Studies of thermal-hydraulic phenomena in EBR-II

Nuclear Engineering and Design 62 (1980) 219-232 © North-Holland Publishing Company

STUDIES OF THERMAL-HYDRAULIC PHENOMENA IN EBR-II * Ralph M. SINGER, Paul R. BETTEN, Eric M. DEAN, Jerry L. GILLETTE, Dale MOHR, John E. SULLIVAN and John V. TOKAR EBR-H Project, Argonne National Laboratory, Argonne, Illinois and Idaho Falls, Idaho, U.S.A.

Received 10 July 1979

An experimental and theoretical program has been undertaken during the past several years with the objective of developing a well-documented understanding of steady-state and transient thermal-hydraulic behavior in EBR-II. The results of this effort have provided reactor designers and system modelers with needed integral-type demonstrations of important phenomena. This paper will discuss the particular problems of steady-state and transient hot channel peaking factors and plant operational characteristics impact upon natural circulation dynamics. Direct in-core experimental measurements have demonstrated that factors used for the prediction of peak coolant temperature rises at normal rated plant conditions may not be conservative due to pin-bundle distortions or inlet flow maldistributions, while those applied during loss-of-flow transients are most likely overconservative due to inter- and intrasubassembly phenomena. The importance of somewhat controllable parameters such as the sequence of primary and secondary pump trips and reactor scram, primary pump rundown times, and nominal operational power-to-flow ratio upon the dynamics of the transition from forced to natural convective flow are also presented.

1. Introduction

to thermal bowing, wire-wrap movements, bundle-duct mechanical interactions, inlet flow maldistributions, or combinations of these mechanisms. In any case, there is now some question as to the adequacy of standard hot-channel analyses to conservatively account for this apparently increased uncertainty in local peak coolant temperatures, On the other hand, local and core-wide temperature measurements made over a wide range of operating power and flow rate levels have conclusively demonstrated that local temperature peaking factors markedly drop as the coolant flow rate decreases. This result has been shown to be valid during both steadystate reduced flow operation as well as during loss-offl0w transients. The phenomena responsible for this favorable result are the in-bundle and core-wide flow redistribution resulting from frictional and buoyancy variations, and transverse, intra- and intersubassembly heat transfer, both of which become increasingly important as the flow rate decreases. The response of a fueled subassembly, the reactor core, and the whole plant to various natural circulation events were also studied. These efforts included the investigation of loss of forced flow events from a fuel

During the past several years, an extensive thermalhydraulics program has been conducted utilizing the only operational LMFBR power plant in the United States, the Experimental Breeder Reactor No. II (EBR-II). The experimental portion of this program has generated steady-state and transient data obtained from in-core and plant sensors, while the analytical portion has resulted in the development of new computational tools and the testing of existing codes. This paper will present a summary of some of the most important conclusions which have resulted from these studies. Measurements of local transverse and axial coolant temperature profiles within an operating, fueled pin bundle have demonstrated consistent non-symmetries. Substantial discrepancies between local temperature profiles predicted by standard rod-bundle thermal-hydraulic codes and corresponding experimental measurements have been obtained. These discrepancies are apparently caused by either pin-bundle distortions due * This work was supported by the US Department of Energy. 219

220

R.M. Singer et al. / Thermal-hydraulic phenomena

handling configuration (low decay power and auxiliary forced flow), from a complete loss-of-site electrical power incident, and from various operational sequences which ultimately lead to natural convective decay heat removal. Among the many conclusions obtained from this work were the observations that during a natural circulation event, it is not possible to predict the in-core thermal-hydraulic transient unless the whole plant response is modeled due to strong system coupling, that peak in-core temperatures would be substantially overestimated unless intra- and intersubassembly flow redistribution and heat transfer were taken into account, that piping heat capacity and heat losses significantly affect component temperature transients, and that flow reversal, stagnation, or chugging may occur in certain types of subassemblies or piping loop configurations during particular types of transients. Since it is not possible to adequately discuss all of these points in this paper, only several of the more important conclusions will be presented.

NoK

PRIMARY PUMP.S(2)

SHUTDOWN COOLER (2) SECONDARY SODIUM SYSTEM

. . . . .

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Fig. 1. Schematic diagram of the EBR-II primary heat transport system.

2. Description of reactor and instrumentation

Since the EBR-II reactor and instruinentation have been extensively reported elsewhere in the literature, e.g., [1,2,3], only a very brief description will be presented here. The reactor normally operates at 60.0 MWt at a coolant flow rate of 0.516 m3/s and contains a complete nuclear-steam-electric system. A simplified schematic drawing of the EBR-II primary heat transport circuits is shown in fig. l. From this figure, it is apparent that EBR-II is of the pool design, with coolant being drawn into the two primary pumps directly from the primary vessel, separated into high pressure and low pressure zones for core and blanket cooling purposes, discharged into a common outlet plenum, piped to a single intermediate heat exchanger (1HX), and finally dumped back into the primary vessel. Temperature and flow rate sensors are located at appropriate locations along the flow paths. In addition to the normal plant sensors, special subassemblies containing extensive instrumentation are usually located in-core. The particular subassembly used for much of the work reported in this paper was designated XX08 and is fully described in ref. [4]. A simplified sketch of this probe, which contains 58 fuel and 3 dummy pins, is shown in fig. 2 where the essen-

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221

R.M. Singer et al. / Thermal-hydraulic phenomena

tial arrangement of sensors are indicated. The abbreviations BTC, 4TC, 7TC, TTC, 15TC, and OTC refer to coolant thermocouple locations at core bottom, 0.4 core height, 0.7 core height, top of core, 1.5 core height, and outlet. Fuel centerline thermocouples are also located at the TTC position. The XX08 flowmeters were calibrated in a special out-of-pile sodium loop and have an accuracy of about +-1% over almost the entire flow range. The thermocouples are periodically calibrated in-place by high flow rate, isothermal operation of the reactor and have an estimated accuracy of +-0.5 °C.

3. Subassembly and core temperature distributions (steady state) 3.1. Full power and flow conditions

In order to either design or to ascertain the adequacy of an existing core restraint system, to calculate structural reactivity feedbacks during an ascent to power, to design an optimum subassembly orificing scheme, to determine the proper thermal input to fuel lifetime codes, to estimate local hot-spot factors, and other related thermal problems, it is necessary to predict coolant temperature distributions both within individual subassemblies as well as over the entire reactor core and blanket regions. A number of computational tools have been developed with varying levels of sophistication depending upon the specific application. In the particular case of the calculation of the thermal-hydraulic performance of a fueled subassembly with the objective of determining the operating margin to previously specified design limits and requirements, the usual approach is to apply a rodbundle subchannel mixing code with appropriate uncertainty factors stiperimposed. An explicit assumption in all of the presently available subchannel codes is that the fuel pins and wire wraps are uniformly spaced within the hex-can wrapper; intra-assembly flow and mixing maldistributions caused by modest variations in subchannel flow areas are accounted for by the use of suitably chosen hot-spot factors. To date, hot-spot factors of this type have been determined by comparing subchannel code predictions with measurements obtained from out-of-pile electrically heated experimental rod bundles. In this section,

equivalent comparisons will be made with actual inreactor measurements from an instrumented fuel subassembly. The experimental data used are from coolant thermocouples located within the XX08 experimental subassembly described previously, and the subchannel codes used are COBRA-IV [5], THI-3D [6], and CLUSTER [7]. The input to all three codes was prepared from the nominal dimensional data available for the XX08 subassembly [4], the standard EBR-II Project physics codes [8], and the experimentally measured coolant flow rate. The boundary conditions used in these calculations (as set by the particular code limitations) were adiabatic hex-can wrapper (COBRAIV), heat transfer to thimble bypass flow (THI-3D), and heat transfer to thimble bypass and six neighbor subassemblies (CLUSTER). The results of these calculations and comparisons to the measured coolant temperatures are shown in figs. 3 and 4, where the temperature rises are measured from the bottom of the fuel column to the top of core position. The inlet temperature is 371 °C. In all cases examined, the codes predict the usual approximately symmetric temperature profile, while the experimental measurements indicate a significantly skewed profile, with the peak temperature displaced from the subassembly center. This skewed profile was observed immediately upon reaching full reactor power following insertion of the instrumented subassembly and did not vary during the approximately two years residence in-core, except due to the gradual

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R.M. S&ger et aL / Thermal-hydraulic phenomena

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power reduction accompanying burnup of fissile material. Since the predicted mixed-mean overall temperature rises in XX08 agreed to within 2% of the measured value, while the local calculated and measured temperature rises at core top differ by about 15%, it appears that only the details of the intra-assembly phenomena modeled differ from that actually existing rather than the overall bundle behavior. Quite similar behavior has been observed in electrically heated pin tests [9] and in an instrumented subassembly in the Rapsodie reactor [10]. Several postulated mechanisms have been examined which could offer an explanation for this result, the most intuitively satisfying one being an 'abnormal' subchannel flow distribution caused by an inlet flow maldistribution or an internal bundle distortion, or both. The former effect could be caused by the design of below-core components (in XX08, an 88 mm long diffuser with an opening angle of 15.3 ° was used), or the latter by differential thermal bowing of the pin bundle and hex-wrapper causing an off-nominal subchannel flow area distribution. As a qualitative test of the effect of subchannel flow area distributions upon the transverse temperature profile in XX08, the COBRA-IV code was used to 'fit' the measured data by arbitrarily adjusting interior and edge channel flow areas while maintaining a constant total subassembly flow area. The results are given in fig. 3 where a maximum area change of 25% was used; this magnitude of change does not seem likely unless the wire-wraps are displaced from their

original locations. These results are encouraging in the sense that an a posteriori calculation can result in reasonably good agreement with the measurements, but in no sense resolve the problem. More data on different size subassemblies in various thermal-hydrualic environments are needed before any substantiable conclusions can be reached.

3.2. Reduced f l o w conditions As the coolant flow rate in a heated system is reduced, tile importance of buoyant forces increases relative to the other system hydraulic forces. The effect of buoyancy in a multi-channel flow system, such as exists in a reactor, is to increase the flow rates in channels which tend to be hotter than the average and to decrease flow rates in cooler than average channels. This flow redistribution directly acts to reduce the transverse temperature profile in a reactor as well as within individual subassemblies. Transitions from turbulent to laminar flow, which can occur at different flow rates in different subassemblies, will also cause a redistribution of flow from that existing under fully turbulent conditions. In addition, the increased transit time of coolant within subassemblies at low flow rates permits an increased propensity for interassembly heat transfer; this effect also directly reduces transverse temperature gradients across a reactor. Earlier reported data obtained from an in-reactor instrumented subassembly [3] indicated that these

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phenomena, which become of increasing importance as the coolant flow rate is decreased, can significantly reduce the local and core-wide transverse temperature gradients in a reactor. Additional data have been obtained since publication of that paper which has reinforced this conclusion. As an example of the flattening of the core-wide temperature distribution which results when the coolant flow rate is reduced from 100% to 5.7% of full value (i.e., from a fuel pin bundle Reynolds' number of 34 000 to 1938), the temperature rises measured at the outlet of a number of instrumented subassemblies in EBR-II are plotted in fig. 6. The quite large scatter in full power and flow temperature rises, which is primarily due to the wide variety of experimental, driver, reflector, and blanket subassemblies in EBR-II, is seen to be markedly reduced at the lower flow rate. A measure of the flattening is the 20 peak channel factor for subassemblies 1, 3'SOT, which drops from 1.476 to 1.296. Similar data under steady-state natural circulation conditions are shown in fig. 7 for a variety of power-to-flow ratios where an equivalent flattening is observed. Similar behavior is observed within an individual subassembly where the buoyancy forces generated by transverse temperature gradients cause a redistribution of subchannel flow. An interesting example of this result is shown in fig. 8 where the top-of-core temperature profile in the XX08 instrumented subassembly is shown for cases of primarily forced convection and t Defined as 3' = ( A T a v g +

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totally natural convection. The mean temperature rise in both cases is almost identical, so that the power-toflow ratio, and hence the Grashof-to-Reynolds' number ratio are also identical. The forced convection temperature gradient is seen to be significantly larger than that under natural convection, and the resulting 2o temperature peaking factor is 1.189 as opposed to just 1.072 under natural convection. Again, with a flow completely buoyancy driven, temperature gradients tend to be destroyed relative to equivalent forced flow situations. 30

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224

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where the flow rate, Q, is defined as the ratio of the actual flow rate to the full flow (100%) value, r 2 the correlation coefficient, and "/TOT the product of 7SOT and ')'TTC. These results imply that if the peaking factor obtained at full flow conditions (TWOf = 1.960) is used to predict the peak coolant temperature rise (within 20) at a lower flow rate, e.g., at Q = 0.05 where "/TOT = 1.485, an overprediction of about 32% will result. These results demonstrate the importance of intraand interassembly phenomena to the determination of local peaking factors under steady-state reduced flow conditions. However, it remains to be shown whether these phenomena have any relevance to the potentially more important transient events which involve the toss of reactor coolant flow. This will be the subject of the following section.

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An integral part of an LMFBR design is the provision of some form of emergency pumping which can be relied upon to continue reactor cooling in the event of a loss of electrical power. These emergency systems usually can supply about 5 to 10% of the full rated flow of the plant and are designed to operate with an extremely high reliability since this event must be assumed to occur a number of times during tile operating life of the plant. In the EBR-II design, a primary systenr auxiliary pump, whose power is normally supplied by the site power system and is backed up with a diesel generator as well as storage batteries, can supply about 5 to 6% of rated flow. During a particular loss of off-site electrical power event, transient temperature data obtained from the XX07 instrumented subassembly [ 11] were recorded. This event consisted of an initial loss of electrical power to the primary pumps, followed by a reactor

R.M. Singer et al. / Thermal-hydraulic phenomena scram automatically triggered on a low-flow signal. The auxiliary pump was in continuous operation during this time. In order to compare the dynamic response of an average coolant subchannel in a driver fuel subassembly to a hot subchannel, two instrumented locations in XX07 were studied, both of which were at the top of the core. The first location corresponded to a subchannel whose temperature tise almost exactly matched that of the total subassembly rise, i.e., could be considered an 'average' subchannel. Thus, this subchannel temperature rise corresponded to that calculated from the total subassembly power and flow rate at steady-state. Using these average subassembly values during the transient, the average temperature as a function of time was computed using an EBR-II simulation code, EROS [12], and compared to the measured value in fig. 11. Excellent agreement is noted throughout the transient. Now, a particular subchannel whose steady-state temperature rise was greater than the average was also chosen. In this case, a peaking factor for this subchannel at full power and flow was chosen so that the initial predicted temperature rise agreed with the measured value (i.e., a factor of 1.43 on the average temperature rise was used). This peaking factor was used throughout the transient to calculate the response at this location; a comparison of this predicted value to that measured is shown in fig. 12. A significant (about 31%) overprediction of the peak transient temperature at this 'hot' subchannel location is apparent. Clearly, the full power and flow temperature distributions

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markedly flatten out on time scales significant to lossof-flow events so as to reduce local peak temperatures.

4.2. Loss-of-flow to natural circulation In the extremely unlikely event that during a loss of site power event an emergency pumping system does not operate upon demand, it is necessary to rely upon natural circulation of the coolant to remove the reactor decay heat. Initial testing of the EBR-II heat transport systems has been completed; one particular transient will be described here with the objective of demonstrating the reduction in peak channel factors during such natural circulation events. This test was initiated from a power of 28.5% and a flow of 32.1% of normal, corresponding to an overall reactor temperature rise of 90 °C. The primary system auxiliary pump power supply was disconnected, and the plant operated at steady-state for approximately 3 hours. The transient was caused by an interruption of the power supply to the primary pumps, resulting in an immediate flow coastdown and an automatic reactor scram within about 1.8 s on a low-flow signal. Coolant temperatures at the outlet of each of the instrumented subassemblies and fuel and coolant temperatures and flow rates within the XX08 subassembly were continuously monitored. The flow rate reached a minimum value of about 2.7% within 45 s of the pump trip, and then increased due to buoyancy forces. In order to provide a graphic illustration of the

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R.M. Singer et al. / Thermal-hydraulic phenomena

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collapse of the initial at-power transverse temperature gradient during this loss-of-flow/natural circulation transient, all of the measured top-of-core coolant temperatures in XX08 are superimposed on fig. 13. Prior

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R.M. Singer et al. / Thermal-hydraulic phenomena the peak transient temperature is reached in XX08, the spread from maximum to minimum is only about 9 °C, or in other words, the 2o peaking factor dropped to 1.114. The steady-state peaking factor for the flow rate at this condition (from fig. 10) is 1.13. A similar plot of the subassembly outlet coolant temperatures in the core region is shown in fig. 14 where the transverse temperature gradient is also shown to drop. For these particular conditions, the initial at-power 20 total peaking factor (7TOT) was measured to be 1.419 × 1.330 = 1.887;however, at the time the transient peak temperature was measured at core-top in XX08, the equivalent value was 1.395. Therefore, if the peak transient temperature in XX08 was calculated by using the at-power peaking factor of 1.887 in conjunction with the predicted behavior of an average subchannel in an average subassembly, the calculations would overpredict the measured value by 35%.

4. 3. Summary o f measured transient peaking factors The two examples described in the preceding sections served to illustrate the observation that the significant flattening of transverse temperature profiles

227

measured during low-flow, steady-state conditions in EBR-II also occurs during loss-of-flow transients. Several additional loss-of-flow transients comprising situations where the auxiliary pump was either on or off were also similarly analyzed to determine the experimental transient 2o peaking factors. These results are summarized in table 1. The results shown in this table, which are representative of those obtained from many other transients, demonstrate the decreased value of the coolant peaking factor during a loss-of-flow event. The transients which involved continued operation of the auxiliary pump and therefore minimum flow rates of about 5 to 6% resulted in peak transient top-of-core temperature rises of about 53% over that in an average subchannel, while the corresponding initial at-power peak was about 97%. Similarly, during the natural convection tests, the peak transient top-of-core coolant temperature rises were about 38% above an average subchannel, while the corresponding initial value was about 87%. These measurements seem to consistently demonstrate that the use of peaking factors on coolant temperature rises developed from full flow and power conditions will overpredict the peak local transient

Table 1 Summary of transient peaking factors on coolant temperature rise during loss of flow events in EBR-II Initial conditions Data source

Rundown time a (s)

Flow (%)

Power (%)

7SOT

7TTC

LOF/Scram of 10/01/77 (aux. pump on)

100

100

1.357

1.415

1.920

LOF/Scram of 1/10/78 (aux. pump on)

100

100

1.440

1.410

LOF/Scram of 1/01/75 (aux. pump on)

100

100

1.380

Transient coolant peaking factors

7TOT "/'SOT

7TTC

b ")'TOT

23

1.203

1.205

1.557

2.030

31

1.281

1.193

1.544

1.429

1.972

52

1.271

1.108

1.498

XX08 Test 7A (nat. circ.)

32.1

28.5

1.330

1.419

1.887

36

1.183

1.114

1.395

XX08 Test 8 (nat. circ.)

36

40

1.326

1.398

1.854

36

1.150

1.120

1.362

a Defined as the time required for the primary coolant flow to drop to 10% of the initial value. b Measured at time of peak top-of-core coolant temperature.

228

R.M. Singer et al. / Thermal.hydraulic phenomena

coolant core temperature tise during loss-of-flow events by about 30%. Since these peaking factors are mechanistic as opposed to statistical or random, they can be predicted a priori from first principles. To do so, however, requires the proper modeling of buoyancy-induced flow redistribution both within and among subassemblies in a thermal-hydraulics code. This directly implies that it is necessary to model a significant fraction of the subassemblies in a reactor in order to accomplish this objective. It appears that a reduction in the calculated hot-spot temperature rise during a protected loss-of-flow event (especially to natural circulation) .of about 30% is well worth this effort.

5. Reactor response to natural circulation transients

Prior to this point, the emphasis has been upon EBR-II experiences in determining realistic peaking factors for steady-state and transient coolant temperature rises. The remainder of this paper will be devoted to a discussion of the experience gained in conducting and analyzing whole plant natural circulation events. Since a recent publication [ 13 ] reported uponnatural circulation testing in EBR-II which addressed conditions of relevance to the plant conditions preparatory to fuel handling, only events initiated from significant fission power levels will be discussed here. Furthermore, since most of the tests and analyses have been directed toward power and flow levels below the normal full ratings of EBR-II, the following material will be similarly restricted. This should not have an undue effect on its relevance since, in all cases, fully prototypical ranges of plant thermal conditions were employed.

energized. The transient is caused by a simullaneous trip of the primary and secondary pumps and a manual scram of the reactor by removing $ 4.00 of reactivity from the core, resulting in a transition from forced to natural convective flow cooling. The reaclor plant response was calculated using the primary system module of the NATDEMO code [ 14] which has been developed to describe the system dynamics of EBR-II and can be used for related plants. Comparisons between experiments and predictions of this code have been reported earlier [14,15]. The calculated thermal and hydraulic response to this event is summarized in figs. 15 and 16. The primary coolant flow rate through an average driver subassembly is seen to drop to a minimum value of about 1.3% of its rated value, at which time the instantaneous power-to-flow ratio rises to about 1.7. This increase in power-to-flow ratio causes an increase in coolant temperature which directly increases tire buoyancy of the coolant and thus accelerates it. Therefore, the flow rate is observed to steadily increase, and within about 65 s of the start of the transient, the instantaneous power-to-flow ratio in the driver drops to 1.0, and steadily decreases thereafter. The thermal response of an average driver to this transient is illustrated in fig. 16, where tire coolant temperatur e at core top and subassembly outlet locations are shown (the inlet temperature is constant at

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The basic purpose of this section will be the illustration of the importance of various plant operational characteristics upon its dynamic response to a complete loss-of-forced-flow event with reactor scram. [n order to provide a base from which the effects of these characteristics can be measured, a reference transient will be defined. This reference event is initiated from steady plant operation at 34% power and 34% flow with the auxiliary pump de-

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R.M. Singer et al. / Thermal-hydraulic phenomena

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371 °C). The rapid initial drop in temperature corresponds to the initially more rapid decrease in power relative to that of the flow. As the rate of power decrease slows down and the flow continues to drop, the coolant temperature increases. This temperature increase continues until sufficient buoyancy is developed to cause the natural convective flow to increase sufficiently to remove the generated heat. For this base case, the maximum coolant temperature in an average driver (at core top) reaches 500°C at 52 s; the temperature at the outlet of this subassembly peaks at about 466 °C, but is delayed by an additional 68 s due to transport lags. With this background, it is now possible to discuss the influence of various system parameters upon the reactor thermal-hydraulic dynamics, and specifically the peak in-core temperature.

5. 2. Effects o f initial po wer and flow levels

40% results in an increase in peak transient temperatures from 486 °C to 516, or stated differently, a 33% increase in power level results in a 26% increase in peak transient temperatures. Variation in the initial power-to-flow ratio (accomplished by fixing the flow rate at 34% and varying the power level) are expected to affect the system response since the initial temperature level and distribution, and hence buoyancy, in the primary heat transport circuits are changed. These results are shown in fig. 18.

5.3. b~ffeet o f primary pump randown times The power to the motor-generator sets driving the EBR-II primary pumps can be interrupted by several i

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flow and result in a somewhat higher peak temperature. This effect is also noticeable m fig. 19 by observing that the mode Ill pump trip results in the smallest value of the minimum flow of any of those studied. Mode 1 results in the fastest initial flow rate decrease so that although the thermal heads are not as diminished as in tile other modes, the minimum flow occurs at a somewhat earlier time, and therefore at a higher power, resulting in a higher maximum temperature.

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different modes, each one of which results in a different pump run-down time. Thus, the length of time and the level at which forced flow is maintained can be adjusted by the procedure with which the power to the primary pumps is interrupted. Tile thre.e modes examined here resulted in the flow coastdown characteristics shown in fig. 19, and the corresponding peak temperatures are shown in table 2. It is apparent that the flow rates during the initial critical 100 s differ substantially depending upon the mode of primary pump trip. Both the initial rate of flow decrease as well as the ultimate minimum flow rate reached are affected by the technique used. An examination o f the flow transients in fig. 19 indicates that although the mode III trip maintains the forced flow at a higher value initially than by mode I1, this increased flow tends to sufficiently reduce the thermal heads in the reactor to adversely affect the subsequent transition to natural convective

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5.4. Effect o f sequence o f primary~secondary pump trips and reactor scram The sequence of reactor scram and primary pump trip also plays an important role in determining the maximum transient coolant temperatures. If the reactor is scrammed first, and the pumps tripped later, the continued forced flow following the reduction in power will reduce the thermal heads in the primary heat transport system. This will delay the ultimate establishment of natural convective flow since higher in-core temperatures are required to provide the necessary buoyancy. Conversely, if the pumps are initially tripped, and the reactor scrammed later, the coolant temperature in the core will initially increase, providing an increase in the initial buoyancy, and resulting in lower transient temperatures during the transition to natural convective flow. These results are summarized in fig. 20, where both the first and second temperature peaks 2 are shown as a function of the delay of the trip of the primary pumps after reactor scram. From this figure, it appears that the minimum transient temperatures result when the pumps are tripped first, and the reactor scrammed 2 to 4 s later. For these conditions, neither first or second peak transient temperatures exceed 490 °C. In the secondary system, the only parameters which influence the maximum in-core coolant temperature in this type of transient are the timing of the secondary pump trip, the asymptotic natural convective flow rate in the secondary system, and the temperature of the secondary coolant as it enters the IHX. The importance of the first of these parameters 2 The first peak is caused by the delay in scram following a loss-of-flow, while the second is caused by the later powerto-flow ratio mismatch.

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is shown in fig. 21. The base case transient assumed that the primary and secondary pumps were simultaneously tripped; here, a delay in the secondary pump trip is allowed. The results in fig. 21 illustrate that the maintenance of secondary flow is an extremely important factor in reducing the core coolant temperatures. For example, if the secondary flow can be held at its initial value of 34% for 30 s by some sort of emergency power system, the maximum core temperature is reduced by 30 °C, a reduction in temperature rise of about 23%. The continued strong cooling of the primary coolant in the IHX by a high secondary flow during the transient provides sufficient negative buoyancy (which is favorable in the down-flowing primary

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The preceding analyses investigated the importance of several operational parameters upon the dynamic response of an LMFBR of EBR-II design to a complete loss-of-forced-flow event. A basic assumption in this study was that the basic plant design was fixed, e.g., relative heights of components could not be varied, or flow circuit layouts adjusted. Thus, only non-structural-type design variations were permitted. Several of the more interesting observations obtained from this effort related to pump run-down times and pump tripreactor scram sequences. Although it might intuitively appear advantageous to design primary pumps with large inertial loadings (such as with fly-wheels) to protect a plant during lossof-flow events, in some plant designs this action might actually result in somewhat higher transient core temperatures if the ultimate coolant flow is buoyancy driven. This results because of the initial cooldown of the heat transport circuits and the resulting increased difficulty in accelerating the more dense fluid. Of course, the combination of primary and secondary pump rundown times must also be examined as was pointed out in [16] ;the present study used a fixed secondary pump rundown. The timing of the reactor scram relative to the start of the primary pump rundown is also a critical parameter. The usual response of a plant to a loss-ofsite power is a flow reduction simultaneous with or followed by reactor scram. This is an advantageous sequence in terms of natural circulation due to the initial small heat-up of the primary sodium and a resultant increase in buoyancy. If the reactor scram occurs first, followed somewhat later by a loss-offlow, the resulting peak core temperatures may actually be larger than the latter case due to the initial cooldown of the primary coolant. The effect of a short continuation of secondary flow following a loss-of-power event was also shown to be quite important. The addition of an emergency power supply of relatively small capacity to the secondary pumps could significantly increase the transient decay heat removal capability of a plant.

232

R.M. Singer et al. / Thermal-hydraulic phenomena

6. Summary and conclusions Direct experimental measurements of local in-core and core-wide coolant temperatures over a w i d e variety o f power and flow conditions in EBR-II have been obtained and analyzed. The resulting conclusions from this study included the identification of the existence o f a stable skewed temperature profile within a fueled (61-pin) subassembly which can not be explained by the use of standard rod bundle the> real-hydraulic codes unless a distortion of the subchannel flow areas from nominal conditions is assumed. This observation, which when coupled with several.recent similar results in out-of-pile and inreactor tests, requires a reassessment of the conservatism now embodied in currently accepted hot channel analyses at rated power and flow conditions in an operating reactor. During low flow rate events, both at steady conditions as well as during transients, extensive evidence is presented demonstrating the flattening of transverse temperature profiles resulting from buoyancy-induced flow redistribution and other inter- and intrasubassembly phenomena. Experiments have shown that a reduction in the predicted hot channel temperature rise of about 30% is possible if these phenomena are properly modeled in reactor thermal-hydraulic codes. A study o f the effect of a number of operationally controllable parameters (as opposed to original design parameters) upon the dynamics of a transition from fission power, forced flow conditions to decay power, natural convection conditions in an LMFBR has been conducted. The essential conclusions of this effort were that the precise sequential timing of the primary and secondary pump trips (as well as rundown times) with reactor scram can have a significant effect upon the peak in-core temperature. For example, the continuation of high levels of primary flow for short times following a reactor scram may be detrimental, while a similar continuation of secondary flow can be beneficial.

Acknowledgement These data could not have been successfully obtained without the direct technical support of many EBR-II personnel. We would like to especially acknow-

ledge Roland Smith, William Perry, Gary Lentz, and all of the reactor operating crews for their eflk~rts related to operational problems, William Booty for his work in developing the data acquisition programs, and Nadine Allen and Lavar Welker for their work in procedures and scheduling.

References [ 1 ] L.J. Koch et al., Argonne National Laboratory, report no. ANL-5719 (May 1957) and ANL-5719 (Addendum) (June 1962). [2] R.M. Singer et al., Nucl. Sci. Engrg. 63 (1977) 78 -82. [3] R.M. Singer and J.L. Gillette, Chem. Engrg. Prog. Symp. 73 (164) (1976) 97-104. [4] A. Smaardyk et al., Argonne National Laboratory, report no. ANL-78-9 (1978). [5] C.L. Wheeler et al., BNWL-1962, UC-2 (March 1976). [6] W.T. Sha and R.C. Schmitt, Argonne National Laboratory, report no. ANL-8112 (December 1975). [7] L.K. Chang, Nucl. Engrg. Des. 42 (1977) 223-231. [8] Internal Project Codes, unpublished. [9] J.F. Dearing, Thermal-hydraulic effects of pin bowing in THORS bundle 3C, presented at the Fourth Annual Meeting of the Core Thermal-Hydraulics Technical Group, Argonne National Laboratory, April 1979. [ 10] D. Leteirturier and L. Cartier, Theoretical and experimental investigations of the thermohydraulics of deformed wire-wrapped bundles in nominal flow conditions, presented at the IAEA Specialists' Meeting on Thermodynamics of LMFBR Fuel Subassemblies Under Nominal and Non-Nominal Operating Conditions, Kernforschungszentrum, Karlsruhe, Fed. Rep. Germany, February 1979. [l 1] J.L. Gillette et al., Argonne National Laboratory, report no. ANL-76-78 (1976). [121 J. Kaganove, E. Dean and A. Campise, Argonne National Laboratory, report no. ANL-7981 (1974). [13] J.L. Gillette et al., Experimental study of the transition from forced to natural circulation in EBR-II at low power and flow, ASME paper 79-HT-10, presented at the 18th National Heat Transfer Conference, San Diego, CA, August 1979. [14] D. Mohr and E. Feldman, A dynamic simulation of the EBR-II plant with the NATDEMO code, Proc. Specialists' Meeting on Decay Heat Removal and Natural Circulation in FBR's, Upton, NY, Feb. 28-29, 1980. [ 15 ] R.M. Singer et al., Studies related to emergency decay heat removal in EBR-II, Proc. Internat. Mtg. Fast Reactor Safety Technology, Seattle, Washington, August 1979, Vol. III, pp. 1590-1598. [16] M.E. Durham, Optimization of reactor design for natural circulation decay removal in a pool-type LMFBR, Proc. Internat. Conf. Optimization of Sodium-Cooled Fast Reactors, London, 1977, pp. 67-76.