Two-phase flow phenomena along an adiabatic riser – An experimental study at the test-facility GENEVA

Accepted Manuscript Two-phase flow phenomena along an adiabatic riser - an experimental study at the test-facility GENEVA Tim Cloppenborg, Christoph Schuster, Antonio Hurtado PII: DOI: Reference:

S0301-9322(15)00024-5 http://dx.doi.org/10.1016/j.ijmultiphaseflow.2015.02.003 IJMF 2158

To appear in:

International Journal of Multiphase Flow

Received Date: Revised Date: Accepted Date:

17 June 2014 9 December 2014 2 February 2015

Please cite this article as: Cloppenborg, T., Schuster, C., Hurtado, A., Two-phase flow phenomena along an adiabatic riser - an experimental study at the test-facility GENEVA, International Journal of Multiphase Flow (2015), doi: http://dx.doi.org/10.1016/j.ijmultiphaseflow.2015.02.003

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(Draft) Tim Cloppenborg*1, Christoph Schuster2, Antonio Hurtado²

Two-phase flow phenomena along an adiabatic riser – an experimental study at the test-facility GENEVA Abstract Long adiabatic riser geometries with low system pressures are present in a lot of energy and petrochemical processes. Natural-circulation systems are an appropriate solution to save operating and maintenance costs. Under certain circumstances natural-circulation systems tend to unstable mass flow, especially in the riser section. The pressure gradients can stress the construction materials and affect the heat transfer. This paper focuses on GENEVA testfacility and its natural-circulation circuit with heat input by steam condensation. The GENEVA test-facility is explained in detail with focus on the local void fraction measurement system in the adiabatic riser section. The differences to former natural circulation test-facilities are particular emphasized. Therefore a transient experiment is presented and analysed. Moreover the influence of flow restrictions at the downcomer outlet is explained and an experimental method is presented, to determine the maximum natural circulation mass flow of the natural-circulation circuit. Besides, a comparison between the two different riser inner diameters, which were used during the experiments, is presented. The convective heat transfer is analysed by taking into account different dimensionless numbers. A variety of experiments were performed up to 100 kW el input power from the evaporators. Flashing and geysering as two types of occurred instabilities are stated and discussed in comparison to former test-facilities. Further phenomena like water hammer and counter current liquid flow are investigated. Based on these analyses constructive solutions can be derived, to stabilize the presented natural-circulation two-phase flow system.

1 Introduction A lot of research test-facilities like DANTON, CIRCUS, DESIRE and PUMA based on naturalcirculation have been constructed in the last decades (see e. g. Schuster 1991, de Kruijf et al. 2003, Furuya et al. 2002 and Kuran et al. 2006). The main aim of these test facilities was the boiling water reactor- and natural-circulation-stability analysis. Particular mention was on the enhancement of numerical tools (e. g. system or computational fluid dynamic codes). In the 1990s the focus moved more and more to passive decay heat-removal systems for nuclear power plants (NPP). Therefore large scale test facilities like PANDA in Switzerland and INKA in Germany were erected (see e. g. Dreier et al. 1999 and Leyer and Wich 2012). At PANDA test-facility different experimental investigations have been performed to analyse the operating behaviour of passive decay heat-removal systems for the KERENA™-reactor and the ESBWR-reactor described by Paladino and Dreier (2012). To perform integral tests with passive heat-removal systems, designed for next generation NPP, the INKA test-facility was built and is operated by AREVA GmbH in Karlstein (see e. g. Leyer et al. 2010). Especially noteworthy are transient experiments at both integral test facilities. For the investigations the maximum containment pressure was chosen as the ultimate safety limit because the containment represents the last barrier to the environment, which was mentioned by Dreier et al. (1999), Leyer et al. (2010) and Yang et al. (2013). For generic analysis of two-phase natural-circulation flow behaviour in dependency of the heat source and the geometric impact factors, GENEVA test-facility was built and 1

* Corresponding author: [email protected]; Technische Universität Dresden, Institute of Power Engineering, Chair of Hydrogen and Nuclear Energy, 01062 Dresden; +49 351 463 33898 2

Technische Universität Dresden, Institute of Power Engineering, Chair of Hydrogen and Nuclear Energy, 01062 Dresden

commissioned at the Chair of Hydrogen- and Nuclear Energy Technology (WKET), TU Dresden, in 2013. It is a downscale of the containment cooling condenser (CCC) but with several unique design and measuring features to analyse the influence on natural-circulation stability by changing the inner diameter of the downcomer and riser, the heat transfer surface, implementation of orifices, etc. (see Cloppenborg et al. 2013). Transient as well as steady state working point experiments were performed up to 100 kWel input power to investigate natural-circulation instabilities. Because of the generic character the main focus, during the investigations, was on the adiabatic vertical riser section, where a self-developed local void fraction measuring system (VFMS) with high time and spatial resolution was installed. Conversely to other natural-circulation test facilities no electrical heating elements were installed in the heated section of GENEVA. The heat input was realised by condensation of saturated steam at the outside of the slightly inclined condensation tubes. The strict modular construction of GENEVA allows generic experiments with different riser inner diameter variants (currently 20.0 mm and 38.0 mm). Both riser inner diameter variants and the influence of flow resistances at the downcomer outlet were experimentally investigated. Stability maps have been derived from the experimental data, where the subcooling was plotted as a function of the input power. From these steady state working points, three natural-circulation flow states were derived: stable single-phase flow (sSPF), unstable two-phase flow (uTPF) and stable two-phase flow (sTPF). During uTPF mass flow amplitudes which are up to six times higher than sSPF can occur (Cloppenborg et al. 2014a). In the riser section of the GENEVA flashing and geysering have been observed as the main instabilities for this kind of system. Even reversed flow has occurred during experiments in the whole test circuit. The main aim of the investigations is to understand the occurring instabilities like flashing and geysering in detail and to derive constructive improvements to decrease the amplitudes of the natural-circulation instabilities. Beside the optimisation of passive residual heatremoval systems, the experimental data can be used to develop efficient facility air conditioners, natural-circulation electrolysers or mobile cooling units. For the described semiopen loop system also other fluids may be suitable.

2 GENEVA test-facility In Figure 1a a scheme of the KERENA™-CCC and in Figure 1b a 3D-CAD scheme of the GENEVA test circuit are shown. In both schemes the direction of natural-circulation is counter-clockwise. The CCC is connected to the shielding and storage pool (1) and is located above the flooding pool (3). It consists of a downcomer (3), slightly inclined condensation tubes (4) and a riser (5). The GENEVA test circuit is approximately 9.0 m high which is three metres less than the respective CCC design of the KERENA™. In the schemes similarities between both circuits are indicated by numbers. The test circuit consists of the heat sink (1), which is located approximately 7.0 m above the steam chamber (2). The steam chamber contains the condensation tubes (4). The heat sink and the steam chamber are connected by the downcomer (3) and the riser (5).

Figure 1:

Schemes of the KERENA™-CCC and natural-circulation direction (black arrow) and the similarities with GENEVA test-facility

2.1 Design GENEVA is a semi-open natural-circulation loop test-facility with water as working fluid. It consists of three main circuits: -

Heat source circuit (to provide the saturated steam produced by the evaporators), Test circuit (to investigate different natural-circulation flow states) and Cooling circuit (to enable steady state working points and to subcool the steam in the heat source circuit).

These three circuits are shown in Figure 2, which is a simplified flow diagram of the testfacility. The numbers for the different components are the same as those in Figure 1. In Table 1 the explanation for the different measuring positions are presented, which are shown in Figure 2. The heat source circuit consists of two evaporators, which are connected to the steam chamber. The produced saturated steam condenses at the tubes in the steam chamber and flows into the condensate drum via a heat exchanger (HX). From the HX the condensate flows back through the return line to the evaporators. Due to the latent heat transfer through the condensation tubes, natural-circulation is initiated counter-clockwise in the test circuit (ref. to Figure 2). The condensation tubes are connected via the downcomer and the riser with the heat sink, which has a water inventory of approximately 400 litres. The heat sink is equipped with a condenser, an electrical heating element with a nominal power input of 3.5 kWel, and a plate heat exchanger (PHX). The PHX is installed at the heat sink to cool down the test facility and to perform steady state working point experiments. Those experiments are performed at constant evaporator power and constant fluid temperature at the inlet of the condensation tubes (CTI).

Figure 2:

Simplified flow diagram of GENEVA including the main components and main measuring positions

A crucial point for this kind of two-phase natural-circulation system is to determine the heat, which is transferred from the steam chamber into the test circuit, due to the different flow cross sections and unstable two-phase flow. An indirect method to determine the transferred heat is presented. As shown in Figure 2 the fluid temperatures are measured at the inlet and outlet of the PHX on both sides as well as the volume flow rate in the cooling circuit. Moreover the mass flow rate of the pump between the heat sink and the PHX is also recorded. When the pump between heat sink and PHX is switched on, the fluid in the heat sink is mixed thoroughly. This leads to a homogenous fluid temperature in the heat sink. The transferred heat can be calculated as a function of the volume flow rate and fluid temperatures in the cooling circuit, by the first law of thermodynamics. A comparison with the energy balance, on the other side of the PHX, leads to similar values. During a steady state working point experiment, the fluid temperature at CTI equals the homogenous temperature in heat sink. This fluid temperature is measured in front of the distributor (ref. to Figure 3), and is chosen to derive the subcooling during the experiments.

Table 1:

Exemplarily measuring positions at the riser

Measuring positions TC

Thermo-couple (typ K)

∆p

Differential pressure

prel

Relative pressure

pabs

Absolute pressure

V3

Volumetric flow

ṁ

Mass flow

ε

Local void fraction

In contrast to the INKA-CCC it is possible to vary the tube diameters of the downcomer and the riser, the number of condensation tubes and to implement orifices at GENEVA testfacility (ref. to Table 2). Referring to Figure 3 the steam chamber consists of four parts connected by flanges. The downcomer is connected by the distributor to the steam chamber and the four individually lockable condensation tubes, while the riser is connected by the drum to the condensation tubes in the steam chamber. Each condensation tube is approximately 4.0 m long with dimensions of 38.0 mm x 2.0 mm. Tube supports are installed inside the steam chamber. The fluid, which is condensed at the outside of the condensation tubes, is accumulated in the drum and flows upwards into the riser. To ensure the predominant flow direction the steam chamber and the condensation tubes are slightly inclined (approximately 6° angle) in the direction of the riser. Table 2:

Fact sheet about GENEVA test and heat source circuit

Test circuit

Heat source circuit


Test, max

[kg/ s]

~ 0.46

Tmax

[°C]

117.0

Pressureabs

[bar]

1.80

N tube

-

1…4

Downcomer-Ø

[mm]

10.0, 15.0

Riser-Ø

[mm]

20.0, 38.0

Riser length

[m]

6.5


steam

[kg/ s]

0.044

Tsteam

[°C]

144.0

Pressureabs

[bar]

4.0

Input power

[kWel]

120.0

Eight inlet nozzles for the saturated steam are installed at the bottom of the steam chamber. These inlet nozzles have an inner diameter of 19.0 mm. The steam velocity is reduced by the steam distributor tubes. Those are installed above the inlet nozzles to prevent a blow-down of the condensate from the condensation tubes. At the outside of the condensation tubes saturated steam condenses and latent heat is transferred into the test circuit. The condensation process can be observed by high speed and high resolution cameras through the windows, which are located at the top and also on one side of the steam chamber. The fluid in the condensation tubes is heated up and flows along the riser, which is approximately 6.5 m long, into the heat sink. In the condensation tubes as well as in the riser the fluid can reach saturation conditions. Therefore in both sections two-phase flow is possible.

Figure 3:

Sectional view of the steam chamber with interior parts and connections to the heat source and test circuit

All components are insulated with rockwoll, to ensure a minimization of heat losses. Therefore an insulation thickness of approximately 100 mm was chosen for the steam chamber. Due to the boundary condition of a quasi-adiabatic riser section an insulation thickness of approximately 40 mm was selected. Other parts of the test-facility were insulated with a minimum of 30 mm insulation thickness.

2.2 Instrumentation To monitor the occurring instabilities in detail, 12 type K thermocouples (TC) are installed along the adiabatic riser, as shown in Figure 4a. Moreover, a self-developed VFMS is installed in the riser. Due to the small inner riser diameters of 20.0 mm and 38.0 mm the measurement system should affect the flow as little as possible. This is the reason why needle shaped conductivity probes where selected. The developed triaxial probes have an outer diameter of 3.0 mm and an inner electrode diameter of 0.5 mm. The electrodes are made of stainless steel. These are isolated with PTFE-shrinking tubes towards each other. The low pressure and the moderate temperature in the riser allow a filigree design. As supply voltage a square-wave signal with 5.0 V and 10 kHz is chosen, in order to avoid electrolysis at the electrodes. The measuring voltage is measured over a shunt of 1.0 kOhm and can be calculated by Ohm’s law. The riser is divided in nine measuring levels, whereby each measuring level has three mounting points as shown in Figure 4a. These are arranged in a

120° angle to each other (ref. to Figure 4b) and in a 45 angle to the vertical axis of the riser (ref. to Figure 4c). In Figure 4b the third probe is indicated in a brighter colour to underline this as an optional position on each level. For the analysis of instabilities and the flow pattern in the riser, 16 conductivity probes are installed, whereof nine probes are installed in the riser centre and seven are installed at approximately 2/3 of the riser inner diameter (see Figure 4b). The measuring frequency per probe is 1000 samples per second and it is recorded simultaneously for all probes with a high-speed PCI-card. Accordingly the self-developed VFMS has a high time and spatial resolution and is capable to analyse, in connection with video analysis at the modular window section, the flow pattern along the riser. To allow a continuously installation of the probes in different depths, PTFE-seals are used instead of cutting rings. Eight probes use a type K TC as inner electrode to enable simultaneous local void fraction and temperature measurement in the riser (see Cloppenborg et al. 2014b). Table 1 shows the explanation for the different measuring positions, which are shown in Figure 4a. The VFMS was evaluated at a small transparent test-facility with different air-water flow mixtures. The high speed video recordings were compared with the synchronized high-speed measuring data for different flow patterns like bubble, slug and churn flows. Based on this, an uncertainty of approximately 15.0 % was derived. The uncertainty decreases with decreasing flow velocity. Furthermore six pressure transducers are connected to GENEVA test-facility. In each case one differential pressure transducer is connected along the downcomer, the condensation tubes and the riser, respectively. An absolute pressure transducer is connected at CTI (lowest point of the test-facility) and one relative pressure transducer is connected at the riser and at the steam chamber. In the test circuit the mass flow is measured with a Coriolis mass flow meter. The used measuring components (location, range and uncertainty) are listed in Table 3.

Table 3:

Overview about the used measuring components at GENEVA

Measuring devices and location

Range

Unit

Uncertainty

0 – 1000

°C

±1.5 K

- Downcomer

-1000 - 500

mbar

0.01 % of the measured value

- Condensation tubes

-500 - 500

mbar

0.01 % of the measured value

- Riser

-1000 - 500

mbar

0.01 % of the measured value

-2200 - 1000

mbar

0.01 % of the measured value

- Steam chamber

-1000 - 4000

mbar

0.01 % of the measured value

- Riser

-1000 - 1000

mbar

0.01 % of the measured value

0 …100

%

±15 % of the measured value

0 – 0.044

kg/ s

0.5 %of the measured value

0 – 1.0

kg/ s

0.1 % of the measured value

TC– type K - In the whole test-facility Pressure transducer - differential

Pressure transducer - absolute - CTI

Pressure transducer - relative

VFMS - Along the riser Vortex mass flow meter

- Heat source circuit Coriolis mass flow meter - Test circuit

a) b)

c)

Figure 4:

Self-developed local void fraction measuring system and installation in the riser with sectional view A-A and B-B on the right side

2.3 Comparison between GENEVA and other natural-circulation systems A crucial question is how the heat source affects the instabilities in a two-phase naturalcirculation system. Former natural-circulation test facilities like DANTON, DESIRE, GENESIS, or CIRCUS have electrical heating rods inside the vertical channel with a constant heat flux and with a chimney (adiabatic riser section) on top, to simulate a nuclear heat source (see e. g. Knorr et al. (2000), de Kruijf et al (2003), Manera and van der Hagen (2003), Marcel et al. (2008)). The heat transfer can be described by Fourier’s law. Manera and van der Hagen (2003) observed the same three operational regions as mentioned in the introduction. For their presented natural-circulation system, they report that flashing is the main instability and especially that the occurrence of flashing is independent of the flashing boundary in the riser. Moreover, their investigations showed that an increase of the system pressure reduces the dimension of the unstable two-phase flow region and the magnitude of the mass flow amplitude during unstable two-phase flow. Further experimental investigations of Manera et al. (2006) at the CIRCUS test-facility were focused on bubble size distribution measurements with wire mesh sensors. Experimental data was used for the development of a steady state model to estimate the radial void fraction profile in the riser. At the CIRCUS test facility experiments have been performed by Demaziére et al. (2008) with four heated

channels and two risers. They presented the influence of a high inlet restriction on the natural-circulation behaviour. Analysis showed that at low inlet subcoolings reversed flow occur in one channel. Under consideration of the investigations by Demizière et al. (2008) and by Marcel et al. (2009) it has to be pointed out, that an increase of the friction losses at the inlet of the heated section do not always lead to a stabilization of the two-phase naturalcirculation system. In contrast, one main unique characteristic of the presented natural-circulation test-facility GENEVA is the convective heat transfer to the test circuit, which is realised by condensation of saturated steam at the outside of the condensation tubes. This can be described by Newton’s law of cooling. The saturated steam is produced in the heat source circuit by electrical evaporators. As described before the steam chamber is composed of four individually lockable condensation tubes, which are slightly inclined (ref. to Figure 3). The inclination angle supports the predominant flow direction towards the riser. According to this the saturation parameters like pressure and temperature in the steam chamber are very important for the heat transfer to the condensation tubes (Cloppenborg et al. (2014a)). Thus the heat transfer to the condensation tubes is not constant, even during a steady state working point. This will be examined in the following sections. At the presented test-facility water is used as working fluid, whereas Freon-12 was used at DESIRE and Freon-134a was used at GENESIS. Another unique feature is the instrumentation of the riser in both, high time and spatial resolution for local void fraction and temperature distribution as described in the previous section. Furthermore the adiabatic riser section is two times longer in comparison to e. g. CIRCUS or DANTON, which has a significant effect on system dynamics, e. g. bubble growth and velocity. During the uTPF counter current liquid flow (CCLF) was observed in the riser. CCLF describes the change in flow direction, when the upwards flow velocity (in this case the steam velocity) is not sufficient to prevent a downwards flow from the fluid located above (in this case the subcooled liquid in the heat sink). One reason for that is the shortened retention time for the fluid in the condensation tubes, caused by the mass flow acceleration. CCLF is an important phenomenon in two-phase flow systems and especially in process and also in nuclear industry. Wallis et al. (1981) described CCLF for parallel vertical tubes analytically as well as experimentally. In the same year Richter (1981) presented an improved correlation for the CCLF phenomenon. Furthermore Katto (1994) presented an analytical study for limiting conditions of CCLF. Issa and Macian (2011) published a review about CCLF with focus on nuclear applications e. g. CCLF during LOCA in the hot leg of a PWR.

3 Natural-circulation and occurring instabilities In contrast to former test-facilities with electrical heating rods, e. g. inside the riser section, where flashing occur, it has to be pointed out how the uTPF affects the heat transfer of the presented passive heat-removal system and vice versa. Besides geometric influence factors this is very important to the overall stability of the natural-circulation two-phase flow. Driving force for natural-circulation is the density difference between downcomer and heated section/ riser. When the steam chamber is filled with saturated steam latent heat is transferred to the fluid inside the condensation tubes, which is heated up. Owing to this, density of the fluid decreases and the heated up fluid flows upwards in the riser, based on the inclination of the condensation tubes. Since the heated section has no significant height difference between inlet and outlet (approximately 0.5 m), its influence on the pressure in the test circuit can be neglected. Due to the reduction of the hydrostatic head along the riser saturation temperature is decreasing in the same manner. In dependency of the heat input and the subcooling of the fluid at CTI, saturation conditions can be reached at the end of the adiabatic riser. In this case the fluid evaporates, which is called adiabatic boiling or flashing. This causes an uTPF in the riser. The main steps are shown in Figure 5, which is described in this section.

The natural-circulation stability depends mostly on the inlet subcooling and the input power. The performed experiments at low inlet power and high subcooling showed that sSPF is existing, while sTPF is present at inlet temperatures near saturation conditions (Cloppenborg et al. 2014a). The result is a large region of uTPF for both riser variants.

Figure 5:

Flow diagram of an flashing instability in the uTPF region

As a consequence of the decreased density in the riser the mass flow increases in the test circuit. This leads to higher Reynolds (Re) and higher Nusselt (Nu) numbers inside the condensation tubes, which are the crucial parameters for the heat transfer coefficient αin. The influence of the Pr number is less, compared to the Re number, which increases up to a factor of six as a result of the higher mass flow rate during two-phase flow. The transferred heat from the steam chamber to the test circuit depends on the mass flow rate in the test circuit as well as on the saturated steam parameters in the steam chamber, which are temperature and pressure. Assuming almost constant parameters in the steam chamber, the latent heat transferred at the outer condensation tube surface Atube,out can be described by the following equation as convective heat transfer.

     , ,  ∗  ,  ∗ ∆
, 

eq. 1

In this case ∆Tm,out is the mean temperature difference between saturated steam (Tsteam) and condensation tube wall (Twall,out). Under the boundary condition that the wall temperature is constant in longitudinal direction (Ltube) of the condensation tubes, the following equation can be derived for the conducted heat through the condensation tube wall. Further assumptions are constant thermal conductivity λ wall and radius rtube of the condensation tubes

   ∗

,  − ,  ,  1 ∗ ln 2 ∗  ∗    ,

eq. 2

The second convective heat transfer, in this case from the inner wall of the condensation tubes to the fluid, can be described by the following equation.

 # %    , &,  ∗  , ∗ ∆
,

eq. 3

For the reason that inlet and outlet temperature of the condensation tubes are different, the mean temperature ∆tm,in can be calculated with the following equation:

∆',()

*+,(-,.,/ − *+,(-,() 01++,() − *+,(-,() ln  01++,() −*+,(-,.,/

eq. 4

Thus, both convective heat transfer terms (eq. 1and eq. 3) are connected by the conductive heat transfer through the wall of the condensation tubes. Accordingly, a decrease of the inner wall temperature, e. g. through an increase of the natural-circulation mass flow rate, leads to a higher heat transfer and reduction of the outer wall temperature. Subsequently, the condensation rate of the saturated steam is increased in the steam chamber, which is followed by a reduction of the steam chamber pressure. As a consequence, the temperature of the saturated steam is reduced and also the convective heat transfer to the outer condensation tube surface. Due to inertia effects e. g. heat conduction and capacity of the tube wall, this effect results time delayed in a decrease of the transferred heat from the inner tube wall to the fluid. Moreover, as a result of the increasing mass flow rate the retention time for the fluid in the condensation tubes decreases. With respect to the decreasing terminal temperature difference from steam chamber to test circuit, saturation conditions cannot be reached in the riser anymore.

4 Experiments 4.1 General behaviour of the test circuit Natural-circulations systems have three characteristic flow states; these are sSPF, uTPF and sTPF. Associated with natural-circulations systems with electrical heating rods articles about these flow states have been published more and more in past years, e. g. Demezière et al. (2008) and Manera and van der Hagen (2003). To understand the general behaviour of the presented natural-circulation system, for passive decay heat-removal, transient experiments are essential. It has to be clarified how the fluid temperature in the heat sink affects the natural-circulation and occurring instabilities. Moreover it has to be examined if all three flow states occur under the chosen boundary conditions. Therefore an experiment with constant evaporator power of 20 kWel with a DN20 riser is presented in Figure 6. Exemplarily, the riser inlet and outlet temperature (Tr,in and Tr,out ) as well as three temperature signals on different heights in the heat sink (Ths,1…3) and at CTI (TCTI ) are presented to investigate the behaviour of the test circuit, while the temperature in the heat sink increases. The temperature sensors Ths,1 to Ths,3 are installed from top to the bottom at the heat sink. Moreover the mass flow signal of the test circuit is shown in the middle diagram of Figure 6. For higher accuracy the mass flow is measured with a Coriolis mass flow meter in the test circuit, which is installed in the horizontal part of the downcomer. In the lowermost diagram of Figure 6 differential pressure signals along the downcomer (∆pdc), condensation tubes (∆pct ) and riser (∆pr) are shown. Figure 6 can be divided into four main stages, which are: start-up (A, 0 s to 400 s), sSPF (B, 400 s to 7000 s), uTPF (C, 7000 s to 10170 s) and sTPF (D, 10170 s to 11500 s). 4.1.1 A: start-up (0 s … 400 s) During the first 400 s the fluid in the condensation tubes is heated up, by condensing saturated steam on the outer tube surface. No natural-circulation is detected in the test

circuit. Due to the occurring density differences creeping flow is possible in the condensation tubes, which is too small to be detected by the Coriolis mass flow meter. An indication for this is the increase of the riser inlet temperature. The differential pressure signals are constant. 4.1.2 B: sSPF (400 s … 7000 s) In consequence of the heat input, temperatures in the riser increase, whereby a constant time delay can be recognized between riser inlet and outlet. A stable single phase naturalcirculation flow rate of approximately 0.04 kg/ s is established and increases continuously up to 0.07 kg/ s. Hence the water inventory of the heat sink is warmed up constantly. The temperature increase in the riser slows down. Parallel to this temperature stratification appears in the heat sink, due to the fluid density differences. The stratification is very stable, which is indicated by the time delayed and concurrent increase of the temperature signals in the heat sink. At last, temperature a CTI increases in the same manner. For this reason the diagram is split from approximately 3900 s to 6020 s. From approximately 400 s to 1500 s a low single-phase mass flow rate is established and differential pressures slightly decrease in the test circuit. The reason for this is the redistribution of the fluid in the condensation tubes. The heated up fluid with lower density flows upwards towards the riser inlet, while the less heated up fluid partially flows in the other direction. When the mass flow rate is increased and stabilized, the absolute amount of the differential pressure along the downcomer is slightly higher than the absolute amount of the differential pressure along the riser. This equals almost the absolute amount of the difference between the actual differential pressure value along the condensation tubes and its initial value at t=0 s. 4.1.3 C: uTPF (7000 s … 10170 s) Based on the semi-open loop characteristic of the test-facility the hydrostatic pressure is reduced from the riser inlet to outlet. When the fluid at the riser outlet reaches saturation conditions it evaporates and decreases the hydrostatic pressure further. This phenomenon is described in the literature as flashing (Manera and van der Hagen (2003) and Furuya (2006)). The increased density difference between downcomer and riser leads to an increase of the natural-circulation mass flow rate. Due to the correlation between mass flow rate and pressure loss, differential pressure signals along the downcomer and riser show increasing amplitudes. Subsequently, the retention time of the fluid in the condensation tubes is decreased. In the following, saturation conditions are not fulfilled anymore at the riser outlet and that is why the evaporation stops. Mass flow rate decreases back to stable singlephase flow rate. Every mass flow peak leads to a mixture of the stratified fluid inventory in the heat sink. As a result, the fluid temperature in the heat sink is homogeneous. Time delayed fluid temperature, in the heat sink, and fluid temperature at CTI are the same. The further increase of the fluid temperature at CTI leads to an increasing amplitude (up to 0.32 kg/ s) of the natural-circulation mass flow rate in the test circuit. This can be explained by applying the first law of thermodynamics for an energy balance around the condensation tubes. Therefore an almost constant heat input during a constant single phase flow rate and increasing enthalpy at CTI are preconditions. As a consequence the enthalpy at the outlet of the condensation tubes (equals riser inlet) has to increase, which is indicated by higher riser inlet temperatures. Subsequently, the flashing front develops further downwards the riser, when saturation conditions are reached at the riser outlet. This leads with every mass flow peak to a higher density difference between downcomer and riser. Moreover it is interesting that after approximately 8900 s, when fluid temperature in the heat sink and along the downcomer is almost homogeneous, differential pressure fluctuations appear along the condensation tubes during a mass flow peak. The amplitude of these fluctuations also increases with every mass flow peak. This behaviour can be explained with water hammer in the riser after a mass flow peak.

Time [s]

Temperature [°C]

0

1000 2000 3000

7000

8000

11000

80 60 40

A

Mass flow [kg/ s]

10000

100

20

Pressure [mbar]

9000

120

Ths,1

Ths,2

Ths,3

TCTI

Tr,in

Tr,out

C

B

D

0.4 0.3 0.2 0.1 0.0

0 -100 -200 -300

∆pdc 0

Figure 6:

∆pr

1000 2000 3000

∆pct 7000 8000 Time [s]

9000

10000

11000

Transient experiment with DN20 riser and constant input power

4.1.4 D: sTPF (10170 s … 11500 s) At approximately 84.0 °C the constant heat input is sufficient enough to generate a sTPF along the whole riser. The riser outlet temperature stays constant, while the fluid temperatures in the heat sink and at CTI increase. Also the riser inlet temperature increases slightly. Subsequently, the stable two-phase mass flow rate increases up to approximately 0.22 kg/ s. Hence the differential pressure along the riser and the downcomer decrease, while the absolute amount of the signal along the downcomer continue to be higher than the absolute value of the signal along the riser. The minor fluctuations of the riser signal are higher in comparison to the fluctuations of the downcomer signal. The reason for this is the varying ratio between liquid-phase and steam-phase of the two-phase flow along the riser. The differential pressure along the condensation tubes is constant with comparable fluctuations, which have a higher frequency than the mass flow oscillations during uTPF. When the fluid temperatures in the heat sink and at CTI reach approximately 90 °C the evaporators are shut down and then the test-facility is cooled down.

4.2 Influences on the flow stability In this section two main constructive influence factors will be examined affecting the naturalcirculation mass flow rate. These are an inlet restriction in front of the heating zone (at the end of the downcomer) and the riser inner diameter. An approach to calculate the maximum mass flow rate of a natural-circulation system will be discussed and compared with experimental data sets. 4.2.1 Downcomer One of the most important variables, characterising a natural-circulation system, is the mass flow rate. Owing to the higher mass flow rate two-phase natural-circulation flow is preferred for most systems, but tends to unstable behaviour under specific circumstances. Different authors have published papers about the positive effect on the two-phase flow stability by increasing the inlet restriction in the single-phase region of a natural-circulation system. Marcel et al. (2009) investigate the effect of an inlet restriction experimentally and with a numerical model, which considers the homogeneous equilibrium model and friction losses concentrated at the riser inlet and outlet. The reduced mass flow rate promotes at the same heat input an increase of the fluid enthalpy at the outlet of the heating zone. Due to this boundaries of uTPF and sTPF are shifted to higher subcooling.

But a crucial question has not been discussed in the literature so far, concerning the determination of the maximum natural-circulation mass flow rate of a specific system. A theoretical approach was done for GENEVA test-facility by using Toricelli’s law. This leads to the assumption that in the described natural-circulation system the maximum fluid velocity is limited by the outflow rate of the downcomer. The fluid velocity c can be determined as shown in eq. 5. This equation is valid for short outflow cross sections, a constant water column h above the outflow cross section, acceleration due to gravity g and neglect of friction losses. % 22 ∗ 3 ∗ 4

eq. 5

Referring to Figure 7 the maximum mass flow at the downcomer outlet can be calculated to approximately 2.20 kg/ s, without taking into account the pressure losses, especially due to friction. With respect to pressure losses the mass flow at the outlet of a DN15 downcomer can be calculated to approximately 0.49 kg/ s, which is four times less, without considering pressure losses. Most important influence factors are bend losses and the tube roughness for the small inner downcomer diameter.

Figure 7:

Comparison of the different calculated and experimentally proven mass flows of the DN15 downcomer

Experiments with a ball valve, installed at the downcomer outlet, are in good agreement to the calculated maximum outflow rate through the downcomer with an inner diameter of 15 mm. If the ball valve is fully open its inner diameter equals exactly the downcomer inner diameter. Open and close operations of valves are shown in Figure 8. The maximum mass flow rate is approximately 0.46 kg/ s, as shown in Figure 7 and Figure 8. This is in good agreement with the experimental data.

Figure 8:

Outflow experiment at GENEVA to determine the maximum mass flow through the DN15 downcomer and in the test circuit, with a ball valve

Another outflow experiment was performed with a needle valve. The maximum mass flow rate is approximately 0.29 kg/ s. The decrease of the mass flow rate is caused by the smaller inner cross-section diameter (approximately 8.0 mm) and the flow diversions inside the needle valve. Due to the unique design of the presented test-facility it was possible to calculate and experimentally prove the maximum natural-circulation mass flow rate under consideration of Toricelli’s law. 4.2.2 Inner diameter of the riser Both, transient and steady state experiments have been performed at the GENEVA testfacility up to 100 kWel. To underline the influence of the riser inner diameter on the naturalcirculation flow, eight experimental mass flow time traces from the GENEVA test-facility are presented in Figure 9. The four diagrams in the left column (Figure 9a to Figure 9d) show experimental data sets for a DN20 riser at different steady state working points. To compare both riser variants with each other, Figure 9e to Figure 9h show experimental data sets for a DN40 riser at approximately the same temperatures at CTI and for the DN20 riser. For all the experiments, presented in Figure 9, a constant input power of approximately 35.0 kWel was chosen.

For the DN20 riser variants, presented in Figure 9a to Figure 9d, the temperatures at CTI are constant. In comparison to experiments presented by Manera and van der Hagen (2003) at the CIRCUS test-facility mass flow fluctuations occur at the GENEVA test-facility even at higher inlet subcoolings. The frequency of the mass flow peaks increases from Figure 9a to Figure 9b. In Figure 9a the mass flow fluctuates between stable single phase mass flow rate of 0.07 kg/ s and a two-phase flow rate of approximately 0.13 kg/ s. With increasing temperature at CTI up to approximately 60 C (ref. to Figure 9b), a sinusoidal oscillation with a higher average mass flow rate occurs. A further increase of the temperature at CTI leads to a stable two-phase mass flow rate of approximately 0.20 kg/ s (ref. to Figure 9c). In Figure 9d

the mass flow rate fluctuates with higher amplitude than in Figure 9c, while the frequency is lower in comparison with Figure 9b. Time [s] DN20 0 100 200 300 400 500 100

TCTI

100 e) 80

80

0.5 0.4 0.3 0.2 0.1 0.0

b)

0.5 0.4 0.3 0.2 0.1 0.0

c)

60

60

40

40

80

80

60

60

40 100

TCTI

80

0.5 0.4 0.3 0.2 0.1 d) TCTI 0.0 0 100 200 300 400 500 DN20

Figure 9:

100 f)

100

TCTI

Time [s]

TCTI

0.5 0.4 0.3 0.2 0.1 0.0

TCTI

0.5 0.4 0.3 0.2 0.1 0.0

TCTI

0.5 0.4 0.3 0.2 0.1 0.0

40 100 g) 80

60

60

40

40

100

100

80

80

60

60

40

40

0.5 0.4 0.3 0.2 0.1 0.0

h) TCTI

0

Mass flow [kg/ s]

a)

Temperature [°C]

Mass flow [kg/ s]

0.5 0.4 0.3 0.2 0.1 0.0

Time [s] DN40 0 100 200 300 400 500

100 200 300 400 500

DN40

Time [s]

Experimental mass flow time traces in dependency of temperatures at CTI and riser inner diameter for constant input power

With respect to experimental data presented in Figure 9e to Figure 9h for the DN40 riser variant, three main differences between both riser variants are obvious. These are the increasing frequencies and the increasing amplitude of the mass flow peaks as well as the fluctuations of the temperature at CTI in Figure 9f to Figure 9h, after a mass flow peak. Hereinafter, the reasons and phenomena for the different behaviour need to be clarified. Figure 9e shows the mass flow rate at high subcooling at CTI for a DN40 riser. The fluctuation periods with rising and sloping amplitudes are based on flashing and CCLF at the riser outlet. When saturation conditions are reached at the riser outlet, flashing starts. In this case large bubbles condense in the high subcooled fluid in the heat sink near the riser outlet. The higher mass flow rate leads to a decreasing retention time of the fluid in the condensation tubes. The velocity of the gaseous phase decreases. At this point the flow pattern changes and Taylor bubbles occur, which do not cover the whole riser inner diameter. Due to this an annular CCLF appears and subcooled fluid from the heat sink flows downwards the riser wall. The mixture of the downwards and upwards flowing fluids has the consequence that no saturation conditions are anymore possible at the riser outlet. Subsequently, mass flow rate decreases back to stable single-phase flow conditions. The

retention time of the fluid in the condensation tubes increases again. This process is repeated several times, Owing to the fact, that the riser is filled with heated up fluid (near saturation conditions) below the mixture volume at the outlet. When the heat transfer from the steam chamber is not sufficient enough, longer periods of stable-single phase flow occur between the fluctuation periods. For the CIRCUS test-facility Demazière et al. (2008) investigated the influence of a high inlet restriction and have observed reversed flow in one of the two risers, but at low subcooling. Due to the shorter riser length and lower subcooling the thermal shock for the construction material was not as high as observed at the GENEVA test-facility. This will be examined further in the sections 4.3 and 4.4. In Figure 9f the influence of the decreased subcooling at CTI is presented. As explained in section 4.1.3 the mass flow rate amplitude increases because of higher fluid temperatures in the riser and thus the length of the flashing front increases. Consequently, the increasing density difference between riser and downcomer as well as the reduced friction losses, according to the increased riser inner diameter, leads to higher mass flow amplitude. The stated experimental data of both riser variants allow the conclusion, that the amplitude increases almost proportional to the riser inner diameter up to the maximum mass flow rate of the natural-circulation system. The main reason is the strong inverse dependency between the friction pressure loss and the riser inner diameter. As mentioned for Figure 9e CCLF is also present for Figure 9f to Figure 9h after a mass flow peak. Due to the increased length of the flashing front the subcooled fluid from the heat sink can even fall down to the riser inlet. As a result heavy water hammer appear in the riser. The occurring impulse is the reason for the rise in temperature at CTI after a mass flow peak. The impulse leads to a local reversed flow in the condensation tubes, so that heated up fluid reaches the measuring sensor at CTI. This is supported by the short mass flow rate stagnation after the peak. Because of the heat, which is transferred to the condensation tubes, single-phase mass flow rate is established again. As shown in Figure 9f to Figure 9h fluid temperature returns back to the initial value, which underlines the homogenous temperature of the fluid inventory in the heat sink and along the downcomer. The duration of sTPF periods increases, while the subcooling at CTI decreases (ref. to Figure 9h). The relationship between system variables can be analysed in more detail by deriving the phase portrait. In a phase space the chosen variables are not plotted in dependency of the time, but with a constant time shift. By plotting one variable along the other, the dynamic behaviour of a system can be further examined. For Figure 10a to Figure 10h the same datasets are used as shown in Figure 9. In this case the differential pressure along the downcomer and the mass flow rate in the test circuit are chosen. In general, for the chosen heat input the differential pressure amplitude is lower for the DN20 riser variant in comparison to the DN40 riser variant. Referring to the left column of Figure 10, all four phase portraits have a flat shape in vertical direction. With increasing temperature at CTI mass flow rate increases, while the differential pressure amplitude decreases (ref to Figure 10a to Figure 10c). As shown in Figure 10c sTPF is characterised by a stable mass flow rate and small differential pressure differences. But a further increase of the temperature at CTI leads to a different behaviour, as shown in Figure 10d. In this case the relationship between mass flow rate and the amplitude of the differential pressure along the downcomer seems to be inversely proportional. A reason for this can be that high void fraction is even present in the condensation tubes and decrease those inner heat transfer coefficient. Subsequently, void fraction is reduced and flow pattern in the riser changes from annular to churn flow, which leads to a decrease of the mass flow rate. Due to this, CCLF is promoted in the upper part of the riser. When saturated liquid covers the inner wall of the condensation tubes again, the heat transfer coefficient increases, and the procedure is repeated.

-200 0.3

∆pressure [mbar]

-150

-100

-50

DN40 0

∆pressure [mbar]

-600 -400 -200

0

200 400 600 0.5 0.4 0.3 0.2 0.1 0.0

a) TCTI = 40 °C

e) TCTI = 35 °C

b) TCTI = 60 °C

f) TCTI = 60 °C

0.5 0.4 0.3 0.2 0.1 0.0

c) TCTI = 78 °C

g) TCTI = 78 °C

0.5 0.4 0.3 0.2 0.1 0.0

0.2 0.1 0.0 0.3

Mass flow [kg/ s]

0.2 0.1 0.0 0.3 0.2 0.1 0.0 0.3

0.5 0.4 0.3 0.2 0.1 0.0

0.2 0.1 0.0

h) TCTI = 90 °C

d) TCTI = 89 °C

-200 DN20

Figure 10:

-150

-100

-50

∆pressure [mbar]

0

-600 -400 -200 DN40

Mass flow [kg/ s]

DN20

0

200 400 600

∆pressure [mbar]

Phase portraits of mass flow rates and differential pressures along the downcomer in the test circuit, at different inlet temperatures and with two riser variants

In contrast to this, the system behaviour with a DN40 riser is different at the same input power and comparable temperatures at CTI. As shown in Figure 10e, for high subcooling at CTI the relation between both variables in the phase portrait seems to be comparable to the behaviour at similar subcooling at the DN20 riser variant (ref. to Figure 9a). But with decreasing subcooling at CTI the behaviour is quite different between both riser variants. This is shown in Figure 10f to Figure 10h, where the direction of the trajectories is clockwise. After flow stagnation in the test circuit absolute amount of both, mass flow rate and differential pressure, increases and reaches maximum values. Comparing Figure 10f to Figure 10g, it is obvious that water hammer occur, but not after every mass flow peak. The careful analysis of the experimental data shows, that the pressure peak amplitude, induced by water hammer, decreases proportional to subcooling at CTI. This is caused by the reduced temperature difference between the CCLF from the heat sink and the upwards flowing fluid in the riser. Furthermore, during the flow stagnation small pressure peaks (in positive direction) are noticeable. These small pressure peaks support the assumption, that heated up fluid from the condensation tubes is pushed into the distributor and flows upwards in the downcomer.

Frequency analyses based on FFT have been performed to examine the relationship between mass flow rate oscillation frequency and subcooling at CTI. The main frequencies are shown in Figure 11, for both riser variants. During the transition from sSPF to uTPF conditions, both riser variants show nearly the same frequency. The smaller riser shows a clear increase in frequency with decreasing subcooling. During sTPF no significant frequencies occur. This happens when temperatures at CTI are between 70 °C and 85 °C (ref. also to Figure 9c and Figure 10c). Above approximately 85 °C two-phase mass flow oscillations are noticeable, with a lower frequency. An increase of the riser inner diameter shifts the described relationship between subcooling at CTI, mass flow rate frequency up to higher frequencies and lower subcooling. Temperatures above approximately 80 °C causes lower frequencies of the mass flow rate, which can be explained by the increase of the sTPF periods, as shown in Figure 9h. This has to be examined by an analysis of the heat transfer on the inside and outside of the condensation tubes.

DN20 DN40

90 80 TCTI [°C]

70 60 50 40 30 -3

10

-2

10

Frequency [Hz]

Figure 11:

Maximum and Minimum oscillation frequencies for different TCTI and two different riser inner diameters, at constant input power

To take a closer look on how the riser inner diameter affects the heat transfer, different nondimensionless numbers are necessary. At first the Re-Number at CTI is calculated for low and high mass flow rates in the test circuit. The presented experimental data is taken from steady state working points at approximately 35.0 kW el for both riser variants. An extract of this experimental data is shown in Figure 9. The Re-Number is calculated with the following equation.

& '15;789 

:;<= ∗
/>?/ '15;'() @ ;<=  ∗ A;<= ∗ ;<=

eq. 6

The relation between Re-Number at CTI and the inlet subcooling is presented in Figure 12. All calculated Re-Numbers are higher than the critical Re-Number, which is approximately 2300. For the presented data, the Re-Numbers indicate turbulent natural-circulation flow in the test circuit. For the maximum Re-Number of the DN40 riser a clear trend in dependency of the inlet subcooling can be derived. In general, the Re-Numbers of the system with the DN40 riser are higher in comparison to those of the system with the DN20 riser at the same subcooling at CTI. For the DN40 riser, the low Re-Number is not calculated, because of the flow stagnation after a mass flow peak. But for the transition from sSPF to uTPF, low Renumbers for both riser variants are nearly the same. In contrast to the DN40 riser, no flow stagnation occurs at experiments with the DN20 riser. In the uTPF region the difference between low and high Re-numbers increases for the DN20 riser. With respect to the DN20 riser, sTPF conditions are established, when the subcooling at CTI decreases, and the

difference between low and high Re-Numbers decreases. Above 85 °C the high Re-Number is constant, while the low Re-Number decreases for the DN20 riser.

90 80

DN20, low DN20, high DN40, low DN40, high

TCTI [°C]

70 60 50 40 30 10

4

10

5

Reinlet

Figure 12:

Reinlet--Number for low and high natural-circulation mass flow rate, for different T CTI, and two different riser inner diameters at constant input power

Due to the high Re-Numbers forced convection predominates in the test circuit, regardless of the riser inner diameter. The following equation was chosen to calculate the Nu-Number for the heat transfer on the inside of the condensation tubes. The friction factor k has to be taken into account and is calculated as a function of the Re-Number. The Nu-Number for the inner heat transfer is described as a function of the Re-Number, Pr-Number, and the inner diameter of the condensation tubes di,ct, as well as the length lct. The Pr-Number is a dimensionless number, which describes the fluid properties. In this case, the referenced values are calculated with Tr=(Tsc+TCTI )/2.



with

 ()?(->

@ B  ∗ A B  ∗ %C B   B 

M E D G ∗ & ∗  :(,Q/ N 8 ∗ P1 + L O R M Q/ E 1 + 12.7 ∗ KD8 G ∗ L N − 1O

E 1.8 ∗ logUV & − 1.5XM

eq. 7 eq. 8

eq. 9

In Figure 13 the relationship between the Nu-Number and the Re-Number is shown for both riser variants. As one would have expected, Nu-Number is proportional to the Re-Number. The higher mass flow rates in the DN40 riser lead to higher Nu-Numbers, in comparison to the DN20 riser.

Nuinside

DN20, low DN20, high DN40, high

100

10

4

10

Figure 13:

Reinlet

10

5

Relationship between Nuinside-Number and Reinlet-Number for two different riser inner diameters at constant input power

In contrast to the heat transfer on the inside of the condensation tubes, free convection predominates the heat transfer on the outside of the condensation tubes. To characterize the influence of the subcooling at CTI on the convective heat transfer from the steam chamber into the condensation tubes, a modified Gr-Number Gr+ was derived. The characteristic diameter d is calculated with d=π/2*di,ct.

Y

Z B  ∗ : N ∗ 3 ∗ ?Q − ;<=  @ B M

eq. 10

A more accurate description of the convective heat transfer is possible by deriving the RaNumber, which depends on the Pr-Number and Gr-Number (Ra+=Gr+*Pr). To underline the influence of the Pr-Number both, a modified Gr-Number Gr+ and a modified Ra-Number Ra+ are plotted for different fluid temperatures at CTI and both riser variants, shown in Figure 14. The Ra+-values are higher as the Gr+-values. It has to be taken into account that the PrNumber decreases, while the fluid temperature increases. As a consequence, Ra+ decreases proportional to the subcooling. The Gr+-values and the Ra+-values of the two riser variants getting closer, while the subcooling decreases.

10 90 80 TCTI [°C]

Gr+

8

9

10 DN20, Ra DN20, Gr DN40, Ra DN40, Gr

70 60 50 40 30 8 10

9

10 Ra+

Figure 14:

Modified Gr+-Number and Ra+-number for different TCTI and two different riser inner diameters at constant input power

With the Ra+-Number and Pr-Number it is possible to derive the Nuoutside-Number for free convection from the steam chamber to the condensation tubes. In this case the following empirical correlation is chosen for the horizontal tube geometry. The factor f considers the influence of the Pr-Number.

with

U M

 .,/?(-> [0,752 + 0,387 ∗ ^&Y ∗ _ `a b _  P1 + L

Ua d Xd Ua

0,559 O R 

eq. 11

eq. 12

The modified Gr+-Number describes the free convective heat transfer as a function of the temperature difference between steam chamber and the condensation tubes. In contrast to this, Nuinside depends on the Re-Number, and thus on the density difference between downcomer and riser. In Figure 15 the Nuinside-Numbers are shown as a function of the Nuoutside-Numbers for both riser variants. For the DN20 riser Nuinside is shown for low and high Re-Numbers. The Nuinside-Numbers are higher as the Nuoutside-Numbers regardless of the riser inner diameter. For the DN20 riser, sTPF is established, when the Nu-Numbers on the inside and outside are nearly the same. This promotes equilibrium between the heat transfer on the inside and outside of the condensation tubes. As a consequence of the reduced maximum mass flow rate the Nu-Numbers for the DN20 riser are smaller in comparison to the Nu-Numbers of the DN40 riser at the same subcooling.

10

3

Nuinside

DN20, low DN20, high DN40, high

10

2

115 120 125 130 135 140 145 150 Nuoutside

Figure 15:

Relationship between Nuinside-Number and Nuoutside-Number for two different riser inner diameters at constant input power

The analyses show, that the oscillation frequencies increase with increasing riser inner diameter, because the convective heat transfer is proportional to the riser inner diameter. This has been pointed out by deriving the dimensionless numbers for the convective heat transfer on the inside and outside of the condensation tubes.

4.3 Natural-circulation at low power input In Figure 9e to Figure 9h the influence of the riser inner diameter and the subcooling on the natural-circulation mass flow rate is presented for constant input power of 35 kW el. To examine the behaviour of flashing in connection to the CCLF in more detail, experimental results with steady state working points have been evaluated and are presented in this section with DN40 riser configuration for three different temperatures at the CTI (55 °C, 77 °C and 85 °C) and with constant input power of 25 kWel. To analyse the axial distribution of the local void fraction in the riser centre, in relationship with the axial temperature distribution along the riser, two diagrams were included in Figure 16, and Figure 18 to Figure 19. The upper diagram in each figure shows the local void fraction signal for the probes in the riser centre, over riser length and time. In the diagram below the temperature distribution from riser inlet to outlet is shown for the same time interval, whereby isothermals (dotted lines) for the upward flow provide further details of the dynamic system behaviour in the uTPF region. Furthermore the mass flow in the test circuit is shown as black graph in each lowermost diagram. 4.3.1 Medium subcooling with low amplitudes (TCTI = 55 °C) Referring to Figure 16, starting at 40 s in both diagrams, it is clear that saturation conditions are not fulfilled along the riser. Due to this no void is indicated and sSPF is approximately 0.07 kg/ s. At approximately 51 s saturation conditions are reached at the riser outlet and the fluid, with almost 102 °C, evaporates. This leads to a further decrease of the hydrostatic pressure along the whole riser and causes further evaporation. The pressure reduction in the riser causes an increase in the mass flow rate up to 0.28 kg/ s. Due to the increasing ReNumber, the heat transfer increases inside the condensation tubes. At approximately 57 s the length of the flashing front is approximately 4.0 m. Meanwhile the mass flow reaches its peak, while the pressure and temperature in the steam chamber decrease slightly. This is shown in Figure 17, which provides an overview about the relationship of different measuring signals in the test circuit and the steam chamber (heat

source circuit). Measuring signals for the probes in the riser centre and near the wall were presented (ref. Figure 17 L8 to L9). These show an increase of the void diameter along the riser. The local void fraction time traces of levels L8 and L9 show that the void diameter covers more than 0.44 % of the riser cross section, with respect to the signals of the eccentric probes. If the pressure in the steam chamber is reduced, the heat transfer to the test circuit decreases too. Time delayed saturation conditions are not fulfilled at the riser outlet and evaporation stops. The temperature signals indicate at approximately 55 s cold fluid at the riser outlet. Time [s] 100

50

150

200

[%] Local void fraction

Riser height [m]

0 6.0 4.5 3.0 1.5 0.0

100 80 60 40 20 0

Mass flow 102

102

4.5

105

105

3.0

102

102

0.4

105

105

0.3 0.2

106 1.5 0.0

[°C]

0.5

0.1 0.0

0

50

100

150

-0.1 200

Mass flow [kg/ s]

Riser height [m]

6.0

110 100 90 80 70 60

Time [s]

Figure 16:

Local void fraction, temperature and mass flow signal in the DN40 riser with 25 kWel and temperature of 55 °C at CTI

Although void is present at the riser outlet, the velocity of the gaseous phase is not high enough anymore to prevent a CCLF from the heat sink into the riser. The CCLF of subcooled liquid into the riser is followed by heavy water hammer at the location where the subcooled liquid is mixed with steam bubbles. At approximately 65 s the mass flow in the test circuit decreases back to 0.07 kg/ s. Hence the heat transfer on the inside of the condensation tubes is reduced and the temperature and pressure in the steam chamber increase again. After 55 s a cold water column stands above the warm fluid from the condensation tubes. Until 70 s this cold water column is continuously pushed out of the riser by the warm fluid. After that, this process slows down until 90 s. During this process, with a constant mass flow rate of 0.07 kg/ s, the retention time of the fluid in the condensation tubes increases. If saturation conditions are reached at the riser outlet, the fluid evaporates again at approximately 100 s. Due to this, mass flow rate increases in the same manner, as described before. The period of the flashing lasts about 50 s, while void can be measured at the riser outlet for approximately 5 s. During the presented instability the frequency and the mass flow amplitude are constant, as shown in Figure 16 and Figure 17. After a transient the next steady state working point is established as shown in Figure 18. During experiments droplet condensation has been observed, through the windows at the steam chamber, on the condensation tubes.

Mass flow [kg/ s]

0

150

200

150

200

0.4 0.2 0.0 downcomer

riser

condensation tubes Test circuit

200

∆p [mbar]

Time [s] 100

50

0

Temperature [°C]

120

Temperature [°C]

110

Rel. pressure [mbar]

-200

250

riser,in

riser,out

CTI

100 80 60 TSC1

TSC2

TSC3

TSC4 Steam chamber

108 106

200 150 100

L9

50

centre eccentric

0 100

L8

50 0 100

L7

0 100

L6

50 0 100

Local void fraction in the riser

Local void fraction (center) [%]

50

L5

50 0 100

L4

50 0 100

L3

50 0 100

L2

50 0 100

L1

50 0 0

Figure 17:

50

100 Time [s]

Overview on measuring signal during flashing, at 25 kWel evaporator input power and temperature of 60 °C at CTI

An important parameter to characterise a natural circulation system is the period or frequency of the mass flow peaks. When CCLF occurs only in the riser and at the upper end of the condensation tubes, an equation for the flashing period τflashing can be derived.

e*+1?f()g

B(?>B ∗ Ah+,(∗ 4-B.C
?ijh

eq. 13

The flashing period is determined by the sSPF velocity in the riser and the drop height hdrop of the cold water with the following equation to approximately 46 s. Referring to Figure 16 and Figure 17 this is in good agreement with the presented experimental data. 4.3.2 Medium subcooling with high amplitude (TCTI = 77 °C) In Figure 18 the steady state working point at 77 °C is presented with a time axis, increased by the factor two up to 400 s, to show both, the increased flashing period and the constant mass flow amplitude. At 0 s the riser is filled with water of approximately 102 °C, from the inlet at 0.0 m up to approximately 4.0 m. Upwards to the riser outlet a cold water column is present above the warm fluid. At this moment saturation conditions are not fulfilled in the riser. At approximately 53 s fluid with 102 °C reaches saturation conditions at the riser outlet. Due to this evaporation starts from the riser outlet down to the riser inlet at approximately 58 s. For almost 25 s only void is present in the riser centre. The mass flow increases up to 0.43 kg/ s. Time [s] 200

100

300

400

[%] Local void fraction

Riser height [m]

0 6.0 4.5 3.0 1.5

100 80 60 40 20 0

0.0 Mass flow

4.5

0.4

104

106

1.5 108

0

Figure 18:

0.3

104

106

3.0

0.0

[°C]

0.5 102

102

200 Time [s]

0.2 0.1

110

102

100

108

0.0 300

-0.1 400

Mass flow [kg/ s]

Riser height [m]

6.0

110 100 90 80 70

Local void fraction, temperature and mass flow signal in the DN40 riser with 25 kWel and temperature of 77 °C at CTI

If saturation conditions are not fulfilled at the riser outlet at approximately 82 s, no further void is produced and CCLF occurs. As a consequence cold fluid can enter the riser from the heat sink above, as shown in Figure 18, at approximately 82 s. As a consequence of the CCLF high temperature gradients occur in the riser. Due to this subcooled liquid falls into the riser. This causes a stagnation of the test circuit mass flow for approximately 5 s and water hammer. After that mass flow rate increases up to 0.07 kg/ s, which equals sSPF.

With respect to the upper diagram of Figure 18, it has to be examined, if the flashing front is extended into the condensation tubes. The period between the two mass flow peaks can be calculated with respect to the length of the cold water column in the riser and in the condensation tubes, as well as the sSPF of approximately 0.05 kg/ s. If the length of the condensation tubes is not considered, for the experiment presented in Figure 18, the period can be calculated to approximately 144 s. Considering that the cold fluid also enters the upper end of the condensation tubes and the slower fluid velocity in the condensation tubes, in comparison to the velocity in the riser, the period can be calculated to approximately 230 s. This period equals nearly the period between the two mass flow peaks shown in Figure 18. The presented calculation of the flashing period proves that CCLF even reaches into the upper end of the condensation tubes. 4.3.3 Low subcooling with increasing two-phase flow period (TCTI = 85 °C) A further increase of temperature at CTI, up to 85 °C, leads to a longer duration of the flashing (ref. to Figure 19). As described for the steady state working points before, if local saturation conditions are fulfilled at the riser outlet, the fluid evaporates at this point immediately. This is shown at approximately 39 s in Figure 19.

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Figure 19:

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Local void fraction, temperature and mass flow signal in the DN40 riser with 25 kWel and temperature of 85 °C at CTI

At approximately 101 s flashing stops, CCLF occur, and fluid from the upper heat sink falls down into the riser. The increase of the inlet temperature at CTI leads to wider mass flow peak. A reason for that is the lower subcooling of the fluid at CTI and the increased temperature of the fluid in the heat sink. Therefore more transferred latent heat is used for evaporation of the fluid in the condensation tubes and not for heating up to saturation conditions. After saturation conditions are not fulfilled in the riser anymore, mass flow decreases down to sSPF. No flow stagnations occur because the fluid from the heat sink does not fall as deep into the riser as shown for the steady state working point before. This can be explained as follows: The minor subcooling in comparison to the steady state working

point, described before, provides the opportunity that the CCLF is heated up and flows upwards with the two-phase flow mixture. Due to the high velocity of the gaseous phase in the riser centre, CCLF flows along the tube wall as a kind of annular flow. If the two-phase flow cannot heat up the CCLF anymore (e. g. decrease of steam chamber temperature, and gaseous-phase velocity), the film thickness of CCLF increases and blocks the upward moving two-phase flow. Subsequently, no saturation conditions are possible at the riser outlet.

4.4 Natural-circulation at high power input The influence of a high power input at low subcooling is very important during start-up of a natural-circulation system. In this case the fluid velocity is very low, while the pressure of the saturated steam in the steam chamber is high. This can lead to saturation conditions in the condensation tubes. Subsequently, steam bubbles flow the riser upwards, whereby they grow and accumulate with each other. In this case, large slugs can cover almost the whole riser length and inner diameter. As a consequence high void fractions occur from bottom to top of the riser. The high void fraction in the condensation tubes leads to a decrease of the inner heat transfer coefficient. According to this, the velocity of the gaseous phase in the riser is reduced and CCLF occur. Due to this, subcooled fluid can fall down into the riser and can enter the condensation tubes. The downwards falling fluid creates an impulse, which is proportional to the drop height. During geysering, this impulse leads to mass flow rate stagnation and even reversed flow in the whole test circuit. 4.4.1 Geysering The geysering phenomenon was described by Wissler (1956), Chiang et al. (1994) and Baars and Delgado (2007). Lu et al. (2005) presented a literature review about geysering in NPP, rocket engines and other fluid test facilities in comparison with natural geysers. Aritomi et al. (1993) described the geysering in more detail during natural and forced convection in a single channel and in parallel channels. Their experiments were performed with heating elements for BWR-stability analysis. They divided the geysering in four stages: 1) bubbles which grow and coalesce along the riser, 2) mass flow acceleration due to decrease of the hydrostatic head, 3) rapid condensation of the slug bubbles in the upper plenum, which is filled with subcooled water and 4) reversed flow in the test circuit followed by flow stagnation. In parallel channels this phenomenon alternately occurred. They observed no difference between geysering in a forced or natural-circulation loop. Lisowski et al. (2014) observed periodic flow excursions at their test-facility, which lead to high mass flow amplitude and reversed flow. In comparison to the GENEVA test-facility electrical heating rods were used inside three parallel riser sections. Moreover only one dualtip optical probe was installed in the chimney between outlet plenum and storage tank. For passive heat-removal this phenomenon has to be taken seriously into account, to maintain the structural integrity of the heat removal system. 4.4.2 Detailed overview of system parameters To examine the behaviour of geysering in connection to the CCLF in more detail, experimental results for a steady state working point have been evaluated. The presented experiment was performed with a DN40 riser. In Figure 20 a compilation of important measuring signals from the test circuit and the steam chamber (heat source circuit) are shown, similar to Figure 17. Furthermore all nine measuring levels for the probes in the riser centre are shown, to underline the high spatial and temporal resolution of the presented VDMS. The compilation allows the investigation of the source of geysering and the time dependency between the test circuit and heat source circuit. Moreover a comparison with other systems or code validation is possible, based on accurate time dependent variables.

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Figure 20:

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Overview on measuring signal during geysering, at 100 kWel evaporator input power and temperature of 55 °C at CTI

Because of the mass flow rate, the differential pressure signals along the downcomer and the riser decrease proportionally, while the differential pressure along the condensation tubes

increases. The increase of the differential pressure along the condensation tubes indicates that saturation conditions are reached in the condensation tubes and fluid evaporates on the inside. The mass flow signal is quite time delayed in comparison with the pressure transducer signals, which can be explained by the position of the Coriolis mass flow meter, which is installed in the horizontal part of the downcomer. High temperature peaks occur at CTI after a mass flow peak. The CCLF seems to be responsible for this. The impulse of the CCLF is also quite proportional to the average void fraction in the riser. Referring to the local void fraction diagram of level L1, high local void fraction is almost the whole time present at the centre and near the wall of the riser inlet. All probes in the centre show in-phase behaviour. The partial reduction of local void fraction along the riser, indicated at approximately 118 s on level L1, and time delayed on level L9, can be a result of the mixture between CCLF at the outlet and the upwards flowing fluid. The probes in the riser centre and near the wall on levels L2 to L9 are in good agreement to each other. These probes in the centre and near the riser wall indicate that the gaseous-phase covers almost the whole riser inner diameter. The average saturation conditions in the steam chamber seem to be independent from the mass flow peaks during the geysering (ref. to Figure 20). This means, that despite the aperiodic mass flow behaviour, the heat is transferred continuously from the steam chamber into the heat sink via the adiabatic riser. Time [s] 40

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Local void fraction, temperature and mass flow signal in the DN40 riser with 100 kWel evaporator input power and temperature of 55 °C at CTI

Figure 21 shows an extract of Figure 20 (from 120s to 200s), to underline the high temperature gradients due to geysering and CCLF at high subcoolings. In the upper diagram of Figure 21 local void fraction is shown from the probes in the riser centre. In the lowermost diagram the temperature distribution along the riser with isothermals in steps of 2 K, and the mass flow rate in the test circuit, are presented. During geysering void is detected at first at the riser inlet and time delayed at the outlet. The velocity of the gaseous-phase is high enough to blow the subcooled water out of the riser. When the velocity of the gaseous phase

decreases, CCLF appear. Due to the high impulse of the downwards falling fluid, reversed flow and water hammer occur in the test circuit. The reversed flow and high input power lead to high fluid temperatures at the riser inlet of approximately 114 °C.

4.5 Comparison between flashing and geysering The main difference between flashing and geysering is the point of origin. In general, flashing arise due to the reduction of the hydrostatic head at the end of the adiabatic riser section. In contrast to this, geysering starts in the heating zone (here in the condensation tubes) when local saturation conditions are fulfilled there. The analyses of the presented experimental data in the previous sections showed that a low fluid velocity in combination with a high heat input promote geysering. Despite the point of origin of the two types of instabilities, flashing and geysering, further differences are pointed out and analysed in this section. The analysis of the PSD and the phase portrait (mass flow in the test circuit as a function of differential pressure along the downcomer) is presented in Figure 22. For flashing, data from the previous section “natural circulation at low power input” are shown for three presented steady state working points. For geysering, the data from the previous section “natural circulation at high power input” are presented. The inverse flashing period is in good agreement with the PSD of the three steady state working points. For geysering a frequency of approximately 0.04 Hz seems to be significant.

Figure 22:

Comparison of PSD and phase portrait between geysering and flashing

All phase portraits in Figure 22 are plotted clockwise. The phase portrait in Figure 22e shows an interesting behaviour for the differential pressure. The relative constant peaks, in positive direction, are a result of water hammer subsequent to the CCLF. Figure 22f and Figure 22g show a similar behaviour as discussed for Figure 10. In comparison to that, the phase portrait for geysering looks quite chaotic. Reversed flow in the test circuit and short periods of two-phase flow at maximum flow rates are characteristic for geysering. But the differential pressure amplitude is not higher than during flashing. Despite some outliers and the chaotic behaviour, the shape in the phase space is comparable to those during flashing. Based on these evaluations a characteristic phase portrait for the presented natural-circulation system seems to be existing for the DN40 riser variant. This has to be clarified by taking into account CCLF. The previous presented analysis of the experimental data show that at a certain stage, flashing leads to CCLF in the DN40 riser and even to reversed flow in the condensation tubes. This causes flow stagnation and an increase of the temperature at CTI. During geysering, the high impulse of CCLF seems to be the reason for the reversed flow in the whole test circuit. To examine this from another point of view, two phase portraits are presented for natural-circulation, one with low heat input and the other with high heat input. For this analysis a coefficient was derived to describe the influence of CCLF to the dynamic of the system. This CTI coefficient is the quotient between the temperature at CTI and the outlet temperature of the riser.The mass flow rate in the test circuit is shown as a function of the CTI coefficient. For both heat input cases, measuring data of three steady state working points are shown in Figure 23 and Figure 24. The phase portrait for flashing, shown in Figure 23, can be divided into four main parts, starting in the right corner at the bottom: 1) Temperature increase at the riser outlet, 2) Acceleration of the mass flow rate, because saturation conditions are reached at the riser outlet, 3) Velocity of the gaseous phase decreases, than CCLF occur and temperature at the riser outlet decreases and 4) Mass flow rate decreases. During stage 1), the constant slopes of the three presented steady state working points indicate an equal heat transfer. With decreasing subcooling, the CTI coefficient of stage 2) shifts to higher values. Simultaneously, the mass flow rate amplitude increases. During stage 3) the mass flow rate is almost constant and the CTI coefficient increase up to approximately one. This indicates that CCLF occur and the temperature at the riser outlet decreases. During stage 4) the mass flow rate decreases, while the CTI coefficient is almost constant. With decreasing subcooling, a region occur, where the impulse of the CCLF is high enough to push heated up fluid from the condensation tubes into the downcomer. In this case the CTI coefficient can reaches values above one (ref. to Figure 23 TCTI = 68 °C). This indicates a very critical region where CCLF has an important influence on the stability of the presented natural-circulation system.

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Figure 23:

Phase portrait of three steady state working points at different temperatures at CTI for low input power

In comparison to this, Figure 24 shows the phase portrait for natural-circulation with high heat input for three different steady state working points. The stages 1) to 4) are not as clear as presented in Figure 23, which indicates chaotic system behaviour. Only stage 2) and stage 3) are quite good to identify, which present mass flow acceleration and onset of CCLF at the riser outlet. During high subcooling, the reversed flow in the condensation tubes and the high heat input lead to a high CTI coefficient. During stage 2), CTI coefficient shifts to higher values, while the subcooling decreases. At low subcooling sTPF is possible and can be characterised by a constant mass flow rate as well as an almost constant CTI coefficient at approximately 0.85. This point is similar to the transition point between stage 2) and stage 3) in Figure 23 for TCTI = 88 °C.

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Figure 24:

Phase portrait of three steady state working points, at different temperatures at CTI and high input power

5 Conclusion To analyse the natural-circulation behaviour of single components, used in passive residual heat-removal systems, the GENEVA test-facility has been built and commissioned at TU Dresden (WKET) in 2013. The unique modular concept of GENEVA allows the analysis of several flow parameters in mid-scale, at low system pressures and with heat input by condensation of saturated steam. As one of the main unique design features, the convective heat transfer has been pointed out in comparison to former test facilities. Moreover, a transient experiment was presented to show similarities and differences to former test facilities, where natural-circulation was investigated. As described in the literature all three natural-circulation flow states occurred. An experimental method to determine the maximum natural-circulation mass flow rate of the system was presented and is in good agreement to the calculated values under consideration of friction losses. The analyses of the presented experiments are in agreement with the literature. The presented analysis of the experimental data showed that an increase of the riser inner diameter leads to a higher natural-circulation mass flow rate. Experimental data of two different riser variants were analysed and compared with each other to underline the influence of the riser inner diameter on the convective heat transfer, and therefore on the stability of the natural-circulation system. In these analyses different dimensionless numbers were used to describe the convective heat transfer on the inside and outside of the condensation tubes. Main instabilities like flashing and geysering have been presented and analysed in detail. A self-developed VFMS with high time and spatial resolution was installed in the riser. Also the temperature distribution along the adiabatic vertical riser and the mass flow signal were analysed. The performed experiments showed that flashing is the dominant instability up to input power of 75 kW el. During flashing a constant period between the instabilities, in dependency of the subcooling, was investigated. This period was calculated by taking into account the riser length and sSPF velocity. The increased riser inner diameter promotes the accumulation of huge bubbles or slugs. When these bubbles condensate rapidly at the outlet of the riser (in the subcooled water of the heat sink) counter current liquid flow can occur in the riser. The resulting cold water column blocks the upwards flow and leads to an increased retention time of the fluid in the condensation tubes. Due to this no constant mass flow is established in the test circuit. This phenomenon promotes at higher input power geysering in the riser. During geysering reversed flow appears in the test circuit and results in an aperiodic mass flow signal. The detailed analyses showed that geysering has a chaotic behaviour by taking into account different fluid parameters in the test circuit. In the presented system counter current liquid flow occurs in the riser for both instabilities, flashing and geysering. This plays an important role for the appearing reversed flow from the heat sink into the riser. For the first analysis a coefficient as function of the fluid temperature in front of the heating zone and the fluid temperature at the riser outlet was derived, to describe the influence of counter current liquid flow on the instabilities. A large database with steady state working points was generated, with measuring data for e. g. axial local void fraction and temperature distribution along the adiabatic riser. This database is very important for the validation of codes and models. Furthermore with its help enhanced models can be implemented into appropriate codes. All tests conducted at GENEVA test-facility have provided a sustainable contribution to the understanding of natural-circulation systems for decay heat-removal, especially due to its unique modular design. Main instabilities were substantiated and proven characterised. Naturally, the evaluated generic data can and should, in future, be used to design passive heat-removal systems. Further investigations should focus on constructive improvements at the riser outlet. These have to be tested further to avoid on the one hand the rapid condensation of large bubbles and on the other hand the reversed flow phenomenon. In addition more experiments on heat

transfer coefficient of the condensation tubes are necessary to decrease the region of uTPF. This is important to stabilize the two-phase natural-circulation system.

Acknowledgment This work is financially supported by E.ON Kernkraft GmbH. Co-operation partners are AREVA GmbH and Helmholtz-Zentrum Dresden-Rossendorf (HZDR).

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Generic investigations of two-phase flow instabilities Stationary working points have been established for constant inlet temperatures and input power up to 100 kWel Two different riser diameters were tested as well as inlet restrictions at the downcomer Flashing is the dominant instability in the adiabatic riser Geysering occurred at high input powers and low fluid velocity with reversed flow