Intentional Depressurization of Steam Generator Secondary Side during a PWR Small-Break Loss-of- Coolant Accident

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1 Journal of Nuclear Science and Technology ISSN: (Print) (Online) Journal homepage: Intentional Depressurization of Steam Generator Secondary Side during a PWR Small-Break Loss-of- Coolant Accident Hideaki ASAKA & Yutaka KUKITA To cite this article: Hideaki ASAKA & Yutaka KUKITA (1995) Intentional Depressurization of Steam Generator Secondary Side during a PWR Small-Break Loss-of-Coolant Accident, Journal of Nuclear Science and Technology, 32:2, , DOI: / To link to this article: Published online: 15 Mar Submit your article to this journal Article views: 236 View related articles Citing articles: 4 View citing articles Full Terms & Conditions of access and use can be found at

2 journal of NUCLEAR SCIENCE and TECHNOLOGY, 32[2], pp (February 1995). 101 Intentional Depressurization of Steam Generator Secondary Side during a PWR Small-Break Loss-of-Coolant Accident Hideaki ASAKA and Yutaka KUKITA Japan Atomic Energy Research Institute* (Received October 26, 1993) The consequence of intentional depressurization of the steam generator (SG) secondary side during a small break loss-of-coolant accident (LOCA) in a pressurized water reactor (PWR) was studied using the ROSA-IV Large Scale Test Facility (LSTF). The LSTF is a 1:48 volumetrically scaled full-height model of a Westinghouse-type PWR. The experiment simulated a 0.5% cold leg small-break LOCA with total failure of high pressure injection (HPI) system. The SG secondary-side atmospheric relief valve (ARV) was latched open after the core top region was uncovered. Then, the primary pressure closely followed the secondary pressure and dropped to the accumulator (ACC) injection pressure of 4.51MPa before the core became severely overheated. A post test analysis, which was performed using the RELAP51 MOD2 code with modifications made by the authors, predicted well the primary and secondary side responses during this experiment. KEYWORDS: PWR type reactors. reactor, steam generators, safety, thermal hydraulics, system depressurization, operator action, loss-of-coolant, ROSA-IV, LSTF, safety analysis, RELAPSIMOD2, computer codes, pressure dependence I. INTRODUCTION During a loss-of-coolant accident (LOCA) in a pressurized water reactor (PWR), the emergency core cooling system (ECCS) is actuated automatically to replenish the reactor coolant inventory. For a Westinghousetype PWR, the ECCS consists of three injection systems, which are, in the order of the injection setpoint pressure : high pressure injection (HPI) system ; accumulator (ACC) injection system ; and low pressure injection (LPI) system. These injection systems are actuated successively as the primary pressure decreases. There is, however, a possibility that the reactor coolant inventory continues to decrease if HPI system should totally fail to start during a small-break LEA (SBLOCA) with a break discharge flow rate too small to depressurize, by itself, the primary system to the ACC injection pressure (typically 4.5 MPa). Thus, this no-hpi SBLOCA scenario, though very unlikely, deserves careful examination. For such a no-hpi SBLOCA event, an intentional depressurization of the steam generator (SG) secondary side, by relieving steam through the atmospheric relief valves (ARVs), is one of the promising means to lower the primary pressure to the ACC injection pressure. The advantage of this procedure over the primary-side steam relief, which also results in depressurization, is that it does not cause loss of the reactor coolant inventory. When auxiliary feedwater (AFW) is available, the secondary side depressurization would be performed in the feed-and-bleed mode, i.e., concurrent steam relief from and cold water supply to the SG secondary side. The effectiveness of the SG secondary side depressurization procedure was tested in the SBLOCA experiments conducted in the Semiscale facility( ). These experiments showed that the depressurization efficiency increased with decrease in the primary side inventory. A more recent experiment at the BETHSY facility simulated a depressurization initiated * Tokai-mura, Ibaraki-ken I. -15-

3 ~~ - ~ LSTF Nucl. Sci. Technol., after the onset of core uncovery'*). This experiment has been analyzed using several computer codes in an international framework. The analyses indicated weakness of certain codes in predicting the primary-side water inventory redistribution and SG heat transfer associated with the secondary-side depressurization. The present paper describes experimental results and RELAP5 code analysis results for an experiment conducted in the ROSA-IV Large Scale Test Facility (LSTF). The experiment, conducted about two years before the above BETHSY experiment, also represented a depressurization which was initiated after the onset of core uncovery. The major difference between the LSTF and BETHSY experiments was that the former depressurized only one of the SGs. This resulted in greater heat loads to the affected SG and more complicated, asymmetric responses in the primary side than the concurrent depressurization of all the SGs as simulated in the BETHSY exper imen t. A post-test analysis was performed using the RELAP5/MOD2 code. The objectives of this analysis were: to assess the code's capability in predicting the thermal-hydraulic responses to the simulated secondary-side depressurization procedure, and to evaluate analytically those quantities which were not measurable in the experiment. An extensively modified version of RELAP5/MOD2, including revised break flow models and interfacial-drag calculation scheme developed by the authors, was used. These experimental results and post-test analysis results are presented and discussed in the following chapters. II. EXPERIMENT 1. Experimental Facility The LSTFt3) is a 1/48 volumetricallyscaled, full-height, full-pressure model of a Westinghouse-type 4-loop (3,423 MWth) PWR. Figure 1 shows a schematic of the LSTF. The four primary loops of PWR are represented by two equal-volume (2/48 scale) loops. LOOP - A LOOP - 8 Steam Generator Accumulator Coolant Pump. Upflow Side \g, Relief Valve I ARV 1 Fig. 1 Schematic of ROSA-IV Large Scale Test Facility Each loop contains an active SG which consists of 141 U-tubes of 19.6mm I.D. and 25.4 mm 0. D. The elevations of major components are preserved full-scale to simulate natural circulation phenomena. Major design characteristics of the LSTF are compared to those of a typical PWR in Table 1. Table 1 Major characteristics of LSTF PWRt PWR/LSTF Fluid volume (m') Number of fuel rods Core height (m) Downcomer gap (m) Hot leg diameter D (m) length L (m) LIDO.: volume (my) Number of loops Number of SG U-tubes 141 3,382 1/24 Average length of SG U-tube (m) t l'vpical Japanese-built 4 loop PWR (LSTF reference I'WR) 16 --

4 Vol. 32, No. 2 (Feb. 1995) Experimental Conditions The depressurization procedure was simulated in a cold-leg break LOCA experiment. The test initial steady state was established for a primary pressures of 15.6 MPa and hot leg and cold leg fluid temperatures of 598 and 562K, respectively. These pressure and temperatures are representative of PWR nominal operating conditions. The break was simulated using a 7.2mm I. D., sharp-edged, 1-mm-thick orifice plate. The orifice was mounted on the bottom of the cold leg pipe, flush with the pipe inner surface. The orifice provided a break area of of the 1/48-scaled PWR cold leg cross-sectional area, representing approximately a 2-inch break in the reference PWR. The break flow rate is estimated from the liquid level in a catch tank. The uncertainty of time-integrated break flow measurement is estimated to be less than 1%. The secondary ARV had opening and closing setpoint pressures of 8.02 and 7.82 MPa, respectively. The ARV was simulated using a 19.4 mm I. D., sharp-edged 1-mm-thick orifice plate. This orifice provided a steam flow rate of 2.8 kg/s at a secondary pressure of 8.0MPa. This flow rate corresponds to about of the 2/24-scaled total capacity ARV Opened ARV Cbnd l2mx)rl Core Powu Trippd of the ARVs on one typical PWR SG. The test represented total failure of the HPI and auxiliary feedwater (AFW) systems, as well as loss-of-offsite power concurrent with scram. 3. Experimental Results ( 1 ) Before Initiation of Secondary Depressurization Operation (0 to 2,140s) The pressures in the primary system (pressurizer) and in the secondary system (SG steam 16 a P e a Data ---- ELAP5(Modifiedl RELAP5 (OrlpMl I loo ooo ooo Time [s) (a) Secondary side pressures Fig. 2 Measured and predicted broken loop (SG-B) secondary side pressures and primary side pressures The break flow, shown in Fig. 3, changed from a single-phase liquid flow to a two-phase flow at about 700s, immediately after the appearance of vapor in the cold legs. This transition was known from the decrease in the break flow. The two-phase discharge 17 - PJN opened (2140s) I Tb AN Cbd Core Par 45MPa Braolr Mlve cbsed ( continued until the cold leg became empty of liquid, at 1,880 s, after loop seal clearing (i.e. clearing of liquid in the crossover leg by steam toward the cold leg) occurred in the broken loop but not in the intact loop.

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6 Vol. 32, No. 2 (Feb. 1995) 105 SG condensation was continuing since the primary pressure remained higher than secondary pressure, even after high-enthalpy vapor discharge from the break began as a consequence of the loop seal clearing. The condensate in the crossover leg was forced into the upflow leg by the steam flow toward the break. That is, the steam flow through the crossover-leg, that was initiated after the loop seal clearing, was not blocked completely by the condensate. Instead, this steam flow passed through the liquid condensate in the upflow leg, forming a bubbling two-phase mixture as shown in Fig. 7(a)-(c). 101 Experiment I bl RELAP 5 [Modified 1 I c 1 RELAP 5 I Original ) ( 2 ) After Initiation of Secondary Depressurization Operation (2,140 to 2,600 s) The secondary ARV was latched open for the broken loop SG at 2,140s with an intention to depressurize the primary system to the ACC injection pressure. The steam discharge through ARV depressurized and cooled the SG secondary side at a rate of about 300 K/h. The primary pressure closely followed the secondary side pressure drop and reached ACC injection pressure of 4.51 MPa at 2,600 s as shown in Fig. 2(b). The secondary depressurization resulted in enhanced steam condensation rates on the SG U-tube primary side. Consequently, the broken-loop crossover leg was refilled by condensate, in both the upflow and downflow sides, as shown in Fig. 4. Meanwhile, the water inventory in the intact-loop crossover leg decreased slowly, probably because of flashing of water associated with the depressurization. Also, there was liquid accumulation on the broken-loop SG U-tube upflow side as shown in Fig. 8, since counter-current flow limiting (CCFL) controlled the draining of the condensate into the hot leg. This situation is illustrated in Fig. 9(a), (b). E n - Data RELAP5 (Modified RELAP5 (Oripinal I Fig. 7 Measured and predicted primary side water distributions immediately before secondary depressurization operation (Time : 2,140 s) The crossover-leg differential pressure, created by this condensate accumulation, accelerated the progression of the second drop in the core collapsed liquid level. The core boiloff resulted in a continuous core uncovering which started at 1,930s, while the primary pressure was still maintained at about 8 MPa. c n Due to CCFL Depretruriiotion. Time Is1 Fig. 8 Measured and predicted SG-B U-tubes upflow side differential pressures Overall, thus, there was a relocation of the primary coolant inventory from the pressure vessel towards the broken-loop crossover leg and SG primary side. This relocation, together with the continuing loss of coolant through the break, caused the vessel mass inventory to decrease with time. 19

7 106 J. Nucl. Sci. Technol., (a1 Experiment (b) RELAP 5 ( Modified) Fig. 9 Measured and predicted primary side water distributions during secondary depressurization operation (Time : 2,400 s) The secondary depressurization also increased the reflux flowrate from the brokenloop hot leg into the core. This reflux flow slowed, or quenched, the heatups at the core top regions, as shown in Fig. 6, whereas the core level drop was progressing as shown in Fig. 5 and dryout was extending to the lower core elevations as shown in Fig. 6. At 2,600 s after break (i.e. 460 s after the initiation of secondary side depressurization), the primary pressure was close to the ACC injection pressure of 4.51 MPa. However, the test was terminated at this time, by turning off the core power and closing the break valve, before the injection started, to avoid thermal shock to the heated pressure vessel internal structures. In summary, this test evidenced the effectiveness of the secondary side depressurization operation for depressurizing the primary system and thereby promoting the injection from ACC tanks during a cold-leg small break LOCA. At the same time, the test has shown that this operation leads to complicated coolant redistribution in the crossover leg and SG primary side, concurrent with a decrease in the vessel mass inventory. Further detail of such coolant redistribution is discussed in the next section based on the code calculation of steam flow through the loop. fl. POST TEST ANALYSIS WITH RELAP5IMOD2 CODE 1. Code Modifications The RELAP5/MOD2 code"), a computer code for best-estimate simulation of thermalhydraulic behavior of light water reactors, is based on one-dimensional, two-fluid, nonhomogeneous, non-equilibrium models. For the analysis of a small-break LOCA, it is essential to predict accurately the loss of reactor coolant inventory from the break as well as the spatial distribution of the residual coolant inventory in the reactor coolant system. These are crucial for prediction of the reactor core cooling. The RELAP5/ MOD2 code has been modified by the authors based on analyses of both integral and separate effect experiments. The code changes pertinent to the present analysis include : (1) Use of the Bernoulli equation for singlephase liquid discharge through a sharpedged orifice(6) (2) Use of the maximum bounding critical flow model for two-phase discharge through a sharp-edged orifice(') (3) Use of vapor pull-through and liquid entrainment models for sharp-edged orifice developed by Yonomoto & Tasaka'?) (4) Use of an interfacial drag calculation scheme in which the junction interfacial drag coefficient is equated to that of the adjacent upstream The modifications (1) through (3) above are related to prediction of break flow rate, and (4) to prediction of the coolant inventory distribution. These code changes, detailed in referen~e'~), enhanced considerably the predictive capability of the code for small break LOCA transients('o). The purpose of this analysis was twofold : to assess the code's capability, and to evaluate quantities which were not directly measured in the experiment. In this paper, the results related to the second objective above are mainly discussed

8 Vol. 32, No. 2 (Feb. 1995) Comparison of Experimental deck. The pump coast down speed and core Results and Predictions power decay after scram signal were deter- The LSTF nodalization model of RELAP5 mined according to experimental data. Singleis shown in Fig. 10. The model consists of phase steam discharge from the break and 192 volumes, 199 junctions and 83 heat slabs. ARV was calculated using the RELAP5 crit- The initial steady state condition in the ex- ical flow model(4) with a discharge coefficient periment was calculated well using this input of 0.84'") Junction mzza Heat Slab Fig. 10 RELAP5 nodarization model for LSTF ( 1 ) Before Initiation of Secondary Depressurization Operation (0 to 2,140s) The analysis was performed using both the original and modified RELAP5/MOD2 codes. Figure 2(a) and (b) show comparisons of measured and predicted primary and secondary pressures. The primary depressurization became slower as the pressure approached the secondary side pressure that cycled between the ARV opening and closing setpoints (8.02 and 7.82MPa). The predicted changes in the secondary side pressure were smaller than the experimental data since the above difference between the ARV opening and closing setpoints was not modeled. This difference does not affect the calculation of the timeaveraged heat transfer from the primary to the secondary side. The primary pressure remained slightly higher than secondary pressure as was observed in the experiment. Figure 7(a)-(c) show the predicted and calculated primary coolant inventory distributions at 2,140 s, immediately before the forced opening of the ARV in the experiment. The original (unmodified) code overpredicted the primary mass inventory and also failed to predict its distribution at this time, while the modified code predicted both reasonably well. Thus, the analysis of the subsequent phase, after 2,140s, was performed by the modified code alone. The major problems seen in the originalrelap5 code calculation was a general underprediction of the break flow rate and a delay in the draining of SG primary side liquid, in particular for the SG upflow -21-

9 108 /. Nucl. Sci. Technol., side (i.e. the hot-leg side). In the calculation with the modified RELAP5 code, the ARV in the broken-loop SG was latched open at 2,140s as was done in the experiment. At this time, the pressure vessel upper regions, hot legs and SG primary side were all steam filled. The intact-loop crossover leg was liquid-full up to the cold leg level. The broken-loop crossover leg was only partially filled as has been mentioned. The secondary side level in the broken-loop SG was 6.6 m (vs. 7.2 m in the experiment). Reflux condensation was continuing for both the intact- and broken-loop SGs in the experiment, since the steam relief through the break of about 0.42kg/s shown in Fig. 3 was less than the calculated steam production rate in the core of about 0.83 kg/s. The core power at this time (1,200 kw) corresponded to 1.7% of the 1/48 scaled PWR rated power. The steam condensation occurred on both the upflow and downflow sides of the U-tubes. The condensate thus drained to both the inlet and outlet of the U-tubes, forming a countercurrent two-phase flow in the U-tube upflow side. The modified code predicted well the coolant inventory distribution at the beginning of depressurization (2,140s), as shown in Fig. 7, including the crossover leg water levels. The predicted void fraction in the two-phase mixture in the upflow side of the broken-loop crossover leg, which was filled by two-phase mixture up to the top as shown in Fig. 7, was about The calculated steam flow rate through the mixture (not shown) was equal to the break flow rate, since the system conditions were nearly steady at this time. This steam flow rate corresponded to a Kutateladze parameter Ku of 2.67, in the crossover leg piping and was lower than the flooding limit for large-diameter vertical pipes of Ku=3.2. Here, it should be noted that the LSTF represents two PWR loops by one ; the crossover leg cross-sectional area is about 1/22-scaled whereas the break area is 1/48 scaled. Therefore, in the 4-loop PWR, if one of the loops is cleared, the steam velocity in the crossover leg would be higher than the present experimental results, and thus the crossover leg void fraction would be higher. This void fraction affects the core collapsed liquid level during this phase of the transient. Therefore, it is an interesting problem to predict void fractions in the plant crossover leg which has a diameter of about 0.7m, i.e. beyond the range of most of the available void fraction data. (2) After Initiation of Secondary Depressurization Operation (2,140 to 2,600 s) The discharge rate shown in Fig. 11 through the ARV was underpredicted by 17 % ; however the secondary side depressurization rate was underpredicted only slightly as shown in Fig. 2(a). The primary pressure, which followed closely the secondary side depressurization, was predicted well as shown in Fig. 2(b). It is most likely that the secondary side pressure was predicted well, despite the underprediction of the steam relief rates, because the secondary mass inventory at the beginning of the depressurization was underestimated by 8.3% in the calculation. The underprediction of the secondary inventory may have been caused by an overprediction of the primary-to-secondary heat transfer which was caused by the underprediction of the time-integrated break flow rate, as shown in Fig. 3, and the system heat losses in the experiment which is estimated to be about 160 kw during the initial steady state but modeled to be 39 kw in the calculation. Although the good prediction of the secondary ---- RELAP 5 I Modified ) ARV Opened (2140 5) z u ? Time ($1 Fig. 11 Measured and predicted ARV flow rates

10 Vd. 32, No. 2 (Feb. 1995) 109 pressure may have been achieved by the compensation of these errors, it provided a correct boundary condition to the calculations of SG heat transfer and primary side responses which are of primary interest in the present study. The steam flow into the broken-loop SG U-tubes increased after the initiation of depressurization in the analysis. The differential pressure between the U-tube inlet and top, which was nearly zero before the depressurization, increased in the experiment because of liquid holdup which resulted from CCFL as has been mentioned. Although counter-current flow was present also in the SG inlet plenum and hot leg, there was no indication that CCFL occurred at these locations either in the experiment or analysis. Steady-state U-tube CCFL experiments have been conducted at the LSTF(12). In these experiments, the primary pressure was kept constant and the core power was changed in small steps to define the onset conditions for CCFL. For pressures ranging from 1 to 7.3 MPa, the onset of CCFL was observed for.jil/' of The calculated value of jil'' ranged from 0.59 to 0.66 during the depressurization, i.e. the predicted steam flow rate was close to, or higher than, the CCFL threshold. The calculated U-tube differential pressure was almost zero during the depressurization transient whereas a liquid holdup occurred in the experiment as shown in Fig. 8. This resulted from that the RELAP5/MOD2 interfacial drag models do not have any special consideration for the CCFL phenomenon. Experiments conducted at the LSTF, in steady-states, have shown that CCFL in the hot leg bend, beneath the SG inlet plenum, occurs at ji (based on the hot-leg diameter of 0.207m) of about 0.3 when the hot leg is uncovered above the vessel upper plenum mixture level. The maximum value of the calculated steam flow rate in the hot leg corresponded to j; of about 0.17 and thus did not reach the above threshold value. After the initiation of the depressurization, the refill of the upflow side of the broken- loop crossover leg progressed at a higher rate in both the experiment and analysis. The steam flow through the two-phase mixture was calculated to stop by about 2,330s, because the SG condensation rate exceeded the core steam production. This decrease in vapor flow rate caused the liquid in the crossover-leg upflow side to flow back into the downflow side, starting at about 2,200s as shown in Fig. 4. In the experiment, as liquid condensate was held up in the U-tube upflow side, the crossover-leg downflow side level increased to counter-balance the U-tube differential pressure created by the liquid holdup as shown in Fig. 9(a). Thus, the differential pressure between the vessel hotleg and cold-leg nozzles, which had an effect to depress the core liquid level relative to the downcomer level, decreased with time. The core level drop was slowed after the initiation of the secondary depressurization, since the hot leg-to-cold leg differential pressure was reduced as the crossover-leg differential pressure decreased. In the calculation, the coolant inventory in the broken-loop crossover-leg and U-tube was underpredicted after the secondary depressurization as shown in Fig. 9(b) because of not predicting the occurrence of U-tube CCFL. This resulted in overpredicting the core coolant inventory as shown in Fig. 5. A suitable CCFL model applicable to the geometry of SG U-tubes inlet is required to predict correctly the coolant inventory distribution. IV. CONCLUSIONS An experiment on intentional secondary system depressurization during a 0.5 % cold leg break LOCA in a PWR was conducted by using the integral test facility LSTF. This experiment was analyzed with use of the RELAP5/MOD2 code. The major results obtained from the experiment and analyses are as follows: (1) In the experiment, the primary-side pressure closely followed the secondaryside depressurization that was initiated by opening the ARV on the broken-loop SG, and dropped to the accumulator in

11 110 J. Nucl. Sci. Technol., jection pressure of 4.51 MPa before the core became severely overheated. (2) Also, the secondary depressurization led to a relocation of the primary mass inventory ; the vapor produced in the core condensed in the SG U-tubes and the condensate was accumulated in the crossover leg as well as in the SG U-tubes in the affected loop (broken loop). Thus, pressure vessel mass inventory decreased and core dryout progressed with time. However, meanwhile, the core heatup was slowed by an increase in the reflux flowrate from the broken-loop hot leg onto the core. (3) The RELAP5/MOD2 code, which included modification to break flow models and to interfacial drag model, predicted well the primary side depressurization behavior. (4) The analytical results show that the U-tube steam flow reached the CCFL condition during the depressurization. This was consistent with the experimental results which indicated liquid holdup on the U-tube upflow side. However, such liquid holdup was not predicted by the code due to lack of capability to predict the CCFL phenomena. [NOMENCLATURES] D : Hydraulic diameter g: Acceleration due to gravity Volumetric flux (=Vkak) Jk: ji : Nondimensional volumetric flux (=J& d (ApgD) P2) Ku: Kutateladze number (=Jg[p4/(goAp)]'/4) V: Velocity uk: Void fraction dp : Density difference u: Surface tension (Subscripts) g: Gas phase 1: Liquid phase k: g or 1 (The original version of this paper was presented at the ICONE-1, Tokyo, 1991, and is published in this Journal after formal review by our Editorial Committee.) -REFERENCES- (I) LOOMIS, G. G. : NUREG fcr-4945, EGG-2509, (1987). (2) CLEMENT, P., CHATAING, T., DERGAZ, R.: OECD f NEA/CSNI/R(92)20, (1992). (3) ROSA-IV Group : JAER I-M , (1985). (a) DIMENNA, R. A., et al. : NUREGfCR-5194, EGG-2531, (1988). (5) FAUSKE, H. K. : Heat Transfer-Cleveland, AIChE Symp. Ser., 59, Vol. 61, (1965). (6) ARDRON, K.H., FURNESS, R.A.: Nucl. Eng. Des., 39, (1976). (i) YONOMOTO, T., TASAKA, K.: J. Nucl. Sci. Technol., 25[5], (1988). (8) KUKITA, Y., et al. : Heat Transfer-Pittsburgh (LYCZKOWSKI, R. W. ed.), AIChE Symp. Ser. 257, Vol. 83, (1987). (9) ASAKA, H., et al.: Exp. Therm. Fluid Sci., 3, (1990). (10 ASAKA, H., et al.: Proc. 1st Int. Conf. on Supercomputing in Nuclear Applications, Mito, Japan, Mar , p (1990). (11) SALmr, D.W.: Nucl. Sci. Eng.. 00, (1984). (14 YONOMOTO, T., et al.: Proc. ANS Int. Topical Mtg. on Safety of Thermal Reactors, Portland, Oregon, USA, July 21-25, p (1991)