Thermal hydraulic phenomena related to small break LOCAs in AP1000

Progress in Nuclear Energy 53 (2011) 407e419

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Review

Thermal hydraulic phenomena related to small break LOCAs in AP1000 W.W. Wang, G.H. Su*, S.Z. Qiu, W.X. Tian State Key Laboratory of Multiphase Flow in Power Engineering, Department of Nuclear Science and Technology, Xi’an Jiaotong University, Xi’an 710049, China

a r t i c l e i n f o

a b s t r a c t

Article history: Received 6 August 2010 Received in revised form 18 February 2011 Accepted 18 February 2011

Since the TMI accident in 1979, a lot of attention in the nuclear engineering field has been drawn to the small break LOCA issue, around which plenty of work has been done both experimentally and theoretically. Subsequent reactor designs have also been greatly influenced. As a Generation III þ reactor that received Final Design Approval by U.S. NRC, AP1000 employs a series of passive safety systems to improve its safety. However, the thermal hydraulic phenomena related to small break LOCAs in AP1000 have not been fully understood and further studies are still required. This paper investigated the available literature and information on thermal hydraulic phenomena that occur during small break LOCAs in AP1000, which included the critical flow, natural circulation, countercurrent flow limiting, entrainment, reactor vessel level swell, direct contact condensation and thermal stratification. In particular, the physical phenomena, theoretical and experimental research conducted in the past few decades, and prediction models as well as their comparison and evaluation for the thermal hydraulic phenomena related to the small break LOCAs in AP1000 were concluded. Ó 2011 Elsevier Ltd. All rights reserved.

Keywords: AP1000 Small break LOCA Thermal hydraulic phenomena Nuclear safety Passive safety

1. Introduction

2. AP1000 description

The small break LOCA issue was originally presented in public hearings by former U.S. Atomic Energy Commission in early 1970’s, which resulted in ECCS Rule 10 CFR 50.46 and associated Appendix K (Nusret, 2007). Since the TMI-2 accident in March 1979, the small break LOCA issue and related nuclear power plant safety have attracted attentions of researchers in the nuclear engineering field (Kawanishi et al., 1991). Up to now, plenty of integral test facilities have been built to simulate PWRs and to acquire experimental data during small break LOCA for safety analysis, including SEMISCALE, LOFT and MIST in USA, LSTF and EOS in Japan, BETHSY in France, SPES in Italy, PKL in Germany and ATLAS in Korea (Choi et al., 2008; Kawanishi et al., 1991). Meanwhile, related safety analysis codes (RELAP, TRAC, TRACE, RETRAN, CATHARE, CANAC, ATHLET, NOTRUMP, etc.) have been developed or updated for small break LOCA transient analysis. Although a great deal of work related to traditional PWRs has been done in the past few decades, the mechanism that controls small break LOCA in AP1000 has not been clearly understood. In this paper, thermal hydraulic phenomena occurring in the course of AP1000 small break LOCA, related physical models, and the experimental and theoretical studies will be discussed as follows.

AP1000 is a two-loop, 3400MWt Westinghouse-designed PWR which received Final Design Approval by U.S. NRC (Nuclear Regulatory Commission) in 2004 and certified in 2006 as a Generation III þ reactor (Wright, 2007). Its design is similar to AP600, but has a number of component changes in size and capacity to realize a significant promotion in the core power. Fig. 1 shows the AP1000 passive core cooling system schematically (Schulz, 2006), which consists of the following aspects (Wright, 2007):

* Corresponding author. Tel./fax: þ86 29 82663401. E-mail address: [email protected] (G.H. Su). 0149-1970/$ e see front matter Ó 2011 Elsevier Ltd. All rights reserved. doi:10.1016/j.pnucene.2011.02.007

(1) Two core makeup tanks (CMTs) provide relatively high flow borated water in a long time at any pressure. (2) Two pressurized accumulators (ACCs) provide high flow borated water in a short time after system pressure drops below 4.83 MPa (700 psia). (3) An in-containment refueling water storage tank (IRWST) provides low flow borated water in a longer time after system pressure drops to near the containment pressure. (4) A passive residual heat removal heat exchanger (PRHR HX) with C-shape tubes is located inside the IRWST to provide passive energy removal through a natural circulation path connecting the pressurizer-side hot leg and the steam generator exit plenum. (5) The automatic depressurization system (ADS) consists of two trains of depressurization valves with the first three stages coming from the pressurizer and the fourth-stage valves coming from the hot legs (Kemper et al., 1992). Steam from the

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Fig. 1. AP1000 Passive Core Cooling System.

first three stages is quenched in the IRWST as heat sink. Fluid (single phase water or steam or two-phase mixture) from the fourth-stage valves is vented directly to the containment. During a hypothetical small break LOCA, these valves open subsequently to provide a controlled depressurization rate of the primary system.

3. Confirmatory investigations

focusing on the high pressure and depressurization phase and the initiation of IRWST injection. Separate effect tests included PRHR HX tests (Yonomoto et al., 1998), ADS tests, core makeup tank tests (Yonomoto et al., 1997), passive containment cooling system tests, DNB tests and in-vessel core melt retention tests. All the tests mentioned above were conducted during the development phase of AP600. The data obtained from those tests can be used for the code validation of AP1000, since each of them focuses on a certain physical phenomenon which has no connection with the core power.

3.1. Experimental research 3.2. Numerical simulation To assess the design maturity of AP600 and AP1000, Westinghouse conducted a large number of tests on integral and separate effects to obtain data for safety analysis codes. In connection with small break LOCAs in AP600 and AP1000, a series of integral system tests were undertaken. Full-height and full-pressure tests were conducted at the SPES-2 (simulatore per esperienze di sicurezza) facility located in Piacenza, Italy (Friend et al., 1998). The purpose of SPES-2 tests was to obtain thermalehydraulic data at high pressure for the validation of safety analysis codes. The second integral facility APEX (advanced plant experiment) located at Oregon State University emphasized the late phase depressurization, transition to IRWST injection and long-term cooling (Bessette and Marzo, 1999; Welter et al., 2005) during small break LOCA. Besides, U.S. NRC conducted independent integral test programs at ROSA (rig of safety assessment) in Tokai-mura, Japan,

Westinghouse used a one-dimensional, general network code NOTRUMP (Westinghouse Electric Company, 2004), which consisted of a number of advanced features, to analyze small break LOCAs. The version used for AP600 was updated for AP1000 and was further validated against applicable passive plant test data. The main weakness of the code was that it had limited ability in modeling upper plenum and hot leg entrainment and could not predict the core collapsed level accurately. Several typical small break LOCA transients, including inadvertent ADS actuation, 2-inch cold leg break, 10-inch cold leg break and double-ended rupture of the direct vessel injection line, were analyzed using NOTRUMP. Results indicated that smaller break LOCAs exhibited a greater safety margin for core uncovery, and the cladding heat up did not occur.

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U.S. NRC developed its own safety analysis codes to assist in identifying safety issues, such as RELAP5 and TRACE (Stephen, 2007). Liquid entrainment of upper plenum and hot leg were taken into consideration in code development phase. Data obtained from the integral test facility APEX and the separate test facility ATLATS (Air-water Test Loop for Advanced Thermalehydraulic Studies) at Oregon State University were used for codes development and assessment. Calculation of U.S. NRC supported the conclusion that AP1000 would not have core uncovery and cladding heat up in design basis small break LOCAs. Since RELAP5 was designed for short-duration of transients in high pressures (Hassan and Banerjee, 1996), the effectiveness of the prediction in lower pressures with new thermal hydraulic phenomena should be considered carefully.

power drops rapidly to the decay heat level. When the pressurizer pressure further drops below 11.72 MPa (1700psia), the safety systems actuation signal (“S” signal) is activated and causes the opening of the CMT and PRHR HX isolation valves. Thus two natural circulation paths are established immediately, through which the decay heat is removed effectively. The reactor coolant pumps trip after a short delay, and the primary system is cooled by natural circulation. The pressurizer is completely empty of water but full of steam at the end of the blow-down phase. The liquid in the upper plenum and upper head may flash and the upper head may begin to drain (Friend et al., 1998; Wright, 2007). It should be pointed out that all of the phenomena mentioned above besides the core makeup tank and PRHR HX behavior are essentially the same as those of conventional PWRs (Wright, 2007).

4. Small break LOCA chronology

4.2. Natural circulation phase

Small break LOCA transient in AP1000 can be divided into four different phases (Friend et al., 1998; Muftuoglu, 2004; Wright, 2007), namely the blow-down phase, the natural circulation phase, the ADS blow-down phase and the IRWST injection phase. A typical pressure transient for an AP600/AP1000 small break LOCA is shown in Fig. 2 (Banerjee et al., 1998). Each phase of AP1000 small break LOCA chronology characterized by different thermal hydraulic phenomena will be discussed in the followings.

After the reactor coolant pumps trip, the primary system is cooled down by various types of natural circulation that depend on the system inventory, including single-phase natural circulation, two-phase natural circulation and reflux condensation (reflux boiling). During the natural circulation phase, the primary system is at a quasi-steady-state with the steam generator secondary-side. The decay heat is removed through the break, steam generator, PRHR HX and core makeup tank recirculation flow. Once the primary U-tubes of the steam generator begin to drain, the PRHR HX becomes the main heat sink of the reactor coolant system. During the time of the steam generator isolation and ADS activation, the PRHR HX together with the break flow, provides a large heat removal rate (Wright et al., 1996).Heat transfer from the PRHR HX tubes to the IRWST occurs either by free convection or boiling depending on the tube outer wall temperature, tank water temperature and pressure in the vicinity (Friend et al., 1998). The core makeup tanks remain in the recirculation mode until the cold

4.1. Blow-down phase The primary system undergoes depressurization from the operating pressure to the secondary-side pressure of steam generator and then reaches a stable primary pressure. As mass is lost through the break, the pressurizer pressure falls continuously. When the pressure falls below 12.41 MPa (1800psia), a reactor scram occurs (Westinghouse Electric Company, 2004) and the core

Fig. 2. AP1000 small break LOCA pressure transient.

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legs or pressure balance lines begin to flash or drain, after which the core makeup tank level drops continuously. The ADS blow-down phase initiates after the water level drops to its actuation setpoint (67.5%). 4.3. ADS blow-down phase At the beginning of the ADS blow-down phase, the primary system is depressurized through the first three stages of the ADS valves which open subsequently to provide a controlled depressurization. When water level of either core makeup tank drops to 20%, the fourth-stage valves attached to hot legs open and vent directly to the containment. After the ADS valves on the top of the pressurizer open, the pressurizer pressure drops to just above the containment pressure by a sparger head (Takeuchi et al., 1999). The coolant in the primary system begins to flash due to the increasing depressurization rate which causes the rise of the mixture level in upper plenum. After that, the coolant refills the pressurizer and the water level in the pressurizer rises markedly. Once the system pressure drops to 4.83 MPa (700psia), the accumulators begin to inject through the direct vessel injection (DVI) lines. Flow from the core makeup tanks may be reduced or even temporarily stopped due to the backpressure increase of the direct vessel injection line. 4.4. IRWST injection phase After the primary system pressure drops to near the containment pressure, injection from the IRWST initiates, marking the end of the small break LOCA transient and the beginning of the longterm cooling (Wright, 2007). 5. Thermal hydraulic phenomena 5.1. Critical flow The critical flow is important for the safety assessment of a water-cooled nuclear reactor (Ardron and Furness, 1976) because it determines the system inventory loss. The discharge flow makes a great impact on the system depressurization rate, heat transfer in the core and the containment pressurization behavior. It is generally accepted that the critical flow rate depends on the stagnation condition of the fluid, the entrance geometry and the length-todiameter ratio of the test section (Henry, 1970). The homogenous equilibrium model (HEM) treats the twophase mixture as a pseudo-fluid and is used in several system codes, such as early version of RELAP for LOCA analysis (Wallis, 1980). The critical flow rate can be obtained by decreasing the downstream pressure until the maximum flow rate is reached. It agrees well with experimental data for upstream subcooled liquid and long pipe condition where there is sufficient time for achieving thermal equilibrium between phases. Because the homogenous and equilibrium assumption is not physically realistic, the HEM has a poor prediction on the blow-down pipe outlet pressure (Arto, 2008). It overestimates the critical flow in a larger break and underestimates the critical flow in a short pipe (Yoon et al., 2006). Slip flow models including the classical Moody and Fauske models take into account the relative motion between the gas and liquid phase. Among them, Moody model (Moody, 1966) is widely used in reactor depressurization analysis and accepted by U.S. NRC for evaluating the discharge rate after the discharge flow becomes two-phase mixture (U.S. NRC, 1992). It should be added that Moody model overestimates the critical flow after comparing with available blow-down experiment data and thus Moody multiplier

(always less than unity) is used to represent the difference (Ardron and Furness, 1976). Thermal non-equilibrium effect is particularly important in the prediction of critical flow from nozzles, orifices and short tubes (Henry and Fauske, 1971). Henry-Fauske model is based on the condition of short flow path, low inlet quality, high pressure and ignoring the effect of friction force. An empirical parameter N is used to describe the degree of the deviation from equilibrium condition. The main weakness of Henry-Fauske model is that it merely correlates experimental data but does not reflect the nonequilibrium mechanism essentially. However, it is a good way to predict the critical flow in a long pipe (length-to-diameter ratio greater than 12). In the NOTRUMP computer code, Henry-Fauske model, homogenous equilibrium model and Murdock-Bauman model (Murdock and Bauman, 1964) are used for the calculation of ADS critical discharge (Westinghouse Electric Company, 2004). HenryFauske model and homogenous equilibrium model are selected for low and high quality two-phase flow, respectively, with a transition at 10-persent static quality. As for the superheated critical flow vented from the ADS flow path after the core uncovers, it is inappropriate to treat the superheated steam as an ideal gas. MurdockBauman model is used for superheated critical flow modeling by introducing a critical flow function f which is given in the form of a table or a figure. The critical flow function is defined as:

f¼

GT0 P0

(1)

where G is the critical mass flow per unit area, and T0 and P0 stand for the upstream stagnation temperature and pressure respectively. Note that Henry-Fauske model, homogenous equilibrium model and Murdock-Bauman model are also adopted by Korean small break LOCA analysis code CEFLASH-4AS/REM under critical flow condition (Young et al., 1995). The reactor transient analysis codes RELAP5/MOD3.3 and RELAP5-3D are all equipped with Henry-Fauske model and TrappRansom model for critical flow calculation (RELAP5/MOD3.3, 2001; RELAP5-3D, 2001). But Chung (2005) was of the opinion that the Trapp-Ransom model based on the thermal equilibrium might not be applicable to the thermal non-equilibrium condition and showed discontinuity at the transition point from single phase liquid to two-phase mixture. 5.2. Natural circulation Natural circulation provides an efficient way for decay heat removal. In both single-phase and two-phase natural circulation, the heat removal rate is mainly controlled by the mass flow rate. On the contrary, vapor condensation dominates the heat transfer during the reflux condensation phase (Nusret, 2007). The steam generation rate in the core is calculated based on steady-state energy balance (Kukita et al., 1990):


Wcore ¼ Q = hfg þ Dhin

(2)

where Q is the core decay power which can be calculated from the proposed ANS (American Nuclear Society) decay heat standard for light water reactors, hfg is the latent heat under specified pressure conditions, and Dhin is the core inlet subcooling. Francesco and Monica (2002) established a natural circulation map according to the data from the integral test facilities SEMISCALE, SPES, LOBI, BETHSY, PKL and LSTF, in which two-phase natural circulation instability was taken into consideration. The natural circulation flow regime map is shown in Fig. 3. The core

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Since it is difficult to model steam condensation in the steam generator U-tubes, a series of condensation experiments were conducted in the APEX facility to provide data for code simulation (Woods et al., 2009; Woods and Collins, 2009). A catch tank and a cyclone separator were used to collect the condensate in the uphill and downhill side of the steam generator U-tubes. Furthermore, the experimental results were compared with the RELAP-3D simulating results. RELAP-3D has two condensation correlation options, namely the Nusselt correlation for turbulent condensation and the Shah correlation for laminar condensation. The comparison indicated that RELAP-3D overestimated the condensation rate, while the Shah correlation predicted the condensation rate and film Reynolds number closer to the APEX data. 5.3. Counter-current flow limiting (CCFL)

Fig. 3. Natural circulation flow regime map.

flow rate was plotted against the primary inventory. It can be seen from the figure that there exists a peak value of the core flow rate in the two-phase natural circulation region. That is to say, a maximum core flow rate would occur at intermediate system inventory, which was observed in integral test facility experiments and RELAP5 simulation (Duffey and Sursock, 1987). According to Francesco and Monica (2002), the mechanism could be explained as follows: the core flow rate was determined by driving force and resistant force, both of which increased with decreasing primary system inventory. The effects of the driving force and resistant force dominated over the other one in a small and large system inventory, respectively. For that reason, the core flow rate displayed a trend of firstly increasing and afterwards decreasing with the decrease of the primary system inventory. Such a ‘peak’ phenomenon was also observed in the APEX-CE test facility (José, 2004) and the SEMISCALE test facility (Duffey and Sursock, 1987) mentioned previously. According to José (2004), the flow decrease was attributed to the decrease of the distance between the core and the steam generator thermal centers due to drain in long steam generator tubes. As the small break LOCA process continues, flashing and boiling in the reactor coolant system produces sufficient steam (Burchill, 1982).After that, the steam flows into the steam generators and accumulates in the top of the steam generator U-tubes in which the natural circulation eventually ceases (No, 1983; Tasakaa et al., 1988; Kawanishi et al., 1991), because the driving force cannot overcome the resistant force in the natural circulation loop. After the flow stagnation, the mixture levels in both the uphill and downhill side of the steam generator U-tubes drop steadily and the natural circulation turns into the reflux condensation mode (Naugab, 1987). The experimental results of the APEX-CE test facility showed that the long tubes in steam generators drained much earlier than the short tubes did in the course of small break LOCA transient. However, currently in many safety analysis codes, the steam generator U-tube bundle is lumped into a single tube, increasing the prediction uncertainty (José, 2004). During the draining phase, steam from the core will be condensed in the uphill and downhill side of the steam generator U-tubes. The condensate in the uphill side will drain back to the core with the draining water (Naugab, 1987) and the condensate in the downhill side will flow to the downcomer over the reactor coolant pumps. The reflux condensation is characterized by a small mass flow rate and primary to secondary temperature difference (Nusret, 2007), and serves as an important decay heat removal mechanism during the small break LOCA transient.

As small break LOCA processes, the two-phase natural circulation ceases and turns into the reflux condensation mode as mentioned above. During reflux condensation, counter-current flow limiting (CCFL) or the onset of flooding occurs if the steam velocity is sufficient to induce a strong steameliquid interface disturbance. CCFL is characterized by large waves on the steameliquid interface, chaotic flow and increased gas pressure drop (RELAP5/MOD3.3, 2001). CCFL phenomenon determines the maximum velocity of the one phase relative to the other one when the velocity of neither of the two phases can increase further without flow regime change (Wongwises, 1996). CCFL may occur in the following structures where the flow direction or flow area changes: the upper core tie plate, the downcomer, the hot legs, the entrance of the steam generator inlet plenum, the uphill side of the steam generator Utubes, and the pressurizer surge line (RELAP5/MOD3.3, 2001; Azevedo et al., 2007; Solmos et al., 2008). CCFL is an important issue related to the safety analysis of PWRs. During a small break LOCA transient, the horizontal portion of hot legs, the inclined portion of hot legs connected to steam generators, the uphill side of steam generator U-tubes and the pressurizer surge line should be carefully investigated (Naugab, 1987). Over the past few decades, many researchers have investigated CCFL phenomenon experimentally and theoretically (Solmos et al., 2008). Various empirical or semi-empirical correlations were obtained by a numerical fit to experimental data, while in theoretical analysis mechanical modeling (de Bertodano, 1994; Wongwises, 1996; Chun and Yu, 2000) based on certain assumptions were widely developed. 5.3.1. CCFL in hot leg geometry The early research of CCFL phenomenon focused on a single vertical, horizontal or near horizontal pipe. Wallis correlation for small diameter tubes, Kutateladze correlation for large diameter tubes and a form between these two correlations are widely used in current safety analysis codes, such as RELAP5 and TRAC. However, the correlations above cannot accurately reflect the real thermal hydraulic behavior of the hot leg geometry consisting of both a horizontal section and an inclined section. CCFL in the hot leg geometry will limit the reflux flow rate to the reactor vessel and has an unfavorable impact on core cooling and reactor safety. In addition, the CCFL occurrence may increase the pressure drop across the hot legs, and thus increase the upper plenum pressure which may result in the decrease of core/plenum mixture level and fuel heatup (de Bertodano, 1994; Jeong, 2002). According to Ohnuki et al. (1988), the behavior near the hot leg bend dominated the CCFL characteristics of the hot legs. CCFL in a horizontal pipe connected to an inclined riser which is analogous to the real hot leg geometry of a nuclear reactor has been receiving special attention recently

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(Deendarlianto et al., 2008). The typical CCFL test configuration is shown in Fig. 4 (Kim and No, 2002). From the literature survey, it can be inferred that mechanisms governing CCFL phenomenon in the hot leg geometry are roughly classified into two types by the liquid flow rate. At low liquid flow rates, a hydraulic jump is observed. The hydraulic jump is characterized by a sudden transition from supercritical flow (local Froude number greater than unity due to gravity acceleration) to subcritical flow (local Froude number less than unity). In that case, CCFL is initiated by the sudden growth of waves in the bended region which already exist in the pipe before CCFL (Choi and No, 1995). For high liquid flow rates, no hydraulic jump is observed before CCFL, and CCFL is initiated by big waves in the horizontal or inclined part of the pipe at higher gas flow rates (see Fig. 5). Similar experimental results were reported by de Bertodano (1994), Wongwises (1996), Deendarlianto et al. (2008) and Christophe et al. (2009). The parameters of the hot leg geometry (including the inclination angle of the bend q, the curvature radius of the bend R, the horizontal length LH, the inclined length LI and the pipe diameter D), the inlet and outlet conditions, the fluid properties, the system pressure and even the experimental procedures may influence the CCFL behavior. Up to now, the effects of many factors have not been clearly understood. The CCFL correlations available in literature reviewed by Navarro (2005) and Deendarlianto et al. (2008) are mostly correlated by the form of the Wallis correlation considering the influences of the length-to-diameter ratio (LH/D) of the horizontal section, the incline angle q and the inclination length LI. The effect of the length-to-diameter ratio (LH/D) has been investigated by many researchers. It is generally recognized that a larger length-to-diameter ratio will slow down the liquid flow and increase the liquid level near the bend due to an increase in friction. In that case, an early formation of unstable waves will take place in the bent region for a fixed gas flow rate (Ohnuki, 1986; Wongwises, 1996; Kim and No, 2002). As for the effect of the inclination angle of the bend q, researchers have different opinions. Wongwises (1996) concluded from experimental results that at low liquid flow rates with a larger inclination angle, the change in flow direction would cause greater turbulence near the bend which would promote the growth of unstable waves at the crest of the hydraulic jump. As a result, an

earlier CCFL would occur for a larger inclination angle. In the case of high flow rates, the flow in the horizontal section would become more supercritical due to gravity acceleration in the inclined section. The flow in the horizontal section would depress the growth of unstable waves. However, Kim and No (2002) compared CCFL data available in literature with different inclination angles from 35 to 90 and concluded that the effect of the inclination angle was not as clear as the LH/D effect. The CCFL data with different inclination angles showed a similar trend. In addition, an airewater CCFL experiment in a horizontal rectangular channel connected to an inclined riser was performed (Deendarlianto et al., 2008; Christophe et al., 2009). An interesting result was obtained that the system pressure had a strong influence on the CCFL behavior. At higher pressure, the gas density increases and the air superficial velocity decreases at the given mass flow rate. As a result, a higher air mass flow rate is needed to induce CCFL. Furthermore, comparison between CCFL correlations obtained for the hot leg geometry displayed a remarkable difference and indicated that each correlation was greatly influenced by the test section geometry including the curvature radius of the bend. 5.3.2. CCFL in pressurizer surge line As mentioned previously, the pressurizer becomes completely empty at the end of the initial blow-down phase. With the opening of the first three stages of the ADS valves, a refilling process of the pressurizer will take place and the mixture level will remain nearly constant during a period of time. After the opening of the fourthstage ADS valves attached to the hot legs, water in the pressurizer and steam in the hot leg flow downward and upward respectively in the surge line and thus a CCFL may occur. A typical mixture level curve during a small break LOCA is shown in Fig. 6 (Westinghouse Electric Company, 2004). In the TMI accident, the pressurizer was indicated to be full of water due to CCFL in the surge line. Then operators stopped a high pressure injection system and increased damage (Kawanishi et al., 1990). Therefore, the prediction of the pressurizer mixture level and draining rate are of importance and what is more, the gravity head caused by the mixture level affects the pressure in the downcomer and further the initiation of the IRWST injection (Takeuchi et al., 1999).

Fig. 4. Typical CCFL test configuration.

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Fig. 5. CCFL flow behavior at low (left) and high (right) liquid flow rates (Christophe et al., 2009) (a) Before CCFL (b)During CCFL.

The prediction of CCFL in the AP1000 pressurizer surge line is difficult due to its complex geometry. By comparing the limiting velocity in the vertical section, in the slightly inclined section and in the horizontal and vertical elbows of the pressurizer surge line, Takeuchi et al. (1999) concluded that the vertical section was the most limiting section and which decided the pressurizer liquid drain rate. However, according to the literature review of Solmos et al. (2008), although a great deal of work had been done on CCFL, there still exited large uncertainty in analysis, and no flooding model was generally applicable to varying conditions. Note that current CCFL models are obtained under experimental conditions

where the fluid property is different from that in high temperature and high pressure conditions during the small break LOCA transient. It was suggested that the effects of steam condensation and the pipe inclination angle should be taken into consideration. Furthermore, Vierow (2008) conducted airewater testing in a test tube with a wide range of inclination angles to investigate the effect of the inclination angle on the surge line CCFL behavior. However, the single test tube with different inclination angles cannot reflect the complexity of the surge line geometry and the water configuration into the surge line, which needs further discussion (Vierow, 2008). 5.4. Entrainment Due to slow depressurization rates compared with large break LOCAs, the liquid and steam phase in the reactor coolant system will be separated and various phase separation effects will influence the thermal hydraulic characteristics during the small break LOCA transient (Burchill, 1982). Among these effects, the liquid entrainment issue including the branch entrainment (from the hot legs to the fourth-stage ADS) and pool entrainment (from the upper plenum to the hot legs) drew much attention during the AP1000 safety analysis. The data of the branch entrainment and pool entrainment obtained from the APEX and ATLATS are used for code assessment. Liquid entrainment during a small break LOCA is shown in Fig. 7 (Welter, 2003). Note that in AP1000 design one steam generator is connected with one hot leg and two cold legs (cold leg-1 and cold leg-2).

Fig. 6. Typical mixture level curve of the pressurizer during a small break LOCA.

5.4.1. Branch entrainment After the opening of the fourth-stage ADS valves, the steam is vented directly to the containment. If the steam velocity is sufficiently high, branch entrainment will occur. Generally speaking,

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Fig. 7. Liquid entrainment phenomenon.

liquid entrainment in the hot legs influences the flow quality and critical flow rate out of the fourth-stage ADS, and eventually determines the depressurization rate and the system inventory (Welter, 2003). Airewater branch entrainment tests at atmospheric pressure were conducted for a wide range of conditions in the ATLATS facility which consisted of a prototype reactor vessel, steam generator, hot leg and the fourth-stage ADS line (Welter, 2003; Welter et al., 2004). Welter et al. (2004) reviewed previous investigations on the branch entrainment at a tee junction with an

upward-oriented branch. During a small break LOCA transient, various flow regimes may exist in the hot leg depending on the hot leg entrance submerged conditions. However, most of the previous experimental data were obtained under stratified flow conditions and the corresponding correlations were employed by RELAP5, namely the horizontal stratification entrainment/pull through model (RELAP5/MOD3.3, 2001). It was pointed out that the RELAP5 underestimated the hot leg entrainment rate during a small break LOCA (Welter, 2003). Furthermore, Welter proposed a new entrainment onset criterion and developed a new model for the

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steady-state entrainment rate based on previous theories. The new entrainment onset criterion is expressed by:

W32



h ¼ K b d d5 rG Drg

3 

   2 #1 2 " h hb 0:22 b þ 1 1 d d

(3)

where WG3 is the gas mass flow rate in the branch, d is the branch diameter and hb is the gas chamber height at the onset of entrainment. Compared with available experimental data of the branch entrainment, it indicated that the new entrainment onset criterion collapsed the available experimental data within 20% and the new entrainment model was able to collapse data from different test conditions which could not be done by previous correlations. Discrepancies still exit between the new entrainment rate model and the available experimental data in the low gas flow quality, which is caused by the intermittency of the hot leg flow. Further experimental work is required to consider the effects of the integral components, the high pressure and high temperature conditions and the hot leg flow regimes on the branch entrainment behavior (Welter et al., 2004). Since air and water are used in the ATLATS facility as the working fluid, phase change between steam and water and its impact on the branch entrainment rate should be treated carefully. 5.4.2. Pool entrainment A new phenomenon was observed during the branch entrainment onset tests in the ATLATS facility that the reactor vessel inventory decreased continuously even after the water level dropped below the hot leg evaluation (Welter, 2003). That was attributed to liquid entrainment from the boiling water surface in the reactor vessel into the hot legs and then into the vertical branch, which was referred to as pool entrainment (Wu et al., 2005). The pool entrainment is dominant compared with the branch entrainment in a double-ended rupture of the direct vessel injection line because of a relatively low mixture level in the reactor vessel. A series of detailed pool entrainment tests were conducted in the APEX facility to provide data for code assessments. The pool entrainment correlations proposed by Kataoka and Ishii (1984) were used for scaling analysis. According to the theory, there exist three different regions where the liquid entrainment is controlled by different mechanisms. They are the near-surface region, the momentum-controlled region and the depositioncontrolled region. The results of the APEX scaling analysis indicated that the pool entrainment behavior below the midpoint of the upper plenum was conservatively simulated (Welter et al., 2005). Wu et al. (2005) studied the liquid entrainment in the reactor vessel with and without reactor internals installed, comparison of which showed that the liquid entrainment with internals installed was several times smaller than that without internals installed mainly due to droplet deposition or de-entrainment in the internal region (such as the vertical guide tubes). However, comparison between the experimental data and Kataoka-Ishii pool entrainment correlations presented large discrepancies. Detailed analysis indicated that the liquid entrainment in the reactor vessel was not likely pool entrainment but may be just a side branch entrainment in a container and the entrainment rate was decided by the pressure difference between the water level and the branch outlet. Relatively little information is available currently on the pool entrainment (or the side branch entrainment). The entrainment mechanism in the reactor vessel and the effects of the internals on the entrainment rate are still not clearly understood and thus an intensive studies are needed.

415

5.5. Reactor vessel level swell The level swell in the reactor vessel is caused by steam formation under the two-phase mixture surface because of boiling and flashing during the small break LOCA transient. The swell level is calculated by the equation given below (Welter et al., 2005):

Hswell ¼ Hcollapsed =ð1  < ac >Þ

(4)

where Hswell is the swell level (or the mixture level), Hcollapsed is the collapsed liquid level, and is the average core void fraction. The level swell is influenced by many factors, including the specific volume of the steam (the pressure), the axial power shape, the bubble rise velocity and the steam disengagement from the two-phase surface (Burchill, 1982). As for conventional PWRs, a crossover leg (i.e. the loop seal) exists between the steam generator exit plenum and the main coolant pump. During the small break LOCA transient, a minimum core/upper plenum mixture level is reached before the loop seal clearing by manometric head balance between the reactor core and the crossover leg (Naugab, 1987; Tasakaa et al., 1988; Chung et al., 1994). In that case, core uncovery and cladding heatup may occur. A primary improvement in AP600/AP1000 design compared with conventional PWRs is the elimination of the crossover leg between the steam generator and the coolant pump (Muftuoglu, 2004). Therefore, during the small break LOCA of AP600/AP1000, the minimum core/upper plenum mixture level does not appear and the core is fully covered and cooled by twophase mixture. The core/upper plenum mixture level is related to the downcomer mixture level by manometric head balance. The prediction of the reactor vessel level swell is important for the core/upper plenum mixture level, and the downcomer mixture level determines the thermal hydraulic phenomena of the hot leg side and cold leg side, respectively. The hot leg side phenomena include the CCFL, the branch and pool entrainment, the PRHR HX heat transfer and the pressurizer drain and refill process. While the cold leg phenomena involve the CMT recirculation and drain behavior and the critical flow rate in a small break LOCA of the cold leg. 5.6. Direct contact condensation (DCC) The steamewater direct contact condensation (DCC) phenomenon may be encountered at several locations during a small break LOCA, such as cold legs, hot legs, downcomer, upper and lower plenum (Cho et al., 2004). The DCC phenomenon is also expected to occur in the IRWST of AP1000 during the ADS blow-down phase in a small break LOCA. After the subsequent opening of the first three stages of the ADS valves, the steam is vented through the ADS lines and quenched directly in the IRWST. The steamewater direct contact condensation in the IRWST is worth investigating because it influences the heat transfer of the PRHR HX and thus the safety injection and passive cooling feature of the primary system. Furthermore, it will impose considerable condensation loads on the tank structures. The system safety analysis code, RELAP5, employs a horizontally stratified condensation model which is oversimplified and large discrepancies exist between the predictions and the test data of separate or integral tests (Choi et al., 2002). In addition, current commercial CFD codes generally have no models to describe the steam jet condensation in the IRWST (Song et al., 2007). The published work of the DCC phenomenon can be roughly divided into three parts (Gulawani et al., 2006): (1) evaluation of the plume (a small vapor cavity generated by the injected steam next to the nozzle exit) shape and length, (2) estimation of DCC heat

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transfer coefficient, and (3) development of condensation regime map. However, most of the current experimental investigations on the DCC phenomenon and the condensation regime maps focus on a single nozzle sparger connected to a horizontal or vertical steam pipes submerged in a subcooled water pool (Park et al., 2007). Relatively little attention has been drawn to the DCC phenomenon of the AP1000-type sparger, namely one vertical pipe connected with four multihole branches submerged in the IRWST. Investigation on the multihole I-type sparger with or without a bottom hole used in the APR1400 reactor (Song et al., 2007) can provide useful information for the DCC estimation of the AP1000-type sparger. A so-called condensation region model (see Fig. 8) in which the momentum and energy of the steam were conserved in the condensed water was proposed to simulate the steamewater direct contact condensation in the IRWST of the APR1400 (Song et al., 2007; Kang and Song, 2008). The steam was assumed to be perfectly condensed in the steam penetration length and the thermal mixing in the IRWST was treated as an incompressible flow (Song et al., 2007). Meanwhile, condensation loads and thermal mixing induced by the direct contact condensation were

experimentally investigated at the test facility Blow-down and Condensation Loop and a condensation regime map used for a multihole sparger was developed (Park et al., 2007, 2008). The condensation regime map mainly focused on the effects of steam mass flux and pool temperature on the development of condensation modes. The results showed that the condensation region model together with the commercial CFD code CFX4.4 could simulate the thermal mixing in the IRWST reasonably well (Kang and Song, 2008). It should be noted that many problems still remain due to the configuration differences between the AP1000-type sparger and the APR1400 I-type sparger. The effects of the sparger hole pattern (parallel type or staggered type), the pitch-to-hole diameter ratio P/D, the branch inclination angle and other factors on thermal mixing in the AP1000 IRWST, and interaction among adjacent discharge holes need to be further discussed. Moreover, the thermal mixing in the IRWST of the APR1400 and the temperature distribution obtained from the experiment cannot be used directly in current system analysis codes. Therefore relatively simple empirical or semi-empirical correlations for multihole spargers based on the DCC mechanism are required to be developed.

Fig. 8. Condensation region model for the DCC of the APR1400 sparger (Kang and Song, 2008).

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Fig. 9. Thermal stratification in CMTs of the ROSA-AP600 (Yonomoto et al., 1997).

5.7. Thermal stratification Thermal stratification refers to the phenomenon that two fluids do not get mixed but separate from each other due to the density difference caused by temperature difference (Kim et al., 2005). It usually occurs in the following components: pressurizer surge line, steam generator feed-water system, safety injection system and chemical and volume control system (Kim et al., 2005; Boros and Aszódi, 2008). The thermal stratification may lead to fatigue failure due to periodic stress distribution. As for AP1000, apart from the components where thermal stratification occurs in conventional nuclear power plants, the estimation of thermal stratification in CMTs is of significance because it greatly influences the progress of small break LOCA chronology. After the opening of the CMT isolation valves, a recirculation path is formed by each CMT, a direct injection line, a part of the reactor vessel downcomer (between the direct injection line and the cold leg inlet), a part of the cold leg (between the downcomer and the pressure balance line) and a pressure balance line (Kukita et al., 1996; Yonomoto et al., 1997). Cold water at the top of the CMTs will be replaced by relatively hotter water from the cold legs which will reduce the driving force of the recirculation path. A clear thermal stratification was formed in the CMTs of the ROSA-AP600 test facility (see Fig. 9). The water in the CMTs formed three regions: a cold liquid region at the bottom, a saturated liquid layer at the top, and a thermally stratified layer between the cold and saturated water. Similar results were also observed at the PACTEL (parallel channel test loop) facility used for the study of passive safety injection system performance during small break LOCAs (Tuunanen, 1998). A transition from the recirculation mode to the draining mode is controlled by two mechanisms: flashing in the core makeup tanks and the pressure balance lines for a smaller break or the uncovering of the balance line inlet due to the draining of the cold

legs for a larger break (Yonomoto et al., 1997; Chang et al., 2006; Sibamoto and Yonomoto, 2006; Woods et al., 2007). The prediction of the thermal layer growth and the onset of CMT draining is challenging and little information is available on that issue. In addition, flow oscillation may exist between the recirculation phase and the draining phase which will increase the predicting difficulties (Tuunanen, 1998). Currently, most of the experimental data is not available for data proprietary and plenty of work remains to be done for the verification of system safety analysis codes. 6. Conclusions Thermal hydraulic phenomena related to small break LOCAs in AP1000 have been investigated, which include the physical description, theoretical and experimental research work, and prediction models as well as their comparison and evaluation for the critical flow, natural circulation, counter-current flow limiting, entrainment, reactor vessel level swell, direct contact condensation and thermal stratification. The investigation shows that mechanisms that govern the thermal hydraulic phenomena have not been fully understood in current research, and discrepancies still exist between different models. In addition, it should be pointed out that current available models used for critical flow, counter-current flow limiting and entrainment are based on experimental data obtained under those conditions which are considerably different from reactor operating conditions. Therefore, investigation in high pressure and high temperature conditions is needed, new models that can reflect detailed mechanisms are also required for the development of system safety analysis codes, and some new features associated with the passive facilities of AP1000, such as the entrainment, direct contact condensation and thermal stratification, should be carefully treated in the future development process of advanced passive reactors.

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