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Loop seal clearing and refilling during a PWR small-break LOCA
Nuclear Engineering and Design 121 (1990) 431-440 North-Holland
431
LOOP SEAL CI,EARING AND REFII,LING DURING A PWR SMALL-B~_AK Y. K U K I T A , J. K A T A Y A M A , H. N A K A M U R A
LOCA
a n d K. T A S A K A
Japan Atomic Energy Research Institute, Tokai-Mura, Ibaraki-Ken, Japan
Received 7 September 1989
During a cold-leg small-break loss-of-coolant accident (LOCA) in a Westinshous~type pressurized water reaclor (PWR), the core liquid level is depressed temporarily before steam dears liquid out of the primary-loop loop seals. The loop seal clearing and associated core level depression may occur repeatedly if the emergency core coolant (ECC) refills the loop seals after clearing. This paper presents an interpretative description of experimental results on loop seal clearing and refilling phenomena obtained at the ROSA-IV Large Scale Test Facifity (LSTF).
1. Inereduelion During a certain class of cold-leg small-break lossof-coolant accidents (LOCA) in a Westinghouse-type pressurized water reactor (PWR), the steam volume in the primary system may continue expanding until steam blows liquid out of the U-shaped pump suction legs (loop seals), opening a path for the steam to be either relieved from the break or condensed in the cold leg in contact with the emergency core coolant (ECC). This steam volume expansion would involve concurrent, manometric liquid level depressions in the loop seal downflow leg and in the reactor core. Fig. l a shows
(o)
schematically the primary coolant inventory distribution at the moment when the loop seal level reacbes the bottom, enabling steam to blow the loop seal upflow leg. A minimum core collapsed liquid level, located at the elevation of the loop seal bottom leg, is takea at this time. Since the loop seal bottom leg is located below the core top for typical Westinghouse-type F W R plants, the level depression may uncover the core upper regions until the core level recovers with the progress of blow~rtg of the loop seal upflow leg. There is a possibility that the loop seals refill with water after the above clearing process. Then, this water may be cleared by steam, if core boiling persists, caus-
J~ STEAM
(b)
Fig. 1. Schematic of coolant distribution at: (a) initiation of clearing of the loop seal upflow leg, and (b) initiation of refilling of the loop seal upflow leg. 0 0 2 9 - 5 4 9 3 / 9 0 / $ 0 3 . 5 0 © 1990 - Elsevier Science Publishers B.V. ( N o r t h - H o l l a n d )
432
K Kukita et a L / Loop seal clearing and refilling
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observed [2] at the ROSA-IV Large Scale Test Facility (LSTF) [3] during tests conducted for a scaled break area of 0.5% (scaled 2-in. break). For (ii) to occur, i~ would be necessary that the loop seal steam flowrate is small enough to allow ECC to drop into the loop seal or that the ECC attains sufficient momentum due, for example, to steam hammer in the cold leg, to penetrate into the loop seal against the steam counterflow. The U.S. Nuclear Regulatory Commission (USNRC) recently identified the possibility of core uncovering, following loop seal refill by ECC, during post-blow-down long-term cooldown for which the core steaming rate may become sufficiently small, due to core power decay with time, to allow the loop seals to refill. This issue was studied by Fletcher and Callow [4] by analyzing PWR plant responses using a simplified thermalhydraulic code for steady-state analysis and the RELAP5/MOD2 code [5]. Subsequently, a series of system-effects tests addressing this issue were conducted at the ROSA-IV LSTF for low-power low-pressure conditions representative of LOCA post-blowdown phases. This paper presents an interpretative description of test results obtained at the LSTF.
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FLOW CONTROLVALVI Fig. 2. Schematic of the ROSA-IV Large Scale Test Facility (LSTF). ing a core level depression to occur again. The two possible loop seal refill mechanisms identified in previous studies [1,2] are: (i) refill with steam generator (SG) condensate, and (ii) refill with ECC, which is injected into the cold legs, penetrating through the reactor coolant pump as shown schematically in fig. lb. For (i) to occur, condensation should persist on the SG primary side after loop seal cleating, and this requires that the primary saturation temperature remains higher than the secondary temperature. This refill mechanism has been
2. Test facility The LSTF [3], shown in fig. 2, is a I :48 volumetrically-scaled, full-height, full-pressure simulation of a
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Y. Kukita et al. / Loop seal clearing and refilling Westinghouse-type 4-loop (3423 MW(0 ) PWR. It has two symmetric primary loops each representing two loops in the reference 4-1oop PWR. Each loop includes an active steam generator (SG) consisting of 141 Utubes, and a reactor coolant pump. The LSTF core comprises 1064 indirectly-heated heater rods and 104 unheated rods having the same diameter (9.5 mm), pitch (12.6 mm) and heating length (3.66 m) as the PWR 17 x 17-bundle fuel rods. The core is divided into three radial regions having core power peaking factors of 1.51, 1.00 and 0.61, respectively. The axial profde of the core power is represented by a 9-step chopped cosine with a peaking factor of 1.495. The LSTF hot legs (207 mm i.d.) are sized to conserve the length to square-root-of-diameter ratio L / V ~ , as well as the scaled (1/24) volumes, to simulate the flow regime transitions in the horizontal legs [6]. The cold legs have the same diameter and elevation as the hot legs. The loop seals (168 man i.d.) are sized to scale the volume, while preserving the height, and have a cross-sectional area of about 1/22 of that of the PWR. The loop seal and cold leg geometries are shown in fig. 3. In this figure, the locations of test instrumentation are shown only for those mentioned in this paper. The analyses of Fletcher and Callow [4] show that the upper pressure-vessel bypasses can affect the loop seal behavior for those plants having large bypass capacities. These bypasses provide steam flow paths between the core outlet and the broken cold leg, parallel with the primary loops. The LSTF bypasses include upperhead spray nozzles, connecting the upper head to the upper downcomer, and simulated hot leg nozzle leakage between each hot leg to the downcomer. The spray nozzle and hot leg nozzle leakages allow bypass flowrates of about 0.3~ and 0.2~ (for the two loops) of the total core flowrate, respectively, during singie-phase steady-state operation. These bypass flow capacities are representative of Japanese-built Westinghouse-type PWRs. Larger spray nozzle bypass capacities, ranging typically from 1 to 4%, are employed in the U.S. Westinghouse PWRs.
3. Test conditions and test procedure This paper focuses on one of the LSTF tests conducted on the loop seal clearing and refilling phenomena. This particular test, designated Run SB-HL-07-FO, was designed to reduce the ambiguity in system energy balance as far as possible. Farrier tests at the LSTF [7] had shown that the ~ balance, both local and global, was affected considerably by system internal
433
heat transfer which could be comparable to the core power at such low core power levels as represented in these tests. To reduce such heat transfer effects, the test was performed for non-standard test boundary conditions and non-standard test geometry as follows: (1) The test was performed with the SG secondaries empty of liquid, for both loops, to eliminate the SG heat transfer effects on the system energy balance. Realistically, the SG primary-to-secondary (or reversed) heat transfer alters the net steam production rate in the system by condensing steam or boiling water within U-tubes. The SG heat transfer effects have been studied in separate LSTF tests [12]. (2) The ECC was injected into the broken-loop cold leg alone using a low-pressure injection (LPI) pump. This was done to reduce the heat loss from the upper plenum to the downcomer, across the core barrel Since the LSTF core barrel heat transfer area is oversized by a factor of about 4 ~ from a scaling point of view, the heat loss can be of considerable magnitude if the core barrel outer surface is kept cooled by the crossflow of the ECC. (3) The test utilized the broken loop (Loop B) alone, with the intact loop (Loop A) inactivated by closing the flow control valve at the loop seal bottom leg, whereas this valve was fully open for the broken loop. This test geometry was chosen to avoid transient vessel crossflow of the ECC which could have occurred if the intact loop were active. The break was located between the vessel and the ECC injection port (see fig. 3), and was simulated using a 32-mm i.d. sharp-edged orifice which was mounted at the end of a horizontal pipe (87.3 mm i.d., 1.4 m long) branching from the cold leg. The orifice flow area corresponded to 10% of the 1/48-scaled PWR cold-leg cross-sectional area. The break flow was routed into subcooled water retained in a open catch tank to condense the steam component. The break flowrate was evaluated by numerically differentiating the catch tank level with respect to time.
4. Test results
4.1. Test ooeroiew The test was performed by manually changing the core electric power in steps, as shown in fig. 4, while keeping the ECC injection rate constant at 5 k g / s and the ECC temperature at 310 K. The loop seal behavior
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was monitored by looking at the upflow- and downflow-leg differential pressures shown in fig. 4. After the test was initiated with a core power of 400 kW (0.56% of the 1/48-scaled PWR nominal power), the loop seal differential pressure indicated cyclic responses. Each cycle was initiated by a gradual level drop in the downflow leg, followed by a quick blowing of the upflow leg, which kept the loop seal clear until it was refilled by dumping of water through the pump. The cyclic changes continued, while the interval of cycle decreased with core power, until the core power was raised to 700 kW (0.98%). Thereafter, the loop seal kept clear until the core power was lowered to 400 kW (0.56%). That it to say, the limiting core power for refilling was dependent on the history of the preceding transient. The core collapsed liquid level was depressed to the loop seal bottom height (about 1.8 m above the core bottom for the LSTF, see fig. 3) every time the loop seal level dropped to the bottom, as shown in fig. 5. These level drops uncovered the core upper regions causing rod surface heat-ups, as shown in fig. 6. The lowest dryout level was dependent on core power as will be discussed.
drop of the ECC through the pump. The loop seal steam flowrate was measured in the test using drag disc transducers located in the downflow leg (see fig. 3). Data from these transducers are shown in fig. 7 in comparison with the core steam production rate, WCORE, calculated from the steady-state energy balance: WCORE= Q/(hFG + AhiN ), where Q is the core power, h FG is the latent heat of vaporization, and AhIN is the core inlet subcooling. The measured steam flowrate deviated from the energy balance prediction considerably, even with the loop seal kept clear steadily, i.e. after 3500 s, due probably to system response delay in reaching equilibrium in terms of coolant inventory distribution and pressure. An approximate agreement between the data and energy balance calculation was reached after 4500 s. The data also contained rapid fluctuations which may have been induced, at least partly, by non-steady steam con-
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The conditions for the onset of loop seal refill are of particular interest. Counter-current flow limiting (CCFL) between the loop seal and cold leg is one of the controlling factors for refill. Namely, refdl may occur only when the steam flow is insufficient to prohibit the
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Y. Kukita et a L / Loop seal clearingand refilling
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densation on the ECC in the cold leg, including occasional steam hammer. After 4000 s, the core power was decreased manually in steps in an attempt to define the flooding limit, i.e. the steam flowrate at which fiquid downfiow begins. For this time period, the subcooled ECC filled the cold leg to the top covering the pump outlet, as the steam coming from the loop seal was condensed totally. This can be seen in fig. 8 where the cold leg water level, derived from the densitometer data, and fluid temperature measured at the pipe top between the ECC injection port and pump outlet are shown.
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The first refill following the power reduction occurred at 5010 s, immediately after the core power was lowered stepwise from 500 to 400 kW, as shown in fig. 4. More precisely, the refill was initiated during a steam flowrate undershoot below the steady-state steam production for this core power (0.19 kg/s), in a transient response to the power reduction. The onset of refill was signaled by a slow increase in the upflow leg differential pressure, which was initiated when the steam flowrate dropped to about 0.1 k g / s ( + a measurement uncertainty of 0.05 kg/s), and this was followed by a quick dumping of water into the loop seal. This large deviation of the transient steam flowrate from the steady-state core steam production rate is suggestive of the flowrate sensitivity to system transient parameters. As shown schematically in fig. 9, the loop seal steam flowrate, WL.S~is controlled, first of all, by the differential pressure between the steam source (the mixture level where steam'leaves the two-phase mixture) and the steam sink (pump outlet where steam is condensed), A p C O L D -- ApHOT, such that
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ature measured at leg pipe top immediately downstream of pump outlet (bottom, uncertainty = + 2 K).
where A is the loop cross-sectional area, 0o is the steam density, and K is the overall pressure loss coefficient for the steam-t'filed portion of the loop. The above equation neglects the "capacitance" effects of the steam-filled volumes in the primary loop upper regions.
436
Y. Kukita et al. / Loop seal clearing and refilling
Such effects may be important during fast transients. An ideal steady state of the system would be represented by
WCORE = WL.S q- WBYPASS, where WBYPASS is the steam flow through the bypasses which was negligibly small as compared with WCORE for the present test, due to the relatively small bypass areas represented in the LSTF. The steam source-to-sink pressure drop, A p C O L D - ApHO-r, is dependent on the coolant inventory distribution in the primary system. The loop hot-side differential pressure, APHoT , consisting mostly of liquid static head, varied largely during the test, responding to the changes in the hot-side coolant inventory. The hot-side mixture level moved between the core upper region and SG inlet plenum. Meanwhile, the cold-side differential pressure, ApCOLD, remained almost constant, as the downcomer liquid level was kept almost stationary at the cold-leg nozzle. The steam flow through the pressure vessel bypasses kept the upper downcomer steam-filled. The measured steam flowrate responded sensitively to the changes in the hot-side coolant inventory, since the steam source-to-sink pressure drop, ApcoLD -ApHOT, was considerably smaller than either ApHOT or APcoLD. The measured loop seal differential pressure (the total of the upflow leg, downflow leg, and pump differential pressures), shown in fig. 7, was about 3 kPa for a steam flowrate of about 0.22 kg/s. The pressure drop occurred mostly in the pump into which liquid may have had penetrated from the cold leg. This sensitivity of the steam flowrate to loop differential pressure appears to explain the observed unstable behavior of the pump CCFL, i.e. the fact that counter-current flow occurred only in a brief transient at the beginning of the refill, and the main part of the refill occurred by a fashion of water "dumping". Coupled system behavior may be responsible for such unstable CCFL breakdown behavior. That is to say, when water begins to pour into the loop seal, this water creates an incremental loop-seal pressure drop, and this reduces the steam flow thereby increasing the liquid downflow rate. Each refill, once started, progressed until the loop seal was completely blocked, probably because of this mechanism. No cases were observed in which the loop seal upflow leg stayed half-filled, being voided steadily as suggested by analysis of Fletcher and Callow [4]. This seemingly unstable nature of CCFL, as well as the considerable fluctuations seen in steam flowrate, made it difficult to define precisely the conditions for the onset of refill, not to speak of the fact that
the critical steam flowrate was comparable to the drag disc measurement uncertainty. The CCFL characteristics are dependent on specific facility (or plant) geometry. For the LSTF, there are two flow area constrictions between the loop seal and cold leg: (i) the reactor coolant pump, and (ii) the venturi flowmeter in the pump suction leg. The pump and venturi geometries are peculiar to the LSTF. The pump area takes a minimum of 0.0090 m2 at the hori-. zontal discharge which opens at the centerline of the cold leg (0.0336 m2 in area). The venturi flowmeter (minimum cross-sectional area = 0.0057 m2) is located in the vertical loop seal upflow leg (0.0222 m 2 in area) beneath the pump suction. Although the venturi flow area is smaller than the pump minimum flow area, there was no indication that CCFL occurred at this location rather than at the pump. As has been mentioned, the first refill following the power reduction initiated at 5010 s when the steam flowrate dropped to about 0.1 kg/s (_+ a measurement uncertainty of 0.05 kg/s). On the other hand, no refilling was seen for steam flowrates above 0.2 kg/s. For instance, after the last loop seal clearing during the power increase (at 3500 s), the loop seal remained clear although the steam flowrate dropped temporarily to about 0.2 kg/s. The following flooding parameters [8] are employed to discuss further the refill onset conditions. Kutateladze parameter:
Ku=;o[p a2 // (go ap)] 1/4 Non-dimensional superficial oelocity o f steam:
J~ =Jo [ oo/( gD Ap ) ]'/2, where )G is the superficial velocity of steam, g is acceleration due to gravity, o is surface tension, At) is the density difference between phases, and D is diameter. The steam flowrates of 0.1 and 0.2 kg/s corresponds respectively to j ~ = 0.54 and 1.1 for the pump minimum cross-sectional area (horizontal flow), and to K u ~ 3.2 and 6.4 for the venturi flowmeter (vertical flow) for a fluid saturation temperature of 400 K typical of this time period. The flooding limit reported in the literature for a horizontal pipe depends on the pipe length-to-diameter ratio, L / D . However, it is not likely that the limiting value of j ~ exceeds 1.0 for any value of L / D . T h e flooding limit for large-diameter vertical pipes reported by Richter [9] is K u = 3.2. Thus, the observed limiting steam flowrate, between 0.1 and 0.2
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kg/s, appears to be reasonable. The effects of the cold-leg fluid subcooling on the limiting steam flowrate is unknown.
4.3. Cyclic loop seal clearing and refilling The repetitive loop seal clearing and refilling observed for lower core powers involved cyclic variations of many parameters. These parameters included, in addition to the core level and Steam flow rate which have already been mentioned, the primary pressure, break flowrate, coolant mass inventory and its spatial distribution, as well as core inlet subcooling. The difference between the ECC and the break flowrates, i.e. the net changing rate of the coolant inventory, is shown in fig. 10 in comparison with the cold-leg pressure. The clear correspondence between the break flowrate and the primary pressure occurred since only a small system overpressure (50 kPa) was needed for the present break area of 10~ in order to obtain a subcooled-liquid break flow in the balance with an ECC injection rate of 5 kg/s. At the same time, the primary pressure was controlled by the balance between the core steam production and the steam relief through the loop seal. The loop seal steam flow was responsive to the vessel coolant inventory, as has been discussed. This interdependence between the loop seal steam flow and the coolant inventory, coupled by the primary pressure, induced cyclic variations of these quantities. A typical cycle can be divided into the following five stages which are characterized by the behavior of the loop seal steam flow rate, WLS., the time derivative of primary pressure, d p / d t , and the time derivative of the primary coolant inventory, d M / d t.
437
(1) Downflow leg level depression (WL.s.= 0, d p / d t > O, d M / d t < 0). After the loop seal refilled, the coreproduced steam, being unable to escape from the break, pregsurized the primary system. The primary pressurization increased the break flowrate to exceed the ECC injection rate (fig. 10) initiating a net reduction in coolant inventory. The inventory loss led to level depressions in the loop seal and core (figs. 4 and 5), which were followed by clearing of the loop seal upflow leg. (2) Upflow leg clearing (WLs.> WCORe, d p / d t < O ) . After the inception of clearing of the upflow leg, the loop seal steam flowrate increased quickly (fig. 7). The steam relief through the loop seal caused a system depressurization (fig. 10). It is noteworthy that liquid in the loop seal upflow leg was cleared completely (or almost completely) once clearing started (fig. 4). For the peak steam flowrate following clearing (0.8-1.1 kg/s), which was two to three times larger than the steady-state core steam production, the value of Ku for the vertical upflow leg (0.168 m i.d.) ranged from 6.5 to 9.1, i.e. well above the Ku = 3.2 flooding-limit criterion [9]. Likewise, the value of j~ for the horizontal leg (0.168 m i.d.) ranged from 0.79 to 1.1. Thus, the loop seal horizontal leg may have been cleared of liquid almost completely. (No level measurement was done for this leg.) (3) Decay m the loop seal steam flow (WE.s. < WcoltE). The loop seal steam flow decayed quickly from the peak value, which was taken immediately after clearing, and almost stopped within 100-200 s after clearing (fig. 7). This flowrate undershoot below the steady-state steam production occurred primarily because of a decrease in the steam source-to-sink pressure difference, APcoLD --ApHOT , due to an overfeed to the loop hot side which was associated with a decrease in the break flowrate after loop seal clearing. Since the break flowrate became smaller than the ECC flowrate after loop seal clearing (fig. 10) due to system depressurization, the ECC was forced into the vessel core region, and eventually to the hot-leg and SG inlet plenum, increasing the loop hot-side static head, and thereby limiting the loop seal steam flow. Also, core boiling suppression, due to the penetration of subcooled ECC into the core (fig. 6), may have played a secondary role in this decay in loop seal steam flowrate. (4) Initiation of refill (WLs.< Wcop.E, d p / d t > O, d M / d t > 0). Refill initiated after the primary pressure had started recovering from the minimum taken after the loop seal clearing (fig. 10). This suggests
438
}I. Kukita et aL / Loop seal clearing and refilling
that refill occurred after the steam relief through the loop seal had become smaller than the core steam production. This limitation in steam flow was due to the above-mentioned transient overfeed to the loop hot side, and was not due to loss of condensation potential in the cold leg; the cold leg was filled with subcooled water at this stage of the transient (fig. 8). (5) End of refill (WL.s. = 0, d p / d t > 0). The refill ended typically within 20 s after the onset of "dumping'" and the steam flow was blocked completely. The loop seal steam was replaced smoothly by liquid, without exhibiting any non-steady condensation, although the liquid was subcooled by up to 80 K. The loop seal refill resulted in a quick reversal of core flow which was detected from a rise in the core inlet fluid temperature (fig. 6). The loop seal downflow-side level continued rising until the loop seal water head balanced with the vessel riser (core and upper plenum) head. Since the latter was smaller at higher core powers, the loop seal maximum level was also lower. As described above, the cychc loop seal behavior included a transient overfeed to the loop hot side occurring after loop seal clearing. This overfeed resulted in an undershoot in loop seal steam flowrate which allowed the loop seal to refill, even when the core steam production rate was large enough, if it could flow steadily through the loop seal, to keep the loop seal clear. This explains the considerable difference observed in the limiting core power for the onset of loop seal refill between the power increase and reduction cases. In a LOCA transient, the core power decays monotonously with time after scram. However, the loop seal steam flowrate may deviate from the core steaming rate, from time to time, being affected by system transient responses as well as automatic and manual interventions. Thus, there is a possibility that all the loop seals become liquid-filled even at a core power for which the steady-state steam production rate exceeds the limit for loop seal refill. Once this situation is realized, then the subsequent loop seal clearing, which may occur if core boiling persists, may tend to initiate a cyclic loop seal behavior. 4.4. Core dryout behavior During loop seal clearing, the core collapsed level drops to the loop seal bottom height. The minimum core mixture level during this level depression depends on void fractions in the core two-phase mixture. Fig. 11
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shows the minimum mixture level, as a function of primary pressure and core power, calculated using the Cunningham-Yeh void fraction correlation [10] which is known to be applicable to PWR rod bundle geometry. The time scale in this figure is based on the ANS (1979) decay heat power curve [11]. The data points shown in this figure are the plots of primary pressure and core power measured at the moment of minimum core collapsed liquid level. These data were taken from two sets of LSTF tests. The first set consists of three loop-seal-refill tests all conducted using similar test procedures for break areas of 5%, 10% (the present test) and 75%, respectively. The second set includes two cold-leg-break blowdown tests that were initiated from a primary pressure of 16 MPa, for break areas of 0.5% and 5%, respectively. For these two full-pressure blowdown tests, the first loop seal clearing occurred at a primary pressure of about 8 MPa. For the 5% break test, both loops were cleared and remained clear until the end of test at 2400 s. However, for the 0.5% break test, only the broken-loop loop seal was cleared, and it was refilled with the ECC. This loop seal was cleared again at a primary pressure of 1.8 MPa. For all the data presented in this figure, the primary loop upper portions were empty of liquid when loop seal clearing occurred, and thus the minimum core collapsed level agreed with the loop seal bottom height. The minimum rod dryout levels in these tests, which were determined on the basis of rod-surface tempera-
Y. Kt&ita et al. / Loop seal clearingand refilling
tures measured at nine discrete elevations along the core, are in reasonable agreement with the predicted mixture level. However, the &Tout level on each rod depended on the linear heating rate, which differed between the three core heating regions. The rod temperature excursions were generally small because of low core power. However, it is interesting to note that the maximum rod surface superheating clearly was not dependent on core power. Because the core level depression was slower and deeper for lower core power, the core upper regions were uncovered for a longer time before the core level recovered (fig. 11). There is a possibility that the core level depression occurs almost statically due to: (i) the bypass flow which may relieve most of the core steam production for extremely low core power [4], or (ii) condensation of steam on the SG primary side. The latter case has been simulated in one of the LSTF tests [12], and comparatively large heat-ups (up to 180 K above the saturation temperature) have been observed.
5. Discussion The present test results, although valuable for understanding the LOCA loop seal behavior qualitatively, are not applicable directly to real plant responses because of the pump geometry and loop seal leg cross-sectional area which were not scaled properly. The CCFL characteristics of full-size pump and legs, including liquid subeooling effects, should be understood to predict F W R loop seal behavior. However, it seems encouraging that the p ~ s e n t test results indicate that the loop seal refdl is not dictated by static CCFL but largely affected by system coupled, dynamic responses which subscale facilities would simulate. For a quantitative prediction of real plant responses, and for a defmition of the possible worst conditions for PWR core cooling during post-blowdown phases, computer codes which are able to deal with those phenomena observed in the present tests are needed. Existing LOCA analysis codes generally have limitations in calculating low-pressure phenomena, including condensation on the steam-water interface. Data from the present tests will provide a basis for verification of codes. One of the most important differences between the real PWR and present test geometries is in the number of primary loops. However, it is not likely that the multi-loop geometry of real plants makes the loop seal responses much more complicated than observed in the test. At such low core power level as represented in the
439
present test, loop seal clearing would occur for only one of the loops, and the other loops would remain sealed without contributing to steam relief from the systen~ It is unknown, however, whether or not any particular loop (broken or intact) would be cleared preferentially. The present paper focused on results for a 10~ break test, as no significant qualitative difference has been found in the test results obtained so far for break areas ranging from 5 to 75~. The primary effect of the break area was that the time-averaged primary pressure was lower for a larger break area, provided that the ECC injection rate was the same.
6. Smamary The basic mechanisms of the loop seal refilling phenomenon during a PWR small-break LOCA have been described by analyzing the results of LSTF cold-leg break tests conducted for low-power, low-pressure conditions representative of LOCA post-blowdown phases. These tests were the first system-effects tests ever conducted on this phenomenon. They demonstrated that the loop seal behavior is governed by the time-dependent integral system responses. Inadequate core cooling situations following loop seal refilling, as predicted by a previous computer code analysis for a PWR, were actually observed in the tests. The major observations from the tests are summarized as, follows. The loop seal refilled with the ECC when the steam flowrate through the loop seal was insufficient to prohibit this from occurring. The onset of ref'fll was affected by several factors including the sensitivity of the loop seal steam flowrate to the coolant inventory distribution in the primary system. The most interesting phenomenon observed in the tests was cyclic repetition of loop seal clearing and refilling. This occurred because of the coupled responses of the system parameters, inelliding: the coolant inventory and its spatial distribution; the steam release rate through the loop seal; and the primary pressure. The loop seal clearing phase of each cycle involved a core liquid level depression which uncovered the core upper regions temporarily.
A
~
t
The authors gratefully acknowledge the cooperation of the USNRC in performln$ this work.
440
Y. Kukita et al. / Loop seal clearing and refilling
References [1] H. Tuomisto and P. Kajanto, Two-phase flow in a fullscale loop seal facifity, Nucl. Engrg. Des. 107 (1988) 295-305. [2] K. Tasaka et al., The results of 0.5% PWR small-break LOCA tests in ROSA-IV LSTF break location parameter tests series, Proc. 15th Water Reactor Safety Information Mtg., Gaithersburg (Oct. 1987). [3] ROSA-IV Group, ROSA-IV Large Scale Test Facility (LSTF) System Description, JAERI-M 84-237, Japan Atomic Energy Research Institute (1984). [4] C.D. Fletcher and R.A. Callow, Long-term recovery of pressurized water reactors following a large break loss of coolant accident, Nucl. Engrg. Des. 110 (1989) 310-328. [5] V.H. Ransom et al., RELAP5/MOD2 Code Manual, Vols. 1 and 2, NUREG/CR-4312, EGG-2396, Idaho National Engineering Laboratory (1985). [6] N. Zuber, Problems in modeling small break LOCA, NUREG-0724, U.S. Nuclear Regulatory Commission (1980).
[7] K. Tasaka et al., Loop seal clearing and refilling during a PWR small-break LOGA, Proc. 16th Water Reactor Safety Information Mtg., Gaithersburg (1988). [8] GB. Wallis, One-Dimensional Two-Phase Flow (McGraw-Hill Book Co., New York, 1969). [9] H.J. Richter, Flooding in tubes and annuli, Int. J. Muhiphase Flow 7 (1981) 647-658. [10] J.P. Cunningham and H.C. Yeh, Experiments and void correlation for PWR small break LOCA conditions, Trans. Am. Nucl. SOC. 87 (1973) 453~ [11] American Nuclear Society, American National Standards for Decay Heat Power for Light Water Reactors, ANSI/ANS-5.1-1979 (1979). [12] Y. Kukita et al., Intentional ccx~lant system depressurization: experimental studies in the ROSA-III and ROSA-IV Programs, Presented at the CSNI Specialist Meeting on Intentional Coolant System Depressurization, Garching, F.R.G. (1989).
431
LOOP SEAL CI,EARING AND REFII,LING DURING A PWR SMALL-B~_AK Y. K U K I T A , J. K A T A Y A M A , H. N A K A M U R A
LOCA
a n d K. T A S A K A
Japan Atomic Energy Research Institute, Tokai-Mura, Ibaraki-Ken, Japan
Received 7 September 1989
During a cold-leg small-break loss-of-coolant accident (LOCA) in a Westinshous~type pressurized water reaclor (PWR), the core liquid level is depressed temporarily before steam dears liquid out of the primary-loop loop seals. The loop seal clearing and associated core level depression may occur repeatedly if the emergency core coolant (ECC) refills the loop seals after clearing. This paper presents an interpretative description of experimental results on loop seal clearing and refilling phenomena obtained at the ROSA-IV Large Scale Test Facifity (LSTF).
1. Inereduelion During a certain class of cold-leg small-break lossof-coolant accidents (LOCA) in a Westinghouse-type pressurized water reactor (PWR), the steam volume in the primary system may continue expanding until steam blows liquid out of the U-shaped pump suction legs (loop seals), opening a path for the steam to be either relieved from the break or condensed in the cold leg in contact with the emergency core coolant (ECC). This steam volume expansion would involve concurrent, manometric liquid level depressions in the loop seal downflow leg and in the reactor core. Fig. l a shows
(o)
schematically the primary coolant inventory distribution at the moment when the loop seal level reacbes the bottom, enabling steam to blow the loop seal upflow leg. A minimum core collapsed liquid level, located at the elevation of the loop seal bottom leg, is takea at this time. Since the loop seal bottom leg is located below the core top for typical Westinghouse-type F W R plants, the level depression may uncover the core upper regions until the core level recovers with the progress of blow~rtg of the loop seal upflow leg. There is a possibility that the loop seals refill with water after the above clearing process. Then, this water may be cleared by steam, if core boiling persists, caus-
J~ STEAM
(b)
Fig. 1. Schematic of coolant distribution at: (a) initiation of clearing of the loop seal upflow leg, and (b) initiation of refilling of the loop seal upflow leg. 0 0 2 9 - 5 4 9 3 / 9 0 / $ 0 3 . 5 0 © 1990 - Elsevier Science Publishers B.V. ( N o r t h - H o l l a n d )
432
K Kukita et a L / Loop seal clearing and refilling
Loop-A .
.
.
.
Loop-B .
.
.
.
.
observed [2] at the ROSA-IV Large Scale Test Facility (LSTF) [3] during tests conducted for a scaled break area of 0.5% (scaled 2-in. break). For (ii) to occur, i~ would be necessary that the loop seal steam flowrate is small enough to allow ECC to drop into the loop seal or that the ECC attains sufficient momentum due, for example, to steam hammer in the cold leg, to penetrate into the loop seal against the steam counterflow. The U.S. Nuclear Regulatory Commission (USNRC) recently identified the possibility of core uncovering, following loop seal refill by ECC, during post-blow-down long-term cooldown for which the core steaming rate may become sufficiently small, due to core power decay with time, to allow the loop seals to refill. This issue was studied by Fletcher and Callow [4] by analyzing PWR plant responses using a simplified thermalhydraulic code for steady-state analysis and the RELAP5/MOD2 code [5]. Subsequently, a series of system-effects tests addressing this issue were conducted at the ROSA-IV LSTF for low-power low-pressure conditions representative of LOCA post-blowdown phases. This paper presents an interpretative description of test results obtained at the LSTF.
] r
STEAM GENERATOR
ACCUMULATOR(~
FLOW CONTROLVALVI Fig. 2. Schematic of the ROSA-IV Large Scale Test Facility (LSTF). ing a core level depression to occur again. The two possible loop seal refill mechanisms identified in previous studies [1,2] are: (i) refill with steam generator (SG) condensate, and (ii) refill with ECC, which is injected into the cold legs, penetrating through the reactor coolant pump as shown schematically in fig. lb. For (i) to occur, condensation should persist on the SG primary side after loop seal cleating, and this requires that the primary saturation temperature remains higher than the secondary temperature. This refill mechanism has been
2. Test facility The LSTF [3], shown in fig. 2, is a I :48 volumetrically-scaled, full-height, full-pressure simulation of a
4246m STEAM GENERATOR
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Fig. 3. LSTF loop seal and cold-leg geometry.
Y. Kukita et al. / Loop seal clearing and refilling Westinghouse-type 4-loop (3423 MW(0 ) PWR. It has two symmetric primary loops each representing two loops in the reference 4-1oop PWR. Each loop includes an active steam generator (SG) consisting of 141 Utubes, and a reactor coolant pump. The LSTF core comprises 1064 indirectly-heated heater rods and 104 unheated rods having the same diameter (9.5 mm), pitch (12.6 mm) and heating length (3.66 m) as the PWR 17 x 17-bundle fuel rods. The core is divided into three radial regions having core power peaking factors of 1.51, 1.00 and 0.61, respectively. The axial profde of the core power is represented by a 9-step chopped cosine with a peaking factor of 1.495. The LSTF hot legs (207 mm i.d.) are sized to conserve the length to square-root-of-diameter ratio L / V ~ , as well as the scaled (1/24) volumes, to simulate the flow regime transitions in the horizontal legs [6]. The cold legs have the same diameter and elevation as the hot legs. The loop seals (168 man i.d.) are sized to scale the volume, while preserving the height, and have a cross-sectional area of about 1/22 of that of the PWR. The loop seal and cold leg geometries are shown in fig. 3. In this figure, the locations of test instrumentation are shown only for those mentioned in this paper. The analyses of Fletcher and Callow [4] show that the upper pressure-vessel bypasses can affect the loop seal behavior for those plants having large bypass capacities. These bypasses provide steam flow paths between the core outlet and the broken cold leg, parallel with the primary loops. The LSTF bypasses include upperhead spray nozzles, connecting the upper head to the upper downcomer, and simulated hot leg nozzle leakage between each hot leg to the downcomer. The spray nozzle and hot leg nozzle leakages allow bypass flowrates of about 0.3~ and 0.2~ (for the two loops) of the total core flowrate, respectively, during singie-phase steady-state operation. These bypass flow capacities are representative of Japanese-built Westinghouse-type PWRs. Larger spray nozzle bypass capacities, ranging typically from 1 to 4%, are employed in the U.S. Westinghouse PWRs.
3. Test conditions and test procedure This paper focuses on one of the LSTF tests conducted on the loop seal clearing and refilling phenomena. This particular test, designated Run SB-HL-07-FO, was designed to reduce the ambiguity in system energy balance as far as possible. Farrier tests at the LSTF [7] had shown that the ~ balance, both local and global, was affected considerably by system internal
433
heat transfer which could be comparable to the core power at such low core power levels as represented in these tests. To reduce such heat transfer effects, the test was performed for non-standard test boundary conditions and non-standard test geometry as follows: (1) The test was performed with the SG secondaries empty of liquid, for both loops, to eliminate the SG heat transfer effects on the system energy balance. Realistically, the SG primary-to-secondary (or reversed) heat transfer alters the net steam production rate in the system by condensing steam or boiling water within U-tubes. The SG heat transfer effects have been studied in separate LSTF tests [12]. (2) The ECC was injected into the broken-loop cold leg alone using a low-pressure injection (LPI) pump. This was done to reduce the heat loss from the upper plenum to the downcomer, across the core barrel Since the LSTF core barrel heat transfer area is oversized by a factor of about 4 ~ from a scaling point of view, the heat loss can be of considerable magnitude if the core barrel outer surface is kept cooled by the crossflow of the ECC. (3) The test utilized the broken loop (Loop B) alone, with the intact loop (Loop A) inactivated by closing the flow control valve at the loop seal bottom leg, whereas this valve was fully open for the broken loop. This test geometry was chosen to avoid transient vessel crossflow of the ECC which could have occurred if the intact loop were active. The break was located between the vessel and the ECC injection port (see fig. 3), and was simulated using a 32-mm i.d. sharp-edged orifice which was mounted at the end of a horizontal pipe (87.3 mm i.d., 1.4 m long) branching from the cold leg. The orifice flow area corresponded to 10% of the 1/48-scaled PWR cold-leg cross-sectional area. The break flow was routed into subcooled water retained in a open catch tank to condense the steam component. The break flowrate was evaluated by numerically differentiating the catch tank level with respect to time.
4. Test results
4.1. Test ooeroiew The test was performed by manually changing the core electric power in steps, as shown in fig. 4, while keeping the ECC injection rate constant at 5 k g / s and the ECC temperature at 310 K. The loop seal behavior
}I. Kukita et al. / Loop seal clearing and refilling
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(s)
Fig. 4. Core power (top, uncertainty = + 30 kW) a n d loop seal differential pressures (bottom, uncertainty = + 2 kPa).
was monitored by looking at the upflow- and downflow-leg differential pressures shown in fig. 4. After the test was initiated with a core power of 400 kW (0.56% of the 1/48-scaled PWR nominal power), the loop seal differential pressure indicated cyclic responses. Each cycle was initiated by a gradual level drop in the downflow leg, followed by a quick blowing of the upflow leg, which kept the loop seal clear until it was refilled by dumping of water through the pump. The cyclic changes continued, while the interval of cycle decreased with core power, until the core power was raised to 700 kW (0.98%). Thereafter, the loop seal kept clear until the core power was lowered to 400 kW (0.56%). That it to say, the limiting core power for refilling was dependent on the history of the preceding transient. The core collapsed liquid level was depressed to the loop seal bottom height (about 1.8 m above the core bottom for the LSTF, see fig. 3) every time the loop seal level dropped to the bottom, as shown in fig. 5. These level drops uncovered the core upper regions causing rod surface heat-ups, as shown in fig. 6. The lowest dryout level was dependent on core power as will be discussed.
drop of the ECC through the pump. The loop seal steam flowrate was measured in the test using drag disc transducers located in the downflow leg (see fig. 3). Data from these transducers are shown in fig. 7 in comparison with the core steam production rate, WCORE, calculated from the steady-state energy balance: WCORE= Q/(hFG + AhiN ), where Q is the core power, h FG is the latent heat of vaporization, and AhIN is the core inlet subcooling. The measured steam flowrate deviated from the energy balance prediction considerably, even with the loop seal kept clear steadily, i.e. after 3500 s, due probably to system response delay in reaching equilibrium in terms of coolant inventory distribution and pressure. An approximate agreement between the data and energy balance calculation was reached after 4500 s. The data also contained rapid fluctuations which may have been induced, at least partly, by non-steady steam con-
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The conditions for the onset of loop seal refill are of particular interest. Counter-current flow limiting (CCFL) between the loop seal and cold leg is one of the controlling factors for refill. Namely, refdl may occur only when the steam flow is insufficient to prohibit the
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Y. Kukita et a L / Loop seal clearingand refilling
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densation on the ECC in the cold leg, including occasional steam hammer. After 4000 s, the core power was decreased manually in steps in an attempt to define the flooding limit, i.e. the steam flowrate at which fiquid downfiow begins. For this time period, the subcooled ECC filled the cold leg to the top covering the pump outlet, as the steam coming from the loop seal was condensed totally. This can be seen in fig. 8 where the cold leg water level, derived from the densitometer data, and fluid temperature measured at the pipe top between the ECC injection port and pump outlet are shown.
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The first refill following the power reduction occurred at 5010 s, immediately after the core power was lowered stepwise from 500 to 400 kW, as shown in fig. 4. More precisely, the refill was initiated during a steam flowrate undershoot below the steady-state steam production for this core power (0.19 kg/s), in a transient response to the power reduction. The onset of refill was signaled by a slow increase in the upflow leg differential pressure, which was initiated when the steam flowrate dropped to about 0.1 k g / s ( + a measurement uncertainty of 0.05 kg/s), and this was followed by a quick dumping of water into the loop seal. This large deviation of the transient steam flowrate from the steady-state core steam production rate is suggestive of the flowrate sensitivity to system transient parameters. As shown schematically in fig. 9, the loop seal steam flowrate, WL.S~is controlled, first of all, by the differential pressure between the steam source (the mixture level where steam'leaves the two-phase mixture) and the steam sink (pump outlet where steam is condensed), A p C O L D -- ApHOT, such that
WL.SffiA~2~(ApcoL v -- APHOT)/K, TIME
Is)
Fig. 8. Cold-leg l i q u i d level n ~ w i t h 3-bean) s a m m a - r a y dee~itometer (top, u o c ~ t y - + 0 . 0 0 5 m), a n d f l u i d t e n , per-
ature measured at leg pipe top immediately downstream of pump outlet (bottom, uncertainty = + 2 K).
where A is the loop cross-sectional area, 0o is the steam density, and K is the overall pressure loss coefficient for the steam-t'filed portion of the loop. The above equation neglects the "capacitance" effects of the steam-filled volumes in the primary loop upper regions.
436
Y. Kukita et al. / Loop seal clearing and refilling
Such effects may be important during fast transients. An ideal steady state of the system would be represented by
WCORE = WL.S q- WBYPASS, where WBYPASS is the steam flow through the bypasses which was negligibly small as compared with WCORE for the present test, due to the relatively small bypass areas represented in the LSTF. The steam source-to-sink pressure drop, A p C O L D - ApHO-r, is dependent on the coolant inventory distribution in the primary system. The loop hot-side differential pressure, APHoT , consisting mostly of liquid static head, varied largely during the test, responding to the changes in the hot-side coolant inventory. The hot-side mixture level moved between the core upper region and SG inlet plenum. Meanwhile, the cold-side differential pressure, ApCOLD, remained almost constant, as the downcomer liquid level was kept almost stationary at the cold-leg nozzle. The steam flow through the pressure vessel bypasses kept the upper downcomer steam-filled. The measured steam flowrate responded sensitively to the changes in the hot-side coolant inventory, since the steam source-to-sink pressure drop, ApcoLD -ApHOT, was considerably smaller than either ApHOT or APcoLD. The measured loop seal differential pressure (the total of the upflow leg, downflow leg, and pump differential pressures), shown in fig. 7, was about 3 kPa for a steam flowrate of about 0.22 kg/s. The pressure drop occurred mostly in the pump into which liquid may have had penetrated from the cold leg. This sensitivity of the steam flowrate to loop differential pressure appears to explain the observed unstable behavior of the pump CCFL, i.e. the fact that counter-current flow occurred only in a brief transient at the beginning of the refill, and the main part of the refill occurred by a fashion of water "dumping". Coupled system behavior may be responsible for such unstable CCFL breakdown behavior. That is to say, when water begins to pour into the loop seal, this water creates an incremental loop-seal pressure drop, and this reduces the steam flow thereby increasing the liquid downflow rate. Each refill, once started, progressed until the loop seal was completely blocked, probably because of this mechanism. No cases were observed in which the loop seal upflow leg stayed half-filled, being voided steadily as suggested by analysis of Fletcher and Callow [4]. This seemingly unstable nature of CCFL, as well as the considerable fluctuations seen in steam flowrate, made it difficult to define precisely the conditions for the onset of refill, not to speak of the fact that
the critical steam flowrate was comparable to the drag disc measurement uncertainty. The CCFL characteristics are dependent on specific facility (or plant) geometry. For the LSTF, there are two flow area constrictions between the loop seal and cold leg: (i) the reactor coolant pump, and (ii) the venturi flowmeter in the pump suction leg. The pump and venturi geometries are peculiar to the LSTF. The pump area takes a minimum of 0.0090 m2 at the hori-. zontal discharge which opens at the centerline of the cold leg (0.0336 m2 in area). The venturi flowmeter (minimum cross-sectional area = 0.0057 m2) is located in the vertical loop seal upflow leg (0.0222 m 2 in area) beneath the pump suction. Although the venturi flow area is smaller than the pump minimum flow area, there was no indication that CCFL occurred at this location rather than at the pump. As has been mentioned, the first refill following the power reduction initiated at 5010 s when the steam flowrate dropped to about 0.1 kg/s (_+ a measurement uncertainty of 0.05 kg/s). On the other hand, no refilling was seen for steam flowrates above 0.2 kg/s. For instance, after the last loop seal clearing during the power increase (at 3500 s), the loop seal remained clear although the steam flowrate dropped temporarily to about 0.2 kg/s. The following flooding parameters [8] are employed to discuss further the refill onset conditions. Kutateladze parameter:
Ku=;o[p a2 // (go ap)] 1/4 Non-dimensional superficial oelocity o f steam:
J~ =Jo [ oo/( gD Ap ) ]'/2, where )G is the superficial velocity of steam, g is acceleration due to gravity, o is surface tension, At) is the density difference between phases, and D is diameter. The steam flowrates of 0.1 and 0.2 kg/s corresponds respectively to j ~ = 0.54 and 1.1 for the pump minimum cross-sectional area (horizontal flow), and to K u ~ 3.2 and 6.4 for the venturi flowmeter (vertical flow) for a fluid saturation temperature of 400 K typical of this time period. The flooding limit reported in the literature for a horizontal pipe depends on the pipe length-to-diameter ratio, L / D . However, it is not likely that the limiting value of j ~ exceeds 1.0 for any value of L / D . T h e flooding limit for large-diameter vertical pipes reported by Richter [9] is K u = 3.2. Thus, the observed limiting steam flowrate, between 0.1 and 0.2
Y. Kukita et aL / Loop seal clearing and refilling 0.4 ,
,
~
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-
~
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Fig. 10. Cold-leg pressure (top, uncertainty ffi + 10 kPa) and
the difference between the break flowrate and the ECC injection rate (bottom, uncertainty ffi + 0.5 kg/s).
kg/s, appears to be reasonable. The effects of the cold-leg fluid subcooling on the limiting steam flowrate is unknown.
4.3. Cyclic loop seal clearing and refilling The repetitive loop seal clearing and refilling observed for lower core powers involved cyclic variations of many parameters. These parameters included, in addition to the core level and Steam flow rate which have already been mentioned, the primary pressure, break flowrate, coolant mass inventory and its spatial distribution, as well as core inlet subcooling. The difference between the ECC and the break flowrates, i.e. the net changing rate of the coolant inventory, is shown in fig. 10 in comparison with the cold-leg pressure. The clear correspondence between the break flowrate and the primary pressure occurred since only a small system overpressure (50 kPa) was needed for the present break area of 10~ in order to obtain a subcooled-liquid break flow in the balance with an ECC injection rate of 5 kg/s. At the same time, the primary pressure was controlled by the balance between the core steam production and the steam relief through the loop seal. The loop seal steam flow was responsive to the vessel coolant inventory, as has been discussed. This interdependence between the loop seal steam flow and the coolant inventory, coupled by the primary pressure, induced cyclic variations of these quantities. A typical cycle can be divided into the following five stages which are characterized by the behavior of the loop seal steam flow rate, WLS., the time derivative of primary pressure, d p / d t , and the time derivative of the primary coolant inventory, d M / d t.
437
(1) Downflow leg level depression (WL.s.= 0, d p / d t > O, d M / d t < 0). After the loop seal refilled, the coreproduced steam, being unable to escape from the break, pregsurized the primary system. The primary pressurization increased the break flowrate to exceed the ECC injection rate (fig. 10) initiating a net reduction in coolant inventory. The inventory loss led to level depressions in the loop seal and core (figs. 4 and 5), which were followed by clearing of the loop seal upflow leg. (2) Upflow leg clearing (WLs.> WCORe, d p / d t < O ) . After the inception of clearing of the upflow leg, the loop seal steam flowrate increased quickly (fig. 7). The steam relief through the loop seal caused a system depressurization (fig. 10). It is noteworthy that liquid in the loop seal upflow leg was cleared completely (or almost completely) once clearing started (fig. 4). For the peak steam flowrate following clearing (0.8-1.1 kg/s), which was two to three times larger than the steady-state core steam production, the value of Ku for the vertical upflow leg (0.168 m i.d.) ranged from 6.5 to 9.1, i.e. well above the Ku = 3.2 flooding-limit criterion [9]. Likewise, the value of j~ for the horizontal leg (0.168 m i.d.) ranged from 0.79 to 1.1. Thus, the loop seal horizontal leg may have been cleared of liquid almost completely. (No level measurement was done for this leg.) (3) Decay m the loop seal steam flow (WE.s. < WcoltE). The loop seal steam flow decayed quickly from the peak value, which was taken immediately after clearing, and almost stopped within 100-200 s after clearing (fig. 7). This flowrate undershoot below the steady-state steam production occurred primarily because of a decrease in the steam source-to-sink pressure difference, APcoLD --ApHOT , due to an overfeed to the loop hot side which was associated with a decrease in the break flowrate after loop seal clearing. Since the break flowrate became smaller than the ECC flowrate after loop seal clearing (fig. 10) due to system depressurization, the ECC was forced into the vessel core region, and eventually to the hot-leg and SG inlet plenum, increasing the loop hot-side static head, and thereby limiting the loop seal steam flow. Also, core boiling suppression, due to the penetration of subcooled ECC into the core (fig. 6), may have played a secondary role in this decay in loop seal steam flowrate. (4) Initiation of refill (WLs.< Wcop.E, d p / d t > O, d M / d t > 0). Refill initiated after the primary pressure had started recovering from the minimum taken after the loop seal clearing (fig. 10). This suggests
438
}I. Kukita et aL / Loop seal clearing and refilling
that refill occurred after the steam relief through the loop seal had become smaller than the core steam production. This limitation in steam flow was due to the above-mentioned transient overfeed to the loop hot side, and was not due to loss of condensation potential in the cold leg; the cold leg was filled with subcooled water at this stage of the transient (fig. 8). (5) End of refill (WL.s. = 0, d p / d t > 0). The refill ended typically within 20 s after the onset of "dumping'" and the steam flow was blocked completely. The loop seal steam was replaced smoothly by liquid, without exhibiting any non-steady condensation, although the liquid was subcooled by up to 80 K. The loop seal refill resulted in a quick reversal of core flow which was detected from a rise in the core inlet fluid temperature (fig. 6). The loop seal downflow-side level continued rising until the loop seal water head balanced with the vessel riser (core and upper plenum) head. Since the latter was smaller at higher core powers, the loop seal maximum level was also lower. As described above, the cychc loop seal behavior included a transient overfeed to the loop hot side occurring after loop seal clearing. This overfeed resulted in an undershoot in loop seal steam flowrate which allowed the loop seal to refill, even when the core steam production rate was large enough, if it could flow steadily through the loop seal, to keep the loop seal clear. This explains the considerable difference observed in the limiting core power for the onset of loop seal refill between the power increase and reduction cases. In a LOCA transient, the core power decays monotonously with time after scram. However, the loop seal steam flowrate may deviate from the core steaming rate, from time to time, being affected by system transient responses as well as automatic and manual interventions. Thus, there is a possibility that all the loop seals become liquid-filled even at a core power for which the steady-state steam production rate exceeds the limit for loop seal refill. Once this situation is realized, then the subsequent loop seal clearing, which may occur if core boiling persists, may tend to initiate a cyclic loop seal behavior. 4.4. Core dryout behavior During loop seal clearing, the core collapsed level drops to the loop seal bottom height. The minimum core mixture level during this level depression depends on void fractions in the core two-phase mixture. Fig. 11
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Fig. 11. Minimum core mixture level calculated using the Cunningham-Yehvoid fraction correlation in comparison with measured rod-surface dryout level. The time scale is based on the ANS (1979) decay heat power curve.
shows the minimum mixture level, as a function of primary pressure and core power, calculated using the Cunningham-Yeh void fraction correlation [10] which is known to be applicable to PWR rod bundle geometry. The time scale in this figure is based on the ANS (1979) decay heat power curve [11]. The data points shown in this figure are the plots of primary pressure and core power measured at the moment of minimum core collapsed liquid level. These data were taken from two sets of LSTF tests. The first set consists of three loop-seal-refill tests all conducted using similar test procedures for break areas of 5%, 10% (the present test) and 75%, respectively. The second set includes two cold-leg-break blowdown tests that were initiated from a primary pressure of 16 MPa, for break areas of 0.5% and 5%, respectively. For these two full-pressure blowdown tests, the first loop seal clearing occurred at a primary pressure of about 8 MPa. For the 5% break test, both loops were cleared and remained clear until the end of test at 2400 s. However, for the 0.5% break test, only the broken-loop loop seal was cleared, and it was refilled with the ECC. This loop seal was cleared again at a primary pressure of 1.8 MPa. For all the data presented in this figure, the primary loop upper portions were empty of liquid when loop seal clearing occurred, and thus the minimum core collapsed level agreed with the loop seal bottom height. The minimum rod dryout levels in these tests, which were determined on the basis of rod-surface tempera-
Y. Kt&ita et al. / Loop seal clearingand refilling
tures measured at nine discrete elevations along the core, are in reasonable agreement with the predicted mixture level. However, the &Tout level on each rod depended on the linear heating rate, which differed between the three core heating regions. The rod temperature excursions were generally small because of low core power. However, it is interesting to note that the maximum rod surface superheating clearly was not dependent on core power. Because the core level depression was slower and deeper for lower core power, the core upper regions were uncovered for a longer time before the core level recovered (fig. 11). There is a possibility that the core level depression occurs almost statically due to: (i) the bypass flow which may relieve most of the core steam production for extremely low core power [4], or (ii) condensation of steam on the SG primary side. The latter case has been simulated in one of the LSTF tests [12], and comparatively large heat-ups (up to 180 K above the saturation temperature) have been observed.
5. Discussion The present test results, although valuable for understanding the LOCA loop seal behavior qualitatively, are not applicable directly to real plant responses because of the pump geometry and loop seal leg cross-sectional area which were not scaled properly. The CCFL characteristics of full-size pump and legs, including liquid subeooling effects, should be understood to predict F W R loop seal behavior. However, it seems encouraging that the p ~ s e n t test results indicate that the loop seal refdl is not dictated by static CCFL but largely affected by system coupled, dynamic responses which subscale facilities would simulate. For a quantitative prediction of real plant responses, and for a defmition of the possible worst conditions for PWR core cooling during post-blowdown phases, computer codes which are able to deal with those phenomena observed in the present tests are needed. Existing LOCA analysis codes generally have limitations in calculating low-pressure phenomena, including condensation on the steam-water interface. Data from the present tests will provide a basis for verification of codes. One of the most important differences between the real PWR and present test geometries is in the number of primary loops. However, it is not likely that the multi-loop geometry of real plants makes the loop seal responses much more complicated than observed in the test. At such low core power level as represented in the
439
present test, loop seal clearing would occur for only one of the loops, and the other loops would remain sealed without contributing to steam relief from the systen~ It is unknown, however, whether or not any particular loop (broken or intact) would be cleared preferentially. The present paper focused on results for a 10~ break test, as no significant qualitative difference has been found in the test results obtained so far for break areas ranging from 5 to 75~. The primary effect of the break area was that the time-averaged primary pressure was lower for a larger break area, provided that the ECC injection rate was the same.
6. Smamary The basic mechanisms of the loop seal refilling phenomenon during a PWR small-break LOCA have been described by analyzing the results of LSTF cold-leg break tests conducted for low-power, low-pressure conditions representative of LOCA post-blowdown phases. These tests were the first system-effects tests ever conducted on this phenomenon. They demonstrated that the loop seal behavior is governed by the time-dependent integral system responses. Inadequate core cooling situations following loop seal refilling, as predicted by a previous computer code analysis for a PWR, were actually observed in the tests. The major observations from the tests are summarized as, follows. The loop seal refilled with the ECC when the steam flowrate through the loop seal was insufficient to prohibit this from occurring. The onset of ref'fll was affected by several factors including the sensitivity of the loop seal steam flowrate to the coolant inventory distribution in the primary system. The most interesting phenomenon observed in the tests was cyclic repetition of loop seal clearing and refilling. This occurred because of the coupled responses of the system parameters, inelliding: the coolant inventory and its spatial distribution; the steam release rate through the loop seal; and the primary pressure. The loop seal clearing phase of each cycle involved a core liquid level depression which uncovered the core upper regions temporarily.
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The authors gratefully acknowledge the cooperation of the USNRC in performln$ this work.
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Y. Kukita et al. / Loop seal clearing and refilling
References [1] H. Tuomisto and P. Kajanto, Two-phase flow in a fullscale loop seal facifity, Nucl. Engrg. Des. 107 (1988) 295-305. [2] K. Tasaka et al., The results of 0.5% PWR small-break LOCA tests in ROSA-IV LSTF break location parameter tests series, Proc. 15th Water Reactor Safety Information Mtg., Gaithersburg (Oct. 1987). [3] ROSA-IV Group, ROSA-IV Large Scale Test Facility (LSTF) System Description, JAERI-M 84-237, Japan Atomic Energy Research Institute (1984). [4] C.D. Fletcher and R.A. Callow, Long-term recovery of pressurized water reactors following a large break loss of coolant accident, Nucl. Engrg. Des. 110 (1989) 310-328. [5] V.H. Ransom et al., RELAP5/MOD2 Code Manual, Vols. 1 and 2, NUREG/CR-4312, EGG-2396, Idaho National Engineering Laboratory (1985). [6] N. Zuber, Problems in modeling small break LOCA, NUREG-0724, U.S. Nuclear Regulatory Commission (1980).
[7] K. Tasaka et al., Loop seal clearing and refilling during a PWR small-break LOGA, Proc. 16th Water Reactor Safety Information Mtg., Gaithersburg (1988). [8] GB. Wallis, One-Dimensional Two-Phase Flow (McGraw-Hill Book Co., New York, 1969). [9] H.J. Richter, Flooding in tubes and annuli, Int. J. Muhiphase Flow 7 (1981) 647-658. [10] J.P. Cunningham and H.C. Yeh, Experiments and void correlation for PWR small break LOCA conditions, Trans. Am. Nucl. SOC. 87 (1973) 453~ [11] American Nuclear Society, American National Standards for Decay Heat Power for Light Water Reactors, ANSI/ANS-5.1-1979 (1979). [12] Y. Kukita et al., Intentional ccx~lant system depressurization: experimental studies in the ROSA-III and ROSA-IV Programs, Presented at the CSNI Specialist Meeting on Intentional Coolant System Depressurization, Garching, F.R.G. (1989).