-3 + 4.1 x 10-56x2.69x10 ) = 4.1 x 10-7 q rem /sec i For 6 and 6$ not to exceed 0.5 and 3.0 rem /yr, respectively, 41
- the maximum release rates for A are then
-2
- (Total body) h, = 6.1 x 10 Ci/s ec
~I (Skin) hm = 2.3 x'10 Ci/s ec 5
4 - ,a y,.- ,-c,,,. ,- 4 ,
_. - . ~ . . - .. .. .. -. 51. (2) Krypton-88:
-3 mr-m 3 /pci-yr DFS = 2.37 x 10 E = ' I '. 2 2.3 0.19 Mev Y
f = 37 .53 35 Y -
-1 103p =
3.2 2.8 3.2 m f a 3 6.8 m"3 10 u = 5.0 15. K = 1.13 .79 3.7 ph = 37 .27 .81 i pa, = .22 .16 .50 j I j
= .8 1. .45 I
2
=
.55 .6 .4 4
i I +KI 2
= 1.4 1.5 1.9 T 2 u, =
.029 .0245 .029 cm /g T =
.87 .88 .87 e
-u,t 1
= 6.56 x 10 -3
~
Thus, u,f E (I j+KI 2) ' e " p,f E '(I +KI j 2)
= 7.45 x 10 -3 l
.T Hence: D = 5 x 10-5kx6.56x10-3 = 3.3 x 10 -7 *q rem /sec
.S
[ D = 5 x 10-5kx7.45x10-3 + 4.1 x 10-5kx2.37x10-3
- 4.7 x 10 ~7 h rem /sec l
(Total Body) -2. Ci/sec Qm = 4.8 x 10 hm = 2.0'x 10 -l I ~
- (Skin) Ci/s ec l
l I I i
^
- . . . - . v , , _ . _ , _ . - , , , . . _ _ , _ . _ , . . . . . . . . . . _ . . z. . - . . . .
j ' i 52. 2 i i f (3) Xenon-133 f 4
~
DFS = 3.06 x 10 mr-m /pCi-yr i E = .081 .031 Mev Y f = 373 .47 Y 2 -1 10 p = .265 1.8 m a 2 10 = 1.94 s 8.4 m"I K = 6.3 3.7 I ph = 1.05 4.5 o, = .64 2.8 I; = 35 .07 1 = .35 .07 2 I +K = 2.6 .33
. T = .027 .154 cm2 /g 1 p, .
T
-p t = .43 .46 e a u,f E (l +KI y 2)
* "* " ' 3 x t0
~
p,f Y.E (I)+KI2)
= 2.9 x 10 Hence;
*T D = 5 x 10 q
- x 1 3 x 10
-4
- 6.5 x to -9 *q rem /sec
_4 _c . _ 4=
.s _~5 + 4.1 x 10 ' q x 3 06 x 10 D = 5 x 10 q x 2 9 x 10 ,
-8
> = 2.7 x 10 rem /sec 1-
-, e - - -y y .+ - - - -- -,.me-
. . - - - - , . -,3,
.s e 53 1
The corresponding maximum emission rates, Q ,, are-I (Total Body) 'd , = 2.4 Ci/ sec (Skin) d, = 3. 5 Cl/ sec (4) Xenon-135 DFS = 1.86 x 10 ~3 mr-m /pCI-yr E = .25 .61 Mev Y > f =
.91 .03 Y
3 m"I 10 u, = 3.34 3.54 3 -I 10 u a = 13.7 9.6 m l K = 31 1.7 ,
, uh =
74 .52 4 ua z
= .45 32 I; = .5 .65 I = .4 .47 ' I +KI =
1.74 1.46 2 T 2 ~
; u, =
.0303 .0319 cm /g T
e "a t
= .86 .85
= 1.3 x 10 ~3
~
u,f YYE (II+KI2) e "a' ,
- . p f E (I +KI )
2
= 1.5 x 10 ~3 a Y -Y 1 T q rer./sec
- Hence
- D = 5 x 10 -5 *Q x 1.3 x 10-3 - 6.5 x 10 D = 5 x 10-5 k x 1.5 x 10~3 + 4.1 x 10-5 q x 1.86 x 10-3 = 1. 5 x 10-76 rem /sec 4
.(Total l Body) _ 6, = 2.4 x 10"I Ci/sec
~I (Skin) 6, = 6.3'x 10 Ci/s ec l*
?
,y-- ym . - , - - - w -, ,o, . , . , . , . ~ + - -y
54. j (5) Summary for Laurel Site: i l k, = max. stack emission rate (Ci/sec) to give no more than 1.5 , rem /yr to the thyroid (child) fiom inhalation, 0 5 rem /yr to. the . . '- whole body, or 3 0 rem to the skin when averaged over one year. 4 Q (Ci/'sec) . Nuclide Thyroid Total Body Skin 125 7 1.4E-4 - I3I I 1.0E-4 - - i
'A - 6.IE-2 2.3E-1 88 Kr - 4.8E-2 2.0E-!
33 Xe
- 2.4E0 3 5E0-l
- 2.4E-1 6.3E-1 1357 ,
1 e a a E
--,-4 -1 .n.. - -
. - - , w
~^
o , 55 5 REFERENCES FOR SECTION C. RADIOLOGICAL SAFETY ANALYSES
- 1. David H. Slade (Ed) , Meteorology and Atomic Energy 1968, i
USAEC, TID-24190.
- 2. M. E. Meek and B. F. Rider, Compilation of Fission Product Yields, NEDO-12154, 1972.
l
- 3. USAEC Regulatory Guide No. 1.3, t
- 4. C. M. Lederer et al, Table of Isotopes, Sixth Edition, Wiley 1967.
i
- 5. Publication No. 2, International Committee on Radiation Protection.
s 6 f i l i l 1 I s l l l I l I 1
l b l APPENDIX 3 TIME REQUIRED TO EMPTY P0OL WATER For the flow of any fluid, Bernoulli's equation reduces to y = / 2gh where v = velocity, ft. per sec. h = head of fluid flowing, f t. g = acceleration due to gravity, 32.2 ft./sec 2 The rate at which water flows from an open tank through an orifice is: { f = C a / 2gh also dV = A dh where V = volume of water discharged, f t.3 t = time, sec
- C = coefficient of discharge, 0.61 a = area of orifice, ft.2 A = cross-section area of tar.k at water level The time required for the water level to fall from the initial level to the final level can be found by rearranging the above equations and integrating both sides with respect to the variable,
^
dt=- 0.61 a / 2gh
=
2A j _ j . t 0.61 a / 2g __ _ The area, A, of the pool system is not constant. There are two different elevations where the surface area changes significantly as shown in Figure 19 Therefore, the calculation can be made stepwise for (1) the pool volume between the initial water level and bottom of the canal, (2) the bottom of the canal and the shelf in the stall, and (3) the stall shelf to the final water level. The maximum 1.D. of the stainless-steel primary i coolant lines penetrating the pool floor is 10.4". The total-cross-sectional area of the four lines is thus 2.18 f t . 2 The times to drain down to the various levels are: Surface h - ft. Area-ft.2 Vol-ft.3 t-min. , Water Level to Canal Bottom 12.0 804 9,628 3.6 Canal Bottom to Stall Shelf 3.25 600 1,925 09 Stall Shelf to Final Water Level 9 09 460 4,170 3.3 Total 24.34 - 15,723 7.9 The total time to drain the approximately 120,000 gallons of water from the pools is therefore a minimum of about 8 minutes.
.GENERAUZED CONFIGURATIOM OF REACTOR POOL SYSTEM
~~ EE. BO 9.115 '
.- - - - - - - - - - - - - - - - - - - - - - - -l .L -J. J. 1 .1 .L l- d
.1NITIAL WATE-Q LEVEL /
# I
'd $
0 o N o > _EL-------. Z27a26'. - , - - - - - - - - - - - - - - - - - . - - _
/ 1 "O in O N O &
T et.79S.876'_ NNs
- mp e 2
'O O O C 9 n g
a > EC784,lS'(AV) BEAM 1.!Olt ( 7 3 -ORE C q EL.7 8 3. .S6 ' , 1 FINAL WATER LEVEi., tM POOL
/
_E--------. L 7 7 8. Ca ls' . - - - - - - - - - - - - - -
//-
SOTTQhl_QE_iTA\.L. 4. POOL FIG. 19
-3
m . ok
,Qs*fp er s.
9 Y & ,,, IMAGE EVALUATION NNNN TEST TARGET (MT-3) 2 1.0 5 m 014
+= F32 l -g22 l E ne E ; lllll=2.0 ;
I.I 3__
. . . , =
l 1.8 1.25 IA 1.6 N 4 6" > MICROCOPY RESOLUTION TEST CHART l t
#% + //p by , $/
ms
, #$+,@
,, e f, v .
' ;:! C Wy):;Q l ~, :,. '
l ' :).; k,,& 4{$* 0 :l -
}, 's:;: l:;7 ggb t i
-b sv Y , ,
l ,
' b, -,) ', -ji' .
L_ _
%_ _ - u _ < - u2 .o J ( D a -. ,
_- -. . - _ . . - . - - . _~. 1 i 3 APPENDIX 4 I Union Carbide Nuclear Reactor Startup Accident Analysis 1
- 1. Summary I
A new analysis is made of the UCNR'startup accident using revised physics parameters appropriate to the operational core. The most important-revision is the reactivity insertion rate, now 8x10-0 $ b per sec. Both an analysis following the method used in the Final Hazards Summary Report
- (Ref. a) and one based on the results of SPERT I and SPERT IV tests 1
{ (Refs. b,c) show that even with a safety-system trip level of 10 MW, or 200% of full power, the resulting maximum fuel plate surface-temperature is more than 450cc below the melting temperature of aluminum. The' shortest
!_ period expected is more than 40 milliseconds, or well above that at which I ; even minor fuel bowing might occur (15 mi.llisecs.j. It is concluded that a 200% trip level is more than adequate protection for the startup l-
- accident.
i
- 2. Physics Parameters i
t For use in-this re-analysis of the startup accident, some of the core
-physics parameters have been revised to be more appropriate to the opera-
- tional core or to reflect improvements that have been made in basic data.
With one exception, the revisions are not large and do not affect the 1 l results of the analysis to any important extent. The factor of most I importance, and one that is markedly different, is the 5-rod worth, for this affects the reactivity insertion rate. The-original analysis as-given in Ref. a, in some instances used calculated parameters that have since been measured; also some errors were made. The revised values are I as follows: i i I d
Total 5-rod worth: 11.6%f Max. 5-rod differential 0.67%K E/ In. = $0.84/in. (measured) worth: Max. reactivity insertion ra te (R) : .0067x8.33x10-2+2.0x10-4=7.6x10-4f/sec. Core size: 6x7 over-all (28 std. fuel, 7 Partial fuel) Al/H2O volume ratio: 0.49 (entire core) 1 Core volume: 156 liters Core loading: 5.19 Kg U-235 (min) Macroscopic fission cross-sect. Icf
=
.045 cm-1 Macroscopic absorption cross-sect. E ac
=
.070 cm-l j Thermal neutron diffusion length, L2 = 3.43 cm2 c
K. = v = 1.57 Eac Tc = 45 cm2 (Ref. d, p. 132) B2 =
.009 cm-2 i
1 = 5.6x10-5 sec, 8 = .0065 (Ref. d, pp. 17,19) Seff = .0081 (Ref. d, pp. 132, 443) Startup source = 8x107 n/s (50-curie Sb-Be) Shutdown K = 0 942 (50% shutdown margin)
, 3 Startup Accident Analyses The same basic assumptions as stated in Ref. a are used here, with the following exceptions:
The withdrawal of all six rods requires defective rod control circuitry. Operator inattention can result in the withdrawal l I
4 A of only two rods at a time. The maximum differential worth is the measured value quoted pbove. The initial analysis follows exactly the same procedure as in Ref. a, using the Newson inequality to estimate the shortest period for the case of a safety-system limited excursion. This is followed by an analysis of similar but self-limiting excursions performed with the SPERT reactors. 3.1 Safety-System-Limited Excursion The initial power level Po is taken conservatively low. No credit is taken for the fact that the startup source has attained an equilibrium activity of about twice its initial value of 50 curies and that the photo-neutron source has been measured to be about three times this value at the
; usual cycle startup time. Then, following the meti'od of Ref. a, we have:
, Po = ] 8 x 7 5 x 10iu
=
0.92 x 10-2 watt i f the safety system trip aower P1 is taken as 7 5 MW (150%-trip . level), the shortest period T will be:
-5.6 x 10-5 --- 1/2 T> a 42 msec.,
2 x 7.6 x 10'N x23logh__. _ o orre:Iprocalperiod.a=f<24sec-I It can be seen that the presence of the (log ) term makes the estimation
~
o of T (or a) very insensitive to trip level. For example, even if P1 were 10 MW, T would still be about 42 msec. This means that the reactivity at trip level is esse <.tlally unchanged, and likewise the various rod insertion times. The estimates of maximum power and released energy for various trip level pcwers between about 5 and 10 MW can therefore be obtained by simple proportion. This analysis will therefore be performed for P1 = 10 MW. )200% trip). For T = 40 ms, the reactivity at trip level = $1.28 (Ref. d, p. 447). _ . . - - - - -, , y
Rod insertion needed to reduce reactivity to prompt critical is:
, Sj = = 0.33" Time required for rods to drop 0 33" at 1/2 g is:
tj
=(f61 12)
= .058 sec.
Total rod insertion required to reduce reactivity to zero is: S3= ', = 1.52" Time required for 1 52" rod drop at 1/29 is: t3 = ( 6 1^ 1.52 )1/2 Time for reactivity to go from $1.00 to zero is: t t2 = t 3-tj =
.067 sec.
Maximum power mP = 10 e = 130 MW Total energy W = Wj + W2+W3
= PIT + T(Pm-Pj) + Pmt2
,. = Pm (T + t2)
= 130 x .108 = 14 MW-sec.
The values of Pm and W corresponding to a trip level of 7 5 MW (150% trip) are, by simple proportion, 98 MW and 10.5 MW-sec respectively. It should be observed that even with a 200% trip level, the total energy is some 2-1/2 times smaller than the damage energy threshold of 37 MW-sec. cited in Ref. a (p. 77) for BORAX. 3.2 Self-Limiting Excursions The results of the SPERT l and SPERT IV experiments (Refs. b, c) can be applied to the Union Carbide Nuclear Reactor (UCNR). Not onl/ are l.
Y some of these plate-type cores quite similar to the UCNR (Ref. a, Table XIX; Re f. c , Ta b l e s I , 11), but the results of tests made with SPERT cores having a wide spread of shutdown characteristics show a remarkably small spread (Ref. b, pp. 20-24). For ramp reactivity insertions it is found (Ref. b) that the characteristics of the resulting excursions are equivalent to those of step insertions having the same maximum reciprccal period (alpha). Accordingly if the latter can be estimated for a ramp then the great body of data on step-induced excursions can be applied in assessing the consequences. It is found, both by theory and by experiment, that the most sensitive parameter is the ramp rate and that other quantities like initial power (Po above), and core reactivity coefficient are of secondary significance. Foraramprateof9.10-Nf/sec,thepromptandtheHurwltz theoretical models in Ref, b predict reciprocal periods (alphas) of about 19 and 17 sec-1 respectively. The experimental data for core A-17/28 gave about 15 sec-l. These figures are all less than the maximum alpha of 24 sec-1 predicted by the Newson inequality (see above, 3.1). The latter is evidently a conservative (safe) estimate. From these various results is appears that an alpha of 20 sec-I would be a reasonable estimate for a ramp rate of 7.6x10-4 /sec with the UCNR. The results given in Table B-1 of Ref. c show the following excursion characteristics for a core sub.m ged 18' in water: Fuel Plate Cooling Water Energy at Max. Surface Flow Speed a T P max Temp. Rise Pm time (ft/sec) (sec-I) (msec.) (MW) (MW-sec.) (OC) 0 20 50 25 1.8 124 0 34 30 68.5 30 141 2.4 18 56 22.5 2.35 136 6 19 53 22.5 25 110
The UCNR core is submerged 25 ft. and the water flow speed at 2400 gpm is about 4 ft/sec. The effect of increasing water head is to increase the maximum fuet plate temperature. For an alpha of approxi-mately 50 sec-1, Table B-1 of Ref. c indicates an increase of 240C in fuel plate temperature in going from a 2' to an 18' head. This increase is about the same as the increase in saturation temperature of the water, viz. 200C. The additional 7' head in UCNR is therefore not expected to increase this temperature by more than about 100C. The significant quantity as far as core damage is concerned is the fuel plate surface temperature. This will be greatest under no-flow conditions. For a = 20 sec-1 (50 millisec. period), initial water tempera-ture = 400C, and a 25' head the estimated maximum fuel plate surface temperature is then 124 + 40 + 10 = 1740C. For a = 34 sec-I (30 ms period) this temperature is 191oC. Clearly these temperatures are well below the 6580C melting temperature of aluminum. In general the results of the SPERT IV tests (Ref. c) indicate that for a < 70 sec -1 (15 millisec. period) there will be no core damage at all; for a 2 /0 sec-1 minor fuel bowing may occur; and for melting to occur a would have to be about 170 sec-l (6 msec period). The postulated UCNR startup accident will not generate an alpha anywhere near as great as 70 sec-1
- 4. Conclusions The self-limiting tests described above show consequences (peak power, energy) that are smaller than those predicted for the safety-system-limited s i tua t i o.,. It can thus be concluded that the inherent self-shutdown characteristics of these plate-type reactor cores is such as to reduce importantly the power and energy generated in the prompt excursion initiated by a ramp (or step) reactivity insertion. The main effect of the safety-system is then to terminate the trailing edge of the power burst and to prevent subsequent power oscillations, in rough illustration of this, Fig. B-7 of Ref. c shows that if a safety scram were initiated at less than 0.2 sec (c.f. Sect. 3.1 above) the trailing edge of the power trace would 1
O be reduced rapidly af ter falling below about 15 MW, and the accumulated energy would be limited to about 4 MW sec. It is further concluded that a safety-system trip-level of 7 5 MW (150%) or even of 10 MW (200%) is safe insofar as the UCNR startup accident is concerned. 5 References
- a. Union Carbide Corporation, " Final Hazards Summary Report 'JCN C Research Reactor", Nov. 1960.
- b. Phillips Petroleum Co., 1 D0-16528, " Analysis of Sel f-Shutdown Behavior in the SPERT l Reactor". July 1959
- c. Phillips Petroleum Co., 100-17000, " Reactor Power Excursion Tests in the SPERT IV Facility", Aug. 1964.
- d. Argonne National Lab. , ANL-5800, " Reactor Physics Constants",
2nd Edition, July 1963 i-i i i l
2 APPENDIX 5 TOP 0 GRAPHY AND GE0 LOGY OF THE SITE
- 1. BROAD PHYSIOGRAPHIC AND GEOLOGIC SETTING OF THE SITE Three physiographic provinces characterize the eastern borderland of the United States. These provinces extend in belt-like fashion roughly parallel to the Atlantic coast from Canada to Florida. They are: (1) the Atlantic Coastal Plain, a lowland underlain by Cretaceous, Tertiary a:id Recent sediments, that dip seaward at relatively low angles. The strata are unmetamorphosed and, indeed, only semi-consolidated; (2) the Piedmont Province, a low-land belt of highly metamorphosed sediments and igneous rocks, onto which the Coastal Plain sediments overlap along their " Fall Line" v.estern boundary, f rom the vicinity of New York City to southern Georgia. This gently seaward-sloping, dissected plateau is an erosional feature carved out of the mountains. It thus lies at the foot of the third province; (3) a mountainous belt of complexly folded and faulted metamor-phic rocks which in many ways are similar to those of the Piedmont Province.
In geologic age, these hard rocks range from Precambrian through the Paleozoic. Still further west lies the Valley and Ridge Province of folded Paleozoic strata, and yet further west the Appalachian plateau, a highland where rocks are but little deformed. in places the generally highly metamorphosed hard rocks of the Piedmont Plateau are displaced by down-faulted segments of Triassic unmetamorphosed sedimentary strata. Such a down-faulted block occupies much of eastern New Jersey. Its western boundary is sharply marked by a prominent fault scarp trending northeasterly and southwesterly from the town of Suffern. Figures 29, 30, and 31 s'ow the location of the plant site, which lies well within the mountainous welt known as the New Jersey-New York Highlands, or more technically as the Reading Prong of the New England Highlands.
- 2. TOPOGRAPHY AND GE0 LOGY OF THE AREA I
IMMEDI ATELY SURROUNDING THE REACTOR SITE The reactor site is in Sterling Forest, 3-1/4 miles north-northwest of Tuxedo Park, Orange County, N.Y., some 1500 feet southwest of Indian Kill, a small stream flowing southeast for a mile and a half to the Ramapo River. The plant borders Long Meadow Road at an elevation of approximately 800 feet. There is a very low north-south topographic divide between Indian Kill drainage and drainage of Warwick Brook to the south which also flows east to Ramapo River. These two small streams, Indian Kill and Warwick Brook, since they drain into the Ramapo River from the vicinity of the site dominate _j-4-
1 the drainage pattern insofar as it concerns the flow of surface or underground water away from the vicinity. The Ramapo River, cutting athwart the mountain-ous highlands and rising to the north near the town of Monroe, flows in a mountain walled valley to Suffern, and thence to Passaic River at Mountair. View and so on to the sea. The flow of the river varies with the season (Ref. 29). During the " water year", October 1952 to September 1953, the mean daily flow for the month of March was 611 cubic feet per second. During the dry month of September the flow fell to 16.3 cfs. An average of the monthly means for this year was 269 cfs; for 31 years this average was 218 cfs. During a record of 31 years, occasional floods attained a rate of 12,400 cfs for a period of a day or so. During very dry spells the rate fell as low as 7 cfs. Though the " Highlands" are ruggex, and the hillsides steep, relief is not great, only a matter of some 400 to 700 feet from the valley floors to ridge tops. A striking feature of the area, a feature resulting from a past era of glaciation, is the clearly evident clogged drainage system. , Swamps and ponds abound along stream channels, as do a multiplicity of lakes, large and small, all bespeaking the fact that present streams have not, under prevailing gradients and climatic conditions cleared their over-burdened channels of glacial debris: fill, clay, sand, gravel and boulders of every size. These latter, especially, strew the hillsides. The reactor building is placed in a north trending spur of Hogback Mountain, a spur which slopes northward from something over 1500 feet of elevation to the level of Indian Kill Lake at 700 feet elevation. it is at the eastern foot of this spur, along Long Meadow Road, that the plant is located. The detailed geology of this area has been discussed in a report enti tled " Magnetite Deposi ts of the Sterling, N.Y. , Ringwood, N.J. Area"
~
by Preston E. Hotz of the United States Geological Survey (Ref. 30). The area is under!ain by closely folded sedimentary and igneous gneisses. The folds are in general overturned to the northwest, resulting locally in relatively steep isoclinal dips toward the southeast. The middle section of the eastern spur referred to above is underlain by quartz-oligioclase gneiss. This rock is regarded as a highly metamorphosed sediment, parts of which have taken on the aspects of an igneous rock. The " outcrop" width of this rock approximates 500 feet. From the data collected from core drillings covering the construction area, it was found that a drill placed at the center of this outcrop would not penetrate the western foot wall of the gneiss until a depth of nearly 400 feet was reached. Actually, the hole was completed at a depth of 200 feet, the approximate claculated elevation of the floor of the reactor building. To the east of the site folded metamorphosed sediments (quartzites) occur, and beyond these, granitic gneisses. These may or may not be derived from ancient sediments. Similar sequences of gneisses extend eastward to Ramapo River and beyond.
- 3. TOPOGRAPHIC AND GEOLOGIC RELATIONSHIPS AFFECTING THE PLANT
- a. Excavation of the Reactor Building The rock is quartz-oligioclase gneiss. In texture it varies from coarse to fine grained, but these differences are interpreted to mean not so much dif ferences in the grain size of the original sediment, as that metamorphic processes have obliterated, to varying degrees and in various places, the original bedding planes of the rocks. That is, parts of the rocks could hardly, if at all, now be distinguished from an igneous rock.
The rock is very dense and tough and relatively free f rom fractures, though all rocks near the surface are somewhat f ractured. The first drill hole produced nearly 98% of core, the second hole drilled at an angle of 660 from the horizontal produced 99% of core. A compression test conducted at the Brooklyn Polytechnical Institute revealed that a cylinder 4-1/4 inches long and 2-1/8 inches in diameter broke only when subjected to the great load of approximately 8 tons to the square inch, even though the section was cut with planes of schistosity dipping 55o from the axis of the cylinder. l All rocks near the surface are somewhat f ractured, and rain water will enter these fractures at the surface and descend to points of escape or to points where fractures give out. Thus, even in fractured hard rocks a water table is built up. In such rocks water descends to depths where fractures die out, beneath which position such rocks may be essentially dry. In core hole No. I water appears to stand at 85 feet i below the surface. Thus, there is 115 i feet of water in the hole,- which suggests that fracturing at depth is slight,
- b. Drainage i t does not appear that contamination of Tuxedo Lake or any of the connecting smaller lakes from plant operations is possible for reasons noted below.
Surface drainage from the site is exclusively by way of Indian Kill. The Kill enters Ramapo River 1-1/2 miles east of the plant at elevation 463 feet. Tuxedo Lake stands at elevation 560 feet. Wee Wah, the adjoining lake to the north stands lower than Tuxedo Lake to which it is joined by a small stream of high gradient. Wee Wah Lake consists of two segments. The southern, higher segment is separated from the lower northern segment by a stream of steep gradient. This northern segment, in turn, discharges over an earth dam and masonry spillway to a small stream that discharges into Ramapo River. Thus, it may be seen that even if the Indian Kill were contaminated, it is not remotely possible to carry such contamination by surface flow to any of this chain of three lakes. l
Indian Kill presents the only obvious path for contamination by undergroun? flow, that is, through alluvial sand, silts and gravels that lie beneati che stream channel, resting on the gneissoid bedrock of the region. Water passes downstream easily but slowly through these alluvial deposits. Obviously such waters could not possibly ascend into the chain of Tuxedo Lakes. Water passing underground beneath the mountainous ridges, through the fractures in the hard rocks does not seem possible. The mountainous tract which is bounded by Indian Kill, Long Meadow Road, Warwock Brook and Ramapo River naturally contains some ground water within fractures in the rocks. But this water drains outward to the nearest and most accessible exist, namely either Indian Kill, Warwick Brook or Ramapo River. Water cannot pass against this outward flow, across this mountainous tract, and even assuming it could, it could not pass the boundary of Warwick Brook which flows east to Ramapo River. Therefore the possibility of contamination of the Tuxedo Lakes chain by water from the vicinity of the plant may be dismissed.
- c. Earthquake Hazard The New Jersey-New York Highlands have a long record of freedom from violent earthquakes. There is no historical record of earthquakes of intensity Vill or greater occurring in this area.
The swimming pool reactor is constructed in very firm hard rock. Only a very violent shock could affect the reactor in the chamber in which it is placed. Such a structure would vibrate as a unit even under violent shock, which is not expected. In the past 20 years, the number of seismic recording stations in this general area has increased steadily to the point where detection of events adjacent to New York City of magnitudes > 1.8 is practically complete. Coupled with interest in the seismic safety of the power reactors located at Buchanan, N.Y., this has resulted in intense study of the seismic activity in the vicinity of the Ramapo Fault and estimates of the probability of occurrence of strong earthquakes at those reactor sites. In Ref. I to this Appendix a detailed discussion of these matters appears. What follows is excerpted from this reference. The UCNR reactor site is~ interior to a lithospheric plate. In such a case, potential earthquakes tend to occur along major pre-existing faults, with the larger shocks showing a greater tendency than the smaller ones to be located on the major throughgoing fault. A survey of the seismic events that have occurred in and around this region shows that most of the activity is in the Precambrian Hudson Highlands, that earth-quakes in this area occur along pre-existing faults, with the large majority within 1-2 Km of the faults. About 50% of all events are almost collinear and lie along or close to the Ramapo Fault system. Using data on small shocks obtained in the past 20 years or so from the seismic network, a relation for cumulative frequency of occurrence has been obtained. . Extrapolation to larger magnitudes shows _4_
e i excellent agreement with the few larger historical events, of intensity VI and Vll. This relation has been used by the authors (Ref.1) to predict the probability of occurrence of intensity Vil and Vill shocks at the upper end of the Ramapo Fault. The results realistically are as follows: Recurrence Period Probability of Occurrence (years) in 20 year period (%) Vll Vill Vil Vill (a) Excluding events
> 10 Km distant: 630 2870 32 07 l
(b) including all events along Ramapo Fault: 340 1880 5.9 1.1 The UCNR site is about 12 Km northwest from the Ramapo Fault and therefore t the predicted probabilities at this site should be less than those given above, viz. less than 3-6% for intensity Vil, and less than 1% for Intensity Vlli, for a 20 year period in the future.
- 4.
References:
i j 1. Aggarwal, Yash P. , and Sykes, Lynn R. , " Earthquakes, Faults, and Nuclear Power Plants in Southern New York and Northern New Jersey", i Science 200, 425 (Apr. 1978). 1 I I
- - , , -.. ,,,--m . - - - -.
2 APPENDIX 6 CLIMATOLOGY OF THE TUXEDO PARK-STERLING FOREST AREA 1 lNTRODUCTION Airborne radioactive materials are not considered dangerous unless they are carried by. wind and rain in critical concentrations to populated areas. Normal weather conditions tend to cause the dispersal of contaminants throughout the atmosphere by the process of turbulent diffusion so that safe dilutions are reached within a short period of time and before the material has traveled more than a short distance. The meteorological processes which are most important in any study of the pollution of the atmosphere are those which tend to reduce the normal mixing which goes on within the lowest air layers. By preventing diffusion, they serve to concentrate harmful pollutants at or near the ground where they can be dangerous to human beings. Three meteorological elements are of significance in determining the amount of mixing which will occur at y time and what will happen to contaminants in the event diffusion into the free atmosphere is restricted. They are the distribution of temperature with height, the surface wind direction and velocity, and the precipitation. The most important of these elements is the vertical distribution of temperature as it is related to possible low-level temperature inversions which may develop within 1,000-1,500 feet of the ground surface. Under such conditions instead of a normal decrease in temperature with elevation, the air aloft is actually warmer than that near the ground. This results in a stable layering of the atmosphere with a lid-like effect being created at the inversion level. Such conditions occur frequently during clear, calm nights when radiation f rom the surface of the earth is at a maximum and there is little or no atmospheric turbulence. They also occur, however, for. extended periods of time during certain types of weather. When the capping effect of a temperature inversion has concentrated air-borne pollutants near the ground, horizontal movement of the air beneath the inversion becomes very significant. It is important to consider the surface wind direction and velocity at such times. The interrelationship of a third element, precipitation, with the first two, temperature inversion and winds, is necessary as well because falling drops of moisture remove contaminants from the air and concentrate them on the ground or in streams where people may come in contact with them. l I
- 1. SURFACE FEATURES The reactor site is located on the east-facing slope of Hcgback Mountain near Tuxedo Park, New York. It is in a small valley which comprises a part of the indian Kill drainage to the north and Warwick Brook on the south. Both of these small basins lie within the Ramapo River drainage which is less than a mile to the east at its nearest point. The Ramapo, in turn, is tributary to the Passaic River.
Physiographically, the area lies in the southeastern section of the Ridge and Valley province (Fenneman, 1938), often called the Folded Appalachians because the topography is dominated by structures due to folding, it is a part of the Reading Prong, a narrow southwestward extension of the New England Upland province, which crosses the northern tip of New Jersey into Pennsylvania, ending at the Schuylkill River near Reading. The Ridge and Valley province is made up of a combination of valley floors and long, narrow mountain ridges with their tops at nearly the same elevations. The mountains may be widely spaced and isolated, with wide valleys (that of the Hudson River, for example) or they may be close together so that the lowlands are disconnected or entirely absent. The ridges are resistant remnants of folded mountains and the lowlands are the result of the erosion of the weaker rocks. The mountain crests, while seldom very broad or flat-surfaced, are at about the same level. This evenness of crest over a wide area has a definite effect on atmospheric circulation. 4 The ridge and valley terrain has an important influence on atmospheric mixing and hence on diffusion of contaminants at the lower levels. Holland's (1952) studies of atmospheric diffusion showed that ridges with an average height of 300 feet above the valley floor can exert a definite influence by increasing the surface drag on the atmosphere. He observed in the Oak Ridge, Tennessee, area the reduction in wind speed was greatest near the valley floor. This is a most significant fact, for the problem or stratification of air in the valleys and the wind flow around the pronounced topographic features distinguishes the pollution climatology of a hilly or mountainous region from that of a level area. Holland also noted a variation in the local airflow pattern between night and day: at night the ridges had the effect on local wind movement of a massive obstacle, while in the daytime its effect was more that of a shallow wave with appreciably lower surface drag. This condition, coupled with the fact that radiosonde data show that most inversions occurred during night or early morning hours, adds importance to the influence of the steep-sloped ridge and valley terrain of the area as a factor in the dissipation of temperature inversions and the resultant dilution of concentrated airborne contaminants.
- 2. GENERAL WEATHER CONDITIONS The most reliable and longest period surface weather observations for the area have been taken for more than a century at Wes'. Point Military Academy, approximately 18 miles east-northeast of the proposed reactor site.
The Weather Bureau has cooperative observing stctions at Suffern and Warwick, 1 New Yrok, and Greenwood Lake and Ringwood, New Jersey, while the Air Force recently operated an upper air sounding station at Stewart Air Force Base about 20 miles northeast of the reactor site. While the climatology of a hilly region is strongly influenced by local topographic features so that observations from surrounding weather stations may not be entirely representative, still we can obtain from the dense network of observing stations and from field observations a fairly good idea of the climatic picture in the vicinity of the reactor site, in addition, hourly obser-
, vations of wind speed and direction at the reactor stack elevation have been made for the past 18 years during reactor operation. The results i of a sampling of these local observations are given later and agree well with the Stewart Air Base data.
The climate of the Sterling Forest area is predominantly influenced by air mass movement and prevailing winds from an inland direction. The weather across the area from west to east at average velocities of 30 to 35 miles per hour in winter and somewhat more slowly in summer. This is a part of the normal cyclonic circulation in which the usual weather-producing low pressure systems follow paths toward the northeastern United States. About 40 per ent of the low centers pass over or close to south-eastern New York. Most of the others come close enough to exert an influence on the area's weather so that there is a regular change in weather patterns without any consistent periods of stagnation. Centers of high pressure alternate more or less regularly with the lows. In the winter-time, their movement is variable, depending on the strength of cold air outthrusts f rom arctic area to the northwest. This movement is slowest during summer and early fall so that, with the prevailing
, westerlies alof t reaching their most northerly movement at the same time, j high pressure centers often become stationary for several days over the l area during these seasons. The result is stable atmospheric conditions which encourage the formation of temperature inversions. Such conditions i
can occur in any month, but persistent stability of several days' duration j occurs on the average only once in several years. Cold air masses of the continental arctic or continental polar types , dominate the area's weather in the fall, winter, and spring. These are very stable at their northern source, but by the time they have reached southeastern New York, having been heated from below as they moved across l the land, their lower layers are generally unstable. During the summer, l the continental outbreaks of cold a'r are weak and maritime tropical . air masses migrate northward to exert an effect on the weather of the area. L
At this time of year, nocturnal cooling results in frequent temperature inversions, but they are most of ten short-lived because of the heating which occurs during the day, resulting in turbulence and mixing of the atmosphere. Low-level inversions, as defined for this analysis, are those which occur between the ground surface and the 950-millibar pressure level, a height difference of approximately 1700 feet. The average diurnal temperature range at the surface is consistently close to 12 C over the entire year. Under standard conditions, a temperature rise of this amount at the surface would just destroy an inversion of 8 C between the ground and 950-millibar levels. Therefore, it is seasonable to consider as
" marked" a temperature inversion strong enough to persist despite a surface temperature increase of as much as two-thirds of the normal range, namely:
2/3 x 12 C = 80C. The pseudo-adiabatic diagram shows that such an inversion amounts to a rise of 40C from ground surface to the 950-millibar pressure level. Figure 1 shows the frequencies of occurrence of all inversions and
" marked" inversions for the years 1950 and 1951 by month. These data, obtained from the upper air soundings taken at Stewart Air Force Base, 20 miles f rom the reactor site, are summarized by percentages in order to indicate the relative distribution of inversions throughout the year.
There were 95 inversions during 1950 and 101 in 1951. Most of the inversions (85 of the total of 95 in 1950) commenced during the night and few of them persisted through the day. In fact, for the two year period cited above, only 21 inversions held for more than 12 hours and only 6 lasted more than 24 hours. inversions were well distributed through the year with no apparent seasonal preference. The pattern of prevailing surface winds throughout the year is shown in Table 1. Monthly wind roses are included in Figure 2. These wird roses are derived from monthly mean values of observations taken at Stewart Air Force Base from September, 1942, to December, 1952. Only eight points of the compass are employed for ease of interpretation and to emphasize the wind direction values. Four categories of wind speed, other than calm, are considered; they are: 1-3, 4-12, 13-24, and greater than 24 miles per hour. Numbers of observations and average percentage f requencies by month and by wind direction are recorded in Table 11. These data show that the predominant wind direction are southwest and west and that combined with periods of calm, they make up 56.1 percent of the observations. Winds from the north and northwest comprised only 14.4 percent of the yearly total. Wind speeds from the south and southwest averaged 10-11 miles per hour; from the north and northwest 11-15 miles per hour, in March, the north and northwest winds are more frequent than at any other time. During the Summer and Fall, the former directions constitute as high as 45 percent of all winds; in the Summer period the predominant wind speeds tend to be lowest, i.e., in the 4-12 miles per hour range. 1
-4
TABLE I PREVAILING SURFACE WINDS, STEWART AFB, NEW YORK, 1943-1952 PERCENT FREQUENCY Wind Velocity Wind Direction Total (MI .Per Hr.) 0 45 900 135o 1800 2250 2700 3150 Calm Annual Calm 12.4 12.4 1-3 0.4 0.5 0.4 0.3 0.3 0.5 0.4 0.3 3.1 4-12 4.2 6.7 4.3 3.8 5.6 16.1 7.7 35 51.9 13-24 2.3 3.1 0.8 1.3 2.1 8.3 8.3 2.7 28.9
> 0.3 0.1 0 0 0 0.6 1.9 0.8 37 Calms occur slightly more than 12 percent of the time on an annual basis and vary from a low figure of 9 percent in March and November te. 17-18 percent in August and September. The influence of down-valley drif t and the f requent formation of temperature inversions in valleys due to nocturnal coaling would tend to localize the distribution of air contaminants to the genera! vicinity of the reactor at those times.
Tables ill and IV give the direction of the surface winds during all inversions and during marked inversions for the years 1950 and 1951. Again, the winds are indicated to eight points of the compass for convenience, it can be seen that the wind directions found during the periods of inversions were quite similar to the prevailing wind directions at other times of the year. While inversions occurred, surface wind direction was from the south-west 31 percent of the time during both 1950 and 1951. Calms existed with more than 20 percent of the inversions. It is important to recognize that winds from a southerly or easterly quadrant would bear contaminants away from the principal urban centers of the New York City area and into the thinly populated ridge and valley terrain or toward the open spaces of the Palisades Interstate Park. A third factor of importance in evaluating possible contamination of the atmosphere and its effect on neighboring population is the occurrence of rain or snow, especially at the time of a temperature inversion. The ef fect of precipitation would be to wash out of the air the radioactive particles and to concentrate them at the earth's surface. If these particles had already been concentrated somewhat by the presence of an inversion, the occurrence of precipitation would intensify the possibility of increased levels of radioactivity at the surface. The possible effect of precipitation on the pollution hazard has been investigated by: (1) analyzing the general precipitation pattern for the area in terms of amount and frequency; (2) associating precipi-tation with times when temperature inversion exists; and (3) examining the converse situation, i.e., the probability of the occurrence of rainless days. TABLE II AVZP. AGE y:scaCY o? FFEVANG SURFAC3 '4DD D.trzCTICN, STT4 ART AFB, ?rf/ YoE,19k3_1952 Supace Wp Dipeticg Total No. 0 cals of cbs. Month o E5 90 135 180 225 27 0 315 Janua:-r No. of Cbs. 612 903 362 191 347 1818 1788 523 892 7436 Percent 83 12.2 4.8 2. 6- 4.7 24.h 24.0 70 12.0 100.0 February No. of cbs. 4c8 789 h 3o Ek6 371 1382 1895 545 814 6790 Percent 6.0 n.6 50 36 55 20.4 27 9 8.0 12.0 100.0 March 642 486 1598 1571 688; 672 7434 No. of cbs. 878 510 379. Percent 8.7 n.8 70 51 65 21 5 21 1 93 90 1co.o Arr11 No. of cbs. 553 736 h62 447 583 1625 1328 773 686 7193 Percent 77 10 3 6.4 6.2 8.1 ' 22.6 18 5 10 7- 95 100.0 Mar No. of Cbs. 498 879 645 694 675 1567 989 592 897 7436 Percent 67 n.8 87 93 91 21.0 13 3 8.0 12.1 1co.o June No . of Obs . 379 587 354 485 814 218h 1013 E5h 922 7192 Percent 53 8.2 49 6.8 n.3 30.4 14.0 63 12.8 ico.o July . No. of cbs. 3Eo 576 3c6 551 882 2355 lo70 303 1054 7E37 Percent 4.6 7.8 4.1 74 n.8 31.6 14.4 E.1 14.2 100.0 August No. of cbs. 555 694 3E2 442 7Co 2047 870 E50 1338 7438 Percent 75 93 E.6 59 93 27 6 n.7 6.1 18.0 1co.o sente=ber No. of Cbs. 614 906 465 547 776 1787 10h6 hh5 1320 7906 Percent 78 n.4 59 70 98 22.6 13 2 56 16 7 ico.o October No. of Cbs. 774 962 456 h14 7Co 2236 n85 h65 991 8183 Percent 94 n.8 56 50 8.6 27 3 14 5 57 12.1 1co.o November No. of Cbs. 580 783 424 342 6he 2059 1686 692 Tc6 7914 Percent 73 99 53 4 3 , 8.1 26.1 21.E 87 89 ico.o Decenber No. of Cbs 583 714 325 201 371 2353 2107 550 974 8178 Parcent 73 87 4.0 2.4 E.6 28.8 25 7 67 H.9 1c0.0 Total No. of cbs. 6538 9407 5c01 E939 7347 230 n 16548 6h80 n266 90537 Percent 73 lo.h 56 54: 8.1 25. E. 18 3 71 12.4 1c0.0 TABLE III SUP7 ACE 'JINES IURING I3 VERSIONS. STr/ ART A?3. N?J YORK, 1990 All Imrersicns Surface 'Jind Direction Percent Month CO 45 0 900 1350 1800 2250 2700 3150 Calm Total of Total 4 4 3 11 12 Janua:7 Febraary 1 3 1 1 6 6 March 0 0 April 2 2 3 2 4 13 14 Nf 1 2 ' ' 1 , 1 7 7 June 2 2 5 2 11 12 July 1 1 2 4 1 2 11 12 Ausust 3 4 7 7 Septe=ber 1 1 2 2 October 3 1 3 2 1 10 10 Nove=ber 1 3 1 5 5 Dece=ber 1 4 2 1 2 1 1 12 13. Total 1 8 3 8 15 29 9 4 18 95
?erce=t of Total 1 8 3 8 16 31 10 4 19 100 Marked Inversion _s Janua:/ 1 3 1 5 33 February 1 1 7 March 0 0 April 1 1 T Mf C 0 June 2 1 3 20 July 0 0 e August 0 0 Septe=ber - 0 0 October 1 1 7 Nove=ber 1 1 '
2 13 Dece=ber 1 1 2 13 Total 0 1 0 1 3 7 2 0 1 15 Percent of Total 0 7 0 7 20 46 13 0 7 100 i
~7- . . __ _ _
TABIS IV SURFACE WUDS LURriG LWEPSICUS, STEWART AFB, NEW YOPL 1951
~
All Inversions Surfacg Wind Direction ?ercent 0 0 0 Month 0 45 90 135 180 225 270 315 Cab Total of Total January 1 5 8 8 February 1 1 ' 3 2 March 2 1 3 2 AW2 2 1 2 1 6 6 May 2 1 2 4 1 1 11 11 June 1 1 1 1 2 6 6 July 1 3 1 1 1 7 7 August 1 2 1 2 6 12 12 September 1 2 1 4 6 6 20 20 Cetober 1 1 1 1 1 3 8 8 November 1 3 1 2 7 7 December 3 2 2 3 10 10 Total 4 8 0 7 9 31 9 8 25 101 Fercent of Total 3 8 0 7 9 31 9 8 25 100 J Marked Inversions January 1 1 2 13 February 1 1 2 13 MLmh 1 7 2 13 April 0 0 May 1 1 2 13 June 0 0 July C 0 Augast 1 1 7 September 1 1 1 3 20 October 1 1 7 November 0 0 December 1 1 2 13 Total 0 3 0 2 1 3 2 1 3 15
?ercent l
of Total 0 20 0 13 7 20 13 7 20 100 r TABG V NUFRA OF CCCURETCIS OF FRECIPITATION AT TrE OF IT. EES!CNS, STF/ ART AFB, NT4 YORK, 1950 All Imrersions Percent Surface Wird Direction Month 00 45 90 135 0 1800 2250 2700 313 Cals Total of Total 2 2 7 21 January 3 2 1 3 9 February 0 0 March 6 18
- 2. 1 1 2 April 1 1 3 May 1 1 1 3 9 June 4 13 July 1 3 2 2 0
'Au6:St O O September 1 1 3 Cetober 1 3 November 1 2 1 1 1 5 15 December 0 3 1 5 5 8 3 1 7 33 Total ?ercent 100 of Total 0 9 3 15 15 24 9 3 21 Marked I=versiens 9
1 2 3 30 January 1 1 17 February 0 0 March 1 17 April 1 O O May ' 0 0 Ju=e 0 0 July 0 0 August 0 0-September 0 0 October 0 0 Ncvember 16 1 1 December 0 0 0 0 3 3 0 0 0 6 mal
?ercent 0 0 100 l
0 0 0 0 50 50 0 of Total l ( _g_ i
TABL2 VI NUIN"4 0F CCcuesCES OF FRECIPITATICN AT TDfE OF IN7ERSICNS, STf4 ART AFB, NFJ YOHC,1951 All Inversiens Percent Sur' ace Wini Direction 2 nth 00 45 0 900 135 0 1c00 225 0 270 0 3150 Cab Total of Total 2 1 L 4 13 January 0 0 February 6 1 L 2 March 1 1 3 April 2 1 3 9 May 6 1 1 2 ' June 1 1 1 3 9 July 1 1 2 4 13 Au6ust 2 4 13 Septe=ber 1 1 1 1 2 6 October 6 1 1 2 November 2 1 2 5 16 December 0 0 4 2 9 4 2 8 32 Total 3 Percent 100 of Total 0 9 0 13 6 26 13 6 25 Marked I:rmrsions January 1 1 2 29 0 0 February 1 1 14 March 0 0 April 1 1 14 May 0 0 June 0 0 July ' 14 August 1 1 1 14 September 0 0 October 0 0 N<:nrember 1 1 14 December 0 1 0 2 0 0 1 0 y Total a i i
?ercent 100 of Total 0 14 0 29 0 0 14 0 43
West Point is 18 miles east-northeast of the Sterling Forest site ' and only 10 miles f rom Stewart Air Force Base. Rainfall data for this station have been used because the annual averages are about the same as those for Stewart Field but the period of record for West Point is much longer. Specifically, the average annual precipitation at West Point for 105 years is 44.59 inches while that for Stewart Field is 43.29 inches for 13 years of record. The monthly averages of precipitation are also comparable. Figure 3 shows the average number of days per month with precipitation, both for the 25 year period and in the 4 individual years during which radiosonde data at Stewart Field were available. On the average, the Fall months of September and October have the least number of rainy days, 7.5 and 7.9 respectively, while May has 11.0 days, the greatest number. The
- annual average number of rainy days per month is 9.6, or approximately one in three.
Tables V and VI show when precipitation occurred in the area during temperature inversions. In 1950, more than one-third (36 percent) of these occurrences came in the months of December and January; while 43 percent were during the Spring and early Summer. In 1951, 29 percent of the cases of inversions associated with precipitation came in December and January, while 3 5 percent came in the three Summer months. Rain or snow fell at times of marked inversion only on 6 days during 1950 and on 7 days in 1951. There seemed to be no favored time of occurrence of marked inversions and precipitation. Figure 4 sums up in graphical form the results given in Tables V and VI, showing the f requency of wind directions at times of precipitation during all inversions and marked inversions. The predominance of wouthwesterly winds during times of inversions and precipitation is unmistakable. It is actually more important to some analyses to know about the absence of rain. Figure 5 gives the probability of drought at Tuxedo Park, only about three miles from the proposed reactor site, it is based on the probability tnat there wil' be a specified number of days without 0.10-inch of rain in two days or less. Thus the curve marked IC[ shows the varying probability throughout the year that 10 consecutive days will pass without as much as 0.10-inch of precipitation. Both probability and mean recurrence Interval are shown. Probability 1.0 signifies recurrence without exception every year; probability 0.5 indicates that a designated drought occurs on the average once in two years; 0.1, that it occurs once in ten years. Drought probability varies from day to day through the year. For example, on January 1, the probability that the next ten days will be dry is 0.064. The figure rises to 0.110 on April 15, drops to 0.063 by May 20, increases to a high point of 0.158 on the 1st of October and reaches an annual low of 0.056 on November 20. By the same means, it can be concluded that a drought of some particular duration and beginning on a specific date will occur at an indicated probable interval. For example, a ten-day drought i beginning on November 20 will occur only once in 19 years, whereas one ' beginning on October 1 will occur every 7 years. l l
3 LOCAL OBSERVATIONS Ever since the initial startup of the reactor in 1961,, hourly observa-tions of wind speed and direction have been made and recorded by theThe wind reactor operations staff during all periods of reactor operation. vane and anemometer for these observations is mounted on a towerDirections adjacent to and at the same elevation as the exit of the reactor stack. are in 22.50 sectors. A sampling of these observations taken over a 12-month period from June 1967 has been analyzed and the resuits compared with those from Stewart Air Base which formed the basis of the original pre-operational safety analysis. These results, transformed to 45 sectors for comparison with those in the preceding section, are given in Table Vil and Fig. 6. TABLE Vil WINDS AT REACTOR STACK ELEVATION (1967-68) Wind Wind Direction (Degrees; North =0) Annual Speed Calm Total 0 45 90 135 180 225 270 315 (mph) 18.7 18.7 Calm 1.5 33 6.3 7.5 4.6 36.8 1-3 5.7 5.6 2.3 2.5 3.0 15.5 8.4 1.2 34.8 4-12 0.8 2.5 0.9 0.6 0.8 4.7 2.0 0.1 8.4 13-24 0.1 0.1 0.4 0.2 0.5 0.2 1.3
> 24 27.0 18.0 5.9 18.7 100.
TOTAL 6.5 8.2 33 5.0 7.4 The cumulative frequencies for each direction agree well with those In each sector, however, the speeds from Stewart Air Base (see Table 1). tend towards the lower ranges. These lower speeds are used in current safety analyses, e.g., those for localities close to the reactor site as given in Appendix 2, Sect. C4, which therefore are more conservative than the original analyses in this respect. NUMBER OF INVERSIONS BY MONTHS ALL INVERSIONS 20 1950 19 51 / 16 - / JFMAMJJ ASONOJ FMAMJ JASONO MARKED INVERSIONS 8 -1950 1951 4 g 2 21 A 1 I (((([M i A M JFMAMJJ ASON DJ FMAM JJ A SONO Figure 1 l
SURFACE kVINOS TEN YEAR PERIOD, 1943-1952 12.0- 28.cf 12.0 ' 9.o (OV8VM January Februo Mor l 0 e.s in.i iz.s
\* ' \* \* @*% WW s Ju y Au e Sept 38.V n W m V F so c go g a
r December / n n ,, Percent' Frequency p Miles Per Hour l Figure 2 l
NUMBER OF DAYS WITH PRECIPITATION 1949 1950 16 - l 14 - 12 / . " 10 __ 6 25 year overage 6
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s APPENDIX 7 Dis ussion of Operating Lifetime The UCNR has been operational for a period that closely approaches 20 years. It is pertinent to inquire into the possible deterioration of systems that may affect the safety of operation for a like period in the future. Fortunately, the relative simplicity of this pool-type design, i
; where most of the systems and components are accessible or replaceable, ,
is favorable to continued safe operation. In fact, during the past t service life numerous components have been replaced, as surveillance
! checks and routine maintenance have revealed failures or poorer performance.
Nevertheless, a survey of systems and components that are directly or remotely safety-related has been made. The conclusion has been reached that the operational safety and reliability of the Safety Systems of this i reactor are substantially better than when the reactor was new, and that there is no reason why this condition should not continue for another 20-year period. This follows from the surveillance reouirements expressed in the Technical Specifications and from the results of this survey. Details of the latter are given below.
- 1. FUEL:
. The integrity of the fuel cladding is assured through the quality
- assurance required of the fuel manuf acturer, the specifications imposed on primary coolant water cuality, the limit on burnup, and the limited time that irradiated fuel spends a~t the site. Spent fuel is shipped every 2-3 years, with the oldest fuel usually having spent 5 years or less on site.
- 2. REACTIVITY CONTROL SYSTEM:
2.1 Control rods: the " poison" content of control rods is monitored through yearly measurements of reactivity worth. Rods are readily replaceable. Five new rods of the Ag/In/Cd type were installed in 1975-77. Unlike the old style boron carbide rods, no swelling problems have occurred with these rods. 2.2 Control rod drives: the electromagnets are assembled on-site and are easily replaceable. .The release and scram times are measured frequently and would reveal-abnormalities. Magnet failure is in the safe direction. All relays in the rod control system were replaced with new hermetically-sealed units in 1974. The fine rod position indicator was rr.placed with an electronic digital readout in 1974.
- 3. SAFETY CHANNELS:
3.1 Safety Amplifiers: Routine channel checks, tests, and calibrations monitor the performance of these. Both enassis (4 amplifiers) were replaced with rebuilt units using new power supplies for the magnets and amplifiers in 1976 and 1978. Spare amplifiers are in stock and a brano new chassis is scheduled for installation this year. Components are readily available. 3.2 Uncompensated Ion Chambers (3): Three new cnambers with integral mineral insulated cables were installed in the 1976-1978 period. Since installation, none of the cable problems that used to be experienced have occurred.
- 4. NEUTRON MEASURING CHANNELS:
4.1 General
Channel checks, tests, and calibrations required by Tecnnical Specifications continually monitor the performance of these systems and warn of incipient problems. In all power supplies, vacuum-tube type rectifiers have been replaced by solid-state devices. All the original chart recorders were replaced by brand new ones in 1974, except 'or the Linear-N, which was replacec in 1978. 4.2 Log Count-Rate: A new fission chamber was installed in 1979; two spare chambers are available. The cable between chamber and preamplifier was replaced with a double-shielded very-low-capacitance type, with mucn improvement in performance.
4.3 Log-N
Two units (one is a spare) were rebuilt with new components in 1977 and 1979. The two recorders (Log-N and Period) were replaced in 1974. A new compensated icn chamber with low-noise cables was installed in 1978.
4.4 Linear-N
The old electrometer was replaced with a modern, much more stable, picoammeter (Keithley Instrument Co.) in 1978. At the same time, a new chart recorder witn transmitting helipot was also installed. A new compensated ion chamber with low-noise cables was installed in 1978.
- 5. POWER AND THERMAL MEASURING SYSTEMS:
5.1 Delta-T System: A new system was installed in 1978, followed by a duplicate _ in 1979. The system includes RTD probes, differential
-recorder, and digital readout indicator, and is capable of improved accuracy over the original system. The system was fully calibrated and operational in 1979.
5.2 Primary Flow-Rate: A new differential pressure transmitter with electrical output and digital indicator was installed in 1979, and was fully operational in March 1980.
- 5.3 Temperature Indication
- The original switched system for indicating temperatures at various points in the primary and secondary
- cooling systems was replaced by an all solid-state digital system with resistance-thermometer probes in November 1979.
- 6. PRIMARY COOLING SYSTEM:
6.1 Pools
the original design, which included within the concrete pour a 3/16-in. steel liner that is welded to all penetrations, guarantees against leakage even if small cracks develop in the concrete. There is no evidence that this liner is not still performing its intended function. The three exposed outer walls of the pool are in plain view and any leaks therein would be evident. The pool-gutter drain lines, which were embedded in the walls, were replaced in 1977 with lines running outside to the walls to allow inspection for leaks from this source, and easy replacement. 1 6.2 Pool Gates: The original rubber seal on the canal gate, which j performs a seconcary containment function, was replaced by a new inflatable seal in 1968. The seal on the pool gate is due to be replaced in 1980 with an inflatable type. 6.3 Primary Piping: The four 10-inch lines to and from the pools eacn consist of two sections. The embedded section is stainless steel, to prevent corrosion of lines that would be difficult to replace, and an aluminum section. The letter is installed in a sand-filled trench below
- the floor and would be accessible for replacement if necessary. This j section was ins acted in 1976 and no sign of corrosion or leakage noted.
(These results were reported to Region I in September 1976.) 6.4 Holdup Tank: A yearly pressure test is made in the air vent line from the holdup tank to verify that the line is leaktight. To check that there is no leakage of water from the tank, twice-yearly checks of radioactivity are made in the french well located in the pump room. ' tis well would receive drainage from any such leakage through the floor or walls of the holdup tank. 6.5 Heat Exchanger: The two tube bundles in the neat exchanger are stainless steel to obviate corrosion in the primary cooling loop. In 1975, one of tnese was replaced with a new SS bundle due to a vibration problem. Since tnen, there have been no vibration problems with either i bundle. A spare bundle is also available. The heat-exchanging performance is monitored and the bundles are cleaned as needed. No metal corrosion has been found in the stainless steel tubes and sheets. Tube leakage is continuously monitored with a differential resistivity me,asuring system.
6.6 Core Outlet: This comprises the plenum, flapper valve, metal Dellows, and spool piece. There has been no evidence of problems or oeterioration of these components. The occurrence of plenum leakage is monitored with the plenum leak detector during reactor operation. Outlet line pressure, which would be affected by any leakage in the above components, is recorded routinely and has remained constant.
- 7. CONTAINMENT SYSTEM:
7.1 Personnel Airlocks: The outer door of each pair of doors in the two swing airlocks was replaced in 1974 with a new door having an inflatable seal. 7.2 Building Air Valves: Routine containment tests, required by Tecnnical Specifications, assure satisf actcry operation and integrity of these valves and detect any deterioration in building leak-tightness. Components of the valves are replaceable. Valve seats were inspected in 1973 anc. found in excellent condition. A manually-operated backup system for closure of these valves was added in 1974. 7.' Emergency Generator: Startup tests are done weekly, load tests semi-arnually, and complete servicing annually. Total usage to date has been minimal. 7.4 Emergency Ventilation: Routine containment tests are performed to assure continued operational availability. Flow rate is verified annually. The efficiency of the charcoal and HEPA filters is measured annually, with no evidence of deterioration to date.
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