EBR-II unprotected loss-of-heat-sink predictions and preliminary test results

Nuclear Engineering and Design 101 (1987) 57-66 North-Holland, Amsterdam

57

EBR-II U N P R O T E C T E D LOSS-OF-HEAT-SINK P R E D I C T I O N S A N D PRELIMINARY T E S T R E S U L T S * E.E. F E L D M A N , D. M O H R , L.K. C H A N G , H.P. P L A N C H O N , E.M. D E A N an d P.R. B E T T E N EBR-II Division, Argonne National LaboratotT, 9700 South Cass Avenue, Argonne, Illinois 60439, USA

Two unprotected (i.e., no scram or plant protection system action) loss-of-heat-sink transients were performed on the Experimental Breeder Reactor-II in the Spring of 1986. One was initiated from full power (60 MW) and the other from half power. The loss of heat sink was accomplished in each test by essentially stopping the secondary-loop sodium coolant flow. Pretest predictions along with preliminary test results demonstrate that the reactor shuts itself down in a benign and predictable manner in which all of the reactor temperatures approach a quenching (or smothering) temperature at which the fission power goes to zero.

1. Introduction A liquid metal reactor (LMR) and its balance of plant (BOP) should be a single well-integrated system designed so that malfunctions in the BOP do not lead to adverse consequences for the reactor. Among the malfunctions which must be considered are a loss of the power to the intermediate sodium-loop pumps, a loss of the intermediate sodium coolant itself, and a loss of the feedwater leading to a dryout of the steam generator. Each of these mishaps is a form of loss-of-heat-sink (LOHS) accident and creates a situation in which, through no fault of its own, the primary system is unable to transfer heat to the BOP via the intermediate heat exchanger (IHX). Consequently, there is a potential for the reactor core to overheat and cause severe fuel damage. The conventional mode of protection for an LOHS accident is the placement of plant sensors which detect the cause of the LOI-IS or the symptoms it produces. These sensors would be part of a plant protection system (PPS) which would initiate a reactor scram to promptly stop the fission process. The typical PPS is highly reliable and effective. It achieves its reliability by employing diverse and redundant systems, each of which must be of very high quality and independent of the others. This approach to safety is both complex and costly. Perhaps a better approach is to design an "inherently * Work supported by the U.S. Department of Energy, Reactor Systems, Development and Technology, under Contract W31-109-Eng-38.

safe" reactor in which natural phenomena, rather than an elaborate PPS, limit the reactor temperatures during an LOHS accident. In such a design the reactor (and plant) would be able to withstand a complete LOHS accident without a scram or an operator-initiated power reduction. In 1983 the French successfully performed an unprotected (i.e., no scram) LOHS test on their Rapsodie reactor plant from about 50% power, Essig et al. [1]. Two unprotected LOHS tests were successfully conducted on the Experimental Breeder Reactor II (EBRII). On March 29, 1986 an unprotected LOHS test was performed from half power (30 MW) and full primary flow. This transient was initiated by essentially stopping the intermediate-loop flow. Although no scram or protective action occurred, the reactor responded in a rather benign manner. On the following day, the same test was repeated, except that it was initiated from full (60 MW) power. Again a benign response was observed. On the afternoon of April 3rd the full power unprotected LOHS test was repeated for an international audience. For more than a decade the EBR-II Division of Argonne National Laboratory (ANL) has been engaged in natural convection and transient overpower testing as well as extensive analytical modeling of the EBR-II plant, Golden et al. [2]. This experience enables the nature of the unprotected LOHS transients to be analytically simulated with a high degree of confidence. The measured data from these LOHS tests is most conclusive, and in this paper we present pretest predictions for both tests along with some of the measured data. These data contribute to a data base which can be used in the development and validation of analytical

0 0 2 9 - 5 4 9 3 / 8 7 / $ 0 3 . 5 0 © Elsevier Science Publishers B.V. ( N o r t h - H o l l a n d Physics Publishing Division)

58

E.E. Feldrnan et al. / EBR-II unprotected loss-of-heat-sink predictions

as a fast-flux irradiation facility since 1967. The plant produces about 20 MW of electricity when operated at its full nameplate power of 62.5 MW. The corresponding reactor, secondary, and steam system flow rates are approximately 485, 315, and 32 kg/s, respectively. The EBR-II reactor, as shown in fig. 1, is submerged in the primary tank which contain some 340 m3 of liquid sodium. Two main primary pumps draw sodium from this pool, which is normally at 371° C, into the two reactor inlet plena. About 84% of the flow passes through the high pressure plenum and cools the core subassemblies while the remainder passes through the low pres-

models, Future posttest analyses will be needed to delve deeper into the measured behavior and investigate some of the causes of differences between measured and predicted results. Refinements in the interpretation of the measured data should also be possible at that time.

2. The EBR-II reactor plant The EBR-II plant is located in Idaho and was designed by Argonne National Laboratory who operates it for the U.S. Department of Energy. EBR-II has been in operation since 1964 and has served primarily

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E.E. Feldman et al. / EBR-II unprotected loss-of-heat-sink predictions

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sure plenum and cools the radial reflector and blanket subassemblies. The flow through all subassemblies merges in the upper plenum before passing through the outlet pipe to the intermediate heat exchanger (IHX). After passing through the IHX, the flow exits into the primary tank before entering the main pumps. The typical driver subassembly in the reactor contains 91 pins of sodium-bonded metallic fuel. Fig. 2 provides a schematic of the entire EBR-II plant - note the connection of the BOP to the primary system via the IHX. Some 60 m of pipe carries sodium from the IHX to the steam generator which is located outside the reactor building in the sodium boiler building. A similar length of pipe takes return flow to the IHX inlet. An electromagnetic (EM) pump in the return leg drives the secondary sodium flow. The steam generator has two parallel superheaters and seven parallel evaporators. 3. Design of the LOHS tests 3.1. Selection of transients

The NATDEMO code, Mohr and Feldman [3], was used to simulate several potential test transients involv-

ing loss of heat sink. A total loss of feedwater was compared with a complete stoppage of secondary flow in terms of meeting test objectives and operational constraints. Both were initiated near rated conditions and assumed no PPS action. In both cases the primary system responded essentially the same except that in the loss-of-feedwater case the loss of heat sink (as felt by the reactor) was delayed by 2 to 3 min. This time was required first to dry out the steam generator and then for the secondary sodium to traverse the distance from the steam generator to the IHX. While both cases produced the ultimate desired conditions in the primary system, a dryout could be detrimental to the steam generator integrity and would require a considerable amount of time for the plant to be returned to its normal operating condition. Therefore, we focused our attention on the stoppage-of-secondary-flow case. The idealized situation of a total LOHS (i.e., no primary system heat losses) is unattainable in EBR-II with intact heat transport systems. Furthermore, this conditions is not essential for meaningful LOHS tests. The shutdown coolers, fig. 1, remove about 350 kW when the louvers on the air-side heat exchangers are open, but still remove about 120 kW when they are closed. Other sources of parasitic heat losses, such as

60

E.E. Feldman et a L / EBR-II unprotected loss-of-heat-sink predictions

the instrument thimble and shield cooling systems, the primary sodium purification system, and the heat transfer through the walls of the primary tank, remove at least an additional 250 kW. Another means of heat loss is via the lILY, whose secondary-side flow is very difficult to stop completely. There are no valves for stopping the secondary flow and, if total power to the pump is turned off, the natural convective flow remains almost indefinitely at 5-8% of its rated value. If the alternate power supply to the pump is used, the pump head can be reversed to oppose the natural buoyancy in the secondary loop. In order to avoid thermal shock in the plant equipment, reactor plant operating procedures, however, prohibit reverse flow. Potential errors in flow readings therefore limit the minimum indicated flow to about 0.5% of full flow. This assumed minimum value of flow removes about 200 kW via the IHX. The only significant sources of heat gain (except for the reactor) are the primary pumps, which produce about 200 kW, and the tank heaters. The heaters are capable of adding about 600 kW to the primary tank sodium but were secured during the tests. In order to assess the influence of unavoidable primary tank heat losses on an L O H S test, Feldman and Mohr [4] simulated two transients with the N A T D E M O code. Both were initiated from the same assumed conditions which closely approximate the rated conditions of the plant. In Case 1 the secondary sodium flow rate was linearly reduced to 0.5% of its initial value in about the first 20 s of the transient and then held constant. In Case 2 the transient was initiated by simply tripping the pump in the secondary sodium loop. Since this is an E M pump with no moving parts, the flow decay characteristic is almost completely determined by the inertia of the moving sodium. After the coastdown of Case 2, the flow rate is determined by the net loop buoyancy. Case 1 represents a situation where the net power losses from the primary tank are minimal. Case 2 represents a contrast where, because of natural convection, the net heat loss is more than ten times greater than for Case 1. There was no scram or PPS action in either case. For the two cases, temperature histories of the reactor inlet and of the outlets of the average driver, reflector, and blanket regions were compared. Differences tended to be only of the order of 5 o C. After about the first 1800 s, temperature changes in both cases were very gradual and residual differences between the two cases could be discerned. At 2600 s, for example, the very low secondary sodium flow of Case 1 permitted a total of only 0.3 M W to be transferred from the primary system versus 3.8 M W in Case 2. This 3.5 MW dif-

ference in heat removal rate was almost entirely compensated by the difference in fission power which was 1.9 M W for Case 1 versus 5.1 M W for Case 2. As a result, the net power transferred to the primary reactor tank at 2600 s was essentially the same for both cases, being 2.5 M W for case 1 and 2.4 M W for Case 2. Thus, we were able to conclude that even if substantial, unavoidable heat losses are present, a very meaningful unprotected L O H S test can still be conducted with the EBR-II. Case I, which had an initial reactor power of 60 MW, is similar to the 60 M W L O H S tests recently completed. The only major differences was that in the test the reactor inlet temperature was 28°C lower than the Case 1 value. As mentioned in [4], the Case 1 analysis does not take into account the reactivity feedback due to the lower driver reflector and its surrounding sodium. This and other changes to the N A T D E M O model have improved the pretest predictions presented below over those of [4]. Because the analytical simulations clearly indicated that a 60 M W unprotected L O H S would be rather benign, we decided to have only one milder test prior to conducting the 60 MW case. For convenience, both tests were designed to be as similar as possible except for the initial power level which was reduced by half in the first test. Both tests were to have the same (constant) primary flow rate. Another consideration was that the 60 M W case was to be the second part of an international demonstration of the inherently safe nature of the EBR-II reactor. The first part, performed in the morning of April 3rd, consisted of an unprotected loss-of-primary-flow (LOF) from 60 MW. (See the preceding paper in this issue). Time constraints required that, to perform both tests on the same day, the restart of the reactor for the L O H S test be done very quickly. Therefore, the initial conditions of the 60 M W L O H S were made identical to those of the L O F test. 3.2. Description of the tests The initial reactor inlet temperature for the tests was 343°C instead of the normal 371°C. The initial reactor coolant temperature rise for the 60 M W test was 99°C. Reactor operation at 60 M W with the reduced reactor inlet temperature would have required the secondary pump to perform beyond its capacity had the steam header pressure remained at its normal value. In order to sufficiently lower the required secondary flow, we reduced the steam header pressure to 6050 kPa, which in turn lowered the I H X secondary inlet temperature.

E.E. Feldman et aL / EBR-H unprotected loss-of-heat-sink predictions

Normally, when the turbine is offline, as was the case during the tests, the steam header pressure is 8810 kPa at the inlet to the turbine stop valve. The reduction in steam pressure also necessitated a reduction in feedwater temperature in order to keep the feed flow to the steam drum sufficiently subcooled. Therefore, the initial feedwater temperature was reduced from 288°C to 260 o C. During normal operation, the louvers on the air-side of the shutdown coolers are closed. Normally, in the event that the reactor inlet temperature would rise to 377°C (from its normal 371°C value) the louvers open automatically. The total power removed by both shutdown coolers would increase by about 230 kw. As demonstrated in [4] and described above, however, 230 kW is virtually insignificant and would not have impacted the behavior of the tests. However, in order to maximize the purity of the tests, the louvers were kept closed during all LOHS tests. In the design of the tests, as well as in the pretest predictions contained herein, the tests were to be initiated by linearly reducing the secondary sodium flow rate from its initial value to 1.67 kg/s (about 0.5% for the 60 MW test) in 20 s. In the actual tests, it was found necessary to trip the power supply to the secondary pump and to manually control the alternate power supply, which is required to retard the flow toward the desired final value as quickly as practical. Once this value was approached, automatic control could be used to establish and maintain the 1.67 kg/s value. (Had the pump not been used to retard the flow, a natural convective flow of 5 to 8% would have resulted as described above.) The pump trip allowed a rapid coastdown which reduced the flow to near natural convective levels in about 20 s and eliminated over 90% of the IHX heat removal capability. Although the reduction to the final value took several more minutes, the effect on the reactor inlet temperature is expected to be small. Because of uncertainty in predicting the actual rate of secondary flow rundown, the NATDEMO code was used to simulate several cases where a range of constant rundown rates was studied. As expected, all of the effects of rundown rate were found to be near-term and thus not important in designing the tests.

4. Analytical models 4.1. The NA T D E M O whole-plant model

Fig. 2 provides a schematic representation of the NATDEMO model which was used to simulate the

61

EBR-II reactor and plant. Important features of the primary system model include detailed descriptions of power generation, parallel-channel thermal-hydraulics with buoyancy, detailed reactivity feedback, reactor inlet and outlet piping, the IHX, the main primary pumps, the auxiliary pump, and the primary sodium pool. The two superheaters, the seven evaporators, the EM pump, and the hundreds of meters of secondary sodium piping are included in the secondary system model. The steam system portion of NATDEMO includes a thermal-equilibrium drum model, the hydraulics for the natural convection recirculation flow through the seven evaporators, the steam pressure control system, the steamline and turbine, and the turbine bypass line. Experimental data obtained since the original NATDEMO model was published in 1981 has validated several improvements in the code. Of particular importance to LOHS transients is the reactivity feedback model, which originally had six components and currently has nine. The ninth, which takes account of the lower axial reflector sodium expansion, is very significant but was not included in the analysis of [4]. The current reactivity feedback description along with many other aspects of the model can be found in Lehto et al. [5]. A portion of the primary system, which is particularly important to the LOHS transient and difficult to model accurately, is the primary flow circuit from the IHX outlet (through the primary tank) to the two primary pump inlets. The paths from the IHX to each primary pump are not identical and are modeled separately. Each path is represented by two consecutive fluid nodes contained in an imaginary porous pipe, or mixing zone. A recirculating flow component within the primary tank passes through the walls of the imaginary pipes and mixes with the main sodium flowing within the pipes. The ratio of flow through the walls to the main flow and the ratio of the sodium volume of each pipe to that of the primary tank are key model parameters. The heat capacitance of the primary tank (bulk volume) is represented by a single node. The measurements of SHRT Test 26, Mohr and Chang [6], were used to validate/calibrate the NATDEMO primary tank flow mixing model. This test had an initial power-to-flow ratio of 1.0 and an initial power level of about 70% of 60 MW and was initiated by a rapid reduction in the secondary flow rate (without scram) which caused a rise in reactor inlet temperature. Although the design of this test limited the inlet temperature rise to about 16 ° C, there are obvious similarities to the LOHS tests described herein. The data from the SHRT 26 test also showed significant primary tank

62

E.E. Feldman et al. / EBR-II unprotected loss-of-heat-sink predictions

temperature stratification, with the higher temperatures occurring in the upper region of the tank; i.e., upwards from the IHX outlet elevation. Thus, while the model was calibrated to function adequately for the LOHS pretest predictions, we recognize that we must improve upon the single-node primary pool model in order to accurately predict a wide variety of dynamic conditions within the primary tank. 4. 2. The H O C H A N individual subassembly model

The XX09 instrumented subassembly, Messick et al. 1986 [7], was in the reactor core during the LOHS tests. This subassembly occupied a control rod position and has 59 driver-fuel type dements (plus 2 dummy elements). Some of the spacer wires, which are helically wrapped around the elements, contain thermocouples which sense coolant temperatures in the pin bundle. Pretest predictions of the XX09 temperatures were made with the aid of the HOTCHAN computer code. Because of the benign nature of the LOHS tests, the monitoring of XX09 temperatures was not needed to address safety concerns. However, during the test demonstration, the direct comparison of measured and predicted XX09 temperatures was of substantial benefit to observers. The HOTCHAN code was developed to predict the dynamic thermohydranlic behavior of a single subassembly. The HOTCHAN representation of XX09 has two radial concentric regions within the pin bundle. A more detailed description of the HOTCHAN code can be found in [5].

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several more minutes. From 1000 s to the end of the NATDEMO simulations (at 2500 s), the measured flows for both tests tend to be near the desired 1.67 kg/s value. The near stoppage of secondary flow virtually eliminates heat transfer from the primary to the secondary system via the IHX. Without this heat transfer,

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In the figures which follow, selected measured results (represented as unconnected symbols) from the 30 MW test and the April 3rd 60 MW test are superimposed on the pretest predictions. All measured data were recorded at 0.5 s intervals. In order to filter out some of the higher frequency noise, we used trapezoidal integration to time-average the measured data over each 5 s interval before plotting it. Except for this filtering, no corrections were made to the raw data presented in this paper. In fig. 3, 100% flow corresponds to 317 kg/s, which is the predicted initial secondary flow rate for the 60 MW test. In both tests the flow was to be reduced from its initial value to 1.67 kg/s over 20 s. As explained earlier and shown in the figure, in both tests most of the flow reduction did occur within the first 20 s, but the last part (which is in the natural convective range) took

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63

The primary sodium flow exiting the IHX mixes with the bulk sodium in the primary tank before entering the two main coolant pumps, located somewhat above the IHX outlet. Because of the mixing and heat transfer experienced by the primary sodium on its transport from the IHX outlet through the primary tank to the high pressure plenum (HPP), the temperatures at the IHX outlet and HPP inlet are markedly different, see fig. 4 vs. figs. 5 and 6. This difference demonstrates the role of the primary tank mixing discussed earlier. Since the flowing sodium only partially mixes with the pool, the HPP inlet temperature responds much more rapidly than does the tank bulk mixed-mean, fig. 7. Prior to the tests a calibrated control rod had been used to measure the difference in reactivity between zero-power critical at 371°C and zero-power critical at 352°C. From these data it was determined that the isothermal temperature coefficient was 0.0067 $/°C. Since about 0.29 $ of reactivity are needed to take the reactor from zero-power critical to 60 MW (i.e., the power reactivity decrement), about a 43 ° C increase * in reactor inlet temperature will reduce the fission power in the 60 MW test to essentially zero. About half of this increase would be needed in the 30 MW test. Because the reactivity feedback model in NATDEMO was calibrated to provide a 0.00067 $ / o C isothermal temperature coefficient, the predicted long-term increase of

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E.E. Feldman et al. / EBR-II unprotected loss-of-heat-sink predictions 46O

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inlet temperature, fig. 5, needed to make the driver fission power, fig. 8, approach zero, is consistent with the above 43 ° C value. In fig. 8, 100% driver power corresponds to 53.7 MW. Although the ion chamber used to measure power senses only fission power, the output value it provides is calibrated to yield total reactor power (in megawatts).

Since the initial p r e d i c t e d driver fission p o w e r is 50 M W in the 60 M W test, the m e a s u r e d driver fission p o w e r was o b t a i n e d b y m u l t i p l y i n g the m e a s u r e d total p o w e r b y 5 0 / 6 0 . F o r fig. 9, w h i c h c o r r e s p o n d s to the 30 M W test, all p r e d i c t e d initial values are exactly half o f

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E.E. Feldman et al. / E B R - H unprotected loss-of-heat-sink predictions

their counterparts in fig. 8. Thus the measured driver fission power is taken to be 25/30 of the total driver power which corresponds to 26.85 MW. The rising reactor inlet temperature acting alone would tend to increase the outlet temperature. However the falling reactor power, which is due to negative feedback, would tend to decrease it. The predicted behavior for the average coolant outlet temperature in the driver, reflector, and blanket regions is shown in fig. 10 for the 60 MW test and in fig. 11 for the 30 MW test. The relatively high thermal inertia of the blanket region coupled with the usual low blanket coolant velocities eliminates the effect that the rising inlet temperature has on increasing its outlet coolant temperature. The pronounced peak in the reflector outlet coolant temperature is due at least in part to the rise in inlet temperature being large compared to the initial coolant temperature rise in the reflector region. As the power goes toward zero, the difference between the reactor inlet and outlet coolant temperatures must also go toward zero, as demonstrated in figs. 5 and 6 for the 60 and 30 MW tests, respectively. Since the "top-of-core" thermocouples (TTCs) in XX09 were located at about 0.04 m below the top of the fuel column (during the near end-of-life test conditions), coolant temperatures at this axial location were also simulated in the HOTCHAN modeling of XX09. Eight of the top-of-core thermocouples are located within the first three element rows which corresponds to the inner

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of the two concentric pin bundle regions represented in the HOTCHAN XX09 model. Therefore, the readings of these eight thermocouples were averaged in obtaining the measured XX09 TTC (inner region) temperatures for comparison in figs. 5 and 6. The predicted early rise in reactor inlet temperature causes a significant drop in the predicted excess reactivity, fig. 12. The subsequent drop in reactor outlet temperature, however, causes the excess reactivity to recover. At 2500 s, which is at the end of the simulation, the predicted excess reactivity for both tests is near -0.015 $ and the power is about 1 MW, which implies that the reactor is only slightly subcritical. If the simulation were run out much further, the reactor temperatures should decrease about 2 ° C and the reactor should reach a true critical condition ( with excess reactivity equal to zero) and the fission power further approaching zero. This final temperature is commonly referred to as the quenching or smothering temperature and, as explained, earlier, should be about 43°C above the initial reactor inlet temperature in the 60 MW test. Note that the final equilibrium condition implies a steadystate condition in which the total reactor power (fission plus decay) is equal to the primary system heat losses and the tank temperature is steady. The measured excess reactivity is obtained from fission power by an on-line inverse kinetics routine. The comparison between measured and predicted excess reactivity in reasonably good for the first 500 s of each test. For greater times, however, the discrepancy is on the order of 0.03 to 0.05 $. This difference in reactivity is in qualitative agreement with the discrepancy in driver fission power, figs. 8 and 9. The difference between the measured and predicted excess reactivity is due, at least in part, to the inability of the primary tank model to predict the temperature response of the primary pool sodium and structures.

6. Discussion and conclusion

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2000

2500

Figs. 8 and 9 show that after about 500 s the predicted driver fission power declines much more slowly than the measured power. Although the cause of this discrepancy will require additional study, the overprediction in power is consistent with the small overprediction in the difference between HPP inlet and XX09 outlet temperatures at 2500 s, as shown in figs. 5 and 6. In addition to the low power at 2500 s, the measured reactor inlet and outlet temperature, figs. 5 and 6, virtually coincide; thus one would conclude that the quenching temperature had been essentially reached

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E.E. Feldman et al. / EBR-II unprotected loss-of-heat-sink predictions

by 2500 s. However, the first time the 60 M W test was run, the transient was allowed to run about 5½ h instead of the 1½ h in the demonstration test. In the earlier test, at about 1.7 h, the reactor outlet and inlet temperature both increased and separated. A local peak power of about 2.3 ~ was observed (by ion chamber response) before the inlet and outlet temperatures became constant and coincident again. Furthermore, the new equilibrium temperature was about 1 0 ° C higher than the previous value. After this change, which took less than an hour, the reactor temperatures remained essentially constant for the remainder of the 5-~ h. The cause of this behavior is believed to be related to the fact that the bulk sodium and primary tank structures (e.g., tank wall and safety rod drives) required several hours to achieve thermal equilibrium because of complex currents within the tank. In assessing the cause of the benign response of the EBR-II to an unprotected LOHS, one becomes keenly aware of the role of negative reactivity feedback, particularly the negative isothermal reactivity temperature coefficient. Since a reactor inlet temperature increase of only 43°C is needed to shut the reactor down, one could question the reactor response if an inlet temperature decrease were to occur instead of an increase. This could happen, for example, if the steam system pressure were reduced and the steam generator feed flow were maintained to prevent dry out. The drop in steam saturation temperature would eventually be felt at the reactor inlet. This issue was addressed in a series of steam-pressure-reduction tests and is discussed briefly by Planchon et al. [8]. In conclusion, the response of the EBR-II to an unprotected L O H S is both benign and predictable. While there are areas which need further study and portions of the N A T D E M O model which should be

enhanced, the N A T D E M O and H O T C H A N models have been shown to be very good predictors of EBR-II behavior. Moreover, the tests highlight the EBR-II as an outstanding example of how inherently safe characteristics can be used to obviate the need for elaborate safety systems.

References [1] C. Essig et al., Dynamic behaviour of Rapsodie in exceptional transient experiments, Proceedings of the International Meeting on Fast Reactor Safety 2 (1985) 635-641. [2] G.H. Golden et al., Evolution of thermal-hydraulics testing in EBR-II, Nucl. Engrg. Des. 101 (1987) 3, in this issue. [3] D. Mohr and E.E. Feldman, A dynamic simulation of the EBR-II plant during natural convection with the NATDEMO code, in: Decay Heat Removal and Natural Convection in Fast Breeder Reactors, ed. A.K. Agrawal and G,P. Guppy (Hemisphere Publishing Corp., Washington, DC., 1981) pp. 207-223. [4] E.E. Feldman and D. Mohr, Unprotected loss-of-heat sink simulation in the EBR-II plant, ASME Paper No. 84WA/HT-7 (1984). [5] W.K. Lehto et al., Safety analysis for loss-of-flow and loss-of-heat-singk-without-scram tests in EBR-II, Nucl. Engrg. Des. 101 (1987) 35, in this issue. [6] D. Mohr and L.K. Chang, Perturbation of reactor flow and inlet temperature for EBR-II reactivity-feedback validation, Proceedings of the International Meeting on Fast Reactor Safety 2 (1985) 617-627. [7] N.C. Messick et al., Modification of the EBR-II plant to conduct loss-of-flow-without-scram tests, Nucl. Engrg. Des. 101 (1987) 13, in this issue. [8] H.P. Planchon et al., Implications of the EBR-II inherently safety demonstration test, Nucl. Engrg. Des. 101 (1987) 75, in this issue.