Experimental investigation of spray induced gas stratification break-up and mixing in two interconnected vessels

Nuclear Engineering and Design 241 (2011) 3935–3944

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Experimental investigation of spray induced gas stratification break-up and mixing in two interconnected vessels Nejdet Erkan ∗ , Ralf Kapulla, Guillaume Mignot, Robert Zboray, Domenico Paladino ∗ Nuclear Energy and Safety Research Department, Laboratory for Thermal–Hydraulics, Paul Scherrer Institute, CH-5232 Villigen-PSI, Switzerland

a r t i c l e

i n f o

Article history: Received 7 March 2011 Received in revised form 15 July 2011 Accepted 18 July 2011

a b s t r a c t To analyze the effect of containment spray on gas mixing and depressurization, two experiments (ST3 1 and ST3 2) were performed with two interconnected vessels. These experiments were conducted in the frame of the OECD/SETH-2 project using the PANDA facility. The vessels were preconditioned such that a helium-rich layer is formed in the upper section of the first vessel, henceforth referred to as Vessel-1. In the case of the first experiment (ST3 1), the remaining volume of Vessel-1 and the entirety of the second vessel, Vessel-2, were filled with pure steam. For ST3 2, the second experiment presented here, pure steam was replaced with a steam–air mixture instead. Water was injected from the top of Vessel-1 with a spray nozzle projecting downwards. Transient behavior of system pressure, as well as global redistribution of gases is investigated. The results reveal that spray activation is very effective in containment system depressurization. Additionally it is found that the depressurization occurs at a higher rate for the systems containing more steam and less non-condensible gas. The depressurization rate gradually slows down, however, as the steam concentration decreases due to condensation, and non-condensible gases spread over the vessel system. It is also observed that the spray activation initiates the breakup of the helium-rich layer. The composition of the gas atmosphere plays a crucial role in determining the initiation time of the breakup; the presence of large amounts of non-condensible gas such as air delays the beginning of the helium layer breakup by approximately 200 s. The downward component of spray momentum causes the entrainment and the recirculation of the ambient gas atmosphere. Together with the entrainment and condensation effect, spray activation influences the gas mixture density in Vessel-1 and this generates a driving force for inter-compartment flow. As a result of this, an increase of helium–rich gas mixture is observed in the regions far away from the spray, i.e., in Vessel-2. © 2011 Elsevier B.V. All rights reserved.

1. Introduction In the case of a severe accident in a light water reactor (LWR), some amount of hydrogen might be produced by the reactor fuel cladding oxidation and can be spread to the whole containment volume. Together with steam and air, the released hydrogen can create locally flammable or even explosive gas mixtures in the containment. Such a typical post-accident condition in LWR containment is characterized by several physical phenomena including gas (air, steam, and hydrogen) transport within the compartments, gas stratification build-up, condensation and evaporation. Activation of safety systems (e.g., sprays, recombiner, etc.) induces some additional phenomena, such as depressurization, stratification and breakup, which have to be investigated in detail. Therefore, analysis of post-accident phenomena related to the complex

∗ Corresponding authors. Tel.: +41 56 3104150. E-mail addresses: [email protected], [email protected] (N. Erkan), [email protected] (D. Paladino). 0029-5493/$ – see front matter © 2011 Elsevier B.V. All rights reserved. doi:10.1016/j.nucengdes.2011.07.025

transient containment conditions require use of sophisticated analytical tools, e.g., advanced lumped parameter (LP) codes, 3D and CFD codes. Containment water spray is an important safety system used in LWR containments to mitigate the effects of a hypothetical accident by simultaneously preventing containment overpressure, enhancing gas mixing, and removing radiological aerosols. Aside from these advantages, water spray injection into a gas mixture comprised of steam and hydrogen may also lead to a local increase in the hydrogen concentration. This is due to the reduction in the quantity of steam as a result of the condensation of the vapor on the water droplet surfaces. Considering the risk of a hydrogen explosion, the advantages and disadvantages of the spray system must be carefully examined (OECD, 1999). The interaction of droplets with the surrounding gas poses challenges with respect to characteristic time and length scales. Considering the extreme range of the scale between the sizes of the water droplets compared with the size of the containment building, one can imagine the complexity of the exchange phenomena. Accordingly, complex computational efforts are required

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for a detailed prediction of the relevant physics. Hence, large-scale experimental investigations are essential to study the phenomena and to provide data for the validation of analysis codes. There have been some experimental programs conducted in several test facilities (TOSQAN, MISTRA and NUPEC) and relevant analytical studies to investigate the containment spray effects on the depressurization and the distribution of non-condensible gases. Corresponding validation studies with computational codes have also been conducted. Some of the CFD validations most pertinent to droplet dynamics and employing TOSQAN spray test results showed that while most of the calculations could recover the equilibrium droplet velocity approximately 1 m below the nozzle, discrepancies between the experimental results and the calculations were mainly observed in the regions close to the nozzle exit where intensive droplet separation and droplet acceleration take place (Malet et al., 2008). Some computational codes such as GASFLOW (Malet et al., 2008), CFX (Babic et al., 2009) and GOTHIC (Andreani and Erkan, 2010) overestimate the rate of the gas mixing and could not predict transient evolution of helium-rich layer erosion process. That is significant since the alteration in the velocity field, driven by the droplet momentum, and local mixture densities determine primarily the development of stratified layer erosion and the global redistribution of the non-condensible gases in a complex geometry such as the multi-compartment structure of a real containment building. Mimouni et al. (2009) performed simulations of TOSQAN tests (Malet et al., 2008), which were implemented to investigate depressurization and helium-rich layer breakup with and without heat and mass transfers between gas and droplets, using NEPTUNE CFD code. The calculations underestimated the transient pressure decrease. In contrast, the steady-state pressure, reached 2000 s after the spray activation, could be reproduced. Although the time evolution of the helium concentration at the top and the bottom regions of the TOSQAN vessel were in reasonably good agreement with the experimental data in later stages of the test, the transient phase of helium-rich layer breakup and mixing was underestimated and a faster mixing time was predicted. The aforementioned validation studies demonstrate that some computational codes have difficulties predicting the transient phase of gas mixing and depressurization in such a vessel configuration; although this configuration is still much simpler than an actual nuclear reactor containment building. The present experiments, addressing the containment sprays, focus on the containment depressurization and distribution of non-condensible gases. Two experiments (ST3 1 and ST3 2) were performed with two different initial gas compositions in two large interconnected PANDA vessels. In addition to the initial helium layer, which is used to simulate hydrogen collected at the top of one compartment after an accident, the experiments provide significant results about (a) the transient system pressure response to the spray activation, (b) breakup of the stratified helium-rich layer, and (c) the time-dependent distribution of the non-condensible gas within the compartments. Additionally, the use of a two-vessel configuration provides valuable insight into the formation of potentially dangerous hydrogen mixtures with increased hydrogen content in regions far away from the spray. It also provides the community with valuable and detailed data for the validation of computational design tools under complex situations.

2. PANDA facility and instrumentation PANDA is a large-scale thermal–hydraulic test facility designed for the investigation of containment system behavior and related

Fig. 1. PANDA facility.

phenomena for different Advanced Light Water Reactor (ALWR) designs, and for large-scale separate effect tests (Dreier et al., 2008). The containment compartments and the Reactor Pressure Vessel (RPV) are simulated in PANDA by six cylindrical pressure vessels. Various auxiliary systems are available for specifying and controlling the initial and boundary conditions for the test. For the spray experiments, only two vessels, Vessel-1 and Vessel-2, serve as the test section (Fig. 1). The two vessels, each with a height of approximately 8 m and a diameter of 4 m, are connected with a 1 m diameter pipe (IP). The Vessel-1 and Vessel-2 were instrumented with 261 and 87 K-type thermocouples, respectively, measuring fluid and wall temperatures (to 0.7 ◦ C accuracy). Gas molar fractions are measured with Mass Spectrometer (MS) capillaries, which are positioned at 59 positions in Vessel-1, 34 positions in Vessel-2 and 15 positions in IP. Gas is continuously sampled through capillaries and sent to MS systems. All MS measurement positions cannot be available simultaneously during one test, since, the measurement is sequential and only one capillary can be selected at a time via a multiport rotating valve per spectrometer. The number of sampling positions used for the measurements varies in each test and during the test evolution. The scanning frequency for the gas composition is typically in the range of 6–10 s for each capillary location (the time depends on the mixture, i.e., two gas, three gas, etc.) therefore it is decided in each test to measure gas concentration at selected locations. Those measurement positions used for the spray tests are shown in Fig. 2. A few millimeters apart from each mass spectrometer capillary, a thermocouple is placed such that gas concentration and temperature measurements are available at almost the same spatial location. The measurement accuracy gradually depends on the gas mixture composition. For the steam–air mixtures, an absolute error for the measured steam–air molar fraction of about ±1.5% is assessed. All concentration measurement results are normalized with the real initial value (CB He ) measured over a specified time interval at position B in Vessel-1 at the beginning of the test and represented in nondimensional form. The spray nozzle, model HH-30 (SSCO-Spraying System AG), is oriented vertically downward in Vessel-1 and its outlet is located at 6.9 m above the bottom of Vessel-1. The nozzle has 6.4 mm outlet diameter and it produces a conical solid spray pattern with an opening angle of 30◦ .

N. Erkan et al. / Nuclear Engineering and Design 241 (2011) 3935–3944

Fig. 2. Gas concentration measurement locations in Vessel-1, Vessel-2 and IP.

3. Initial conditions and test procedure Two containment spray tests have been performed. Denoted as ST3 1 and ST3 2, these experiments differ only in the composition of the initial gas atmosphere. For ST3 1 a helium-rich layer was created in the upper part of the Vessel-1, while the remaining volume of Vessel-1 and the full volume of Vessel-2 were filled with steam (Fig. 3a). For ST3 2, the helium-rich layer was created in an initial steam–air environment at the top of Vessel-1, and the remaining volume of Vessel-1 as well as the full volume of Vessel-2 was filled with the steam–air mixture (Fig. 3b). The initial concentrations of the steam–helium and steam–helium–air mixtures were adjusted to specific values for both tests according to the project requirements. The variable parameter for the two tests is the initial gas atmosphere composition for which the nominal values are given in Table 1. Nominal initial helium molar fractions in the layer (30%) are equal for both tests. Initial gas compositions were measured at several radial positions on different levels then time averaged values over 60 s preceding the beginning of the test are taken as real initial conditions. Initial helium and helium–steam concentrations along the vertical axis of Vessel-1 are given in Fig. 4 for ST3 1 and ST3 2, respectively. The initial helium content of the gas mixture in the region above 6.8 m is significantly higher than in the region below

Table 1 Initial specified nominal values (actual experimental parameters are those obtained with used test procedures). Conditions

Helium fraction in the layer (%) Steam fraction below the layer (%) Air fraction below the layer (%) Vessels initial pressure (bar) Vessels initial temperature (◦ C) Water injection flow rate (g/s) Water injection temperature (◦ C)

Tests ST3 1

ST3 2

30 100 0 2.6 129 840 40

30 60 40 2.6 129 840 40

Fig. 3. Schematic view of experimental layout and parameters.

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a

8000

Vessel-1 cent.-axis Vessel-1 wall Vessel-2 cent.-axis Vessel-2 wall

7000 6000

4000

Elevation [mm]

Elevation [mm]

8000

ST3_1 (He) ST3_2 (He) ST3_2 (Steam)

2000

6000 5000 4000 3000 2000 1000

0

0

0.5

1

1.5

2

0 0.9

B

C/CHe

1.6 B level(7478mm) C level(6926mm) D level(6926mm) F level(6000mm) G level(5626mm)

1.2

B

C / CHe

1 0.8 0.6 0.4 0.2 0 -2000 -1500 -1000 -500

0

500

1000 1500 2000

Radius [mm]

b 1.6

B level(7478mm) C level(6926mm) D level(6926mm) F level(6000mm) G level(5626mm)

1.4

B

C / CHe

1.2 1 0.8 0.6 0.4 0.2 0 -2000 -1500 -1000 -500

0

500

1000 1500 2000

Radius [mm] Fig. 5. Initial radial distribution of helium concentration for the tests ST3 1 and ST3 2.

b

1.05

8000

1.1

1.15

1.2

Vessel-1 cent.-axis Vessel-1 wall Vessel-2 cent.-axis Vessel-2 wall

7000

Elevation [mm]

6.8 m for both tests, representing the formation of the helium stratification layer prior to the spray activation. Any presence of helium below 6.8 m is a consequence of mainly convective transport caused by the helium injection during the helium layer buildup. The helium gas is almost uniformly distributed along the radial cross-section of the Vessel-1 for both tests. The normalized helium molar fractions are presented in Fig. 5a and b, demonstrating two-dimensional representation of initial radial uniformity. While the gas atmosphere below the layer is composed of pure steam (100%) in test ST3 1, it consists of air and steam (40% and 60%, respectively) in test ST3 2. Initial nominal vessel pressure, which was chosen as a typical post-accident containment pressures, is 2.6 bar for both tests. Initial fluid temperature corresponding to the saturated steam temperature at 2.6 bar is 129 ◦ C. Initial axial fluid and wall temperatures (normalized with the average of initial gas

1.4

1

T / Tg

Fig. 4. Initial axial helium and steam molar fractions in Vessel-1.

a

0.95

6000 5000 4000 3000 2000 1000 0 0.9

0.95

1

1.05

1.1

1.15

1.2

T / Tg Fig. 6. Initial axial temperature profiles in ST3 1 and ST3 2.

temperatures, Tg, in Vessel-1 and Vessel-2) are given in Fig. 6a and b in non-dimensional form. While no significant temperature gradient exists in wall temperatures except the upper region of Vessel-1 (Fig. 6a), some deviations from the average value are observed for the gas temperatures mainly in the helium layer region. It is difficult to control the helium injection temperature due to the sudden expansion or compression of the helium gas while we are injecting into the Vessel-1 that leads to slight temperature deviations (5% of the average in maximum) in the helium layer region. Demineralized water, stored in a water tank, is pumped through the spraying nozzle into the Vessel-1. Storage tank is also pressurized up the same pressure level with the vessels to suppress the effect of pressure variation in the vessels on the injection flow rate. By means of this pressure equivalence, water pump could preserve the flow rate (840 g/s) constant during whole test duration. Nominal water injection temperatures (40 ◦ C) are very low compared to the ambient gas temperature (129 ◦ C) for both tests.

The PANDA vessels and the IP were preconditioned before each test according to the specified initial conditions. The vessels were initially heated up and pressurized with pure steam to reach the pressure and the temperature level specified in Table 1. To create the required steam–air mixture for ST3 2 compressed air at room temperature was injected. After reaching the specified pressure and temperature, the helium was injected at the top through a tube located at 6 m height in Vessel-1 (Fig. 1) until the nominal helium concentration was reached. The resulting helium-rich layer covers up to 0.1 m below the outlet of the spray injection nozzle (Figs. 3 and 4). After the helium layer build-up, the test phase was started by injecting water at the specified flow rate and temperature through the spray injection nozzle. The vessel system was kept closed and isolated from the outside environment to investigate the system depressurization induced by the cold water spray injection. Before starting the test, it was assured that no water accu-

N. Erkan et al. / Nuclear Engineering and Design 241 (2011) 3935–3944

4

Exp LNDF, f(D)

ST3_1 ST3_2

1 0.8

3 P / P0

# probability [1/mm]

3.5

3939

2.5 2

0.6 0.4

1.5

0.2

1 0

0.5 0

0.2

0.4 0.6 0.8 1 Equivalent Diameter [mm],D

1.2

mulation exists in the bottom of the Vessel-1. Spray water being injected from the top of the Vessel-1 accumulated in the bottom of the Vessel-1 and remained there during the whole test duration. 4. Results and discussion 4.1. Droplet size measurements In order to identify the spray droplet size distribution, which may in turn be utilized as a boundary condition for an accurate validation of computational codes, size measurements were performed in separate tests at atmospheric pressure and room temperature using backlight-imaging technique. Demineralized water at room temperature was injected at the same flow rate as for the experiments. Since the droplet population was too high to record high quality droplet images at this flow rate, an obstacle to reduce the droplet density in the imaging plane cut half of the spray cone. Shadow images of the droplets were recorded 120 cm away from the nozzle exit in the center of the spray cone using a CCD camera with a lens having a focal length of 200 mm. 2D projected droplet areas were measured with the image analysis software ImageJ (Abramoff et al., 2004), and equivalent diameters were calculated from the measured droplet areas. Number probability of measured droplet diameters and some statistical diameters are estimated with their errors (typical range in between 10 and 15%, Yoon et al., 2004) originated from the current imaging and image analysis technique. The mean diameter (D10 ) and Sauter mean diameter (SMD, D32 ) is found to be 0.269 ± 0.06 mm and 0.582 ± 0.076 mm, respectively. The volume median diameter (VMD 0.5), which determines the weight of large sized particles on the size distribution, is found to be 0.697 ± 0.07 mm which fits well the data (0.630 mm) provided by the nozzle producer company. This implies that half of the total volume of water injected is made up of droplets with diameters larger than 0.697 mm. A log-normal distribution function (LNDF) (see Eq. (1)) is used to represent the data obtained from the particle size measurements (Presser et al., 1990; Hinds, 1999): 1 f (D) = √ exp 2g D

−[ln(D/Dg )] 2g2

2

1000

1500 2000 Time [s]

2500

3000

3500

Fig. 8. Pressure evolution in the vessels.

Fig. 7. Droplet size distribution.



500

1.4

main liquid core and accompanying droplet interactions, both of which determine the size distribution, completed in very short distances (compared to the large scales of the vessels) away from the nozzle exit. In further downstream regions, droplet diameter distribution does not change gradually due to the mechanical effects and the distribution of droplets becomes substantially more uniform with radial location (Yoon et al., 2004). Considering the measured 0.697 mm VMD 0.5 value, the droplets having larger diameters than 0.697 mm mainly govern mechanical and thermodynamical effect of the spray. Therefore measured size distribution can be a good approximation for initial droplet size as a boundary condition. 4.2. Depressurization In the case of an accident transient in a LWR, the main mechanism of a containment spray in reducing the containment pressure is the condensation of steam on water droplet surfaces (Malet et al., 2005). Although it is likely to a much less extend compared to condensation, the direct cooling of hot gas atmosphere is another effect contributing to depressurization. To analyze the effect of the spray injection on pressure evolution, the normalized pressures (normalized with the real initial pressure, P0 ) are compared for both tests (Fig. 8). As expected, the pressure decays as a function of increasing time at a different rate for the two experiments. This decay of pressure progresses faster for ST3 1 where the gas mixture is composed solely of helium and steam. The pressure of the system decreases by 55% over 2000 s in ST3 1; in ST3 2, where a three component steam–helium–air mixture is used, and the pressure is reduced by only 33% over the same period of time. The reason for a slower depressurization in ST3 2 compared with ST3 1 is the inclusion of lower steam concentration and the existence of the air. The difference in the behavior of pressure decay suggests that the initial steam content plays a significant role in determining the temporal behavior of the pressure in the entire system. The pressure decreases at similar rates in the first 200–300 s for both tests. Afterwards, the slope of the curves starts 1.1



(1)

This functional representation of experimental data exhibits a reasonably good approximation of the empirical size distribution using the relevant parameters, Dg and  g , calculated from the experimental data (Fig. 7). Dg and  g are the geometric mean diameter and geometric mean standard deviation, respectively (Hinds, 1999) which are calculated from the measured droplet sizes as Dg = 0.227 mm and  g = 0.565. High droplet density close to the nozzle makes the measurement impossible due to the image overlapping for almost all measurement techniques. Droplet separation processes from the

0.9 T / T 1w

0

0

0.7 ST3_1 (Wall temp.) ST3_1 (Steam sat. temp.) ST3_2 (Wall temp.) ST3_2 (Steam sat. temp.)

0.5

0.3

0

500

1000 1500 Time [s]

2000

2500

Fig. 9. Comparison of wall and steam saturation temperatures normalized with the average initial wall temperatures in Vessel-1.

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1.3

1.1

ST3_1 ST3_2

0.9 1

1

T / Tg

SR [-]

1.1

B D L1 L2

1.2

0.9 0.8 0.7

0.7

0.5

0.6 0.5

0

200

400

600 Time [s]

800

1000

0.3

1200

0

500

1000

1500 2000 Time [s]

2500

3000

3500

Fig. 10. Saturation ratio of steam in Vessel-1 (ST3 1).

Fig. 12. Global temperature (normalized with average initial gas temperatures in the Vessel-1) variation in Vessel-1.

to differ significantly. The presence and increasing concentration of non-condensible gases (air and helium) lower the local steam partial pressure thereby decreasing the condensation and eventually depressurization rate, making the spray less efficient in the test ST3 2 by the time progress. In order to investigate possibility of steam condensation on the vessel walls, spatially averaged Vessel1 wall temperature and steam saturation temperature variations in time are plotted in Fig. 9. During both tests, saturation temperature of steam remained low compared to the wall temperatures, which indicates the condensation of steam extensively occurred on the water droplets instead of the Vessel walls. A re-evaporation process also exists simultaneously with the steam condensation on water droplets. The spray nozzle had an angle of 30◦ and therefore the drops did not impinge on the sidewalls. Hence, major part of the re-evaporation can occur from the surface of the water pool accumulated at the bottom of the Vessel-1. In spite of this reevaporation, any pressure fluctuation could not be observed during both tests, which proves that the magnitudes of the re-evaporation rate remained lower compared to the condensation. This small magnitude re-evaporation probably shows its effect on the vessel pressure as delaying the depressurization rate. Andreani and Erkan (2010) calculated that the re-evaporation is less than 1% of the condensation during the first 200 s, however, it becomes more than 10% at around 500 s. Despite those predicted re-evaporation rates, they overpredicted the depressurization rates. Consequently, cumulative effects of re-evaporation from the pool at the bottom and from the droplets might have a combined effect of slowing down the depressurization rate. Time dependent local saturation ratios, SR(t), of the steam are calculated (Figs. 10 and 11) employing measured molar fractions at the positions in Vessel-1 depicted in Fig. 2 according to the following equation:

Psteam denotes the instantaneous partial pressure of the steam measured at a position, and Psat is the saturation pressure of the steam. The latter is calculated with international-standard steam tables for the temperature values, T(t), measured at the same position. For experiment ST3 1, below the helium layer in Vessel-1, the saturation ratio is approximately 1 during around first 400 s of the test (Fig. 10). This implies that the gas mixture preserves the steam saturation in this period while condensation of steam takes place together with the associated temperature decrease (Fig. 12). In contrast to this, the initial SR values are much lower than one for ST3 2, the case with lower steam content. After the start of injection, an increase of the saturation ratio throughout all measurement locations is observed; this increase ranges up to approximately 0.9 at 200 s (Fig. 11). This rapid SR increase is caused by the sudden decrease of temperature (Fig. 12) and associating decrease of steam saturation pressure. A high rate of heat absorption by the water droplets from the gas atmosphere creates a cool air–rich gas mixture and that leads to an exponential temperature decrease within 200 s (Fig. 12). The fluctuations are observed in test ST3 2 that has two noncondensible gas components (air and helium) below the helium layer. The steam–reach mixture is entrained into the spray cone area, leaves its steam content inside and it goes out of the spray envelope with poor steam and cold air-rich content. This steam dilution cycle creates transient steam-rich and steam-poor mixture in the volume between the spray cone and the vessel walls, which creates intermittent perception of steam concentration, thereby, SR values.

SR(t) =

Psteam (t) Psat (T (t))

(2)

4.3. Gas mixing in Vessel-1 Mechanical interaction of water droplets with the ambient gas atmosphere causes the breakup of the helium layer and promotes mixing of the helium with the other gases below. The mixing of

1.3

1.1

B (ST3_1) D (ST3_1) GH (ST3_1) L1 (ST3_1)

1.2

1

C / CB He

SR [-]

1.6

B D L1 L2

1.2

0.9 0.8 0.7

L2 (ST3_1) B (ST3_2) D (ST3_2) GH (ST3_2)

L1 (ST3_2) L2 (ST3_2)

0.8

0.4

0.6 0.5

0

200

400

600 Time [s]

800

1000

Fig. 11. Saturation ratio of steam in Vessel-1 (ST3 2).

1200

0

0

200

400

600 Time [s]

800

1000

Fig. 13. Time-dependent helium molar fractions in Vessel-1.

N. Erkan et al. / Nuclear Engineering and Design 241 (2011) 3935–3944

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Fig. 14. Schematic view of gas flow patterns in vessels system.

320–450 s. This helium moves down towards the other measurement positions, indicating an efficient redistribution of the helium content throughout the entire vessel in both tests. Gas mixture densities are calculated (using the ideal gas law for non-condensible gases, steam densities are taken as saturated steam density for the measured partial pressure of steam), normalized with the initial density (0 ) and presented in Fig. 15 for both tests at some positions above and below the density interface. The normalized gas mixture density below the spray nozzle decays continuously for both tests such that they collapse perfectly onto one another with the densities measured above the spray noz-

1.2 B (ST3_1) L1 (ST3_1) B (ST3_2) L1 (ST3_2)

1.1 1

ρ / ρ0

helium and its redistribution in Vessel-1 due to the spray activation is recorded by the variation of helium concentrations measured at the locations given in Fig. 2. Spray activation causes the breakup of helium-rich layer and results in about uniformly mixed Vessel-1. Except for position B, normalized helium molar fractions show similar trends in all positions, irrespective of the location or of the test being measured (Fig. 13). Initial helium concentrations in the lower part of vessel are close to zero since the helium-rich layer is confined to the upper part of Vessel-1 before the spray activation. After the spray is activated, the ambient gas mixture entrained by the spray envelope and it impinges to the bottom of the Vessel-1. While the spray droplets remain at the bottom, the gas recirculates as it would be an upward wall jet. Upward flow streams and entrained flow steams created by the momentum exchange between the droplets and ambient gas in the volume between the spray cone and the vessel wall are shown in Fig. 14 schematically. The flow streams initiate the breakup of helium-rich layer by interfering with the density interface, and promote mixing with the gases below. For ST3 1, helium-rich layer erosion begins earlier than 200 s (position B in Fig. 13), while in the case of ST3 2 the start of the helium layer breakup is delayed and occurs approximately 200 s after the spray activation. After the breakup, some amount of the helium is transported downward and helium content increases with time in the regions below the spray, reaching an asymptotic level of 0.4 for both tests at t = 1000 s. Normalized helium fraction at point B decays very quickly, within the range of approximately

0.9 0.8 0.7 0.6 0.5

0

200

400

600 Time [s]

800

1000

Fig. 15. Gas mixture density versus time in Vessel-1.

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0.35

IP1

IP2

IP4

0.3

IP5

0.3

C/CB He

B

C / CHe

G H1

H2 I1

I2 L2

L1 M1

M2

0.25

0.25 0.2 0.15

0.2 0.15

0.1

0.1

0.05

0.05

0

A C

0

400

800

1200 Time [s]

1600

0

2000

Fig. 16. Time-dependent helium molar fractions in IP (ST3 1).

zle. After equivalent density values are attained at all positions, the densities in the test ST3 1 continue to decrease due to the diluting effect of the condensation and mixing. On the other hand, such a steep decrease cannot be observed in the densities after the mixing completed in test ST3 2 because of having lower steam content compared to the test ST3 1. The difference in the density of the fluid below and inside the helium-rich layer before starting the erosion process is higher in the test ST3 2. As it is mentioned above spray activation induces a strong gas circulation below the layer, and the helium-rich mixture above the spray is very likely to be disturbed by the upward flow streams penetrating the layer boundary. Upward flow streams interfering with the density interface in the test ST3 2 encounters higher negative buoyancy force due to the higher density difference mentioned above, therefore, helium layer erosion develops in test ST3 2 slightly later than it does in ST3 1. Negative buoyancy may not be the only factor effecting the timing of helium layer breakup, but also, the characteristics of the global gas circulation may have a role in the breakup mechanism. As the surrounding gas density increases droplets become more responsive with the surrounding gas and the rate of momentum exchange between the phases is enhanced (Prosperi et al., 2007). Thus, drops submitted to high gas density mediums lost more momentum and have entrained more external gas into the spray envelope (Prosperi et al., 2007). Based on this, some part of the upward flow streams are re-entrained by the spray before reaching the helium layer. The re-entrainment of the upward flow streams would become stronger in high-density gas environment below the helium layer such as in the case of ST3 2. As a result, the recirculation loop is more confined below the helium layer in test ST3 2, whereas, upward flow streams, which have lower densities in the test ST3 1, are less entrained by the spray, so that they can reach the helium layer with more momentum. Consequently cumulative effects of negative buoyancy and re-entrainment managed the helium layer breakup mechanisms and lead to a time delay in the breakup of helium layer for the test ST3 2 compared to the case of test ST3 1.

0

400

800

1200 Time [s]

1600

2000

Fig. 17. Time-dependent helium molar fractions in Vessel-2 (ST3 1).

simultaneously throughout the experiments at specific positions in Vessel-2 and the IP (Fig. 2). For the experiment ST3 1, helium molar fractions measured at vertically aligned positions at the exit of the IP to the Vessel-2 are presented in Fig. 16 in non-dimensional form. At the uppermost position (IP1) normalized helium molar fraction increases faster and reaches a value of up to 0.3, whereas, that increment progresses slower for the lowest three positions. This difference indicates the presence of a stratified gas mixture flow through the IP towards Vessel-2. Since a steam-helium mixture with higher helium content is lighter than pure steam, the steam–helium mixture coming from Vessel-1 travels along the top of the IP. After approximately 320 s, a counter flow of the steam–helium mixture with low helium content starts from Vessel-2 through the lower part of the IP to Vessel-1 in order to compensate for the loss of steam mass in Vessel-1 due to the condensation process. Hence, a corresponding weaker increase of the helium concentration is also visible in the other regions of the IP (IP2 to IP5 in Fig. 16). This bidirectional inter-compartment flow between two vessels through the IP is exhibited schematically with Phase 1 in Fig. 14. The variation of normalized helium molar fraction in Vessel-2, such as at the positions of I1 and H1 (Fig. 17), demonstrates that the helium-rich mixture coming from the Vessel-1 rises upward in Vessel-2 (Phase 1 in Fig. 14). The helium concentration increase is also detected at the C and A positions with relatively lower values compared to the values at I1 and H1. This indicates the dilution of the helium-rich mixture over its path from Vessel-1, mixing with steam as it rises to the upper levels of Vessel-1. It is also found that the mixture densities in the IP and Vessel-2 are continuously decaying with time (Figs. 18 and 19). The mixture density at the highest position of IP (IP1) is slightly lower compared to the lowest position (IP5), which also confirms the existence of a stratified flow in the IP due to the higher helium content of the mixture coming from Vessel-1. As the gas mixture is uniformly mixed with helium in Vessel-2 at all positions, no significant density strat-

1.2

4.4. Helium transport into Vessel-2

IP1

IP5

1 0.8 ρ / ρ0

Multi-compartment geometry (twin vessels interconnected by a horizontal pipe) of the PANDA facility enables to study inherently asymmetric, 3D configurations which can be more representative for the phenomena occurring in real multi-compartment containment structures and the data obtained in such a geometry can pose an extra challenge for 3D simulation tools thereby being extremely valuable for validation purposes. While the spray nozzle induces the gas mixing in Vessel-1, the gas concentrations in Vessel-2 are also affected by a bidirectional gas transport through the IP. To monitor the corresponding phenomena, the variations of gas concentrations are recorded

0.6 0.8

0.4 0.7

0.2

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0 0

300

850

1000

600

1150

900 1200 Time [s]

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Fig. 18. Gas mixture density versus time in IP (ST3 1).

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1.2

1.2 C M1

1

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0.8 ρ / ρ0

ρ / ρ0

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0

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Fig. 20. Time-dependent helium molar fractions in IP (ST3 2).

0.3

A C

G H1

H2 I1

I2 L2

L1 M1

M2

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0

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Fig. 21. Time-dependent helium concentration in Vessel-2 (ST3 2).

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1500 2000 Time [s]

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1.2 C M1

1 0.8 ρ / ρ0

ification of the ambient gas is observed during the spray operation time (Fig. 19). For experiment ST3 2, the normalized helium molar fraction at the highest IP measurement position (IP1) decays during the first 500 s until it reaches zero while the other positions stay close to zero (Fig. 20). Later in the experiment, the helium content at the lowest measurement position (IP5) increases very fast and reaches 0.25 at approximately t = 800 s. Subsequently, the amount of helium rises also at IP4 but with a smaller growth rate compared to IP5. The same holds true for IP2 and IP1 later in time (at t = 1500 s) due to the counter flow of the gas mixture from Vessel-2. In Vessel-2, the normalized helium fraction remains close to zero for all positions during the first 1200 s (Fig. 21), and afterwards it begins to increase. This increase is faster for the lower positions such as M1 and M2. It is also found that helium molar fractions are increasing for position L1 (located above M1), indicating that the helium rich mixture fills Vessel-2 from the bottom upward. As in the case of ST3 1, the mixture densities in the IP and Vessel2 continuously decrease with time (Figs. 22 and 23). In the IP, no density difference exists until t = 800 s; after this point, a density

500

Fig. 22. Gas mixture density versus time in IP (ST3 2).

Fig. 19. Gas mixture density versus time in Vessel-2 (ST3 1).

C / CB He

IP5

1

0.8

C / CB He

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Fig. 23. Gas mixture density versus time in Vessel-2 (ST3 2).

stratification with a slight difference begins to develop (Fig. 22) with the subsequent increase of helium in the gas mixture in the lower part of the IP (Fig. 20). From the trace of helium-rich mixture (Fig. 20), obviously a flow reversal takes place in the IP as it is described in Fig. 14 with Phases 1 and 2. Due to the condensation in Vessel-1, the steam concentration decreases and air concentration increases in the gas mixture. When the condensation and cooling process reached to a level (at around t = 800 s) that the gas mixture in Vessel-1 becomes heavier than the gas mixture in Vessel-2 (Fig. 22), the helium-rich mixture starts to flow through the lower region of the IP (Phase 2 in Fig. 14) by reversing the previous tendency of flowing through the upper region (Phase 1 in Fig. 14). While the helium-rich mixture comes from Vessel-1 through the bottom of the IP, no perceptible stratification in the mixture density exists in Vessel-2 (Fig. 23). Although no significant differences in the density arises until approximately t = 1200 s, the gas mixture begins presenting a stratified tendency as the time proceeds. 5. Conclusion The results of two containment spray tests performed in the PANDA facility utilizing two interconnected vessels were presented in terms of normalized temperature, pressure, molar fractions and densities. The effects of a containment spray on the system depressurization by steam condensation and cooling, the breakup of a helium-rich layer and the mass transport between two vessels were investigated. The droplet size distribution of the spray droplets was determined with a separate cold test using optical techniques and image analysis. The depressurization and thus the efficiency of the spray as an accident mitigation device for a given spraying rate and temperature is strongly influenced by the amount of steam and non-condensible gases present initially in the system. While the very early phase of spray activation produces similar rates of pres-

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sure decay for the two tests, later pressure evolution starts to differ from each other due to the different initial non-condensible gas and steam contents and their different evolutions in time. It can be speculated that upward flow streams created by the spray induced gas circulation interferes with the density interface and erode the helium-rich layer. The characteristics of the helium-rich layer erosion highly depend on the initial gas compositions. The timing of the helium-rich layer breakup and mixing is gradually influenced by the density differences between the helium-rich layer and the gas mixture that is transported in upward direction by the spray induced flow streams. These density differences, which are varying during the operation of the spray, determine the degree of the negative buoyancy applied by the helium-rich layer to that flow streams. Therefore, the breakup of helium-rich layer was delayed in the system that contained air–steam, where initially larger density difference exists between the layer and the mixture below the spray. The gas transport to the adjacent vessel through the IP relies also on the initial ambient gas composition. When only steam is present in the initial gas mixture outside of the helium-rich layer, a counter-current flow with a definite direction throughout the whole test developed. When the air is introduced into the mixture composition, a flow reversal in the inter-compartment flow occurs and the helium-rich mixture flows along the bottom of the IP towards the bottom of the adjacent vessel. The flow reversal is primarily governed by the steam condensation and the corresponding evolution of mixture densities in both vessels. Along with the momentum exchange, mass transfer due to the condensation affected the characteristics of the gas atmosphere’s global convection in both vessels and in the IP. Gas transport between the two vessels was initiated by the spray activation in one vessel and resulted in an increase of the helium concentration in the other vessel. Acknowledgments The authors gratefully acknowledge the support of all of the countries and international organizations participating in the

OECD/SETH-2 project, the members of the Management Board and the Program Review Group of the SETH-2 projects. We would like to thank Dr. Michele Andreani from the analytical group for the fruitful discussions, the staff members Max Fehlmann, Chantal Wellauer and Wilhem Bissels for their engaged support in conducting these experiments, and Kasmira Pawa for the check through. References Abramoff, M.D., Magelhaes, P.J., Ram, S.J., 2004. Image processing with ImageJ. Biophoton. Int. 11 (7), 36–42. Andreani, M., Erkan, N., 2010. Analysis of spray tests in a multi-compartment geometry using the GOTHIC code. In: Proc. of 18th Int. Conf. on Nucl. Eng. (ICONE18), May 17–21, Xi’an, China. Babic, M., Kljenak, I., Mavko, B., 2009. Simulations of TOSQAN containment spray tests with combined Eulerian CFD and droplet-tracking modelling. Nucl. Eng. Des. 239, 708–721. Dreier, J., Paladino, D., Huggenberger, M., Andreani, M., Yadigaroglu, G., 2008. PANDA: a large scale multi-purpose, test facility for LWR safety research. In: Proc. of Int. Conf. on the Physics of Reactors (PHYSOR08), Interlaken, Switzerland. Hinds, W.C., 1999. Aerosol Technology Properties, Behavior, and Measurement of Airborne Particles. Wiley-Interscience Publication. Malet, J., Lemaitre, P., Porcheron, E., Vendel, J., 2005. Water Spray Interaction with Air–steam Mixtures Under Containment Spray Conditions: Comparison of Heat and Mass Transfer Model with TOSQAN Spray Tests. NURETH-11, Avignon, France. Malet, L., Blumenfeld, S., Arndt, M., et al., 2008. Sprays in containment: final results of the SARNET spray benchmark. In: The 3rd European Rev. Meeting on Severe Accident Research (ERMSAR-2008), September 23–25, Bulgaria. Mimouni, S., Lamby, J.-S., Lavirville, J., Guieu, S., Martin, M., 2009. Modelling of sprays in containment applications with a CMFD code. Nucl. Eng. Des. 240, 2260– 2270. OECD, 1999. State-of-the-Art Report on Containment Thermal–hydraulics and Hydrogen Distribution. NEA/CSNI/R (16). Presser, C., Gupta, A.K., Dobbins, R.A., Semerjian, H.G., 1990. Influence of size distribution on droplet mean diameter obtained by ensemble light scattering. In: Liquid Particle Size Measurement Techniques, vol. 2, ASTM STP1083, pp. 93– 111. Prosperi, B., Delay, G., Bazile, R., Helie, J., Nuglish, H.J., 2007. FPIV study of gas entrainment by a hollow cone spray submitted to variable density. Exp. Fluids 43, 315–327. Yoon, S.S., Hewson, J.C., DesJardin, P.E., Glaze, D.J., Black, A.R., Skaggs, R.R., 2004. Numerical modeling and experimental measurements of a high speed solidcone water spray for use in fire suppression applications. Int. J. Multiphase Flow 30, 1369–1388.