Loss-of-primary-flow-without-scram tests: Pretest predictions and preliminary results

Nuclear Engineering and Design 101 (1987) 45-56 North-Holland, Amsterdam

45

LOSS-OF-PRIMARY-FLOW-WITHOUT-SCRAM PRETEST PREDICTIONS AND PRELIMINARY D. M O H R ,

L.K. CHANG,

E.E. F E L D M A N ,

TESTS: RESULTS

*

P.R. B E T T E N , a n d H . P . P L A N C H O N

EBR-II Division, Argonne National Laboratoo;, 9700 South Cass Auenue, Argonne, Illinois 60439, USA

A series of tests in the Experimental Breeder Reactor No. 2 (EBR-II) has been concluded that investigated the effects of a complete loss of primary flow without scram. The development and preliminary study of these events is first discussed, including the test limits and controlling parameters. The results of two of the tests, SHRT 39 and 45, are examined in detail, although a compact summary of all the tests is included. The success in meeting the objectives of the test program served to verify that natural processes will shut down the reactor and maintain adequate cooling without control rod or operator intervention. The good comparison between predicted and measured results confirms that such events can be analyzed without elaborate codes if the basic processes are understood. Furthermore, recent studies suggest that the EBR-II results are characteristic of new innovative LMR designs being pursued in the U.S. that incorporate metallic driver fuel.

1. Introduction One important class of accidents in liquid-metal cooled reactor (LMR) design and safety studies in the total loss of all pumping power to the reactor coupled with a failure of the reactor shutdown system. A series of tests in Experimental Breeder Reactor-II (EBR-II) [1] was completed in February 1986 which successfully demonstrated that the reactor will inherently, safely shut itself down by virtue of its negative reactivityfeedback characteristics. The tests showed that natural processes such as thermal expansion of reactor materials and natural convection of sodium coolant can prevent excessive core temperatures even if a serious accident were to disable the plant protection system (PPS). Because these tests were conducted up to the full reactor power rating (60 MWt), the results directly support several inherent safety features that are part of the advanced L M R plants being designed in the U.S. today under the auspices of the U.S. Department of Energy. During the conduct of the S H R T testing program [2] on EBR-II, the primary objective has been to verify the adequacy of passive decay heat removal in LMRs. This testing experience has contributed greatly toward acceptance of natural convection as a passive, principal means of safety related decay heat removal. In addition, data analysis from a large number of tests has allowed our * Work supported by the U.S. Department of Energy, Reactor Systems, Development and Technology, under Contract W31-109-Eng-38.

overall system simulation code N A T D E M O [3] to be validated for a variety of power levels and transient conditions. This validation process has shown that lumped parameter codes can yield very reliable pre-test predictions, both for protected (with scram) and unprotected (without scram) L M R transients. The important aspects of the model are discussed elsewhere in this issue [4]. The successful demonstration of natural convection in EBR-II as a safe means of cooling the core following loss-of-flow (LOF) with scram led to a culmination of the S H R T program with a series of unprotected L O F (LOFWS) tests. These tests were conducted during 1984-86 within four "test windows". Along with the basic philosophy of performing the test series in a " b o o t s t r a p p i n g " fashion, the tests were gradually escalated in severity and the modeling was continuously validated throughout the test program. In this paper, representative results are taken from two L O F W S tests, S H R T 39 and 45, and compared with their respective pretest predictions. It will be shown that the N A T D E M O and related H O T C H A N codes predict the test transients with very good success and actually required neither model nor fundamental parameter changes as the S H R T series progressed. Although the work is beyond the scope of this paper, the L O F W S test results are currently being closely examined through detailed post-test analyses. Early indications are that some modeling improvements could be made in the areas of bowing reactivity feedback, primary pump friction with stationary impellers, and primary pool mix-

0 0 2 9 - 5 4 9 3 / 8 7 / $ 0 3 . 5 0 © E l s e v i e r S c i e n c e P u b l i s h e r s B.V. (North-Holland Physics Publishing Division)

46

D. Mohr et al. / Loss-of-primao'-flow-without scram-tests

ing/stratification dynamics. These refinements should have only a small impact on predicting the important in-core temperatures, but are expected to improve the analysis of transport and mixing phenomena within the tank and the reactivity feedback dynamics in the latter stages of the transients. Before the outset of the LOFWS tests, an important series of reactivity perturbation tests was conducted in EBR-II. These tests are reported in detail elsewhere [5] and established the response of reactor power to primary flow and reactor inlet temperature changes. These tests completed the model validation we judged necessary before embarking on actual LOFWS tests. The N A T D E M O / H O T C H A N predictions described in this paper are compared with plant data obtained from both standard and special instrumentation with a Data Acquisition System (DAS). The latter category includes signals taken primarily from the XX09 INSAT (described elsewhere in this issue [6]), which included 28 thermocouples and two flowmeters. As was true for all prior INSATs having resided in the core, these data permitted an in-depth analysis of the thermal-hydraulic response of the driver region - thus XX09 was critically important to the successful conduct of the SHRT series.

2. Development of the LOFWS test series An extensive amount of analysis was conducted in preparation for the LOFWS test series. The NATDEMO pretest analyses included evaluation of several parametric effects (plus uncertainties) associated with subassembly thermal-hydraulic behavior, reactivity feedback, and primary coastdown characteristics. A test transient was considered potentially acceptable if the driver and blanket response satisfied temperature and damage criteria established previously - these are described below. In addition to the test transients themselves, we also calculated the response to abnormal (unlikely) LOF events that conceivably could occur during the tests. These transients had to satisfy limiting criteria for peak driver temperatures and clad damage as well as for peak blanket temperatures. For clarity in the discussions below, basic distinguishing characteristics are given for the three test groups; i.e., SHRT Group V, Group VI-A, and Group VI-B. The Group V tests were all conducted at 10 MW, whereas Groups VI-A and VI-B both involved tests in the 30 to 60 MW full power range. The tests in Group V were adjusted in severity by varying primary coastdown rate, auxiliary pump condition, and secondary pump condition. The tests in VI-A were varied by initial

power and flow levels, primary coastdown rate, and secondary flow condition. The Group VLB tests were distinguished by initial power and flow, primary coastdown rate, and auxiliary pump condition. 2.1. Criteria f o r the tests and abnormal events

In addition to meeting many other requirements, an individual test could be considered acceptable only if its predicted peak clad temperatures did not exceed certain limits for the driver and blanket fuel. In the first two groups of LOFWS tests (V and VI-A), the fuel-clad interface temperature (with uncertainties) was limited to 715°C for the driver and 705°C for the blanket fuel. These are temperatures for the incipient eutectic formation involving these fuels and the cladding. Since a great deal of confidence in the predictions was gained from the Group V and VI-A tests, the driver fuel criterion was modified to be twofold for the VI-B tests: the predicted test temperatures could not exceed 815°C and the accumulated damage for any fuel element could not exceed 0.25 (1.0 represents expected cladding failure). The blanket fuel-clad criterion was kept the same as above, however, for all test groups and for normal and abnormal events. For the analysis of abnormal events [4], different criteria from the above were established for the driver fuel, and these in turn differed for the test groups as our testing experience matured. For Groups V and VI-A, the driver fuel temperature could not exceed 815°C, but could exceed 715°C for up to one minute in the hottest element. For the VI-B tests, the 815°C limit was increased to 900°C (in-core sodium boiling temperature). In addition, the accumulated potential clad damage was limited to 0.25, including the tests themselves and any abnormal event - should one occur. For the blanket fuel (for all test groups), the predicted temperature for an abnormal event could not exceed 705°C in the hottest element. For other types of subassemblies in the core during SHRT testing, the temperature limit corresponded to sodium boiling. All of the above criteria were based on established limits for EBR-II metallic fuel and modified with recent information described elsewhere in this issue [7]. The measured results followed the predicted temperatures closely throughout the SHRT program, and the above fuels criteria were not violated for any test. Moreover, the reactor was not challenged with abnormal events during the SHRT test periods. Other criteria associated with the LOFWS testing program were related to large components along the heat transport path. Structural analyses were performed

D. Mohr et al. / Loss-of-primary-flow-without-scram-tests

47

on the reactor vessel cover, reactor outlet piping, IHX, secondary piping and the superheaters. As reported in more detail elsewhere in this issue [4], thermal stress a n d / o r fatigue damage limits were not encroached for any of these components during the LOFWS testing. 2.2. Reference S H R T core loading~configuration

The SHRT testing was planned with core loadings that ideally were to be "identical" from one test series to another. Although identical loadings could actually not be achieved, basic objectives were met in regard to maximum power-to-flow ratios in the driver, reflector and blanket regions at rated reactor power and flow. Of course, a given loading also had to satisfy neutronic limits on excess reactivity, shutdown margin and control rod worth. In the SHRT series, the normal amount of excess reactivity was not required for burnup compensation and more attention was paid to minimizing azimuthal flux tilt and subassembly power-to-flow ratios. Although the SHRT loadings contained fewer experimental subassemblies than usual, they were not greatly different from typical EBR-II loadings which can have significant variations from run to run. Because these tests involved no reactor shutdown during the transients, we were also concerned about reactivity feedback characteristics of the SHRT loadings. Previous experience had shown, however, that almost all of the feedback exhibits only small ( - + 5%) variations from run to run and that the nonlinear (bowing) component undergoes the largest changes. Analyses of the linear components have been made for various loadings, allowing a comparison with the overall PRD * results and with the prompt (linear) components obtained from dynamic tests. Since the calculation of bowing reactivity is both time consuming and inaccurate, special tests were performed to deduce this component by alternate means. These results showed that the reactivity feedbacks exhibited typical PRD results for all the SHRT loadings and that the bowing component was somewhat positive, but stable throughout the test series. After the Group V tests were completed (all at - 10 MW), the relative blanket flow was increased in anticipation of the gradually increasing blanket power generation due to 239pu buildup. This flow change was accomplished by adjusting the opening of the low pressure plenum (LPP) throttle valves, see fig. 1, which * PRD is the power-reactivity decrement, i.e., the reactivity required to bring the reactor from a zero-power, hot-critical condition up to full power.

Fig. 1. EBR-II primary system. resulted in a 17% increase in the blanket flow rate. The new flow configuration then remained fixed for the duration of SHRT testing and caused the peak blanket temperatures to remain well below the established criteria for all transients. 2.3. Parametric computer studies

The NATDEMO code and the related HOTCHAN code (both described elsewhere in this issue [4]) were used extensively to design the Group V and VI tests and minimize peak temperatures. These calculations involved parametric studies of key quantities including flow coastdown characteristics, reactivity feedback, initial tank temperature, auxiliary pump condition, initial reactor power and flow, individual driver subassembly power and flow, and individual blanket subassembly power and flow. Most of the studies were done because of "plant adjustable" effects, e.g., coastdown characteristics having various shapes and pump stopping times. In other cases, the parameter in question was not adjustable but simply had a significant uncertainty. In the first category (adjustable effects), the effects consisted of primary and secondary coastdown characteristics, initial tank temperature, auxiliary pump condition, and initial reactor power and flow. The secondary category (effects with uncertainties only) comprised reactivity feedback coefficients, driver subassembly power and

48

D. Mohr et al. / Loss-of-primarv-flow-without-scram-tests 850

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COASTDOWN TIME, s

Fig. 2. Peak driver transient temperature vs. pump stop time. flow, and blanket subassembly power and flow. Although the optimal coastdown characteristics had been determined, the pump stopping times were also assumed to have a small uncertainty because of random variations. The bulk of the preliminary studies consisted of "tuning" the desired primary and secondary coastdown

characteristics with the aid of NATDEMO. Although two distinct coastdown rates ("fast" and "slow") were employed in the Group V series, coastdown tuning was confined to the Group VI-A and B test series. Because definite test limits had been established for peak clad temperatures of the fuel, this analysis dealt primarily with minimizing the peak driver temperature for the desired initial conditions associated with rated power (60 MWt). It should be pointed out that peak blanket fuel temperatures ceased to be a competing constraint after the decision was made to increase the LPP flow split. Three aspects of coastdown tuning were explored for the VI-A tests: pump stop time, coastdown shape (speed vs. time) and "staggering" of the pump stop times. The primary attribute was found to be the pump stop time which has a profound effect on peak transient driver temperature, as shown in fig. 2. The coastdown shape tends to have a lesser effect, but it is important for short pump stop times. It should be pointed out that the desired decay of pump speed early in the transient was found to be faster than a typical hyperbolic decay in order to minimize driver temperature peaks. Pump staggering herein means that the stop times (for multiple primary pumps) are purposely made un-

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D. Mohr et al. / Loss-of-primary-flow-without-scram-tests

equal. This scheme was found to have a beneficial effect on peak core temperatures due to the nature of the EBR-II "active" coastdown (i.e., that used for Group VI-A). Previous experience had shown that pump speed for this type of controlled coastdown drops rather abruptly into zero (from - 10% of rated speed) regardless of how the pump controller is programmed. This abrupt drop in speed does not occur, however, for the "passive" coastdown mode (i.e., that used for most of Group VI-B) in EBR-II. See fig. 3 for a pump speed control functional diagram including means of active and passive control and section 2.4 for further discussion. Separate analytical studies of these effects were conducted * and confirmed the experimental results - a full discussion of these effects, however, is beyond the scope of this paper. It should be noted, however, that the impact of the "chopped tall" of the active coastdown was a significant temperature perturbation (30 to 40°C) compared with the case of a "smooth tail". Because EBR-II has two primary pumps (and both behave similarly), we decided to stagger (i.e., separate) the stopping times about the average stop time so that the local temperature perturbation was minimized, i.e., about halved. Note also that the overall peak driver temperatures usually occur near the time when the pump speeds drop to zero. The final target for this staggering was about 7% of the average pump stop time thus, the difference in stop times for a nominal case of 300 s was - 20 s. The secondary pump coastdown characteristic was also studied analytically for its effect upon peak driver and blanket temperatures. Because the secondary pump is an electromagnetic type, it does not have a shaft "stop time"; however, the term here is used to describe when the applied voltage goes to zero. Although the secondary pump stop time does affect peak temperatures, it is not as effective as that of the primary pumps. Secondly, we found that the coastdown shape was unimportant if it is reasonably hyperbolic. Two other effects related to secondary coastdown rate should be noted, one being the effect it has on the blanket region peak temperatures. Although an extended secondary coastdown reduces driver temperatures peaks somewhat, it tends to increase blanket temperature peaks. This apparently is due to the increased blanket flow (due to increased loop buoyancy) which in turn causes thermal-hydraulic equilibrium to be more nearly reached at this stage of the transient when the -

* These studies were done with the EBR-II FLOWDOWN code which simulates the M-G sets, controllers, pump motors and power supplies for the main pumps.

49

power/flow ratio is high. Another effect of secondary coastdown rate is the thermal stress condition caused in the IHX. In this case, the IHX stress tends to be excessive in the IHX top end if the secondary pump stop time greatly exceeds that of the primary pumps for an unprotected LOF. Thus, for those tests having an extended secondary coastdown, the stop time was limited to being - 3 5 % longer than for the primary coastdown. Once the need for specifically tuned coastdown characteristics was established, modifications were made to the controllers associated with speed control of the primary and secondary pumps. The functional requirements, features and test results for the new controllers are described elsewhere in this issue [6]. The initial tank temperature and auxiliary pump condition are plant adjustable effects that were also analyzed. The former effect caused peak transient temperatures (in driver or blanket) that were about 1.1°C lower for each I ° C reduction in tank temperature. For the auxiliary pump effect, three different conditions were easily obtainable - these being (1) auxiliary pump on and connected to rectifier/charger, (2) pump on and disconnected from rectifier, and (3) auxiliary pump off. The final parameters studied here are the initial reactor power and flow. As discussed further below, all of these adjustable parameters were incorporated into the LOFWS tests in order to limit the peak fuel temperatures in the progression through a given test group. In the category of parameters involving only uncertainties, we were concerned mainly with reactivity feedback, in particular the bowing component as mentioned above. In assessing the impact of feedback uncertainties, we assumed uncertainties in all components to occur randomly, as indicated elsewhere in this issue [4]. In addition, a NATDEMO analysis was done to evaluate the impact of varying the bowing component alone over a large range, 0 to +4¢, with a nominal value of + 2.5¢. The effect of the above variation on peak driver and blanket temperatures was calculated to be + 35 and + 2 5 ° C , respectively, for the most severe LOFWS transient, i.e., SHRT 45. 2.4. Final test description

The LOFWS tests of the SHRT series were placed into test groups V, VI-A and VI-B. The general approach was to gradually increase the severity of the tests within each group and monitor the predicted vs. measured results during each test. The Group V tests were conducted in May 1985 and consisted of six unscrammed loss-of-flow tests. In each case, the initial conditions were 16% reactor power (10 MWt), 19%

50

D. Mohr et a L / Loss-of-primary-flow-without-scram-tests

Table 1 Summary of the Group VI-A test descriptions and peak temperatures SHRT No.

Initial power (% of rated)

Initial primary flow (% of rated)

Initial secondary flow (% of nominal a)

33

50

100

69

34 b

50

100

69

35 36 37

50 50 100

50 50 100

48 48 106

38 39

100 100

100 100

106 106

Primary pump coastdown conditions

Secondary pump coastdown conditions

Type 1 (300 s, actively controlled) Type 1

Mode 2 (400s, actively controlled) Mode 1 (tripped) Mode 2 Mode 1 Mode 1

Type 1 Type 1 Type 2 (600 s, actively controlled) Type 1 Type 1

Mode 2 Mode 1

Predicted peak cladding temperature of driver (° C) Nominal

w/uncertainties

561

585

569

593

588 588 570

625 625 604

619 644

652 672

Note: Auxiliary pump tripped, bulk sodium temperature = 352°C, and steam header pressure = 6980 kPa for all tests. a Nominal flowrate corresponds to 60 MW, 371°C bulk sodium, and steam header pressure = 8806 kPa. b This test was deleted.

reactor flow and a tank temperature of 338°C. The reactor responded basically as expected, and measured coolant temperatures from XX09 fell within the uncertainty b a n d s predicted for each test. These results are reported elsewhere [8] and show that the m a x i m u m driver fuel-clad temperature reached was about 6 7 9 ° C ( S H R T 32). Because the measured results c o m p a r e d so well with those predicted, these tests formed the basis of our plans to conduct similar tests all the way to 100% power. The G r o u p VI-A tests comprised seven planned

L O F W S transients identified as SHRTs 33 through 39. These tests were conducted from initial conditions of either 50% or 100% power, 50% or 100% flow, and a tank temperature of 352°C. The basic descriptions of these tests are summarized in table 1, and all primary coastdowns are "actively" controlled wherein the main power supply to the p u m p drive trains and clutches is retained. N o t e that the S H R T 34 test was deleted to save time, especially since it was j u d g e d as unnecessary for b o o t s t r a p p i n g to higher powers. The seven test variables c o m p a r e d in the figures below comprise

Table 2 Summary of the Group VI-B test descriptions and peak temperatures SHRT Initial Initial Initial Primary pump No. power primary flow secondary flow coastdown (% of rated) (% of rated) (% of nominal a) condition

Auxiliary pump condition

40

50

100

68

Passively controlled, 95 s

635

41

50

100

68

42 43 44 45

50 70 90 100

100 100 100 100

68 84 98 104

Actively controlled, 200 s Same as SHRT Same as SHRT Same as SHRT Same as SHRT

On battery Trip of 2400-V 598 breaker to M-G set Off Same as SHRT 41 598 Off On battery On battery On battery

676 713 774 802

40 40 40 40

Secondary pump coastdown condition

Same Same Same Same

as SHRT as SHRT as SHRT as SHRT

Note: Plant conditions are 343°C bulk sodium and steam header pressure = 6048 kPa for all tests. a Nominal flowrate corresponds to 60 MW, 371°C bulk sodium and steam pressure of 8806 kPa.

41 41 41 41

Predicted peak cladding temperature of driver ( ° C) Nominal w/uncertainties

643 664 722 747

622

D. Mohr et al. / Loss-of-primary-flow-without-scram-tests

primary pump speeds, reactor (fission) power, excess reactivity, XX09 flowrate, XX09 core coolant temperature, subassembly outlet temperature, IHX secondary outlet temperature and secondary sodium flowrate. In the case of reactor power response, a direct comparison is included between predicted and measured fission power since total power is actually not measured. The total driver power response is thus composed of the fission power plus the fission product decay power components. The Group VI-B tests were composed of six planned LOFWS transients defined as SHRTs 40 through 45. These tests, described in table 2, were conducted from initial power levels of 50 to 100% rated, primary flow of 100% (all cases), and a tank temperature of 343°C. All primary coastdowns (except for SHRT 41) are "passively" controlled wherein main pump driving power is lost but the clutch controller power supply is retained these conditions are similar to those for a station blackout. The resulting pump stop times for the VI-B tests are much shorter than for the VI-A tests. Note also that the auxiliary pump continues to be powered by battery (rectifier tripped) in 4 of the VI-B tests which is the auxiliary pump condition if total station blackout should occur. 3. Conduct of tests and results 3.1. Initiation o f a test

The key steps required to initiate a LOFWS test will be briefly outlined. The first requirement is to establish the proper reactor power and flow rate at stable plant conditions. This step includes setting up instruments and starting data recording with the DAS. Next, the A-C electrical power to the auxiliary pump is disconnected. Depending on the test, the pump is then supplied by battery power only or completely disconnected from all power. Just prior to test initiation, the control rod drives are deactivated to preclude control rod movement during the transient period. This action prevents insertion or withdrawal of the rods by the drive motors, but it does not affect the scram function. In order to protect the reactor from equipment failures during testing, special scram protection is inserted just prior to test initiation. This protection is based primarily on in-core temperatures measured with the special XX09 INSAT. As the XX09 protection is inserted, the normal protection for loss of primary flow is bypassed simultaneously - the latter comprises low primary flow trips and high subassembly outlet temperature (SOT) trips. (Note that reactor overpower protec-

51

tion is retained for the transient.) In addition, the Group VI-B tests incorporated protection based on pump speed. If either pump speed decay fell below a safe envelope, a trip signal from the coastdown test monitor would be generated to scram the reactor with a single rod. The transient forcing functions consist of tripping the primary and secondary pumps. This is accomplished by removing the 2400 V power supply from these pumps. Finally, as the transient ensues, the reactor/plant is closely monitored on various recording devices. Note that power is reduced by natural phenomena - no automatic or operator action is necessary except to terminate each test. If a forcing function is abnormal a n d / o r core temperatures exceed set limits, an automatic or manual scram, however, are available for immediate reactor protection. 3.2. Test predictions and measured results

The selected plant variables for Group VI-A are shown superimposed with those of VI-B for the most severe tests in each group, namely SHRT 39 and 45. The predictions for these variables are shown superimposed in each figure. In addition to the figures that follow, tables 1 and 2 include the predicted peak driver temperatures for all of the Group VI transients. These values cannot be directly compared, however, with the XX09 measured temperatures below because XX09 is cooler than the hottest driver. Note that these tables contain both the "nominal" and "maximum" peak clad temperatures, which are the predicted values without uncertainties (nominal) and with uncertainties, respectively. The basic forcing functions, primary pump speeds, for SHRTs 39 and 45 are compared in fig. 4. As mentioned earlier, the pump stop times for SHRT 39 are purposely staggered. While the coastdown shapes for SHRT 45 were designed to be identical, intrinsic differences between the two pump drive units cause - 5 s difference in stop times. As was true in all active primary coastdowns, the speeds in SHRT 39 dropped abruptly from - 70 rpm while this tendency was essentially absent in SHRT 45. The reactor (fission) power response is given in fig. 5 which also includes the calculated decay heat component. The results show that measured power initially decreases rapidly as predicted due to the negative reactivity feedback associated with increased reactor temperatures. The time lag between pump speed and power decay is due primarily to the delayed neutron response of the neutron kinetics since the relationship between

D. Mohr et al. / Loss-of-prtrnarv-flow-without-scram-tests

52

core coolant temperatures and feedback is much more prompt. The measured power later in the transient, being lower than predicted, is apparently due to more negative feedback than predicted. 120 #1 Code !npu'l' ]

i

100-

The "measured" excess reactivity is obtained by means of inputting the power signal into an inverse kinetics routine solved by the Data Acquisition System (DAS) computer in real time. These results are shown in fig. 6 which shows the minimum DAS values for SHRT 39 and 45 to be - 2 5 ¢ and -35¢, respectively. Note that the excess reactivity directly reflects the core tern-

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53

D. Mohr et aL / Loss-of-prirnary-flow-without-scram-tests

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. . . . . . . . . . . . . .

-1300

o~

650

-=oo~

7

6oo-

-11oo ~ ci. E

E 550-

z~',

L1000 ~

40500-

"".

a ........... ~AAAAA

A

4,50-

. . . . . . . . . . . .

900

AA

AAAAAAAAAAAAAAL • 800

400

O-200

0

200

400

Trne into

600

800

Transient, s

Fig. 7. XX09 flow rate for SHRT 39 and 45 tests.

1000

.... -200

I .... 0

i .... 200 Time

I .... 400

i .... 600

into Transient,

i .... 800

1000

s

Fig. 8b. Temperature of XX09 coolant (inner-region TTC's) and hottest driver cladding for SHRT 45 test.

D. Mohr et al. / Loss-of-prima
54

effects become important. One chief difference between a L O F W S flow trace and one following a L O F with a scram is that the latter tends to have a local minimum just after the pumps reach zero speed, The final asymptotic flows measured for S H R T s 39 and 45 reveal a greater difference than predicted. Since the final flow in S H R T 39 is driven only by natural convection, the predicted flow and core temperatures will tend to be significantly overpredicted in the latter stages by the "excessive" calculated power. As explained above, the XX09 I N S A T has 13 coolant thermocouples located at the same elevation near the top of core (TTC). Eight of these 13 are distributed within the central three rows of elements and are shown averaged in fig. 8, Thus, a direct comparison can be made with the predictions from the H O T C H A N code for the same central region at this elevation. The H O T C H A N predictions and measured results for the remaining two rows of elements are not given in this paper. The other reactor temperature shown is a comparison of predicted and measured subassembly outlet temperatures (SOT's) in fig. 9. The N A T D E M O code predicts this response for an average driver channel, and the measured average comprises SOT thermocouple data that represent the driver region, i.e. the first 6 - 7 rows *

700 -

650

Predicted, SHRT 39

f

1000

A Measured, SHRT 39 525

..F'~

~Predi~fed..$HRT ~_~

950

500

- 900 ~

475-

~- 450-

i,

o

z E

850

\ ~O00©C i

400.

- 750

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1

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. . . .

200

J

. . . .

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400

. . . .

600

I

,

,

r

800

1000

Time into Tronsient, s

Fig. 10. IHX secondary outlet temperature for SHRT 39 and 45 tests. Although flow weighting factors were applied to the measurements, the responses construeted from the individual TTC's and SOT's may differ somewhat from the true average thermal responses at these elevations.

1

?'~

-

I

(J

5 5 0 j]

l:~redicted Driver Outlet. SHRT S9

ii A

Measured

i --~'~i~'d

~ [ 0

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600

!

M. . . . .

] /

SOT AveragQ, SHRT 39 /

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L. . . . . . . . . .

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-J " 1100

i i ~'~.

Measured, SHRT 39 I Predicted, SHRT 45

,i 0

Measured,

j +

Measured, SHRT 38 ',

SHRT

4,5 1

o

2 55O

1000 +~ Q_

E

E

500

900 45O

400

00

.... --200

J .... o

J ....

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[ ....

200 400 600 l i m e into Transient, s

~ ....

soo

~..~.. -e-

~-.~.

rY lO

b_

8oo

1ooo

Fig. 9. Averaged outlet coolant temperatures of driver region for SHRT 39 and 45 tests.

1

....

--200

I ' ' '""

0

I ' '

200

'""

I ....

400

I ....

600

I

800

.... 1000

T i m e into Transienf, s

* SHRT core loadings always contained drivers in the first six rows and, in some loadings, a portion of row 7.

Fig. 11. Secondary loop flow rate for SHRT 38, 39 and 45 tests.

D. Mohr et al. / Loss-of-primary-flow-without-scram-tests Fig. 10 shows the IHX secondary outlet temperature response which is also a good indicator of the IHX primary inlet temperature within 20 s into the transient. This signal was the only thermocouple selected from the IHX because of its veracity. Unfortunately, the primary side inlet temperature is no longer available and the primary outlet thermocouples are not reliable dynamically. Finally, the secondary inlet temperature remains essentially constant during most LOFWS transients. The data from these test series are concluded with the secondary flow response given in fig. 11, which is taken from a reliable electromagnetic flowmeter. In both the SHRT 39 and 45 tests, the secondary pump was tripped at transient initiation. For comparison purposes, however, the programmed secondary coastdown for SHRT 38 is also shown which had a stop time of - 400 s. This extended secondary coastdown enhances the primary system buoyancy and is responsible for peak driver temperatures in SHRT 38 being about 15°C lower than in SHRT 39.

4. Discussion of results

4.1. Adequacy of predictions Test acceptance was based principally upon measured vs. predicted XX09 temperatures after verifying the forcing functions as correct. The XX09 measured temperature transients all fell within the pre-test calculated band except for SHRT 41. In that case, the measured temperature exceeded the predicted upper value (with uncertainties) at the peak by - 1 4 ° C . It should be noted that SHRT 33 was conducted with an erroneous primary coastdown on its first attempt. Because of the good success in the repeat of SHRT 33 and also because of the predicted small increment of severity ( < 10 ° C) to SHRT 34, the latter was omitted in the test sequence. After an examination of forcing functions and results for SHRT 41 upon test completion, the test was labeled successful although a detailed post-test study will be made of the coastdown behavior near zero speed and the frictional behavior of driver subassemblies under purely natural convection. Note that the latter stages of both SHRTs 41 and 42 involve natural convection, i.e., the auxiliary pump is off. We should also mention that a reactor scram occurred during SHRT 42 on its first attempt because the XX09 trip settings had been based on an XX09 power/flow ratio that was somewhat low. After the trip settings were corrected, SHRT 42 (repeat) ran successfully as did all other tests in the VI-A and VI-B groups. Other portions of the

55

NATDEMO model that may also be addressed include the primary pump impedance (near zero speed), the IHX thermal-hydraulic behavior and tank mixing/ stratification effects - particularly at low flows. In one sense the primary system model of NATDEMO was challenged more severely in the longer coastdown tests of Group VI-A than for the VI-B tests. This is true because the tendency for core inlet temperature to change during the time of interest is much greater during the longer coastdown transients. Thus, the tank mixing/stratification phenomena were involved more significantly during most of the VI-A tests. It should also be emphasized that the transport of a temperature wave from the IHX outlet to the core inlet in EBR-II is a complicated process and difficult to describe in sufficient detail with a lumped-parameter model. In SHRT 33 the inlet temperature remained essentially constant while in 35 and 36 it decreased slightly during the transient. Finally, in SHRTs 37 to 39 the inlet temperature increased, especially in 37 and 39. These trends were reasonably well predicted by NATDEMO except for SHRT 38 which will be studied for this effect in a post-test analysis. Because of the large isothermal temperature coefficient of reactivity, inlet temperature changes can (and did) significantly affect the core response during a LOFWS event. Other long-term reactivity effects that may be significant are local pool temperature effects on the control and safety rod drivelines as well as on the primary tank wall. The relative expansion of the tank and these drivelines can obviously lead directly to reactivity effects involving the control and safety rods. This paper does not directly address the two LOFWS tests of Group VI-B that were repeated during the demonstration test week in April, 1986, i.e., SHRTs 43 and 45. An important aspect here, however, is that the results of the repeated tests are very similar to their original counterparts (if adjustments are made for small power differences) which implies good reproducibility.

4.2. Plant impact of the tests The main impact of the LOFWS tests was on the driver fuel which reached temperatures about 85°C over the fuel-clad eutectic value (715 ° C) during SHRTs 45 and 45 (repeat), with uncertainties included. As discussed elsewhere in this issue [9], the calculated driver fuel element "damage" for these tests was found to be small, however, and no element breach was noted during the SHRT testing. The post-test analysis revealed that the maximum potential damage accumulated during these two tests was 0.021 if uncertainties are in-

56

D. Mohr et al. / Loss-of-primarv-flow-without-scram-tests

cluded (to be compared with unity). If uncertainties are not included, then the accumulated damage was found to be only 0.001. An additional analysis was made of a S H R T 45 type of test if the auxiliary pump were turned off completely instead. In this case, the hottest driver is predicted to reach a peak temperature - 1 5 ° C higher than for S H R T 45. The fuel damage associated with such a test is estimated to be 0.025 (with uncertainties) compared with about 0.01 for S H R T 45 as performed. The impact of the L O F W S tests on the balance of plant appears to be small, as predicted. Other than the reactor fuel and upper vessel structure, the I H X experienced somewhat more severe thermal cycling than other large plant componentsl mostly due to large flow imbalances between primary and secondary sides. Because the primary flow continuously decreases during this type of transient, however, the temperature wave leaving the reactor outlet plenum is greatly attenuated as it travels down the heat transport path. This is especially true once the disturbance reaches the I H X and enters the secondary sodium system. Thus, the effect of L O F W S testing on the superheaters, evaporators and steam drum was minimal.

4. Conclusions The main conclusion to be drawn from the Group V and VI tests is the confirmation that natural processes will shut down EBR-II following a L O F W S event and maintain adequate cooling without automatic control rod or operator action. Furthermore, a large body of analysis [9] indicates that these results are characteristic of a range of L M R sizes if metallic driver fuel is used. It was also found that the N A T D E M O and H O T C H A N codes predicted the test results, in general, with very good accuracy requiring neither model nor fundamental parameter changes throughout the test series. We feel that a basic understanding of the phenomena involved is probably more important than the development of elaborate, complex codes. A wealth of information is now available that will enable modelers to validate codes encompassing an entire L M R plant. The tests were designed not only to progress safely through a test sequence, but also to evaluate the effects of coastdown time, initial p o w e r / f l o w ratio, and a small amount of residual

pumping after loss of main pumps. A large amount of data was also accumulated for the first time on transient axial temperature gradients within the primary tank. This should contribute significantly to the understanding of mixing/stratification phenomena within a large sodium pool with an active, low velocity coolant stream. Finally, the test results are significant to EBR-II itself in its response to a station blackout in the absence of a reactor scram. From the lessons learned from S H R T 45, we can state that the r e a c t o r / p l a n t presently has the ability to safely shut itself down should this event actually occur at normal, full-power operating conditions.

References [1] L.J. Koch, W.B. Loewenstein, and H.O. Monson, Addendum to Hazard Summary Report, Experimental Breeder Reactor-II (EBR-II), ANL-5719 (Addendum), Argonne National Laboratory (January 1964). [2] R.M. Singer et al., Decay heat removal and dynamic plant testing at EBR-II, Second Specialists' Meeting on Decay Heat Removal and Natural Convection in LMFBRs, Brookhaven National Laboratory, April 17-19, 1985. [3] D. Mohr and E.E, Feldman, A dynamic simulation of the EBR-II plant with the NATDEMO code, Decay Heat Removal and Natural Circulation in Fast Breeder Reactors, eds. A. Agrawal and J. Guppy (Hemisphere Publishing Co., Washington, DC, 1981) 207-224. [4] W.K. Lehto et al., Safety analysis for the loss-of-flow and loss-of-heat sink without scram tests in EBR-II, Nucl. Engrg. Des. 101 (1987) 35, in this issue. [5] D. Mohr and L.K. Chang, Perturbations of reactor flow and inlet temperature for EBR-II reactivity-feedback validation, Fast Reactor Safety Conference, Vol. 2 of Proceedings, Knoxville, Tenn., April 1985. [6] N.C. Messick et al., Modification of EBR-II plant to conduct loss-of-flow-without-scram tests, Nucl. Engrg. Des. 101 (1987) 13, in this issue. [7] C.E. Lahm et al., EBR-II driver fuel qualification for loss-of-flow and loss-of-heat-sink tests without scram, Nucl. Engrg. Des. 101 (1987) 25, in this issue. [8] L.K. Chang, D. Mohr, H.P. Planchon, Loss-of-flow without scram tests in EBR-II and comparison with pretest predictions, Nuclear Safety (to be published 1986). [9] H.P. Planchon, LI. Sackett, G.H. Golden, and R.H. Sevy, Implications of the EBR-II Inherent Safety Demonstration Test, Nucl. Engrg. Des. 101 (1987) 75, in this issue.