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Forwards Response to Questions Contained in 780922 NRC Ltr Re Core Inlet Helium Temp Vs Time Transients Generated by Tap Code
	ML20062B537
Person / Time
Site:	Fort Saint Vrain
Issue date:	10/25/1978
From:	Justin Fuller; PUBLIC SERVICE CO. OF COLORADO
To:	Gammill W; Office of Nuclear Reactor Regulation
References
	P-78178, NUDOCS 7810310085
	Download: ML20062B537 (82)
	v • d • e

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cotonAoo somoi 2 October 25, 1978 J K. FULLER Fort St. Vrain i v.c. m. 1 Unit No.1 P-78178 i Mr. William P. Gamill N. Director for Standardization and Advan'ced Reactors Division of Project Management 4 U.S. Nuclear Regulatory Commission Washington, D.C. 20555 Docke #50-267 t

Subject:

Fort St. Vrain, Accident Re-analysis, Request for Additional Information Ref.: NRC letter dated September 22, 1978 T. Speis to J. K. Fuller Gentlemen: Attached please find our response to the questions forwarded to PSC in 'tr.:! referenced correspondence. r If you need further information, please let us know. t Very truly yours, I J. K Fuller, Vice President 1 Engineering and Planning JKF/FES:ers Attachment s 781031 00%5 f // Yg\\

-) +... Response to NRC Question 222.007 Question: Your response to question 222.002 was incomplete. 1. You state that core inlet helium temperature vs. time tran-sients generated by the TAP code were verified by hand cal-culations. Provide a description and general methodology for performing these hand calculations. Provide overlay comparisons of helium temperature vs. time calculated by TAP and by hand. 2. You state that' primary system pressure vs. time as calculated by the RATSAM code, has been independently verified by engi-neering calculations. Provide a description and general (, methodology for performing these engineering calculations. Discuss how you account for helium energy redistribution to the PC.tV lower cavity liner and internals during the LOFC. ' Provide overlay comparisons of primary system pressure vs. time calculated by RATSAM and by engineering calculations. Answer: Part 1: -Hand calculations were performed at three selected points in time to independently verify the core inlet helium' temperature vs. time curves utilized in the DSDA and firewater cooldown accident analyses. The general approach was to calculate a counterflow heat exchanger effectiveness for the steam generator based upon the installed surface area; helium flow-rate, inlet temperature, and properties (from RECA); and water side flowrate, inlet temperature, and properties (from performance of available components .under accident cooldown conditions). This method is outlined in Chapter 11 Section 4 of Principles of Heat Transfer by. Frank Kreith, 3rd edition,1973, ( and other heat transfer texts. 4 Figure 1 presents a comparison of core helium inlet temperature , ' values generated by TAP and by hand calculations for the firewater.cooldown. Results from the hand calculations are shown for assumed, firewater tempera-tures of 120*F and 80'F at the steam generator inlet. The initial transient in the TAP curve represents the transition from powered operation core inlet heliua conditions to the accident condition considering the primary system capacity. Throughout the remainder of the transient, the agreement is excel- . lent and the TAP input to RECA is independently confirmed. Figure 2 presents a similar comparison for the DBDA accident case. In this case, hand' calculations indicate that the core inlet helium tempera-ture is conservatively high. Part 2: Primary system helium pressure is governed by the local helium temperature in the various major volumes of the PCRV. The top plenum above the core contains approximately 15% of the helium, the bottom plenum. below the core contains about 5%, the steam generators contain 15%, the lower cavity contains about 50% and the core outer. annulus about 15%. Clearly, the helium temperature in the lower cavity has a major impact upon primary system e r gf . pressure. n a. ..a p ge:+ yy: . cyp.<.%;S.,Q &.iq.cih AiQkV.hlg'd Qt.- w ;&52.C.C..-,--p yf. $;,9.. 'e-G.p.pgjj:lg g. A- .s c.se v::ummu - ! o

w a. 5,* '. [ { ' e Under loss of Forced Circulation (LO'FC) conditions, the lower cavity temperature drifts downward from the operating value is the heat is transferred from the helium to the ' colder masses includinE the liner and concrete. In addition, hot gas from the top plenum above the core is drawn down the annulus to the lower cavity as the helium in this cavity is cooled. Because the RATSAM code includes thermal models for these processes, this code was used to generate the PCRV pressure vs. ti ce behavior under LOFC con-ditions. Results were utilized as an input to the RECA analyses. These same thermal processes modele.d in RATSAM can be represented using simplified engineering calculations. As an independent verification of the RATSAM LOFC pressure vs. time results, a thermal model of the PCRV lower cavity was formulated and solved as a function of time. This lower cavity model included heat transfer from the helium gas to the PCRV lower sidewall, bottom of the PCRV, core support floor columns, bottom of the core support floor, and special internals such as steam generator handling rails. Solid noce heat capacities for these masses were estimated including asso-ciated thermal barrier cover plates and kaowool insulatien. Heat transfer to each mass was calculated including both thermal conductivity and the helium-to-soM d heat transfer film coefficient. (Note that since the total thermal resistance is basically a function of the kaowool properties, the film coef-ficients are relatively insignificant.) Heat addition to the gas in the lower cavity due to temperature induced backflow from the top plenum above the core was approximated by a simplified model. Primary system pressure was calculated using a five factor formula accounting for the temperature in each of th'e major volumes of the PCRV. Lower cavity temperature was taken from the simplified thermal model described above. Core top and bottom plenum temperatures were taken from RECA results. Steam generator temperature was held constant while annulus temperature was conservatively set equal to the top plenum t.emperature. A short digital computer program was employed to carry out these computations. The model was initialized to steady-state operating tempera-tures existing prior to the LOFC. Temperatures and pressure were recalcu-lated 'at every second over the two hour time interval using a simple, first order, finite difference approximation for integration. - Results of this pressure calculation were compared to the RATSAM = pressure transient for the LOFC in Figure 3. RATSAM code results show a more rapid pressure decay early in the LOFC due to helium temperature decay in the steam generator and annulus volumes and an eventual leveling of the pressure trend as natural convection and contraction around the primary system carries hot gas into the lower' cavity. The simplified engineering calculations based on estimat'ed temperature changes confirm the general behavior of the RATSAM pressure transient. ~

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.'[. Response to NRC_ Question 222.008 Question: In your response to item 2 of Q222.004, you stated that " current plant data analysis indicates that a fully closed valve loss coefficient is approximately one half of the FSAR value.. " Provide this data in the form of observed loss coefficients versus orifice valve positions. Briefly discuss how the data was obtained (measurement locations, etc....) Answer: Figure 1 presents th'e Fort St. Vrain flow control (orifice) valve Effective loss characteristics as measured during prototype component testing (FSAR values) and as recently derived from plant operating performance. The FSAR values were derived frem a full scale component test per-fo'rmed in ambient air. The orifice valve was mated to a simulated region plenum and array of coolant channel boles with a blower located downstream. Static pressure was measured in the channel and total pressure was calculated using known gas conditions. Barometric pressure was used as the inlet total pressure. From these measurements, the total irrecoverabio pressure loss from the plenum above the orifice to the coolant holes expressed as coolant hole velocity head was calculated. The curve presented he been corrected i by +0.5 in loss coefficient to account for a region plenum redesign incor-porated after the tests were run. The curve derived from recent FSV plant data was developed using measured values for orifice position and core region outlet temperature. At the full open position, the FSAR valve loss coefficient was taken as correct and a region flow and region peaking factor (RPF) were derived from this data. As the valve was closed, changes in RPF with tempera.ture were accounted for analytically and the orifice valve effective loss characteristics were calcu-lated. The change in the orifice valve loss coefficient from the full open position to the fully closed position was confirmed by the fact that the dif-ference between calculated and measured

region power peaking factors is not a function of orifice position. Setting the valve loss coefficient to the FSAR value at the full open position has been confirmed by the fact that calculated core pressure drops are close to measured values.

Comparison of the FSAR and recently derived curves at the fully closed valve position indicates that the maximum effective loss coefficient from plant data analysis is approximately one half the FSAR value. It was this observation which led to the conclusions presented in the answer to Item 2 of. Question 222.004. ~ l ~

Neasured region powsr peaking factors are in fact calculated from measured values of core inlet temperature, region out.et temperatures, core flow rate and region orifice flow coefficients (energy balance). The calculated region power peaking factors are obtained from core neutron dif-fusion calculations.

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A ~ e Response to NRC Question 222.009 Question: Temperature vs. time comparisons between data and RECA predictions presented.a response to Q222.001 indicate that the code fairly consistently underpredicted temperatures in regions 20 and 32 through 37. What is attributed to this observation? Address any possible connection (i.e., bypass flow and region gaps) between the underpredictions and the oscillations, since both appear to be characteristic of the same region. Answer: The cause of the RECA code underprediction of region outlet tempera-tures in regions 20 and 32 through 37 is not presently undertood. This anomalous behavior during scram transients may be indicative of some phenomena which are also associated'with fluctuations. Such a possibility is being pursued and results will be reported as a part of the fluctuation effort. ( e e m 4 yOg e.g e-e w ee - me-at e m

Response to NRC Question 222.010 ~ Question: The staff has reviewed your response to Q222.006 regarding assump-tions on operator actions for the firewater cooldown accident. It appears this response is inconsistent with statements made in Section 10.3.9 of the FSAR, in that a main loop rundown cannot be assumed for this event due to postulated seismically-induced failure of piping and equipment downstream of the main steam bypass valves. If it can be shown that main loop rundown can be accomplished using seismically qualified equipment, compensacion would exist for the 11/2 hour delay that might accrue in initiation of firewater cool-down in accordance With the procedural sequence outlined in the Manual Class I/ Class II Valve Study. If this cannot be shown, you should provide an analysis of the firewater cooldown event assuming no circulator rundown is available due to seismically-induced failures in necessary equipment and accounting for the maximum delay in initiation of firewater cooldown that can be attributed to carrying out the procedural steps in the " Manual Valve Study". Show temperature versus time traces for fuel temp-erature, helium outlet temperature and thermal barrier cover plates. Secticn 10.3.9 of the FSAR was a postulated worst case event to [ Answer: demonstrate that the firewater cooldown can be accomplished with once through water usage due to seismically-induced downstream pipe failures. Analyses of the.cooldown on one firewater driven circulator submitted to the NRC have followed the previously accepted sequence of events described in Section 14.4.2.1 of the FSAR. A very conservative scenario taking no credit for a main loop rundown following scram and imposing the maximum 1 1/2 hour delay from the " Manual Valve Study" (see Answer to Question 222.006) represents a . ( more severe case than previously analyzed under Section 14.4.2.1. In addi-tion, it 'should be noted that the cooldown on the liner cooling system follow-ing an LOFC envelopes the cooldown on one firewater driven circulator as an available alternate scheme. It cannot be shown that a main loop rundown ca'n be accomplished using only equipment seismically qualified for the Safe Shutdown Earthquake (horizontal ground acceleration of 0.10g). However, considering that all structures in the plant are designed to support the loads due to the Operat-ing Basis Earthquake (horizontal ground acceleration of 0.059) and the i l redundancy of equiptrent and power sources for forced circulation plus redun-dancy in.the' ultimate heat sinks, it is r et unreasonable to assume a main loop rundown on one circulator'for fifteen minutes following a seismic event. This was the basis for the response to Q222.006. In com-pliance with your current request, the accident was analyzed as outlined in the question. ~ The RECA code was utilized to analyze the accident with a reactor scram following a 1-1/2 hour LOFC. At 1 1/2 hours after scram. forced cir-culation is restored using one firewater driven circulator and cooldown is assumed to proceed as in' the firewater cooldo'wn accident. Results are pre-sented as temperature versus time traces for both the conservative equilibrium . 49.-y. -v.;M;. t <--w- +, r + . ~ .n., w ~ +:.e ~ ~ ~ ~, . w... x...._- .1 m_

', i.. i .i l { core conditions (equilibrium core with pelton wheel nozzle differential pres-sure = 175 psid, i.e., boost pumps installed) and the conservative initial l core conditions (initial core with pelton wheel nozzle differential pressure = 103 psid, i.e., currently available). Figure 1 presents maximum fuel temperature vs. time for the con-l servative equilibrium core conditions. Fuel temperature increases during the LOFC period. Following restoration of forced circulation, the curve drops and becomes somewhat irregular as the maximum fuel temperature shifts from l node to node within the core model. Maximum fuel temperature attains a second peak at approximately 11 hours into the transient when the flow in some regions l t 1s reversed. Figure 2 presents average and maximum core outlet helium gas temp-f eratures as functions of time. These curves are relatively flat during the f LOFC portion of the transient-as little hot gas circulates to the core outlet i plenum. Upon restoration of forced circulation, core outlet temperatures rise rapidly to peak values and then decrease as most of the core cools down. i Flow in the remaining regions reverses and hence doe: not contribute to core-i outlet conditions. Maximum thermal barrier temperature at the steam generator i duct inlet was not plotted as a time function because this value has been t conservatively calculated using peak values for both the maximum and average core outlet gas temperaturc even though these peaks occur at different times. The calculated maximum value of 1702*F would be lower ~if computed from values taken at the same instant in time. The safety limit for the steam generator inle.t duct thermal barrier is 2000*F. Figure 3 presents temperatures associated with the core inlet plenum as functions of time. During the LOFC, these temperatures increase as natural convection redistributes energy into the upper plenum. The average temperature of the top head thermal barrier cover plates remains below the average upper 4 ( plenum gas temperature due to conduction to the cold liner and concrete. Local hot spots would be attenuated by natural convection within the upper plenum and by reradiation to cooler surfaces. Thus the maximum thermal bhrrier coverplate l temperature will remain below the 1500 F,1imit. ~ Upon restoration of forced circulation, these' temperatures decrease i rapidly as cold gas floods the upper plenum above the core. As the transient progresses, some regions of the core reverse and the maximum temperature of gas ' exiting from the top of the core climbs to a second peak. The themal barrier remains much cooler, however, since the average gas temperature is low and [ most of the core inlet plenum cavity surface acts as a radiation sink. Note that the av6 rage top head thermal barrier cover plate temperature is slightly hotter than the average gas 'emperature and liner temperature due to radiant } heat inputi from hot core surfaces. Figures 4, 5, and 6 present similar information under conservative t l. initial core conditions and 103 psid pelton wheel nozzle differential pres-sure. In this case the calculated thermal barrier temperature at the steam generator inlet duct is 1607'F. Again, the maximum top heat thermal barrier cover plate will remain below the 1500'F limit throughout the transient. l I ) f O- ~. * -y\\'<c hw, ? ~ 1..:,q,. ...n % o- ~ ~.

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i L PUBLIC SERVICE COMPANY OF COLORADO d P. O. BOX 840 DENVER. COLORADO 80208 i October 26, 1978 1 K. FULLER Fort St. Vrain v.cc..cs or,a Unit No. 1 P-78177 Mr. William P. Gammill Asst. Director for Standardization and Advanced Reactors Division of Project Management U.S. Nuclear Regulatory Commission Washington, D.C. 20555 Docket #50.-267

Subject:

Graphite Surveillance Program Ref.: NRC Correspondence June 8, 1978 W.P. Gammill to J. K. Fuller Gentlemen: In the referenced correspondence the NRC enumerated ten questions and included additional inquiries in the latter part of the " Topics to be Addressed" paragraph. A total of fifteen questions have been identified and are responded to in the attached enclosure. If there are any questions, please let us know. Very truly yours, J. K. Fuller, Vice President Engineering and Planning JKF:ers Enclosure = f m <e .. * * = = - -.. -,...... =w+ -= -- e

.r....... ~ ENCLOSURE 1 RESPONSE TO NRC QUESTI0f1 #1 i QUESTIO!!: Include discussion of means and schedule to include mechanical testing of + Surveillance Specimens. l ANSWER: Mechanical tests of in-reactor PGX surveillance specimens will not be specified in the surveillance test plan. Laboratory bench tests of mechanical properties of oxidized PGX, howevar, will be performed during the course of research on PGX oxidation effects funded by DOE. The laboratory tests will define the me-chanical beha' ior' of PGX graphite in the presence of uniform oxidation, surface v f oxidation, and oxidation gradients. - i The annular in-reactor surveillance specimen configuration does not permit the use of qualified standard mechanical testing methods which form the basis for - the data bank with a known statistical distribution for PGX graphite. The lack of a qualified test method with acceptance criteria, therefore, preclude mechanical testing of the,in-reactor surveillance specimens. Further, we do not believe mechanical' testing is required for the surveillance program since only credit for material with less than 0.1 wt % burnoff will be taken in the structural analysis. There is no reason to suspect the structural adequacy of virgin material. l e e a P i L = J l M*' -. ~. -.. )

,;.l...,-.. i RESPONSE TO NRC QUESTION #2 r QUESTION: If mechanical testing is not feasible, the means for relating the changes to be measured to the' structural integrity.of the core support structure should be fully described. ANSWER: I For the conditions of temperature, pressure,' moisture level and flow rate in the reactor, the graphite has been predicted to oxidize only at the wetted surfaces with little or no in-depth oxidation occurring. Out of core experi-ments have produced oxidation profiles (percent burnoff versus distance from the edge) that are quite steep, thus,. supporting the pred~iction. Therefore, it is believed with a reasonable degree of confidence that oxidation in the core support merely removes material from the wetted surfaces leaving an inner core of virgin material with unchanged mechanical properties. Addition ' ally, it is anticipated that the in-service surveillance (ISS) specimens will oxidize on their wetted surfaces in the same manner as the core support. The oxidation profiles will, be determined for the ISS specimens. Structural integ2 rity of the core support will be evaluated from these ISS oxidation profiles. It has been clearly established in laboratory experiments that no significant decrease in strength occurs for burnoff less than 0.1 wt. %. s O "D ~ 6 a e <= m o

m <. RESP 0tiSE TO flRC OUESTI0tl #3 QUESTI0ft: Include detailed test procedures to assure a known and relatable exposure his-tory. ~ ANSWER: The surveillance program monitoring will use. existing plant instrumentation as described below to provide a known and relateable exposure history. Temoerature Excosure Each fuel region has a permanently in' stalled thermocouple that continuously measures the region outlet gas temperature. The actual outlet temperature of regions that contain surveillance specimens will be used to calculate the temperature at the surveillance specimen. The surveillance specimens will be located in the hottest bottom reflector transition block in a refueling region. This reflector block re.sts. on the core support block as shown in Figure 1. The measured region outlet temperature is approximately 3o inches downstream of the surveillance specimens. Coolant flows past the surveillance specimen through the bottom reflector transition block, and through the core support i block coolant channel. The composite temperature of the six support block cool-ant channels is measured by the region exit thermocouples. The t$mperature at the surveillance specimen locations will be. calculated by utilizing the measured region powers, the calculated column power peaking factors, region flows, and the region outlet temperature. Gas mixing between the specimen and region outlet temperature sensor must also be considered since the specimen will experience a somewhat higher temperature than the sensor. Moisture Excosure The primary coolant moisture level is measured by 2 Dew Point Hygrometers that are connected to sampling lines which sample the coolant in the helium circu-lator exit plenums. , The dew point hygrometer's sampling heads are operated at a primary coolant ~ pressure of 200 psig. O P .em- = i.

a e7 ...e__.____. ..__.e _4 Their electrical output, the dew point temperature of the condensing / frosting mirror, is transmitted to dew point temperature indicators in the control room as shown in Figure 2. The dew point hygrometers are adequate to meet Tech Spec requirements for dew point measurements to -75 F ($1 ppm). Investigations are 1 in progress to improve lower limit measurement to -90 F (0.3 ppm) which is the expected moisture level at 100% reactor power. Data Acquisition i The following plant variables will be recorded on an approximate four hour interval during steady state operation. Region Outlet Temperatures Helium Circulator Inlet Temperatures ( Orifice Positions E Rod Positions Moisture Level Circulator AP (core flow) Primary Coolant Pressure Figure 3 shows a general concept of the data acquisition method. Other Oxidants Measurements of CO, CO, and H2 in the primary coolant are made daily utilizing l 2 a gas chromatograph. These measurements are made manually, and the data is recorded and stored. The C02 contribution to the burnoff calculations will be accounted for by adding 10% to the total burnoff. This 10% addition is based on a conservative estimate-( of the CO2 effect. The effect that low hydrogen to water ratio has on burnoff will be evaluated to determine if the reaction is inhibited by maintaining the iron in the oxidized non-catalytic state. If it is found that the low operating concentration of hydrogen does inhibit the oxidation reaction, it will be included in the calculations. DOE funded experiments are being carried out at GAC to quantify the effect of hydrogen to water ratio on PGX reaction rate. j t. r 5 e i l = e-e-e -== ., - eca a p.

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=._. _._. 4: N, f RESPONSE TO NRC QUESTION #4 QUESTION: Discuss how the PGX specimens tested relate to the actual PGX used for the core support blocks. ~ ANSWER: The PGX surveillan e specimens to be tested in reactor will be taken from two full size (1.14 m diameter by 1.83 m long) PGX logs. These two logs, designated 6484112 and 6484-138 by' GA, have been characterized for mechanical properties and chemical impurities. The ash contents were 0.7 wt % and 0.3 wt % and iron impu.-ity levels were 0.32 wt % and 0.12 wt % for logs 6484-112 and 6484-138, respectively. Data provided GA by Unioit Carbide Company (UCC) show no syste:aatic change'in ~ ash or iron content of PGX graphite manufactured between June 1972 and November 1977. The data, compared in Table 1, indicate no systematic change in ash or iron content of PGX graphite between 1972 and 1977. Data is not available for ash and iron content of PGX graphite production ' prior to June 1972, when the PGX graphite for FSV was manufactured. However, UCC personnel indicate the PGX in Fort St. Vrain should be about the same as the PGX tested and reported above. The ash and iron contents of log 6484-112 lie near the high portion of the band of impurity levels in the 1972-77 PGX production logs, while log 6484-138 con-tained less than the mean ash and iron impurities. Inclusion of specimens from both PGX logs will provide an opportunity to assess the in-service condition of g representative PGX logs from the production population. The ash and iron contents of logs 6484-112 and 6484-138 have been compared statistically to the ash and iron content of the 1972-1977 production PGX graphite population using the data of Table 1. The resulting values for the deviation from the population mean divided by the standard deviation were: 1 x-x Log s Ash Iron c 6484-112 1.2 1.2 6484-138 0.7 0.6 4 These values are not significant at any reasonable confidence level, therefore, there is no basis to believe that the ash and iron contents of logs 6484-112 and 138 did not come from the same populations of ash and iron contents as the production PGX graphite. 1

I _9 1 Mechanical property data available for the PGX logs inserted as components in the Fort St. Vrain reactor show the mean ultimate tensile strength (UTS) to be i l 1314 psi, similar to the 1308 psi mean UTS of the previously published UCC data (Ref.1) (see Table 1). UCC data show a mean UTS of 1437 psi, with no systematic change in UTS from June 1972 to Novmeber 1977. The mean axial UTS at the end of logs 6484-112 and 6484-138 were, respectively~,1450 psi and 1380 psi. The standard I deviations in all test series were in the range 150 to 200 psi. Therefore, all PGX UTS data available to GA indicate no systematic change in UTS from the Fort i St. Vrain cor:ponent material through PGX produced in november 1977. Based upon the purity and ultimate tensile strength data discussed above, it is ~ i concluded that the properties.of PGX graphite have not changed significantly g since the manufacture of Fort St. Vrain components. herefore, the surva test specimens are considered to be reDresentative o PGX material in F3 respect to UTS and impurity levels and, therefore, in oxidation characteristics. / TABLE 1. PROPERTIES OF PGX GRAPHITE .e Property Ash Iron Axial UTS Data Source (wt%) (wt%) (Psi) 1 31 ' Mean Fort St. Vrain Component Acceptance QA Tests (28 logs) 202 Std. Dev. UCC Production Data for 1.14 m 0.51 0.2E 1308 Mean ( diameter PGX (Ref.1)' (5 logs) 0.07 0.10 153 Std..Dev. UCC Data for June 1972 to 0.45 0.19 1437 Mean { Nov.1977 (26 logs) 0.22 0.10 152 Std. Dev. Log 6484-112 (Ref. 2) 0.72 0.31 1450 Mean 174 Std. Dev. Log 6484-138 (Ref. 2) 0.29 0.13 1380 Mean ~ 174 Std. Dev. 1 Ref. 1. "HTGR Fuels and Core Development Program, Quarterly Progress Report for the Period Ending August 31, 1977", ERDA Report GA-A14479, Sept.1977, pp.11-128. Engle, G.D. and L.At Beavan, " Properties of Unirradiated Graphites Ref. 2. PGX, HLM and 2020 for Support and Permanent Side Reflector LHTGR Components". ERDA Report GA-A14646, June 1978, pp. 5-5. 'g' 4m. e =,m.. e. a e e .*****.e-- e - =e. ee e e. * = k

RESPONSE TO NRC QUESTI0ft #5 QUESTION: Discuss provisions for maintaining archive samples of the test specimens. ANSUER: Requirements for archive samples of the test specimens have been included in the surveillance test plan. Twenty (20) archive specimens from each of the' two PGX logs will be retained by PSC. These archive specimens represent adja-f cent material to that of the in-reactor surveillance specimens. 9 y O e o I, 9 9 e 6 ( 4 4 s g e G e 9 4 4 i e %g N*

z.- ~, L l b RESP 0tlSE TO flRC QUESTI0ft #6 i. QUESTI0ft: Include an analysis to demonstrate that all significant portions of the Test Plan, including the number of specimens, are in accordance with requirements to obtain statistical meaningful data. [ i ArtSWER: I The standard deviations observed in laboratory determination.of oxidation weight loss and oxidation depth of PGX graphite average approximately~ 40% of the mean values. Sixteen surveillance specimens will be tested at each time point in the surveillance program. For the sample size n = 16, the standard error of the nean is equal to the standard deviation divided by /5~, or 40%/ /IT = 10% of the mean. An accuracy of 10". on the mean is sufficient for the p,urposes of this surveillance program. b N b b I i J (4 p { l t j i k 4 L ew- ..oe-- a o . e. e s .e .,e.e ..e.,

I i 1 o L RESPONSE TO NRC OUESTION #7 t I } QUESTION: Discuss a specimen withdrawal schedule with contingency provisions similar to the requirements of 10CFR Part 50, Appendix H. I t I L ANSWER: i The in-reactor surveillance specimens will be installed at the first refueling I f - and will be withdrawn at stasequent refuelings as follows: l ) Sample Withdrawn at Approximate Reafon Refueling Number EFPY Exoosure l t 25 2 1 l 30 4 3 t 24 6 5 l 22 9 8 t 16 27 17 8 The above sche'dule would be adjusted to remove samples at a faster rate should i specimens at any withdrawal interval show a burnoff significantly greater than [ predicted. f I r { i t - ? i ? L \\ e e l ~ l 6 I e L e., -,..w..m e-- =+ 9==ee***

r 'W g.. p ~.= : - w i e RESPONSE TO NRC QUESTION 88 QUESTION: [ Discuss the advantages and disadvantages of pre-oxidation of test specimen surfaces. ANSWER: The two PGX graphite logs to be used for the surveillance specimens have been I extensively characterized. Additionally, material taken from a sample log directly adjacent to where the in-core specimens are taken will be well char- ' acterized for oxidation rate in out of pile tests. There is no need 'to per-form pre-oxidation on the surveillance specimens to determine their oxidation characteristics. Due to the large variabiiity in the initial oxidation rate of PGX, burnoff to 1 to 2 wt. would be required to reach a steady-state oxidation rate. Labor-atory pre-oxication of specimens to 2 wt. % burnoff would introduce a bias in oxidation mechanism and distribution which is not compatible with the intended program designed to determine oxidation of PGX a't reactor conditions. Therefore,* the surveillance test plan will not indlude pre-oxidation of specimens. t 8 e 6 i ) l

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RESPONSE TO NRC QUESTION #9 QUESTION: Discuss the similarities of oxidant transport characteristics between out-of-core tests and anticipated surveillance conditions. ANSWER: The primary difference between laboratory tests of PGX oxidation and in-core surveillance conditions is that the ficw regime in tests to date has been laminar while tn2 flow regime is turbulent in the reactor. The flow regime is important since it affects the transport rate of oxidant from the bulk stream to the PGX surface. This transport rate is reflected in the coefficient of mass transfer, whose magnitude can affect the oxidant concentration and consequently the burnoff in PGX during reactor operation. Current burnoff predictions include the. effect of the flow regime upon 'the mass transfer nefficient. Experimental and theore-tical work performed at the Dragon Projer '.Refs. 1 and 2) is considered adequate to provide a valid basis for the evalua+'. of the diffusional resistance repre-sented by the mass transfer coefficient

ng laminar flow conditions achiev-able in a icw-pressure laboratory test apparatus. In addition, the high pres-sure test apparatus scheduled for completion in late 1979, will be capable of simulating the conditions prevailing in the reactor during turbulent flow with lower turbulent flow Reynolds numbers but compensating flow diameters to obtain similar mass transfer of the oxidant to the graphite surface.

The processes and equations for oxidant transport to and into the graphite sur-face are illustrated in Figure 1. For the flow to the surface, the mass trans-fer coefficients are: RePrf (laminar regime) i k = 1,62 c o Re.e Pr." (turbulent regime) k= 0.023 Here, Re a h is the Reynolds number (>4000 for turbulent flow) Pr = he dia-is the diffusional Prandtl number or the Schmidt number (Pr s 1.4), d is t meter, I. is the axial length, 01,2 is the diffusion coefficient calculated using Eq.11-11.from Ref. 3 based on the Lennard-Jones potential function, u is the flow velocity, and u is the kinematic viscosity. The calculational procedure for both reactor and test conditions is to calculate the Reynolds and diffusional Prandtl numbers, knowing the helium flow rate pressure and temperature. Both above equations for mass transfer to the surface have been experimentally veri-fied as described in Ref. 1. For oxidant transport and reaction within the graphite, the mass transport coefficient k' is given by,

15-h k' s f = / D'Kv R hRT s

volumetric reaction rate constant K

= p y HO 2 L a / '/K

diffusion length D

y pg : graphite density R: universal gas constant T: abso. lute temperature PH2O : partial pressure of water in free stream .PH2 : partial pressure of hydrogen.in free stream 'Rs: reaction rate with graphite K P t H0 3 H 3 P s 1+K2 P + HO 2 2 K,K K3 and n are experimentally detemined. 1 2 \\ The ratio of the mass transfer coefficient across the boundary k to the mass transport coefficient in the graphite k' is defined as the Sherwood number Sh. Sh = k/k' The reaction rate decrease factor f is then, Sh ~ f Sh + 1 which is equal to the ratio of the actual graphite burnoff to the burnoff in - the abs'ence of a mass transfer resistance. Th'e above equatiohi'Mve been applied to several representative conditions which could be experienced by a single CSB channel at different times i j duri.ng full power operation. Two,rather extreme conditions were chosen: l i 1 I .e4

G r m a) a relatively high temperature / low flow condition (due to a high regio'n mismatch and/or high local power tilts), and b) a relatively low temperature /

high flow condition. Table 1 illustrates the detailed flow characteristics of the two conditions. It is found that the dominent effect is the pronounced change in graphite reactivity at high temperatures. Note in Table 1 the large increase in oxidant transport coefficient in the graphite compared to the ' smaller decrease in the coolant oxidant mass transfer coefficient calculated going from low to high temperatures. As a result, at high temperatures, the graphite is consumed rapidly at the surface. This causes a large oxidant concentration drop from bulk coolant to graphite surface. This drop must be properly accounted for in caldulating oxidation of the graphite. Results in Table 1 show that the oxidant drop could be 50% for some conditions at full power but only 7% at others. Similar calculations at 50% power indicate the ( range of oxidant drop is 0 to 20%. Below 25% power, the drop is negligible. Experiments have been performed at General Atomic under laboratory conditions at high and low flow and for different configurations, e.g., disks, cylinders, circular channels. All experiments were performed at atmospheric pressurts and ) in the laminar flow regime. The mass transfer correlation in the laminar flow regima, for stream parallel to flat plates, is used for the case of disks at low Repolds number (less than one). Results presented in Table 1 show that for the high flow rate tests (hollow cylinders with small flow diameters), the H O mass transfer to the graphite 2 surface is high relative to the transport into the graphite at 900 C. Only a 2% depletion is predicted at these conditions. Thus, the test design can be considered appropriate for obtaining chemical reaction rates free from mass transfer effects in the temperature region of interest. On the other hand, for low flow rate tests the configurations and Reynolds numbers are such that mass transfer of H O from the bulk coolant to the PGX i 2 j surface is the rate-controlling mechanism. Therefore, the low flow tests can simulate, to a greater extent than expected in the reactor, the H2O con-centration drop from bulk coolant to PGX surface even though the flow regime is laminar rather than turbulent. Future DOE funded, out-of-pile testing using the high pressure test rig being constructed at GA will be capable of Reynolds numbers in the fully established turbulent flow regime. By choice of flow diameter, the desired reactor mass transfer conditions can be simulated. References ' l. M. R..Everett, D. V. Kinsey and E. Roemberg, " Carbon Transport Studies for Helium Cooled High Temperature Reactors", Dragon Project Report DP-491, November 1966. 2. L. J. Valette, "The Influence of Mass Transfer on the Kinetics of Graphite Oxidation by Low Concentration of Oxidizing Impurities in an Inert Carrier Gas". Euratom Report EUR-2762e, 1967. 3. R. C. Reed and T. K. Sherwood, ~The Properties of Gases and Liquids, pub-lished by McGraw Hill Book Co., T966. l l

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( s .i m te 1 ^ APPLIhATION OF BOUNDARY LAYER EQUATIONS FOR PCX ~ ~ Reactor Conditions ull p wer) High flow Low flow Symbol {"*8 (bfig tests tests (b) o Ft0W CONDITIONS Cas temperature. T 975*C 800* C 900 *C 900 *C ~ turbulent 'ttirhulent laminar laminar P.cgime Reynolds number Re 111,000 295,000 263 0.17 v 54 111 664 0.027 Yelocity, ft/sec ft /sec D 2.1x10-4 1.7x10-4 0.~009 0.005 2 ~ biffusion coef ficient, Sciridt number PD 1.44 1.43 1.44 2.1 g), F ow dimension, ft d 0.63 0.63 5.2x10-3 8.2x10-2 !! ass transfer coefficient; ft/sec k 0.10 0.17 7.14 0.027 s CP.APilITE St:RFACE CONDITIONSft /see D' 1.9x10-6 1.5x10-6 8.07w10-5 8.07x10-5 2 Diffusion coefficient, -I K 4009 111 179 425 Volunctrie reaction rate, sec y 0.12 0.19 !! ass transport coefficient, ft/sec k 0.09' O.013 CCliPARISON OF HASS TRANSFER Sherwood number, k/k' Sh 1.11 13 60. 0.14 Reaction rate decrease factor (Sh/Sh+1) f 0.53 0.93 0.98 0.13 (a) flow dimension = diameter for hollow cylinder geometry, characteristic length for thin discs, [ in parallel flow (b) Flow condition due to difference in core helium orifice setting and column power tilts. a h e .-n. -e-. ,,,.,s, ,,.---.--,e--.

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1.~.V..- i .'. 1 RESPONSE T0 t:RC QUESTION #10 P QUESTION: Discuss the possibilities and practicabilities for including surveillance specimens of Type ATJ graphite in the specimen' holders. ANSWER: The ATJ specimens cannot be installed in the same fuel columns as the PGX samples. because there is not sufficient space in the five lower reflector transition blocks to install them without compromising the structural integ-rity of the blocks. Furthermore, the ATJ samples would have to be located differently within the region and fuel column than the PGX samples to reflect core support post maximum temperature. Post temperatures are con'siderably. lower than the core outist plenum maximum temperatures because of gas mixing in the core support block and lower core outlet plenum. It is also important to note that ATJ ISS specimens'are not needed since this graphite has a much lower steam / graphite reaction rate than Type PGX (see Mechanical.& Chemical' Property Changes of ATJ Graphite after Oxidation in Helium-Steam Mixtures, GA Report GA-014525, October 1977). Analysis, using the above experimental data and performed at the Tech Spec conditions, of the effects of oxidation on the core support posts and seats, shows that oxidation is minimal such that required safety factors are met at the end of reactor life (see Question / Answer #11b). f f l l l l

m 20-t i i RESPONSE TO NRC QUESTION #11a i QUESTION: { Discuss the justification for the program (e.g., experimental data on PGX). ANSWER: The steam / graphite reaction rate of Type pGX graphite is found to be consider-ably higher than any other graphite in tb FSV HTGR. Due to the high vari-ability in the out of core, experimenta41y obtained burnoff rates, the upper ' bound of the data is established at a maximum of 1000 times that of Type H-451 fuel block graphite at HTGR operating temperatures. When this reaction rate is used to assess the effect of oxidation on the HTGR core support floor blocks it is found that the useful, life could be as little as 12 years of full power reactor ^ operation at Tech Spec conditions. A prediction based upon expected (reactor operating) conditions at full power operation indicate that 58 years is the time required to reach useful life. Useful life is defined as the time it takes for the operating stresses to reach allowed limits. The plant lifetime design ob- ~ jective was 30 years. Thu.s predicted lives span a time frame shorter and longer-than the design life objective of the reactor. The spread in life predictions is caused by the uncertainty in predicting the operating history of region outlet temperature and helium coolant impurity levels. Uncertainties in temperature measurement accuracy are also considered. In addition, the upper bound of all oxidation rate data, as mentioned above, are used in the analysis,. An in-service surveillance program would provide ( oxidation of samples of core support block material experiencing the actual temperature / coolant impurity levels in the reactor. Periodic inspection of these samples would provide in pile oxidation data to verify the accuracy of the pre-dictions as well as provide necessary information for adjusting the predictions. The most important parameter effecting the lifetime prediction is the PGX steam / graphite reaction rate. The validity of the reaction rate data depends upon how well the experiments were performed to determine the rates. The following text summarizes the important aspects of the experiments performed to give confidence to the rate equation used in the core support oxidation study. Basis of PGX 0xidation Rate Introduction The oxidation of PGX graphi'e by steam is greatly enhanced by catalysis due to metallic impurities. The measured reacticn rates are in the range of 100 to 1000 times greater than that of more pure grades of graphite such as fuel block graphites H-451'or H-327. The concentration of impurities (i.e., l total ash) is around 0.5% (5000 ppm) (Ref.1) while the ash content of lower j oxidation rate pure grades is of the order of 100 ppm. Iron is known to be l l ~

O nu ..u.. k' c.. ~, the most catalytic cf the various impurity species, and concentrations in the range of 3000 ppm a"e not uncommon in PGX. Large variability in the reaction rates are observed from sample to sample and from log to log because of the heterogeneity of such properties as density, pore size distribution, impurity concentration and size distribution of the impurity particles. Because of the variability, only maximum or upper limit reaction rates are used in oxidation calculations. As an interim measure, for integral calculations, the upper bound of the reaction rate data of PGX graphite is assumed to be equal to 1000 times the rate of H-451 graphite, a grade whose oxidation characteristics and rate constants have been thoroughly investigated (Ref. 2,3). It will be shown that this assumed rate represents the approximate upper bound of the available data, and is, therefore, conservative. Work is continuing on PGX graphite oxidation, under a DOE contract, with the specific goal of determining a set of PGX reaction rate constants. ( Theoretical Basis of Graphite Oxidation The reaction of graphite with water vapor involves three basic steps: (1) diffusional transport of reactant across the boundary layer, (2) inpore diffusion of reactants and products to and from active sites within the porous media, and (3) chemical reaction. Three temperature zones have been identified where each of the steps predominate (Ref. 4). Zone I, a low-temperature zone, identifies a temperature range in which the overall reaction rate is controlled solely by the chemical reactivity of the graphite. In this zone, reactants diffuse throughout the graphite virtually undepleted, giving rise to uniform water concentration and burnoff throughout the sample. In Zone II (interme-diate temperatures) the rate i:; controlled by combination of inpore diffusion and chemical reaction. In this zone, the concentration of reactant and, there-fore burnoff, varies exponentially in depth giving rise to an " oxidation burn-off profile". In Zone III, at high temperature, mass transport across the gaseous boundary layer is the limiting step and oxidation occurs predominantely at the surface. Strictly surface oxidation (Zone III) occurs in tne HTGR for ( conditions of high temperature, high He pressure, low water concentration and high chemical reactivity of PGX graphite. Varying one or more of these four critical parameters as is normally done in accelerated laboratory testing can cause transition into Zone II where inpore diffusion occurs, causing oxidation at greater depth. The burnoff profiles for the three temperature zones are shown schematically in Figure 1. In Zone I, the homogenous or chemical regime, the overall reaction rate RT is equal to the ir.trinsic rate, Rs.* In Zone II, the inpore regime, the overall reaction is described by the follow-ing relation: RT = S/V Fj L Rs where RT = overall volume average reaction rate of the sample, s-l, S/V = geometric surface-to-volume ratio, 0.24 cm-l,

The intrinsic reaction rate, Rs, is the chemical reaction rate of steam and graphite as would be d'etennined by a well mixed reaction of guohite powder with most helium.

N, Sb "

i. ; '.=

1 22-I( ) j = efficiency factor, g) r/:. F 7 modified Bessel fur.ctions of the first kind (first and I,I = j 0 zero order, respectively), r, = radius of the sample, r = radial depth of interest. L = diffusion length, or L = / eff/Ky D e RT K = volumetric m ction rate constant Ky = f2P s y H20 where .P = Density of graphite T5 Absolute temperature R = Universal gas constant PH2O = Partial pressure of water in the helium' carrier gas D,ff = effective diffusion coefficient of oxidant in graphite given by D,ff = m Dl,2 D1,2 = gas, phase diffusion coefficient of water in the inert gas m = a geometry parameter related to porosity / tortuosity of the graphite pRT 2 L Rs

T2P 0.2 -

H20 1 (2) F where R, = chemical reaction rate in absence of mass transport given by the standard Langmuir-Hinshelwood equation: i 1000 k P j H0 2 R = 1+kP 75 + k P (3) 8 2 3 where the currently used values of k, k and k are those measured for z 7 H-451 graphite. The multiplier 1000 refTects t$e upper limit of measured PGX reactivity. r M e %g 7,.,,, 4,_ __pme% e ' ~ ' '

......~-.u..... +. y , [ i. In Zone III, at high temperatures, the overall reaction is equal to the rate of arrival of water molecules at the graphite surface which is given by: ,2 (Re Pr f) 1/3 (laminar flow regime) (4) k = 1.62 k = 0.023 Re Pr 0.4 (turbulent flow regime) (5) ,2 0.8 where k = mass transfer coefficient to the surface d = channel d'i& meter Re = Feynolds number Pr = Prandt1 number 1 = axial length It is important to' compare the value of k with that of k', the inpore diffusion coefficient, given by: . k' = / eff Ky (6) D Because PGX graphite is so' chemically reactive, a t temperatures of technical interest (700 to 1000 C), oxidation is characteristic of either Zone II or Zone III or something between the two zones. This has caused considerable experimental difficulty in attempting to measure the intrinsic chemical reac-tion rate Rs at temperatures above 700 C, using traditional low flow conditions because the measured value, RT, is always dominated by transport considerations. At low carrier gas flow conditions considerable depletion of reacta'nt occurs necessitating s at the transport consideration be ac' counted for to properly interpret the data. This has dictated an experimental approach where mass transfer is specifically enhanced so that the intrinsic reaction rate, Rs, is measured directly. At high temoeratures, this is accomplished by performing tests at high flow velo ' city where the rate of mass transport to the sample wall, k, is large compared to the rate of diffusion / reaction in the bulk of the sample, k'. At low tem-peratures, where k' is low more conventional low flow experiments are used and i still meet the requirement of knk'. Excerimenta'l Apoaratus and Procedure Three types of apparatus were utilized in this study. ~The first type is a con-ventional microbalance apparatus which necessarily utilized relatively low flow l rates. For this reason the microbalance was used primarily for parametric type studies, for which is it uniquely suited. Absolute reaction rates were assumed to be valid only at temperatures below 700 C where the ratio k/k' was estimated to be 20 or greater indicating that mass transfer to the surface was 20 times faster than inpore diffusion and reacti'on. I

r. i A second type of apparatus involved the use of a stationary sample subjected to gas velocities in the range of 600 ft/sec and where k/k' = 60, at test tempera-tures of <900 C. The third type of apparatus was a fixed bed of granular graphite subjected to is measured directly by gas relatively high gas velocities. In this test, R3 analysis. Low Flow Acoaratus ~ The test procedure involved the use of automatic r'ecording microbalances with which instantaneous reaction rates were measured by sample weight-loss determinations. A diagram of the apparatus is given in Figure 2. Selected graphite specimens were supported by a platinum wire from the balance beam in ( a nonreactive container at constant temperature and exposed to a measured partial . pressure of water vapor for a known period of time. Water vapor was transported via a carrier gas. composed of either inert helium or a mixture of helium and A flow rate of 8.3 cm3 sec was used in all tests (1.6 cm/sec at ll73K). / hydrogen. Water vapor was added to the carrier gas by passing the carrier gas'over water (or ice) maintained at constant temperature with an FTS refrigeration system. Moisture levels were determined by use of EG&G dew point or MEECO electrolytic type moisture monitors. The sample temperatures were maintained at 25*K with the use of SCR-controlled resistance furnaces. Temperatures were. measured with sheathed Chromel-Alumel thermocouples read with an IRCON digital output amplifier. Thin disk specimens (0.0625 in, thick by 1.25 in. diam.) were used to minimize the effects of mass transfer or inpore diffusion on the reaction rates. A typical experimental run involved preconditioning the sample at the test I temperature by soaking it in the H2-He mixture for a minimum of 24 hours. This initial soaking simulated conditions in the reactor, ensured that the sample was thoroughly outgassed, rendered the metallic catalysts in the reduced state and the H2 had achieved nr.ar sorption equilibrium with the graphite specimen. e After the presoak period, the run was initiated by adding moisture to the test gas until the weight-loss rate (reaction rate) stabilized. After prespecified burnoffs were attained, moisture levels were changed by adjusting the tempera-ture of the refrigerated saturator at the gas inlet. Hich Flow Apparatus ~ The high gas flow rate apparatus is shown schematictlly in Figure 3 (Ref. 11). A' cylindrical specimen of PGX graphite was encase < in a tightly fitting crucible of H-451, a relatively nonreactive graphite. (he crucible was machined to fit the standard taper in the quartz furnace tube. This configuration forced the test gas to flow through the central 1.58 mm diameter circular channel, with a negligible bypass flow. The entrance channel into the H-451 crucible is 38 mm long to ensure that entrance effects are eliminated and fully established flow conditions exist in the test section. At a flow rate of 6 f./ min, the velocity of gas in the channel at 1173 K (900 C) is 200 m/s. (This can be compared to the velocity through the reactor core and support structure of s10 to 30 m/s.) The ratio of mass transfer coefficient to reaction rate coefficient, k/k', was cal-culated to be 60; therefore, the high gas flow rate tests were not limited by mass transfer across the gas boundary layer.

g., ~ The tests were conducted by first passing a helium and 2". hydrogen gas mixture through the sample at the specified test temperature, 973 to 1323 K (700 to 1100 C), for a minimum of 24 hours. This initial soaking simulated conditions in the reactor and ensured that the specimen was thoroughly outgassed, the metallic catalysts were in the reduced state, and the H2 had achieved near sorp-tion equilibrium with the graphite specimen. The tests were initiated by add-ing 200 Pa water vapor to the inert gas by switching the inlet gas stream through a refrigerateo moisture saturator. The inlet and outlet gas streams were monitored for moisture content with an EG&G Model 440 dew point hygrometer. The outlet gas was monitored for reaction products CO, CO2, and CH4 with a Carle-model gas chromatocraph. At'the end of the exposure (1 to 36 hours, depending on temperature) the samples were removed, weighed, and sectioned to determine their oxidation profiles. The sectioning was performed by drilling out the central channel with a series of reamers of increasing diameter. The density of each removed increment was detemined by measuring the weight and volume of the remaining material. The sectioning was continued until the original-bulk density-(unoxidized graphite) was reached. From the density determinations, oxidation profiles (percent burnoff versus depth) were constructe,d.. It is sufficient that in' the high flow test the value of k/k' is 60 at 900 C , indicating that mass transport was not the limiting step. The reaction rate data were obtained by analysis of the oxidation profiles. Concentration profiles wer.e analyzed with the following relations: (7)* = = e or taking the logarythm of both sides of equation (7) yields in B = -x/L + in B (8)* o where C/Co= ratio of oxidant concentration at depth x to that at the surface, and B/B = ratio of' burnoff at depth x to that at O the surface. From Eq. 8, a plot of in B versus distance x yielded a straight line with slope of 1/L and intercept of B. o divided by time. Rs values are determined from the intercept burnoffs, Bo Bo < 100"., D,ff and m values are determined from L using Eq. 2. Granular Bed Accaratus The use of granular beds in graphite oxidation studies has been reported previously (Ref. 3). In a present application where the chemical reactivity of the PGX graphite is high, a granular bed configuration was combined with high gas velocities' to obviate mass transfer limitation. The apparatus is shown in Figure 4. Helium carrier gas containing 200 Pa H O 2 is passed through a fixed bed of PGX granules (4 mm high x 18 m and 2000 Pa H2 j diam.) at relatively high gas velocity. The gas flow rate used was 6 1/ min STP l

Equations 7 and 8 are solutions of infinite slab geometry and are shown to be a close approximation to the hollow cylinder geometry of the present tests.

l

~... -. 9 -n ( L j 4 which is equivalent to a space velocity of 154 cm/sec at 900 C. The actual velocity around and between the granules is several times the space velocity and because of the tortuous nature of the flow path, gas turbulence exists, hence mass transport of oxidants to graphite surfaces is greatly accelerated. With the use of small particle size granules (20 to 30 mesh) it was assumed that inpore diffusion would also not be limiting, therefore, reaction rate measurements were accomplished directly by monitoring the gaseous products CO and CO2 in the effluent gas stream with a gas chromatograph. C Soecimen Preparaticn All specimens were taken from PGX logs 6484-138 and 6484-112 in accordance.with ~ a standard sampling plan shown in Figures 5 through 10. The oxidation rate specimens were either thin disks (32 m diameter by 1.6 m thick) or right solid cylinders (ranging in diameter from 6.4 to 32 m by 76 m long) or right cylinders 1.58 m ID,15.8 m OD by 38 mm long. The granular graphite was of various sieve sizes 10-14 mesh,14-20 mesh, and 20-30 mesh. The PGX logs were standard com- 'mercial-grade product, both having high ash and metallic impurity content. Log (112 was previously reported to average 1500 ppm Fe and log 1.8,200 1 Fe Refs.1,2) when specimens taken at locations throughout the togs were measured by emission spectroscopy. More recent measurements on localized. spec-imens taken specifically for oxidation work on logs 112 and 138 showed Fe con-tents in the range 2900 to 3200 ppm and 1250 to 1280 ppm, respectively, when measured by an atomic absorption method. The bulk densities of the logs 112 and 138 were 1.78 and 1.74 Mg/m3, respectively. All specimens were ultra-sonically cleaned in acetone, oven dried at 100 'C, and stored in desiccators. They were always handled with either foreceps or gloved hands, to preclude external contaminations. Exoerimental Results Results of Parametric Studies Usino the Microbalance at low Flow Conditions Experiments using the low' flow apparatus have studied the effects of burn-and P n the reaction rate. off, temperature, PH2, P H 0, CO2 2 The effect of the H2/H O ratio and percent burnoff on the reaction rate is 2 shown in Figure 11. The upper curves where H /H2O = 6.7 and 10 show reaction 2 rates increasing with percent burnoff, whereas the lower curve where H /H O = 2 2 0.3 shows a decrease in rate versus burnoff. It has been previously reported 1 (Ref. 4) that under the test conditions given in Figure 11 for H2/H2O = 0.3 and at equilibrium, iron impurities exist as an oxide. Assuming iron is the primary catalyst, the data in Figure 11 suggest that during the exposure period iron was slowly oxidized, thereby diminishing its catalytic activty. This be-havior is consistent with work at Brookhaven National Laboratory (Ref. 5) which has shown that iron in the reduced state is much more effective as a catalyst ~ than iron in the oxide form.

Pn denotes the partial pressure of gas n.

r* ,, ~. .u ..... - - _.. - - - - - ~ l.., ".. ( Thermodynamic calculations show that at temperatures in the range 700 to 1000 C iron in the reduced state is stable at H2/H2O i 3 (Ref. 6). This is illustrated in Figure 12 whe'e reaction rate is plotted as a function of PH2/PH0 At 2 Pg2/Pg,0 around 3, the apparent reaction rate is low compared to higher ratios. At P,7Pg20 above about 10, the reaction rate is relatively constant. This is H of importance because in the reactor at steady state the ratio H2/H20 is ex-pected to be 10 or more. The effect of water vapor concentration on reaction rate at two temperatures is shown in Figure 13. The reaction rate is considered to be linear with respect H0 (The linear line is,the dashed line in Figure 13). to P 2 The effect of sample geometry on reaction rate,is shown in Figure 9. At higher temperatures the reaction rate per geometric surface area is essentially con-stant with the surface-to-volume ratio (S/V). At temperatures above 1093 K (820 C), essentially all moisture that diffuses into the specimen reacts. At lower temperatures, however, a definite dependence of surface reaction rate on S/V is observed. This phenomenon is illustrated in Figure 15, which is a plot of the apparen activation energy versus S/V. The data from the thin disks. S/V = 1.38 mm-{, produced an apparent activation energy of 240.kJ/ mole, which is consistent with that expected if the reaction was predominantly chemical in nature. Data from the larger specimens showed apparent activation energies about 1/2 of those for the thin specimens, which is.. consistent with the reaction being controlled by a combination of inpore diffusion and chemical reaction (Refs. 6,7). The effect of' t'emperature on reaction rate for low flow rate tests is shown in Figure 15. In this plot data from other laboratories (closed points) may be compared with that obtained at General Atomic (open points). The solid line in Figure 16 is the reaction rate of H-451 graphite multiplied by 1000, which is the reference used for reactor calculations. (This reference rate will be used in the interim until a set of rate constants specifically for PGX graphite are obtained. ) At temperatures below about 700 C the data exhibit an apparent activation energy ~ of around 240 kJ/ mole which is indicative of Zone I, the chemical or homogeneous regime. At higher temperatures the reaction rates decrease, exhibiting a tem-perature dependence of only a few kJ/ mole, which is indicative of gas phase mass transport control, or Zone III. It is apparent that the high temperature data points obtained under low flow conditions are invalid for measurements of in-trinsic reaction rate, because of the mass transport limitation. This difficulty was obviated by experiments at high flow rate where mass transfer to the specimen surface wd specifically enhanced. The results.of CO2 oxidation of PGX graphite are shown in Figure 17. The pre-dicted line for H O oxidation based on 1000 times the H-451 reaction rate is 2 shcwn for comparison. The oxidation of PGX by CO2 is 0.1 or less than the predicted rate by H 0. Reaction rate data taken from the literature for G-5 2 graphite (Ref. 6) and multiplied by 1000 and in KC graphite (Ref. 8) (multiplied by 500) are superimposed on the PGX experimental points. The reaction rate con-stants for the KC graphite - CO2 reaction are presently incorporated into the GOP (Graphite Oxidation Program) computer code. em

.b I

m

g. l ? I l I Results of Exoeriments at Hioh Gas Velocity r Referring to Figure 16, it is apparent that mass transport limitation j exists for PGX graphite oxidation above about 700-800 C. This is illustrated t in Table 1 which lists calculated boundary layer conditions at 900 C in the reactor as well as in the laboratory tests. Referring to the last column in i Table 1 it is apparent that at 900 C the ratio of mass transfer to chemical rate (k/k') is 0.14 indicating that mass transfer is much slower and, therefore, the controlling step. Under this condition, depletion of moisture at the speci-men surface was 87% which further illustrates the same point. The high flow . tests, described in Section 3; have k/k' of 60 and an apparent H O depletion of 2 { only 2%, which, when compared to conditions in the reactor, indicates that the j high flow tests had even less transfer limitation than the core support region i in the reactor. The conclusion of the calculaticns in Table 1 is that the high flow tests were limited by ' chemical reaction and. inpore diffusion (rather than f mass transfer to the surface) and could, therefore,' be analyzed using Eq.1. The conditions and results of the high gas flow rate tests 'are given in Table 2. In tests 1 through 4 the effect of gas flow rate on rate and moisture depletion was detemined. In the other tests, the temperature was varied. The oxidation profiles are shown in Figure 18. The slopes of all the data plots, and there - fore the L values, show relatively small variation. Rs values were determined from the intercept burnoffs Bo, in those. tests where Bo was <100%. Rs was not calculated in test HF-1 when a surface burnoff greater than 100% was obtained. The Rs values are plotted versus reciprocal temperature in Figure 19. The high-temperature data points in Figure 19 were derived from the high gas flow i ~ rate tests; the low-temperature data were obtained from low gasf flow rate tests using thin disk specimens in the microbalance apparatus as described in Ref.1. The dashed line in Figure 19 was calculated from the reference H-451 reaction rate constants x 1000. The line is considered a conservative upper bound of i the data, particularly at higher temperatures. When sufficient data are obtained l on PGX graphite, a set of reaction rate constants unique to FGX will be developed. f /. Results of Granular Bed Tests "A list of conditions and results of the granular bed tests is given in Table 3. In the first 3 tests, the effect of flow rate.was studied and in tests t 5 and 6 the effect of particle size was measured. The data show that in general the reacticn rate increased with decreasing particle size and/or with increasing flow rate; which in both cases tended toward minimizing mass transport limita-tion. The results of measurements of the product gas, CO, of several runs are shown in Figure 2. In all tests, C0 concentration in the effluent gas (or the i apparent reaction rate) monotonically decreased with time. In some cases, the reaction rates decreased by factors of 10 to 100 during the exposure. The reason for this behavior is unknown. A hypothetical explanation is that during i the exposure to water vapor, the metallic impurities which catalyze the oxidation reaction are changed to chemical foms (such as oxides, silicates, carbonates) l which are noncatalytic. Hydrogen was added to the test gas to simulate the reactor conditons, and to ensure that the iron impurity, which is the major impurity and is also a primary catalyst, is in the reduced or catalytic state during the exposure to steam conditons. It is possible, however, that other impurities, or mixtures of inpurities such as Ti, V, Ca and Si, could become oxidized during the runs and cause the observed effect. I J j- ~

9 .,,a.' t r -29= 1 A plot of initial reaction rate vs 1/T is shown in Figure 20. Also shown in Figure 20 are data reported previously (Ref. 2) using the first two rethods described in the introduction. The dashed line in Figure 20 represents the cur-rent reference reaction rate for PGX graphite which is equal to 1000 times the reaction rate of H-451 graphite (Ref. 4). Conclusions Out of core oxidation burnoff rates for PGX graphite have been es.tablished. These experimentally determined rates have been combined analytically with Tech - Spec and expected reactor op'erating conditions to obtain a broad range of l'ife-time precictions for PGX graphite. The ISS Specimen Program will allow verifi-cation of the laboratory determined burnoff rates and establish the actual reactor operating conditions for validation of the methodology for predicting lifetime for PGX graphite. Results from the program will provide confidence l i in the currently predicted lifetime or will indicate where adjustments need to be made. e 4 h = 4 I L 3 6 t e e e. 4 k d r I O t- ..... ~..,...

c. --....( l j b References t 1. Engle, G. B., " Properties of Unfrradiated HTGR Core Support and Permanent Side Reflector Graphites: PGX, HLM, 2020, and H-440N", ERDA Report GA-A14328. General Atomic Company - May 1977. '2. "HTGR Fuels and Core Development Program Quarterly Progress Report for the Period Ending May 31, 19'77", ERDA Report GA-A14418, General Atomic Company, June 1977. 3. "HTGR Fuels and Core Development Program Qu' rterly Progress Report for the a / Period Ending August 31, 1976", ERDA Report GA-A14046, General Atomic Company, September 24, 1976.

4. ' Everett', M. R., and b. V. Kinsey, "Some Aspects of Carbon Transport in HTGRs" Dragon Project Report DP-365, August 1965.

5. " Reactor Safety Division Quarterly Progress Report for Period April to June 1976", ERDA Report BNL-50559, Brookhaven National Laboratory, June 1976, p. 73. 4 6. Wicke, E., et al., " Corrosion Rate of Graphite by CO2 and H O Vapour 2 Taking into Account In-pore Diffusion and Temperature Gradients", Dragon Project Report DP-391, January 1966. 7. Walker, P.1.., et al., " Gas Reactions of Carbon", Advan. Catalvsis ll, 149 (1959). 8. Giberson, R. C., " Rate Constants for the Reaction of Carbon Dioxide and Water Vapor with KC Graphite", Battelle Pacific Northwest Laboratory, Report BNWL-CC-1381, October 1967. 9. "HTGR Fuels and Core Development Program Quarterly Progress Report for the Period Ending August 31, 1977", ERDA Report GA-A14479. General Atomic Company, September 1977.

10. Price, R. J., " Cyclic Fatigue of New-Isotropic Graphite: Influence of Stress Cycle and Neutron Irradiation", General Atomic Report GA-A14588, November 1977.

11. "HTGR Generic Techslogy Program Fuels and Core Design Development",

Quarterly Progress Report for Period Ending May 31 _1978,.GA-A14953 (Addendum), General Atomic Company, June 1978. eu.m %W ee w

.:c' F 1..... t... i. l t i TABLE 1 BOU'*DARY LAYER CO::DITIONS (Calculations at 900 C) Laboratory High Lov Conditions Symbol Reactor Flou Flow I Regi=e . Turbulent. Laminar tsminsr L \\ Velocity, it/sse V 54 664 0.1 Reynolds number Re 111,000 263 0.17 ~ Mass trans, coeff, it/sec 'k 0.1 7.1 0.027 Reaction rate /inpore diff coeff, ft/see k' 0.09 0.12 0.19 k/k' Sh 1.1-60 0.14 Depletion at surface 47 2 87 e-( 4 t E h O e mm. k j i e s ? .e + ... ~.~ ~. ee

.2 e e e Ie me .F .F P= se O' W tr.l M.

O.

ep a= g e e N

=

C O O r= o L = at w A

=

S q b en e ett e M. .e. M. 4 N. Se s

== g a m .F O w

== C C M O e gg M 3 54 4* N h M O+ .O H E .g C. C. C. O. O. O. g O O O O O O 8 e n m N I N. N. N. N. N. 3 g w O O O O O O O g <O N C O o U

== c c M. w Ou e =. ^ - M. N = g N o O. =- 6 E-O ert m N =r M C O l p g<m

=

Q. u C 0% Y My e d U m e M e D. ^*5 O O m O i.n w

=

on w te O h = ert m M N 0 Pa. m O O.

O y

>w

=

N m N N

=

N g O.E aC w e- .=. 3 O O

.A e art Os art er en g

C

=.

.f.- ad. .f. N. r. E*'.t P*

b. N N3w g

A 3 53 P= Q O O O O O O b O = g .n. .n. .n. .n. M. . a w ED 4 .t art

O M
=

g., m w w 4O 'M"' m s O .e a e a e a .m. c6.O e a e mo no mo no nO mO 60 m Og Ew P= 0 r= 0 P= 0 NO N af% P= C hO N W* 7. w O= w<D C c3 mo u = to W =5 cs -m ,==w g.e hd a e= w se e= es = s= w,= w== w w e.4 .e S a=. A e

A e

e m .m a e e & G 8"5 m O O O O O to De N C.rt u% e'l w O O O 05 = N O C C w N N

=

gw w w w w w w OO C N 3 ** O B

=.

3 .= 0 m.= to se. =a. ,N b = u .= ' N F't er get to e et e a -* 9 I 4 4 4 I t I w eO Ba. b b b b fa. b to. la % = M 2 M 2 E E e l I = i I l \\ = l I L i I I ) e I'*

e.-**.
=--*e--==

e- - ~...... -, _. L

I i Table Summary of Cranular Bed Testt . - ~ Reaction Rate. P p CO S Pa U Flost Sample 2 Outlet Initini Average Test Mesh Tpp Rate Weight Time Average (Initial) W Based on Based on f No. SI c C T htin n lirs Pa Pa Frnetton CO M CD-1 10-12 900' 1 1.7 140 160

5. 7 (-8) 6 250 97 1.1 (-7) i

~ 2 10-12 900 6 1.7 10 260 85 0.36 1.6(-7) 2.2 (-8).- '3 10-12 900 18 1.0 0.75 28'O 28 .0.23 1.5(-7) 3(-7) l 4 10-12 800 6 1.0 2.5 300 19 0'.094 3 (-8) 3.5 (-8) ~ 'S 16-18 900 6 1.0 2.25 280 93 0.36 1.6(-7)

1. 8 (-7) 6 20-30 900 6

1.0 2 260 i 72 0.29 1.4 (-7). ' 1.5(-7) j

3. 5 (-8)

%d .L s 7 20-30 800 6 1.0 4 180 55 0.14 9.5 (-8) 8 20-30 700 6 1.0 1.25 280 8 .017 1.4 (-8) 1.3 (-8) ~~ ' 12 .04 2(-8) 2.3 (-8) l 9 20-30 750 6 1.0 1.75 280 10 11-451 900 6 1.0 16.5 260 0 7.0 (-4) O' 6.9 (-11) +30 .P in all tests = 2400 Pa H 2 s e e e e i -, - -, ~,. _,, -.- - - - - ,n,-. w e- -,----,----.,w.

i* - 4.-

i .l, t ZONE 1: LO'if TE!GERATURE CHEMICAL PIACTION IS SLOli ALLOilING MOISTUPI TO PENETRATE i VIRTUALLY UNDEPLETED Ee ~. + HO 2 .} . - - =,, = = '. T v = =- ZONE 2: I N 1 =-..s,.,- D,i e&,- -. 2------- OVERALL REACTION CONTROLLED BY UO"BINATION OF CHEMIC,AL REACTION AND IN-PORE DII?USION t t. He I + i g HO 2 1 i .4 \\ } ~~~ ~ ZONE 3: HIGH TE!SERATURE s' CHEMICAL P.EACTION IS SO FAST THAT,OVERALL RATE IS CON-TROLLED BY TRANSPORT TO SURFACE OR DIFFUSION THROUGH BOUNDARY LAYER g I.- I He I + \\ HO j 2 ( e Fig. 1 THEORETICAL OXIDATION PROFILES FOR THREE TEMPERATURE REGU!ES s. g. __e

CD 0710?.ALAJf CE i

I l =

t. s,.

i e!{. ~ (IEACTk"X! PROC'JCTS +- , STRIP CHART p* C ' R U '- - ttATER Y POR + C:5.ti~4 4 .c l .4 .v .l N, h.. e..,.,. n,. n,.,P L g a vv .u SAMPLE FUR!iACE t.s i J v FIG. 2 Typical app.2ratus for continuous veight loss determi:ution of graphite oxidatica by coisture in incte gas ~ l -~ - - - t.

l PROCESS GAS (He.H H O) 2 2 }

- R ESISTAt;CE.

WOU:40 FUR!iACE ? i I H 451 G RAPHITE f f \\ j,f, PGX GRAPHITE

'9';

r. L ~ l ? QUARTZ TUBE 4 / STAT:0ARD TAPER = 1 i -THERt'0 COUPLE = SCALE = 1:1 o Tig.3 c'raphite oxidation apparatus fo'r high gas fiev ra:e experiments ~

1

. - - -u. ~ 637. e .s i PROCESS ~ GAS (HE,H2,H 0) 2 mtt ,v i . 1. l. i ~ t 1 ~ l .e QUARTZ TUBE t = C Q u er* ra w o o n-- Q Q w, #. ~ %'.::,7,;,.-;i 'u:24,:S PGX GRANULES trJitE.P.; 41 1-G--l+-+1ESit ~ SINTERED ~ QUARTZ FRIT k ), L r s e t PC 00%T GM r r (kA,%0 i CDco2.) s r c

/

FIG. 4:'HIGH VELOCITY GRANULAR BEo EA?ERIMENT [ m e u. - # *Eum..uu.aus. epee '* M ein...'.un.* #.

m. 8. N.eemme m=me m meaew.. %

,y,,,,,,,,, ,...g,, pwo ~. ~~~~ -

_\\ U

- L ?

.--.=. a... l { ~ 1143 DIAf.1 305 DIAf.1 00 -l DISCARD '^"N SSS .i 7., ~ 2700- - N -:_ cgo ~~ ~ L J i-152 E13 DIA!.! l g - []';.; W. l{ ~ '.~ .?. / 152 l '. N. .-,~ 102 ( 1057 DIAf.1 180o SLAB 1 914 a s 152,,. DISCARD SLAB 2 ~ 152 TOP 51 SLAB 3 . o 152

p..

SLAB 4 I t / 152 SUB5 '/ Y 1829 '/ 152 g i i // ~///// 1 / SLAB 6 1 /////// 152 SLAB 7 ( 102 g A SLAB 8 152 SLAB 9 i 152 i SLAB 10 ~ SLAB 11 SLAB 12 0 f.iElJSI0tJS

DISCARD,

_ FIG.5: slabbing diagra: for oxidation rate and strength versus oxida-tion of support floor block graphite PCX, legs 6484-112 and -138, slab 6 e. e p t 4 e

l i l e i. - [,; / gygo t .y i j I [ l

/-

I l 20 i l. l >y l t to g i=. I ' FATIGUE, I 12 l l .[ [ [l 7, j, l 1 4 l I l . SECTION AX \\ 2 l c w m s, i I, ,/' oC 3 ^\\ TAKE 4 CORES 17 E s s ',s I E 08 152.4 Nr.1 127.0 f.!!.1 Tig. 6: 0xidation rate and mass' transport sanpling diagram for PCX graphite, sinb 6: (a) section Ax, 270* to O' a e 5 0 4

il r -152.4 MM-1 127.0 MM - = . I i f. 1800 ,4#, r I 7 B C 0 E ~ TAKE 4 CORES A 9-15 1 SECTION BX t 1 8 l I I to f 14 18 l 'f 1 o ga A l (b) section Oxidation rate and mass transport sampling diagram for PCX graphite, slab 6: Fig. 7: BX, 90* to 180* I 0 e, ( g l. -l f e,--

5 11 ( i TAKE 4 C0!!ES. 17 23 -l [*. t 28 g @h /dd, /.,d ",' ,oS \\ J I 'r* ' g\\.r\\ Ic 00 to \\ \\ l /,. ' -< ;.' t 2, t i \\ 8 /, ' l t, - \\ t \\ l N ',,l'[ o 'go - e SECTION AE i (c) section Oxidation rate and mass transport sampling diagram for PGX graphite, alab 6: s Fig. 8: - AE, 270' to 0* l l a i i l i

q s ~ g l l TAKE 4 C0llES-25 31 j -l r. c\\'.m. 36 s 40 44 48 1 32 a \\ 30 "o" Y \\ l's.l t/.$' l / i \\d i t g l />(,L't,, 1800 a l' '- a, ~ / '? E IECTION BE (d) section { i Oxidation rate and mass transrort sampling diagram for PGX graphite, slab 6: Fig. 9; BE, 90' to 180* l e t ,,.r-. ,Q. c y

q .c \\ \\ p020 lI,/ ' \\ \\s s 2700' \\ \\ \\ 012 014 016 i ,f 018 f 00 g g // i \\ \\ \\ 04 OG f I e \\ \\ 02 03) 010 ,i ; f g y i t 'g \\ \\ g l s l s SECTION AY t 1 i Fig. 10: sampling plan for oxidation profile mass transport of PCX graphite, logs 6484-112 and ff 138, slab 6, section AY, 270* to O' j o e . m

sy z,. =. j .;,.5 . ~ l 10-5 i OoD00OOO@ OODOO ~ O O Og gg 3 g6 3 A A A A A a a og A O gA T = 1033 K (8200 ) C . dells!TY = 1.72 ;.10/t.;3 SN = 1.38 f. M-1 O P g20 = 300 PA - g g 10-6 _ [ cog p O P OOo H2 p O OO fp (PA) Hg H O -l 2 O O 3000 10 ~ O .,u_ O O A 2000 6.7 t g O 0 100 0.3 g' O G 5 O .= 10-7 o O ~. O O OO 000 .OO O ~ O O p f f I I I 10-S 0 2 4 6 8 to 12 BURfl0FF (%) FI - II.: Effect of P4 and burcoff on PCX g nphite oxidation E 2 S.O

O \\.. - I' I -6 10 O .i. O 1013 K O \\ O 973 x ~ 'l' I .. '3, l 1 ? O 933 x .P. 0 = 100 P.s h., l i J ..,r. = .s n - / g-f . a m y 6 6' a s -] ,a lo v .a. Q O ^ N V4 o ~~... - o-6. s 9 ............. ~. - i. p. __..O s-s- -8 ~ 10 ~ i e e. g...... y O t t t t t 9 1 9 1 1 1 1 1 30 11 3 10 P / Pit 0 H 2 2 FIC. 12 EFFECT OF F /P ON PCX OXIDATION PATE O

.tt .,..j_.__..._....,,__.;..____. _ . _... g, l s i t, 1 i l 1 I 16 ( 3 p-1.785 MG/M ~ /PH O = 10 PH2 2 SN = 0.124 MM*1 1 O 1173 K ($33*C) 75!! 133 O 1073 K(S33'C) 7553143 E W >= " 19 z o 3: o E l f l ~. f 7 / /. gg I I I 1 1 I i / O ~ '01 10 2 10-3 l Pil 0(PA) 2 ' E## 71g. 13: , Effect of PHO 2 g I g

t t P gg-1 0.1033 K (220'C) PH = 1000 Pn. 2 0 '1013 K (7 3*C1 Pg2 = 1C0 PA 6 $33' K (653'C) l ~ 0= 6 l BUP.fiOFF a 1% Q FA8 K(575*C) 1 I t 10-2 l t [ t O O t O t U O l v 1N v g 2, w n C w u.s p i 'sc A m A o p L3 u, f 104 .o. t 1 O I ~ i 39 ~ se 104 i 'l 1 i I i 1 t 0 C.2 0.4 0.6 0.8 1.0 1.2 1.4 l $N (MM*,), r ~ t Fig.14: Oxidation of P0X graphite versus S/V, ~ t '~"~^

48-28 O 24 ~ ~ PH = 1000 PA. 2 PH O = 100 PA ~ t 22 2 B1)RHOFF a 1% 20 s 13 O ~ 3g 'e, 14 ~ s Q ~ 12 'o' x, o u,. O 10 .1 ~ i-8- "' r' 6 4 2 t 1' 1 I f f ' O O 0.2 0.4 0.6 0.8 1.0 1.2 1.4 l SN (MM 1) ~ ng. 15: E, versus s/V for Pcx graphite e 1

. t 1

r. ~ l l l --r--- 7,_ -,y -r

m.. . j .<c l O o_ o9 @ 'O O. i _a 10 7 O E-451 RATE x 1000 - O f .f. o o .b pl l O ~ { G 24 0 0 (d d o -9 10 Q ~ m A .3 o T. A - V.~O - gk,'%, 0- .,. 4 -- ,,, 2 i , y*- g Q g -. r o es bg h ~ g ~ '.o g Q @ 10 ~- Samole I S/V cs KQ 0 7694-36 138 13.8 4 g . d 7694-134 133 13.8 4 0 7694-73 138 13 8 .p g 4 ~~O7694-48 138 13.8 OO />. 6 759S-80 138 13 g 7694-16 138 6. I 7694-3 133 3.4 . 4.\\:. .O t - 10 -- [h 7 69t.-26 138 2.4 -.Q7694-11' 138 1.6~ .s D 7694-21 112 13.8 i 1. 7694-120 112 13.8. ~ B:C Loop 3.4 s 5.3 DP.;CO:t. . A an. ~ 5.3 100,0 C 900 C 800 C 700 C. 600 C 7.0 8.0 9.0 10.0 11.0 12.0 ~ 1/T x 104 (K-1) 1. FIC. 16 PGX CRAPHITE OXIDATION RATE VS.'1/T; COMPARf66:; WITH DATA

FROM OTHER LASORATORIES

( O e. 07.'6 -D j

e., - ..-_...-...a. ,. g r e - t. e . s \\ t ~, .H O REACTION RATE, l 2 ~ 10 H-451 x 1000 1 I 4 DRACOM DATA, i ~- cC-5 Cur:inz) x.1000 t s O s. f _c g 10 ' ^ s s. CDP CODE ~ d g y.Z"M" b s E, t% u v N,z. y 500 g s ..... t.. x 0 g .o F< u I -10 ~~ i s2 10 O g SAITLE LOG O O NO. NO. 6 ~ . O 7694-36 138-O I O 7694-48 138 Q 7694.73 138 i 10 ~ TQ 7694-120 112 ~ g a 1000,o 800.O 700,O 900.o 600.O C C C C C 4 6 6 a 7.0 8.0 9.0 10.0 11.0 12.0 1/T x 104 (K-1) FIG. 17:.oxIDATIcx or PCx CRAP 111TC BY CO ^ 2 l l

I 1 g,- = s 4 l l 100 ~~ c f. VOLUMETRIC AVERAGE 10 , TEST TEMP. BURt!OFF O g fl0. (K(C )) (%) C O ~ E 0 3 1173 (SCO) 0.46 c R O.4 1173 (900) 0.76 C5 1223 (950) 0.49 A 1 1173 (900) 7.1 V'8 1073 (800) 0.44 \\

9 1:23 (1050) 0.35

( bb b N 1 0 I i 1 0 1 2 3 DEPTH (Mt.1) Fig.18: Oxi_dation profiles for high gas flow rate tests e ,d %g .e e

w:s .2: ,cc: l .s o. - 5.2 - g, (K) 1273 1173 1073 973 873. 4 10 9 1 1 1 i L \\ g + 4 ex gg-7 g \\' H 4'51 REACTIO? RATE ~ 2 X 1000 10-8 - \\ ~ \\g \\ 7 -,s g \\g i [ N g2 g 0 = 10 g P 2 D 10-3 sal.'PLE LOG NO. fl0. O O 7598 80 138 \\ \\ O D O 7654-3S 138 g A 7634 48 138 O ~ ' ~ Q 7694-73 138 0 .s k. N 76S4134 138 A 10 \\ 0 763S.120 112 A 7694 21 112 O g\\ $ HF 3,4,5,6,8,9 138 HIGH FLOW TESTS O\\ ~ A g \\. o A \\ i i i i t i

,g t 7

8 9 10 11 12 ~ 1/r x 10 (K-1) 4 l i ( Fig.19: Intrinsic reaction rate of PCX graphite versus 1/T w t emo

2,.

s,.* i.

r. j l ~ (g) 1273 1173 1073 973 873 6 107 6 6 6 L i ~ .\\ l y + t i G 8 2. A[s, 4 10-7 \\ O- - H 4st asAcTiar: P.Are X 1000 h \\ 10-8 \\ \\U. g g= \\ I 7 *. \\U 'P}/PH O = 10 O [- g 2 D 10'f SAMPLE LOG NO. NO.

k

~ O 7533 80 138 0 5 133 O 7634 36 A 7594-18 138 ~ \\ Q 7594 73 133 o ~ 1 l C 7694134 133 A \\ t jo-!O O 7634.120 112 A 7634 21 112 5 O O HF 3,4,5,6,8,9 138 HIGH FLOW TESTS O ce z-*9 ne cie a. s., m a\\ 7 A \\ \\. I 'l I t

t. O e9 k

, gg' 11 '7 8 8 10 11 12 '4 1/T l0 (K-I) Fig.20: Intrinsic reaction rate of PCX graphite versus 1/T 1 e O =

z. * :., -

.=..a.. t l.- j i RESPONSE TO NRC OUESTION 411b OVESTION: Give general information of the Core Support Structure. ANSWER: The core support structure consists of a graphite and conposite metal and con-crete structure under the core and a cylindrical steel core barrel surrounding ( the core. Its function is to support and laterally restrain the fuel and re-flector elements and to direct the helium coolant flow to and from the core. The arrangement o'f the core support structure is illustrated in Figure 1. Each of the 37 refueling regions of the core is supported and located on a graphite ~ core support block of hexagonal cross section that is approximately 37 in. across its flat-faces. The side reflector elements are also supported and located' on graphite blocks which fit together with the hexagonal blocks under the core. Each of these graphite blocks is, in turn, supported by three graphite posts, which have spherical ends to permit them to rock slightly in accommodating dif-ferential expansion between the parts of the structure. The top ends of the posts are located in spherical seats in the core support blocks, and their bottcm ends rest in spnerical seats in insulating bases. Primary coolant leaving the core flows through passages in the core support blocks and through the plenum formed by the core support posts. For further details on the Core Support Structure, see FSV, FSAR, Section 3.3.2. O 9 e' e m m. = - l I 1

l . l rs

.55

.k ]' '[ e f. g E .E f_ Y]. W2 1 3 3 Il i I i il .s .i \\, .\\ \\ \\ \\ \\-

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gto, c.

o F. r -7_ 05

i l

0 m. otoco - ~. u-n -u::. .t ca s = u. ts. og o i N d 'c*0 $ f 5 Y 'h f-f g. ,J 3 i .ce ectorot/ca r.c _ ~. - m 03 3 = - s A.. 8 l

c a 0 0 o; p

t i i , c o o> 0; rd-I I =, n i

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= O^ 0, ~ ~3rcec' o 3=0, t I 2 ) ) . Kj. tl l l, CD 00 b

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' g. I s t 6 3 03 CD 0 f_ , / w W_ _ - . JJ ** / / .n m. 1 . R=m=.:n,M;;) j -, _. /J - m t . - - _ : rm. __._.- - _. m./-.. ~ a\\ v/~_ / D / / V \\ g -l p< R 2I l 11 . E-L 1 c eO e.

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RESPONSE TO NRC QUESTION #12a QUESTION: Give information on consideration of thermal stresses in the core support block. ' ANSWER: The stress analyses of the core support block due to primary loads and steady state thermal loads have been completed and are discussed in the following i sections. The FSV-FSAR states that "the high temperature portions of the structure con-structed of graphite and carbon are designed with a safety factor on ultimate strength of three, under the combined functional and seismic loadings." (Sec-tion 3.2.2.1 ). This means that a safety factor of 3 is required for primary loads such as those imposed on the core support by the deadweight of the core,, the core pressure drop, and seismic accelerations. Secondary stresses (thema]) are not included in this criteria. It is' judged that the safety factor of 3 is ~ sufficiently large to accommodate secondary stresses without compromising the structural integrity of the CSB. The core suppor't block is a hexagonally shaped block made of pGX graphite which supports the weight of 7 fuel columns consisting of reflector blocks, active core fuel blocks, plenum element and orifice assemblies. It is supported over the core support floor by three equally spaced graphite support posts. The space between posts forms the lower plenum for the core outlet. gas. The core suoport block is subjected to internal temperature. gradients as well as the core pressure drop during the operating life of the reactor core. I Steady state power operation produces temperature gradients within the CSB' which are caused by temperature differences of the gas stream passing through six inlet channels. This analysis combines the steady state thermal stresses ~ l with the primary stresses caused by mechanical loads such as the weight of the core, core pressure drop, and seismic accelerations. The same structural models were used for both mechanical load and thermal stress analysis separately and for the combined case as well. The analyses uses the three dimensional finite element method. Thermal Analysis J During steady state operation of the plant, there are power variations (power tilt) across a refueling region which result in fuel column gas temperature differences. The core support blocks, which support each region, will exper-tence different steady state gas temperatures in each of the six coolant channels in the top of the core support block. Heat transfer will take place from the hotter channels to the. cooler ones setting up a temperature field across the block with resultant themal stresses. p

dy. l ! r The differences in gas temperature across the block vary slowly with time. This variation is dependent on fuel burnup and region helium flow rate which are adjusted to maximize plant and fuel efficiency. The thermal analysis was performed by the finite element program HEAT 3 (Ref. 4) which has the same isoparametric 20 node brick finite element as the ccde THREED (Ref. 2). The element is formulated assuming a quadratic temperature field within the element with matching temperatures at element interfaces. The model is a 60' sector of a. 360' solid with adiabatic boundary conditions on 'the two cut sides, as shown in Figure 1. The rest of the boundary conditions are derived from the predicted gas conditions as follows: 1. The gas temperature in one inlet channel is l'520*F with a flow rate ~ I cf 9431 lbm/hr. 2. The gas temperature in the adjacent inlet channel. (post side) is 1722*F with a flow rate of 9003 lbm/hr. 3. The gas temperature and flow rate from 'the central fuel column across-the top of the block is 1616*F and 878 lbm/hr. 4. The gas temperature and flow rate along the outside of the block is ~ ~ estimated to be 1522*F with a flow rate of 545 lbm/hr. Heat transfer coefficients are derived from these conditions. To determine which steady temperature case will result in the maximum thermal stress, the temperature gradient is calc.ulated from the temperature field.. The time at which the maximum gradient occurs is selected as the most probable. Other likely candidates are selected based on judgment. Stress analyses are then carried out to see which case is critical. ( The worst case temperature difference occurs at a time point of approximately 17 years from the beginning of plant life. At this time, the gas temperature difference between' adjacent inlet ducts is a maximum of 200*F at 1005 power out-put in refueling region 21. Typical isotherm at the top of the CSB are shown in Figure 2. This location was selected since it is the area where maximum thermal stresses occurred. 0xidation Analysis Wetted surfaces of the CSB oxidize at different rates depending on flow rate and graphite temperature. The same ratio of oxidation on various surfaces of the CSB was used in this analysis as was used for analysis for primary loads only. Namely, they are 1 A:B:C: = 4:3:1 where A = surface of coolant mixing chamber B = top surface of CSB C = outside surface of CSB stem M %g

%g4 l

-i w.

2. -.

t,- - ] ) l These ratios are then used in the structural evaluation to parametrically reduce the section sizes of the CSB. Structural Analysis Model: The structural model consists of a 60* sector of the CSB. This assumes periodic symmetry of the block supports and coolant channels every 60 degrees. The thermocouple sleeve is also modeled as though it pierced the CSB l .every 60* when, in fact, it ha's symmetry at 180*. This is a conservative approx-imation since the sleeve hole is a stress riser. The finite element used is a 20 node isoparametric brick, nith 3 degrees of freedom at each node. It is formulated with an incomplete cubic disolacement field with element-to-element displacement compatibility satisfied along mating faces and edges (Ref. I through 4). Loads: The loads used in the analysis are the combination of primary loads and the secondary load as depicted previously in the thermal analysis of this report.. The primary loads are shown in Figure 1. Secondary stresses result from the thermal locds induced by the maximum gas temperature difference of 202*F between the adjacent coolant inlet channels at the 100% steady state power condition as previously described. i Boundary Conditions: The following boundary conditons were applied to the stress analysis model: 1. At the O' and 60' cut sides, the boundary conditions are free to move in radial and vertical directions but are fully constrained from t movement in a direction normal to the plane of symmetry. 2. At the interface of the support seats, the nodal points are free to move horizontally but fully constrained from vertical deflection. Although the support post seat does have a finite spring stiffness in the vertical direction, the boundary conditon applied in the analysis makes the post seat rigid. Since the critical stresses are expected to occur in the top portion of the block, this limitation is not expected to affect the essential results even though it is highly conservative in the post seat area. Material Properties: Since this analysis is limited to stresses near 1/3 of the specified minimum ultimate strength, linear elastic material behavior of the PGX g'aphite is assumed. This assumption is semewhat conservative but is r believed not to affect the acceptable accuracy of the result. The material properties used in the analysis are summarized in Table 1. l I l ~ J

j t' 59-I l 6 i Structural Criteria and Stress Analysis: Since the steady state thermal stresses are a result of slowly varying gas temperatures during the normal plant operation, its combination with the primary stresses due to mechanical loads can be treated under fatigue criteria for cyclic loading. The allowable stress i criteria for graphite structures stated in Section 3.2.2.1 of FSV Final Safety Analysis Report (FSAR) does not consider cyclic or secondary (i.e., deformation controlled) stress states. Thus, it is necessary to rely on good engineering practice to set limits on the primary and secondary stresses. Results of Analvsis The maximum tensile stress is 451 psi and occurs in the thin web of material between the coolant channel and the keyway at the top of the core support block I as shown in Figure 3. This stress results from the combination of worst core steady state thermal loading, core weight, core support block weight, core (- pressure drop, and operational basis earthquake' loads at 100% steady state power output. The largest contribution to this stress is from the themal loading which is 425 psi. The other 25 psi is from all other sources. The point cf maximum primary plus secondary stress occurs in a different location than the point of maximum primary stress as shown in Figure 4. The maximum primary stress occurs at point _A in the top surface of the core suppor,t block in the web between coolant holes.' It is to be noted that the maximum secondary stress is highest at the beginning of life when no oxidation of the. ~ CSB has occurred. As oxidation proceeds', and the section sizes become smaller, ~ E the maximum thermal stresses at point B in Figure 4 decrease as shown in Figure 5. However, at point A, where the maximum primary stress occurs, the combination of thermal and mechanical stresses increase as oxidation proceeds'. The increase is caused by the primary component since the thermal stress com-ponent is decreasing only slightly as oxidation proceeds. The net effect is t' to add an increment of secondary stress on top of the primary stress shifting the curve of stress versus oxidation depth upward. Maximum compressive stresses are well below allowables and are not critical in t the core support block. A comparison of the maximum primary plus secondary stress with the minim'u'm [ ultimate strength of 1000 psi shows that the stress is 45% (safety factor of 2.22) of ultimate strength yielding a margin of 55% avail'able between the maxi-t ' mum operating stress and minimum ultimate strength. For primary stesses a saf'ety factor of 3 is required by the FSAR. Looking at point A in Figure 5, this safety factor is reached when the depth of oxidation i reaches 0.61 in. When the secondary stresses are included at a depth of oxida-tion of 0.61 in, the total stress is 410 psi. This results in a safety factor of 2.44 showing that the stress is still substantially below the minimum ulti-l mate tensile strength of 1000 psi. s' _These steady state themal stresses slowly cycle between 0 and the maximum stress and can thus be compared with fatigue data on PGX graphite (for R = 0). Figure 6 is a plot of some push-pull' type fatigue test data for PGX graphite. I The lower curve below the data is the best fit curve that has 99% of the data above it at a confidence level of 95%. This is the fatigue curve used in graphite e m

i : ** " l, - .. ~ r j I I design. It has been estimated by counting the temperature swings in the pre-dicted 30 year temperature history that the equivalent number of thermal cycles is 445. At that number of cycles the allowable cyclic stress range is 680 psi. Using an average cyclic stress range at point B of 430 psi yields life fraction margin of 0.63 or a margin against fatigue of 37%. At 100,000 cycles, or close to the endurance limit, the allowable cyclic stress range is 600 psi. Thus, there is no reasonable amount of cycling which will cause premature failure of the core support block for the mechanical and thermal loads of this analysis. Conclusion The core support block will withstand the steady state thermal stresses up to 0.61 in, of oxidation on the inside surface of the coolant channels and still (.-. maintain its intended str.uctural integrity. O G see e B e ( e e 4 e- = g me 4 +

id [ [ l References r 1. Young, R., "Three D. Finite Element Program MESH 3A Input Reference Manual". f I 2. Young, R., "Three D. Finite Element Program THREED Input Reference Manual". ^ 3, Young, R., "Three D. Finite Element Program PRINT 3 Input Reference Manual".' 4. Young, R., "Three D. Finite Element Program HEAT 3 Input Reference Manual". t 5. Price, R. J., " Cyclic Fatigue of Near-Isotropic Graphite: Influence of Stress Cycle and Neutron Irradiation", GA-A14588, November 1977., M g e e w ' e e P 2 ~ I

TABLE 1 MECitANICAL PROPERTIES OF PCI CRAPHITE DIRECTION 1_ 2 3 YOUNG'S H0DULUS (PSI) Egg = 1.0 X 10 E22 = 1.0 X 10 E =.75 X 106 6 6 33 v 2 = 0.1 v23 - 0.15 v31 =~0.15 . ' POISSON'S RATIO 1 r 6 6 6 SilEAR HODULUS (PSI) 012 =.45 X 10 C23 =.330 X 10 G3 =.330 X 10 TilERMAL EXPANSION STRAINS Along the. third axis or Z axial direction the expression is: E "Th3 = 1.32333 X 10-6 ( T-TREF)+ 0.35333 X 10-9 (T-TREF)2 In the isotropic plane (R-0 plane) the expression.. is j, Th u - 0.92333 X 10-u (T-TREF) + 0.473333 X.10-9 (T-TREF)2 c All temperatures are expressed in units of Fahr nheit. Where: T = temperature of Interest I TREF = 70F is the reference temperature C h = expansion in inches / inch T l I i L I

? f. I::. D *.. Q, g. 's P' Pi-o

r./;.-

, SOURCE / ~,. L (LBS). (PSI) .Q s,s TasitSho*F 'I'- / Core Wgt.' 19',200 .15.9 ini90l*S.Tep5=l72n. s - $40034rfr. sJ - -Q: ,'T,,$=.1!!224. '. g',*d. Pressure Dro'p 10,000 8.3 s d i 1 ?!!4E r.' r un 4 scnot'. (LAf3 .. ~.,, ~_._ , l cc.. tral Seismic (*.133 g) 2,560

2.1 y

E s-G*IUnsts y _ Kl O +Tr p =1s1G*F ^ p( N =an ase. ~ ~ \\ I p' i TOTA 1. - 31,760 26.3 s ]\\ ~. ~ \\., N*,j ,s N \\ g w .. h. \\ N.... . / ~ (- i nra FIGURE 1-i i e i. ' CORE SUPPORT BLOCK LOADING AND GAS TEliPERATURES j .i j .----w .w-,-ar ee-es e-+ e-w -er-1-----w +-

?. FIC. 2 CORE SUPPORT BLOCK STEADY STATE TEMPERATURE DISTRIBUTION AT TOP SURFACE CAUSED BY MAX. CAS TEMPERATURE DIFFERENCE ' LEGEND - IllNs 1522. 0 Ia 1557.C0 1 ea. ?= 1559.00 + 3= 1534.00 S (= 1600.00 5s IGIS.00 ,.., f, E: 15U 00 so- ?s 16(S.G0 - ~ 8a 106{.00 .,.j'.~.. :.... ,.,, 3.:. ' 9s 1000.00 _ 10 s IC9G.00 .

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3 ~* n m .~ FIC. 4,HiXIHtM PRIMARY PRINCIPLE STL.,S.EISTRIBUTION AT TOP SURFACE C IHE CORE SUPPORT BLOCK .. j; g.. (o.5 INCn OXIDATION) l. t LEGEND:. at-- MIN: -(l.7S( * !: -33.000 e to -J ?= .CGO 3 33.C00 ~ ( E'.000 ~ 5: 53.000 l Ge 131.0C3 l s e-.i ~ .

o.

.7 : 165,000 6a 150.000, s-S: 731.000 . 't s ;;- a=== t 10 261.030 1 MAX: 289.0]! ' ~ ~ s... + s4- - A-Max. Primary Strees O j g. Location j. o B-Max. Primary and h -d' ' y st-4 Secondary Stress g, % l Location

i

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t; ' .? n . I L RESPONSE TO NRC QUESTION il2b QUESTION: Give information on design details of the core support posts. 4 ~ ~ ANSWER: The core support post (CSP) is a 38 inch long by 6 inch diameter column of Union Carbide Company ATJ graphite. The post ends are spherital which roll ( in larger radius spherical cups or seats as shown in Figure 1. When a core . support block is moved horizontally relative to the concrete floor the posts tilt. This design feature allows constraint free movement of the floor to adjust to the many thermal movements that can o: cur during. operation of the plant. The CSP is loaded by the weight of the core and core pressure drop. The maximum CSP loads are 10,000 lbs under a refueling region and 17,000 lbs under the per-manent side reflector (PSR) support block. Since the PSR post is identical in dimension and material to the CSP it experiences' the critical loads. In addition,' the PSR support post experiences the greatest horizontal displacement of the top relative to the bottom during all operating conditions. The horizontal displacement is caused by thermal expansion of the steel core barrel during the rise to power. The PSR support blocks are connected to the core barrel by metallic keys, s The peripheral refueling regions experience the greatest horizontal movement of the core support blocks since they are farthest away from the floor center-line. Thus, the critical core support posts under the active core are located under the peripheral refueling regions. This action results in the post load being applied eccentric to the post elastic axis as shown in Figure 2. This eccentricity causes bending in the post which is added to the axial compression. The posts are bathed in the core exit gas, which is then mixed, and is distri-buted to the steam generators in the lower plenum. Gas exiting a peripheral refueling region would be reflected somewhat by the main flow heading towards the steam generator entrance ducts. Flow patterns in the lower plenum are extremely complicated requiring extensive flow mixing tests to determine tem-perature distribution and velocities in the plenum. To simplify the post oxidation analysis, it is a.uumed that the gas exiting from a peripheral region impinges on the core support and permanent side reflector support post. Thus, the temperature history of the gas at the exit plane of the core support block is used in the oxidation analysis of the posts. Since there is mixing of the s s.

_......--l c' I region outlet gas with other gas in the lower plenum, tending to reduce the i temperature, it is felt that oxidation predictions will be higher than expected, making this assumption conservative. 1 The temperature history used in the oxidation analysis is rearranged into a { monotonically increasing temperature curve with respect to time as shown in [ Figure 3. This temperature frequency diagram is representative of extreme .l operation of the plant at Tech Spec limits of operator control and total oxi- [ i dant levels. The basis of this temperature history is outlined jn Table I. l Oxidaticn analysis is carried'out using a solution derived from coupled dif~- fusion chemical reaction rate theory developed.for porous media. The Langmuir-Hinshelwood equation is used to describe the steam-graphite reaction rate of i r-ATJ graphite. All of the coefficients are uniquely determined feom oxidation 4 experimental data on ATJ. The upper bound of the data is used to detemine r the reaction rate for this analysis. The upper bound is 2.3 times higher than j the mean curve through the data and is shown in Figure 4. t The concentration of oxidants in the primary coolant is set at th'e Tech Spec [ l j limit of 10 ppm total oxidants. Chemical equilibrium of the coolant impurities is determined with the code GOP II, Table II gives the results of this equili-j brium analysis. Hydrogen is determined to be at 11.4 vppm from a different equilibrium analysis where radiolytic reactions are accounted for. This value of l I hydrogen is lower from this analysis giving a more conservative result since hydrogen is an jnhibitor in the steam graphite reaction rate of ATJ graphite. { I The results of the oxidation analysis show that structurally significant burn- ,[ l off is confined to a 0.22 inch thick layer on the surface of the post and the oxidation profile is very steep. Figure 5 shows the oxidation profile'in the t f post. i It has been shown that very little mechanical property changes occur up to l' burnoff in ATJ. Thus, depth of structurally significant attack has been defined as the depth where the density reduction is 0.1%. All material with a reduction of 0.1% or less in density is assumed to be unaffected and possesses its original mechanical properties. This assumption is based of experimental data. In the structural analysis, the surface oxidation of the CSP is accounted for by reducing the elastic modulus from the virgin material to the outside surface. 1 The oxidized surface is divided into annular layers of material. The average burnoff across the annuli is determined from the burnoff profile. Then, the appropriate elastic modulus is assigned to each annulus according to the average burnoff of 'each layer. The multimaterial cross section of the CSP is trans-formed into an equivalent cross section of virgin material such that force t l equilibrium and strain compatibility are satisfied. This analytical technique is known as.the Method of Transformed Sections connonly used for beams of two i or more materials. j i The results of the stress analysis show that at the end of life, the posts under [ the active core have a minimum safety factor of 7.46. While the posts under the PSR have a minimum safety factor of 4.31. The FSV-FSAR requires a safety factor of 3. r I

b a;r., 71-l { If all material with a density reduction of 0.1 wt. % or greater is assumed to be removed, the safety factor is still 6.82 for the posts under the active core and 3.95 for the posts under the PSR. It is concluded that the posts are tolerant to the effects of oxidation and will maintain their structural integrity throughout the life of the plant. O e b ( i y e e. S 6 b e G G I i h ~' e d L I l r

~ s TABLE I I

BASIS OF CORE SUPPORT BLOCK, ilEL10M TEMPERATURES l

1. Column power obtained from fuel management calculations where Tech. Specs. are adhered to. - . Cycle repea'ted between Reload 2 arid 7 beyond Reload 7. ~ 2. Operator control band' is Tech. Spec. limit of 200*F above the a'vorage region on'tlet temperature. - 1 3.' The region average temperature is set at 50*F above average to account for hot sector. .4. Region outlet thermocouple unc'ertainty is. assumed to have a systematic erro'r of + 25*F and a normally distributed random error with a 2 o of + 25'F. '5. It is assumed that the core inlet and exit temperatures will vary linearly from beginning-of. 4 life to end-of-life conditions caused by a drop in steam generator efficiency. ~ BOL' EOL - ~ ~ Core inlet 716F 724F. Core outlet 141gF ,1470F + ~ Gas. Temperature Increase [ .Across Core 703'F . 754*F, ~ t 6. As built steam generator characteristics'and planned use of attemperation'at-full power are used. ~

7. Temperatu're excursions resulting* fron transients are modeled as + 40*F for 5% of life and + 15*F'

.i for 15% of 11fe. e

8. ' Monte Carlo techniques are used to combine the statistically varying parameters to produce gas temperature distributions.

t a s.

i~n. . s 73 6 l-1 ( TABLE II Primary Coolant Equilibrium Impurity Concentration HO 1.2 vppa 2 CO 6.3 vppa CO2 2.5 vppm t ~ TOTAL 10.0 vppm O H2 11 4 VPm e e e. ~. e e 9 4 9 e ks e e O e 4 6 9 / s v. 9 e h e t G e O e

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. =..

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I..

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s.

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g s. ...__...a .7s. 3 -} l. i l t BEARING BETWEEN... - . N B LOCK 4.S E AT l ., -J'. c v 'uppgg l . SEAT s __ - ' BEAM COLU MN p . BENDING L e 'A M=Ps. . M. - m I f. C RITIC AL. e& E C L ON.' - l' L e I s g.. ~ g ~ -- + . --F. .r.. g L - a .j ,I .r t = t r

q. p -

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A w w - =. -.. .o j 1B00 --- ( 1700 / W

ag g..

/ 1600 w !h W u, .e s t. 1500 -- 4 t 1400 o 0.0 0.2 0.4 0.6 0.8 1.0 t FRACTION OF LIFE AT TDIPERATURE LESS TRAN T CORE SLTPORT BLOCK EXIT Pb;E TD'2ERATURE H,ISTORY JIG. 3. FSV USED IN CORE SUPPORT POST _DXIDATIm' AALYSIS %g c L

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.S. =- J-s.=- - l. : - - - L- =.t-z-l =.m -R = - = =- =+- - h - = -~ m= - ~=n+

.=~- M VEWC=_-k-ir -i?=~i =4=K== ' \\ EE=-t-i2i:=1=- -E < _ % + i.=1=-i'=' E W K 4==-) 1 s __ =_. _. _; ; _... a :.=.:_== _= u =_=._.=_:.:_ = =_-. l.._=_ =.== \\ =_..l _._. _ = cr_ :.=: I ----- : __ _.u... - u= .-- =_ :c = __-_-._.__i!* i_.. m e e ..1. .._{ 4 -_9,. .g .._y _.-4 -p. e l -W4x ' l'. x O.1 0.2. [ Distance from surface of Post (Inch) FIC. 5. TSV % 6" DIA. CORE SUPPORT POST OXIDATION PROFILE e

u s

)

i RESPONSE TO QUESTION #13 4 QUESTION: Include means for minimizing and monitoring oxidation levels in the structural graphites. ~ ANSWER: Incorporated in the facility are design features and procedural controls that I limit the amount of impurities and contaminants that may enter the primary . coolant. The plant purification system is designed to remove inleakage of. water, carbon dioxide, carbon monoxide and hydrogen. As stated in the response to Question #3, the monitoring program will be employed to follow a.nd record the actual operating level of oxidants. The moisture monitors are used to measure moisture in the core. A standard gas chromatograph is used to monitor concentrations of carbon dioxide, carbon monoxide and hydrogen gas in the primary coolant. e ( j ti t e 6, O e

m, e. e. f -80. RESPONSE TO NRC 00ESTION #14 QUESTION: Discuss procedures to be initiated in the event the surveillance specimens show an abnormal rate of oxidation. ANSWER: As' stated in the answer to Question #7, if the specimens removed. indicate a burnoff,significantly greater than predicted, the specimen removal schedule ( .ould be accelerated and the expected CSB lifetime re-evaluated. w e e em e e b e I ? I \\ t e e l 1 OG e O. g a c t r w

__.2._ t I*s a . RESPONSE TO NRC QUESTION 815 QUESTION: Discuss a comnitment to compare th' computer code CSBB0 results with oxidation e measurements on the surveillance specimens. ANSWER: The CSSBC computer code will be used to correlate measuremente cf burnoff pro-files in the surveillan~e program specimens. This will be done by maintaining c day-to-day records of moisture exposure history and core outlet temperatures in those regions with surveillance specimens. These data records will be input to the CSBB0 code to calculate the burnoff history..The gas tamperature at the specimen location must be calculated from the' region outlet ~ temperature since considbrable mixing occurs between the specimen and cort: outlet temperature sensor. Flow data will also be required to calculate cne drop in oxidant concentraticn across the boundary layer. By comparing rea-sured burnoff' profiles with the computer predictions, adjustments will be made in PGX reaction rate and steam diffusion coefficient until the measured pro-files are matched. If adjustment to match the burnoff profiles is in the direction to decrease reaction rates, confidence will be gained in the conser-vatism of the predictions. If adjustment is required in the other direction, the predicted CSB lifetime will require re-evaluation. t I t OG .:}}

ML20062B537

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