3. CIRCE FACILITY
This chapter is focused on the experimental activities performed on the CIRCE facility [1] to characterize the gas enhanced circulation technique in a HLM pool system.
Following the results obtained during the tests performed in the frame of the former TECLA project [2], new tests have been carried out in the CIRCE facility to complete the studies on the gas enhanced circulation technique.
The gas lift technique has a very good reliability and a medium efficiency (more than 50%) but a low pressure head. The main advantages of this kind of pumping system are the simplicity and the absence of problems concerning erosion and corrosion.
For application in a pool-type nuclear power plant, the gas lift is not applicable when high pressure heads are required and in this case the most viable solution seems to be the installation of a mechanical pump, also owing to the good reliability achieved by these devices. However, for HLM pool facilities, conceived for thermal-hydraulic and component tests, the gas lift technique seems to be a suitable pumping system.
For the application in CIRCE facility, where low pressure heads are required (less than 1 bar), the gas lift technique was preferred, because of the greater reliability and the very low mechanical-structural effort required with respect to mechanical pumping systems.
3-1. DESCRIPTION OF THE CIRCE FACILITY.
The CIRCE facility basically consists of a main cylindrical vessel (named S100) with an inventory of LBE of about 70 tons. The facility includes a LBE storage tank (S200), a small LBE transfer tank (S300) and the data acquisition system. In figure 1, an isometric view of the facility is shown.
A full description of the facility is reported in [2]. In the following, only a short description of the test section used during the experimental campaign is given.
Figure 1. CIRCE isometric view.
3-1.1 TESTSECTION.
The test section used to perform the experimental campaign is shown in figure 2. Its main components are described hereafter.
• Riser: a pipe, geometrically equal to one of the 22 pipes that forms the reference XADS plant riser [3] (diameter=202.7 mm, height=3854 mm). Since this element is scaled 1:1, the flow mapping achievable by tests will be easily applicable to the full scale reference plant. Inside the riser, the gas injection pipe is housed, ending with the injector (see figure 4).
Figure 2. Sketch of the test section.
• Dead Volume: delimited by the test section support ferrule, it maintains the real height of the full scale dead volume, whereas the cross section is scaled 1:22. At present the dead volume does not play an active role in the test runs; so, it is kept empty. However, future tests to study the coupled behaviour of the riser-dead volume system have been already planned.
• Fitting volume: It consists of a cylindrical volume delimited by two circular plates (diameter=982 mm), 225 mm spaced from each other. It joins the feeding conduit with the riser. This element is designed and scaled to represent the upper plenum of the core, connecting the core outlet to the riser inlet.
• Feeding conduit: The Venturi-Nozzle flow meter is installed on the conduit, together with a downstream drilled disk. The hole manufactured in this disk is sized to provide a head loss suitable to simulate the corresponding one in the full scale core of the XADS concept [3].
• Downcomer: is the outer side of the test section, included in the descending path of the LBE. It is scaled to represent 1/17 of the full scale downcomer; so, it is characterized by a cross section of 0.7 m2.
The test section is equipped by all the instrumentation suitable to perform the measurement required by specification during the test, as thermocouples (TEXX), differential pressure (DPTXX) and level meters (LTXX), as indicated in figure 6 and figure 7
3-1.2 GASINJECTIONSYSTEM
The gas injection system is a subsystem of the Argon Recirculation System (ARS) (see figure 3); it consists basically of an injection line connected upstream to the compressor set of the ARS.
The connection is realized by the flow measurement block, consisting of three parallel lines equipped by gas flow meters having different measurement ranges. Downstream this block, the main injection line reaches the upper flange of the main vessel; then, it runs inside the riser. The line ends at the bottom of the riser with a “U” bend on which an injection nozzle is installed (see figure 4).
The different measurement ranges of the flow meters are as follows: the first flow meter (FT208A) works in the range [0.056÷0.56 Nl/s], the second (FT208B) in the range [0.35÷3.5 Nl/s], while the third works in the range [2.2÷22 Nl/s]. Once the gas flow rate to be injected is chosen, it is possible to switch on the appropriate flow meter, opening the associated isolation valve and excluding the other ones (see figure 5).
Figure 4. Drawing of the injection system installed at bottom of the riser. The output signal from the flow meter is acquired and processed by the Data Acquisition System (DAS). This processing represents a correction taking into account the difference between the calibration conditions and the operating conditions in terms of pressure and temperature.
The pressure and temperature compensation with respect to the calibration values is made by the following equation:
eff m T P
where Q indicates the corrected volumetric flow rate in Nl/s, while eff Q indicates m the measured volumetric flow rate in l/s; C and T C are the temperature and P pressure coefficients, defined by the constructor as follows:
t T e T C T = (2) 1 e P t P C P = + (3)
in which T is the calibration temperature, equal to 423 K, while t T is the operating e temperature, measured by the thermocouple TE209 upstream the flow meter; likewise, P is the calibration differential pressure, equal to 4 bar(g), while T P is the e operating pressure measured by the pressure transducer PT207 installed upstream the flow meter (see figure 3).
FT208C FT208A FT208B Bypass Valve V70 V65 V67 V66 From Compressors To Riser
Figure 5. Sketch of the flow meters installed on the main gas injection lines of ARS.
3-1.3 INSTALLEDINSTRUMENTATION
In the CIRCE plant, bubble tubes [2] are installed to transfer pressure signals from the LBE to differential pressure cells operating with gas at room temperature. In particular, the Level Measurement System (LMS) and the Differential Pressure
Measurement System (DPMS) make use of the above instruments to perform their tasks.
A sketch of the test section instrumentation as reported in the control panel of the DAS is shown in figure 6.
Before start with the last series of gas-injection enhanced circulation tests, the DPMS was improved, by installing several pressure transducers intended to measure the pressure drops along the riser. In particular, five DPT cells (from DPT05 to DPT09, as depicted in the figs. 6 7), are used to estimate the two phase pressure losses as well as the void fraction along the riser during the experiment, while the DPT03 is useful to measure the single phase pressure drop through the drilled disk (hole size: Φ =127) [5,6], installed downstream the flow meter, to simulate the pressure loss through the core in XADS [7,8,9] reactor. So, before running the experimental campaign, all the installed instrumentation has been checked and calibrated in order to obtain the gravimetric head between each couple of bubble tubes under static condition (stagnant pool).
Figure 6. Control panel of the DAS reporting a sketch of the test section with the installed instrumentation for the gas injection circulation tests
During the calibration, the distance between the bubble tubes connected to the differential pressure transducers was also checked, converting the differential pressure measured to mm of LBE. In figure 8 and 9, the results obtained for the instrumentation in the riser and for the bubble tube installed across the drilled disk respectively are shown; in the case of figure 8, these distances are not the same between each couple of bubble tubes; indeed, as matter of fact, the measured values are different from the value of 651 mm declared by the manufacturer. For this reason, the values obtained by measurement as reported in table 1 have been considered as the distances between the bubble tubes.
DPT008 DPT009 TE 09 DPT006 DPT005 DPT007 R Z 10 TE 07 TE DPT002 06 TE DPT003 11 TE LT
Figure 7. Locations of the measurement point for the thermocouples (TE), differential pressure meters (DPT), level meters (LT), pressure meters (PT) and
610 620 630 640 650 660 670 680 0 10000 20000 30000 Time [s] ∆ H [m m ] DPT09 DPT08 DPT07 DPT06 DPT05
Figure 8. Relative distance of the bubble tubes which supply the differential pressure transducers installed on the riser, evaluated at no flow conditions in the
test section.
Concerning the bubble tubes which supply the DPT03 transducer, utilized to measure pressure loss through the drilled disk (see figure 6) with calibrated hole [5], also in this case the measured distance is different from the value of 355 mm declared by the manufacturer. Then, for the evaluation of the gravimetric head through the drilled disk, the measured value of 364 mm is adopted.
Concerning the flow meter, it was already tested during a former dedicated experimental campaign, when its characteristic curve had been obtained. The LBE flow rate is then evaluated using the following formulation [2]:
0.5
2.233 8.793
Mɺ = ∆p + 40≤Mɺ ≤350 kg/s (4)
where Mɺ is the LBE flow rate, while p∆ is the differential pressure measured by DPT002 (expressed in Pa), decreased by the gravitational contribution.
However, as for the other instruments, the DPT cell (DPT002) related to the flow meter has been checked and calibrated in order to obtain the gravimetric head between the two taps installed on the Venturi-Nozzle.
0.340 0.350 0.360 0.370 0.380 0.390 0 5000 10000 15000 20000 25000 30000 Time [s]
∆∆∆∆
H [m ] 320°C 250°C 220°C 200°CFigure 9. Distance between the bubble tubes which supply the DPT03 differential pressure transducers, evaluated at no flow conditions.
DPT005 DPT006 DPT007 DPT008 DPT009
H
∆ [mm] 633 626 655 671 649
Table 1. Distances between the bubble tubes which supply the differential pressure transducers installed on the riser.
3-2. PERFORMED TESTS
The tests were performed under isothermal conditions, at different LBE temperatures (200°C, 220°C, 250°C, 320°C); for each test temperature, several steady states of circulation have been obtained injecting different gas flow rates, each state being characterized by a different LBE flow rate. The different gas injection flow rates were selected in the range from 0.5 to 7.0 Nl/s, in order to investigate a wide range of entrained liquid metal flow rate, up to 250 kg/s.
When the isothermal conditions at the selected temperature are reached, a gas flow rate is injected at the bottom of the riser by the appropriate 6 mm ID nozzle (see figure 4), with a gas pressure of about 5 bar, needed to overcome the backpressure of the LBE column. Then, after a quick transient, a steady flow condition for the liquid metal is obtained. During the test, all the performed measurements are stored on a HD by the DAS, in order to make available also data about the transient behaviour of the system.
3-3. EXPERIEMENTAL RESULTS
All the results obtained during the first experimental campaign on gas enhanced circulation formerly performed in CIRCE at 200°C were confirmed by the new tests. Several steady states of liquid metal circulation were obtained in the pool facility throughout the investigated temperature range. Indeed, as in the previous tests [2], instability phenomena occurred at very low gas flow rate.
Moreover, the results obtained by the new tests provide important information on the role played by the temperature and the void fraction in the riser on the behaviour of the system.
3-3.1 FLOWRATES
Figure 10 shows the trend of both gas injection and the LBE flow rates during the test performed at 320 °C. Steady states and transient phases can be clearly recognized. In particular, the transient phases are very short (less then 30 s), indicating a low fluid-dynamic inertia of the system. At every performed test,
steady-state values of circulation flow were represented by the average of the recorded measures. Similarly, average values of all the useful quantities are calculated to characterize the flow.
0 50 100 150 200 250 29000 34000 39000 44000 49000 54000 59000 Time [s] L B E F low Rat e [k g/ s] 0 2 4 6 8 10 G as I n je ct ion F lo w Rat e [ l /s ]
Figure 10. LBE flow rate and argon flow rate vs. time for gas enhanced circulation tests performed at 320°C
The values of temperature, gas flow rate and LBE flow rate, obtained from the all the performed tests are plotted in figure 11. As it can be noted, the flow rate increases with the temperature. In fact, as the temperature increases, the dynamic viscosity of LBE decreases and the viscous pressure losses decrease too. However, this dependence is rather weak and this behaviour suggests that the viscous pressure losses are small in comparison to the singular pressure drops. In fact, increasing the temperature from 200 °C to 320 °C, the dynamic viscosity of LBE undergoes a decrease of about 30%; so, for the same driving force in the systems (that means the same void fraction distribution along the riser) a stronger increment in the flow rate should be expected. However, this increase can be expected to be lower since the singularities play a key role in the system. Therefore, it seems reasonable to consider that the singular pressure drops along the flow path are predominant compared to the viscous ones. So, the LBE flow rate is quite independent from the temperature.
Moreover, the influence of the LBE temperature on the expansion that the argon undergoes when injected in the bottom of the riser is negligible, since the gas can be considered already at the same temperature of the LBE when injected.
In Figure 11 the data obtained during the first experimental campaign on gas enhanced circulation [2] are also shown (mark 200_old); it should be noted as the latest and the previous data are in good agreement, confirming a good repeatability of the results. 0 50 100 150 200 250 0 0.002 0.004 0.006 0.008 0.01 0.012 0.014 Gas Injection Flow Rate [kg/s]
L B E F lo w R a te [ k g /s ] 320 250 220 200 200_old
Figure 11. LBE flow rate as function of the gas injection flow rate at steady-state conditions for all performed tests.
Plotting the same data in bi-logarithmic scale, as in figure 12, it is easy to observe as the two flow rates are related by a “power law”. A best-fit of the data, provides the following equation:
0.3748 1241
LBE Ar
Mɺ = ⋅Mɺ (5)
A similar result was already achieved by the first campaign on gas enhanced circulation. Since the viscous pressure drops constitute a small contribution in comparison to the singular pressure drops along the loop, this kind of relationship between the two flow rates can be approximately justified on theoretical basis [2], [10].
y = 1241x0.3748
10 100 1000
0.0001 0.001 0.01 0.1
Gas Injection Flow Rate [kg/s]
L B E F lo w R a te [ k g /s ]
Figure 12. LBE flow rate as function of the gas injection flow rate and best fit line of experimental data.
3-3.2 FLOWPATTERNS
Introducing the superficial velocities [11] for the two phases, these data can be plotted as shown in figure 13, where the trend of the liquid superficial velocity ( j ) in f the loop is reported as function superficial velocity of the gas phase in the riser ( j ), g at different LBE temperatures.
As expected, j (i.e., the LBE volumetric flow rate) increases with increasing f g
j . In this case the best-fit equation is the following (see figure 14):
0.3787 1.5819 f g j = ⋅j (6) where LBE Ar f g r LBE r Ar M M j j A
ρ
Aρ
= = ⋅ ⋅ ɺ ɺ (7)are the superficial velocities of the two phases and A indicates the riser flow cross r section.
0.1 1 0.001 0.01 0.1 1 jg [m/s] jf [ m /s ] 320 250 220 200
Figure 15. Loop flow as a function of injected argon flow and the LBE average temperature
y = 1.5819x0.3787 0.1 1 0.001 0.01 0.1 1 jg [m/s] jf [ m /s ]
Figure 16. Loop flow as a function of the injected argon flow with the best fit line of experimental data.
Once again, it can be noted as the temperature of the LBE is a relatively low importance parameter to characterize the circulation in the performed tests.
As previously mentioned, instability phenomena occurred at low gas flow rates. In particular, for tests performed with argon injection rates lower than 1 Nl/s, the system shows a pulsed behaviour. However, this phenomenon was already observed during the first tests campaign [2].
Figure 17 shows this pulsed behaviour for the test carried out at 320 °C, with an initial injected flow rate of 0.5 Nl/s; as it can be noted, both flow rates undergoe a periodical and quite regular oscillations, with a frequency of about 30 mHz; the LBE flow rate shows also large magnitude oscillations.
This behaviour is probably related to Argon bubble growth and detachment mechanisms, involving periodic entrance and expulsion of LBE in the injection line at low gas flow rates.
3-3.3 PRESSURELOSSES
Several differential pressure measurements were performed along the riser, as well as along the single-phase part of the test section. In particular, the pressure loss across the drilled disk and the pressure loss along the riser, useful to calculate the void fraction by the “manometric method”, were evaluated.
In Figure 18, the data have been plotted in bi-logarithmic scale; also in this case, by fitting the data a power law is obtained:
2.176 0.2837
frict LBE
p M
∆ = ⋅ ɺ (8)
It should be noted that, for the nominal liquid metal flow rate [3] of 200 kg/s, the pressure drop measured trough the drilled disk is about of 30000 Pa.
As known, the pressure loss is usually expressed as: 2 2 2 1 2 LBE frict LBE LBE
LBE eff M p k v k A
ρ
ρ
∆ = ⋅ = ɺ (9)where k is the loss coefficient associated to the drilled disk and is in general assumed not dependent on the flow rate, since the flow is fully developed.
Using the previous expression, it is possible to obtain the value of k, given the flow rate and the pressure loss.
0 40 80 120 160 200 67600 67650 67700 67750 67800 Time [s] L B E F lo w R a te [k g /s ] 0 0.2 0.4 0.6 0.8 1 G a s In je ct io n F lo w R a te [ l/ s]
LBE Flow Rate Gas Flow Rate
Figure 17. Pulsed behaviour of the system at low gas injection rate (T= 320 °C)
y = 0.2837x2.1764
1000 10000 100000
10 100 1000
LBE Flow Rate [kg/s]
∆∆∆∆
P [ P a ] 320°C 250°C 220°C 200°CFigure 18. Pressure drop through the drilled disk as function of the LBE flow rate with the best fit line of experimental data (steady-state conditions).
y = 6.1349x0.1762 0 5 10 15 20 25 0 50 100 150 200 250
LBE Flow Rate [kg/s]
k
Figure 19. Singular pressure drop coefficient for the drilled disk as function of the LBE flow rate
Figure 19 shows the results obtained from the application of the equation (9) on the available experimental data; as it can be noted k is not exactly constant, but is a weak function of the mass flow rate through the drilled disk, according to the following best fit curve:
0.176 6.135 LBE
3-4. DATA ANALYSIS
The driving force for enhanced circulation derives from the void buoyancy generated by the injected gas in the riser. In order to quantitatively evaluate the phenomena, by estimating the slip ratio, the flow quality and the overall driving force, detailed information about the void distribution in the riser is very important.
An estimation of void fraction is possible by means of the integral momentum equation along the riser. Under static conditions, the differential pressure measured by a generic transducer DPTXX (see figure 20) connected to two bubble tubes can be expressed as:
static
m LBE m
p p+ p− ρ gH p
∆ = − = = ∆ (11)
i.e., since the system is under no flow conditions, the differential pressure between two bubble tubes is due to the gravimetric contribution only.
In the case of steady-state circulation it is necessary to take in account other the contributions of acceleration and friction:
m grav acc fric
p p+ p− p p p
∆ = − = ∆ + ∆ + ∆ (12)
So, equation (12) (see e.g., [11]) expresses the pressure drop in the channel as the sum of three components, due to acceleration, friction and gravity.
Moreover, in the case under analysis the fluid is a two phase mixture, so: grav m p ρ gH ∆ = (13) 2 2 m m acc m m G G p ρ+ ρ+ − + ∆ = − (14) 2 2 2 m fric lo lo ris l G H p f D φ ρ ∆ = (15) where
• ρm =
(
αρg + −(
1 α ρ)
l)
is the mixture average density in the considered riser part;• 1
H dz H
• g g H
dz
ρ
=∫
ρ
is the average argon density in the considered riser part; • H is the length of the considered riser part;• m l g m ris ris M M M G A A + = = ɺ ɺ ɺ
is the mixture mass flux;
• 2 4 ris ris D
A =π is the riser flow area;
• 2
(
)
2 2 1 1 1 g g l l m m v v G ρ α ρ αρ+ = + − is the area averaged dynamic density; • v is velocity;
• 2
lo
φ is the two phase friction multiplier;
• flo is the friction factor for liquid single phase flow at the same mass flow m
G as the total two phase mass flux;
To evaluate the friction term, an estimation of the two phase multiplier 2 lo φ is required. The method of Lockart-Martinelli [13] was adopted to perform this estimation.
According to [13], 2 lo
φ can be expressed as:
(
)
1.8 2 2 1 1 1 lo C x X X φ = + + − (16)where C is a constant equal to 20, since both the phases are turbulent. 2
X is the Martinelli parameter defined as:
(
)
(
)
2 friction pressure gradient in the liquid phase / friction pressure gradient in the gas phase /
l fric g fric dp dz X dp dz = = (17) then: 0.2 1.8 2 g 1 l l g x X x
ρ
µ
ρ
µ
− = (18)g g l M x M M = + ɺ ɺ ɺ (19)
Note as, for the system under analysis, the liquid mass flow rate is four orders magnitude higher than the gas one. Hence we can assume:
g l M x M = ɺ ɺ (20)
For all the tests performed, equation (16) gives values very close to unity. It means that, by equation (14), the two-phase flow in the riser shows a behaviour very close to a single phase flow from the friction pressure losses point of view.
g Riser DPTXX ∆pm = p+−p− H DAS Bubble tube - + LBE flow
Figure 20. Sketch of the DPTXX
installed on the riser for the void fraction evaluation.
The friction factor for single-phase flow can be calculated by the Churchill correlation [11], whose use is allowed both for laminar and turbulent flows:
(
)
1/12 12 3/ 2 8 1 8 Relo f A B = + + (21) where:(
)
(
)
16 0.9 1 2.457 ln 7 / Relo 0.27 / ris A Dε
= + 16 37530 Relo B= Re m ris lo l G D µ =
In the tables 2-5 the estimated values of the friction pressure drops are listed, for the all the obtained steady-states; as already discussed, for each test five values for the ∆pfric were calculated, one for each part in which the riser has been subdivided (see figure 6).
average T = 320 °C LBE Mɺ [kg/s] x Gm [kg/m2s] fric p ∆ 5 [Pa] fric p ∆ 6 [Pa] fric p ∆ 7 [Pa] fric p ∆ 8 [Pa] fric p ∆ 9 [Pa] 101.1 8.74E-06 3290 37.0 36.4 38.0 38.8 37.5 123.4 1.57E-05 4015 55.3 54.2 56.5 57.7 55.6 118.6 1.53E-05 3857 51.1 50.2 52.3 53.3 51.4 161.4 2.23E-05 5252 94.3 92.3 96.0 97.8 94.2 161.7 2.33E-05 5262 94.9 92.8 96.6 98.3 94.7 182.9 2.86E-05 5952 122.1 119.2 123.9 126.0 121.3 178.5 2.95E-05 5809 116.7 114.0 118.4 120.4 115.9 205.8 3.65E-05 6696 156.4 152.4 158.1 160.6 154.4 215.1 4.28E-05 7000 173.2 168.4 174.4 177.0 170.1 209.5 3.97E-05 6817 163.3 159.0 164.7 167.3 160.8 224.2 4.69E-05 7293 189.4 184.0 190.4 193.1 185.4 227.9 4.96E-05 7415 196.9 191.1 197.6 200.4 192.4 87.5 1.04E-05 2847 28.3 27.8 29.0 29.6 28.6 115.5 1.49E-05 3758 48.6 47.7 49.7 50.7 48.9 118.8 1.53E-05 3867 51.3 50.4 52.5 53.6 51.7 161.4 2.24E-05 5251 94.3 92.3 96.0 97.8 94.2 158.6 2.26E-05 5162 91.2 89.3 92.9 94.6 91.2 181.9 2.94E-05 5919 121.1 118.2 122.8 124.9 120.2 181.4 2.98E-05 5903 120.5 117.7 122.2 124.3 119.6 207.2 3.64E-05 6742 158.6 154.5 160.2 162.8 156.5 210.0 3.93E-05 6834 163.9 159.6 165.4 168.0 161.5 231.6 5.20E-05 7537 204.5 198.3 205.0 207.7 199.3
Table 2. Summary of the friction pressure drop in the riser for tests performed at 320°C
average T = 250 °C LBE Mɺ [kg/s] x Gm [kg/m2s] fric p ∆ 5 [Pa] fric p ∆ 6 [Pa] fric p ∆ 7 [Pa] fric p ∆ 8 [Pa] fric p ∆ 9 [Pa] 87.6 1.18E-05 2850 28.7 28.2 29.4 30.0 29.0 111.0 1.66E-05 3610 45.4 44.6 46.4 47.4 45.7 114.7 1.64E-05 3732 48.3 47.4 49.5 50.4 48.7 152.4 2.38E-05 4958 84.7 82.9 86.3 87.9 84.7 151.5 2.34E-05 4931 83.7 82.0 85.3 86.9 83.7 176.9 2.95E-05 5755 114.3 111.7 116.1 118.2 113.8 171.0 3.06E-05 5564 107.4 105.0 109.0 111.0 106.8 198.5 3.59E-05 6459 144.9 141.3 146.7 149.1 143.5 207.1 4.14E-05 6739 159.2 155.1 160.8 163.3 157.0 221.7 5.00E-05 7214 185.1 179.9 186.2 188.9 181.5
average T = 220 °C LBE Mɺ [kg/s] x Gm [kg/m2s] fric p ∆ 5 [Pa] fric p ∆ 6 [Pa] fric p ∆ 7 [Pa] fric p ∆ 8 [Pa] fric p ∆ 9 [Pa] 112.5 1.67E-05 3660 46.7 45.8 47.8 48.8 47.0 151.8 2.40E-05 4938 84.1 82.4 85.7 87.3 84.1 153.6 2.38E-05 4997 85.9 84.2 87.6 89.3 86.0 172.3 3.01E-05 5606 108.8 106.3 110.5 112.5 108.3 175.7 3.15E-05 5718 113.3 110.7 115.0 117.1 112.7 195.9 3.66E-05 6375 141.3 137.9 143.1 145.5 140.0 193.1 3.61E-05 6282 137.2 133.9 139.0 141.3 136.0 206.0 4.27E-05 6704 157.8 153.7 159.3 161.9 155.6 201.8 4.17E-05 6565 151.2 147.3 152.8 155.2 149.3 222.2 5.47E-05 7231 187.3 181.9 188.2 190.8 183.2
Table 4. Summary of the friction pressure drop in the riser for tests performed at 220°C
average T = 200 °C LBE Mɺ [kg/s] x Gm [kg/m2s] fric p ∆ 5 [Pa] fric p ∆ 6 [Pa] fric p ∆ 7 [Pa] fric p ∆ 8 [Pa] fric p ∆ 9 [Pa] 79.8 1.15E-05 2597 24.2 23.8 24.8 25.4 24.5 110.6 1.72E-05 3600 45.4 44.5 46.4 47.4 45.7 148.6 2.42E-05 4836 80.9 79.2 82.4 84.0 80.9 173.1 3.22E-05 5631 110.2 107.7 111.9 113.8 109.6 193.2 3.79E-05 6285 137.7 134.4 139.5 141.8 136.5 201.9 4.32E-05 6568 151.6 147.7 153.1 155.6 149.6 224.3 4.57E-05 7300 186.7 181.8 188.4 191.4 184.0 221.9 5.67E-05 7220 187.1 181.7 187.9 190.5 183.0
Table 5. Summary of the friction pressure drop in the riser for tests performed at 220°C
To estimate ∆pacc, a separated flow model was adopted. In the general two-phase case, the velocities and the temperatures of the two two-phases may be different. However, thermodynamic equilibrium conditions between the phases are assumed.
In the case of a one-dimensional system, for any cross section along the riser, it is possible to express the mass flux of the two phases as:
m g g
xG =ρ αv (22)
(
1−x G)
m =ρ
l(
1−α
)
vl (23) Obtaining v and l v from these two equations and substituting them in the g definition of ρm+
(
)
(
)
2 2 1 1 1 m g l x xρ
+ρ α ρ
α
− = + − (24)So, the acceleration pressure drop in each branch of the riser is given by:
(
)
(
)
(
(
)
)
2 2 2 2 2 1 1 1 1 acc m g l g l x x x x p Gρ α ρ
α
ρ α ρ
α
− + − − ∆ = + − + − − (25)However, this equation requires the knowledge of void fraction in the section where the bubble tubes are installed; to overcome this problem, the value of the void fraction in the downstream section (
α
−) is substituted with the average value obtained for the downstream branch of the riser, while the value in the upstream section (α
+) is substituted with the value obtained from the upstream branch.Using this approximation, equation (25) for the nth part of riser can be written as follow:
(
)
(
)
(
(
)
)
2 2 2 2 2 , 1 1 1 1 1 1 acc n m g l g l n n x x x x p Gρ α ρ
α
ρ α ρ
α
− + − − ∆ = + − + − − (26)In this way, however, it is not possible to estimate ∆pacc in the first and the last part of the riser. Anyway, it is possible to assume that ∆pacc in these two branches has the same order of magnitude as in the other branches.
As it can be observed from the tables 2-5, for all the investigated two phase flow conditions, the friction pressure drop in each part of the riser is less than 250 Pa, while the gravimetric term is always two or three orders of magnitude larger.
As first approximation, it is possible to assume that:
(
)
grav fric acc
p p p
∆ >> ∆ + ∆ (27)
This means that in each branch it is:
m grav m
p p+ p− p
ρ
gH∆ = − ≃∆ = (28)
so, the average void fraction can be estimated as: LBE m LBE g
ρ
ρ
α
ρ
ρ
− = − (29) withm m
p gH
ρ
=∆ (30)By the equations (29) and (30) it is then possible to calculate the void fraction in each branch of the riser and to estimate the acceleration pressure drops by equation (26); the results obtained by applying the above described methodology confirm that, as previously noted, the contribution of friction and acceleration pressure drops is negligible. For all performed tests, the estimated summation of the friction and acceleration pressure drop is about 1% of the measured pressure drop; so, the assumption proposed by the equations (27) and (28) is in good agreement with a posteriori estimates.
So, by the equations (29) and (30) is possible to estimate the steady state void fraction distribution along the riser, for the all performed tests. A summary of the data obtained by this analysis at 320 °C is listed in tables 6.
Figure 21 shows the void fraction trends as function of the injected argon flow rate at different values of LBE average temperature; the data are relevant to the branch where the transducer DPT007 is installed. As it can be noted, the temperature does not play an important role, as the trends are similar for all the tests. 0.01 0.1 0.01 0.1 1 jg [m/s] αααα 320°C 250°C 220°C 200°C
Figure 21. Steady State void fraction evaluated by the transducer DPT07 for the all performed tests.
So, if the same data are considered regardless the average temperature, it is possible to obtain a power law relationship between the void fraction and argon superficial velocity (see figure 22).
In the other parts of riser the trend is similar; the power laws which best fit the data are characterised by different coefficients due to pressure changes along the riser. average T = 320 °C LBE Mɺ [kg/s] Ar Mɺ [kg/s]
α
5α
6α
7α
8α
9 101.1 8.84E-04 0.022 0.016 0.011 0.004 0.002 123.4 1.94E-03 0.037 0.027 0.019 0.008 0.004 118.6 1.82E-03 0.032 0.024 0.017 0.008 0.003 161.4 3.60E-03 0.067 0.052 0.035 0.019 0.007 161.7 3.77E-03 0.067 0.053 0.036 0.021 0.008 182.9 5.24E-03 0.087 0.072 0.045 0.029 0.011 178.5 5.26E-03 0.082 0.070 0.044 0.029 0.011 205.8 7.50E-03 0.112 0.095 0.061 0.043 0.017 215.1 9.20E-03 0.122 0.104 0.069 0.049 0.019 209.5 8.32E-03 0.115 0.098 0.066 0.046 0.018 224.2 1.05E-02 0.132 0.114 0.077 0.054 0.019 227.9 1.13E-02 0.137 0.117 0.079 0.057 0.019 87.5 9.11E-04 0.015 0.011 0.008 0.002 0.001 115.5 1.72E-03 0.031 0.023 0.016 0.007 0.003 118.8 1.81E-03 0.034 0.024 0.017 0.007 0.003 161.4 3.62E-03 0.066 0.053 0.034 0.020 0.008 158.6 3.59E-03 0.065 0.051 0.034 0.019 0.008 181.9 5.36E-03 0.086 0.072 0.045 0.029 0.012 181.4 5.40E-03 0.084 0.073 0.046 0.029 0.012 207.2 7.54E-03 0.112 0.096 0.063 0.043 0.018 210.0 8.25E-03 0.116 0.099 0.065 0.046 0.018 231.6 1.21E-02 0.142 0.121 0.082 0.061 0.019Table 6.Summary of the estimated steady state void fraction distribution along the riser
for tests performed to 320°C.
Figure 23 shows the trend of the steady state void fraction distribution along the riser as function of the argon superficial velocity for the tests performed at 200°C; it should be noted as, for the same injected argon flow rate, the void fraction increases along the riser due to the pressure decrease.
Moreover, in all the five branches in which the riser is virtually divided, the trend seems to be the same, even if for the lowest branch, where the transducer DPT09 is installed (see figure 6), a little deviation can be noted for j greater than g 0.1 m/s.
y = 0.4814x0.8607 0.01 0.1 0.01 0.1 1 jg [m/s] αααα
Figure 22. Steady State void fraction evaluated by the transducer DPT07 and the related best fit curve.
0.001 0.01 0.1 1 0.001 0.01 0.1 1 jg [m/s] αααα 5 6 7 8 9
Figure 23. Steady state void fraction distribution along the riser as function of the injected gas flow rate for the tests performed at 200°C.
Making use of the void fraction data it is possible to estimate the slip ratio distribution along the riser; according to [11]:
1 1 g l l g v x S v x
ρ
α
α ρ
− = = − (31)In tables 7, the results obtained by equation (31) for the tests performed at 320°C are listed; for the all tests, the slip ratio increases with the depth in the molten metal alloy with respect to the free surface; moreover in the upper part of the riser (see fig. 7), S seems to be quite independent from the gas flow rate.
average T = 320 °C LBE Mɺ [kg/s] Ar Mɺ [kg/s] Ar Mɺ [Nl/s] S 5 S 6 S 7 S 8 S 9 101.1 8.84E-04 0.50 2.6 2.6 3.1 7.5 10.9 123.4 1.94E-03 1.10 2.7 2.8 3.2 6.5 12.4 118.6 1.82E-03 1.04 3.1 3.1 3.6 6.6 12.6 161.4 3.60E-03 2.05 2.1 2.0 2.5 3.8 8.4 161.7 3.77E-03 2.15 2.2 2.1 2.5 3.7 8.4 182.9 5.24E-03 2.98 2.0 1.8 2.4 3.2 7.1 178.5 5.26E-03 2.99 2.2 2.0 2.5 3.3 7.3 205.8 7.50E-03 4.27 1.9 1.7 2.2 2.7 5.8 215.1 9.20E-03 5.23 2.0 1.8 2.3 2.7 6.3 209.5 8.32E-03 4.73 2.0 1.8 2.2 2.7 6.0 224.2 1.05E-02 5.98 2.0 1.8 2.3 2.7 6.9 227.9 1.13E-02 6.43 2.1 1.9 2.3 2.7 7.1 87.5 9.11E-04 0.52 4.4 4.8 5.3 17.0 31.2 115.5 1.72E-03 0.98 3.1 3.2 3.7 7.2 14.4 118.8 1.81E-03 1.03 2.9 3.1 3.5 7.2 15.6 161.4 3.62E-03 2.06 2.1 2.0 2.6 3.7 7.9 158.6 3.59E-03 2.04 2.2 2.1 2.5 3.9 8.4 181.9 5.36E-03 3.05 2.1 1.9 2.5 3.3 7.0 181.4 5.40E-03 3.07 2.2 1.9 2.5 3.3 6.8 207.2 7.54E-03 4.29 1.9 1.7 2.2 2.7 5.7 210.0 8.25E-03 4.69 2.0 1.8 2.3 2.7 6.1 231.6 1.21E-02 6.85 2.1 1.9 2.3 2.7 7.8 Table 7. Summary of the steady state slip ratio distribution along the riser
for tests performed at 320°C.
These results confirm the choice to use a separated flow model. In fact, so large values of the slip ratio suggest a considerable mechanical non-equilibrium between the phases.
3-5. CIRCULATION PERFORMANCE
As already discussed, in order to estimate the driving force for the enhanced circulation (∆pDF), it is necessary to evaluate the average void fraction in the riser. As it can be noted from the figure 24 and 25, the trend of the steady state void fraction along the riser shows a linear behaviour; the origin of the axial coordinate being placed at the nozzle injection section.
For each set of data it was possible to obtain a linear best fit equation (see figure 24 and 25), which has been adopted to extrapolate the average value of the void faction in the riser.
T = 320°C y = 3.777E-05x y = 1.153E-05x 0 0.02 0.04 0.06 0.08 0.1 0.12 0.14 0 500 1000 1500 2000 2500 3000 3500 Axial Coordinate [mm] V o id F ra c ti o n 1 Nl/s 2 Nl/s 3 Nl/s 4 Nl/s 5 Nl/s 6 Nl/s
Figure 24. Steady state void fraction trend along the riser for the tests performed at 320°C.
Figure 26 plots all average void fraction values as function of the injected argon flow rate for the all performed tests.
T = 220°C y = 2.596E-05x y = 4.504E-05x 0 0.02 0.04 0.06 0.08 0.1 0.12 0.14 0.16 0 500 1000 1500 2000 2500 3000 3500 Axial Coordinate [mm] V o id F ra c ti o n 1 Nl/s 2 Nl/s 3 Nl/s 4 Nl/s 5 Nl/s 7 Nl/s
Figure 25. Steady state void fraction trend along the riser for the tests performed at 320°C.
0.001 0.01 0.1 0.01 0.1 1 jg [m/s] V o id F ra ct io n 320°C 250°C 220°C 200°C
Then, the driving force can be estimated by the following equation:
(
)
(
)
DF l m ris ris ris l g
p
ρ
ρ
gHα
gHρ
ρ
∆ = − = − (32)
Figure 27 shows as the driving force for enhanced circulation increases with the gas injection flow rate. Under steady-state conditions the driving force is equal to the total frictional pressure drop along the loop; so, figure 27 shows also as the flow resistance increases with the LBE flow rate (see also figure 28).
The pumping power related to the gas lift can be calculated as:
LBE p DF LBE M P p
ρ
= ∆ ɺ (33)The results of Equation (33) are displayed in figure 29; the picture represents the pumping power obtained for each steady-state reached in the system. As can be noted the data are best fitted by a power law relation. Once again, the LBE average temperature play an unimportant role.
0 5000 10000 15000 20000 25000 30000 35000 40000 0 0.002 0.004 0.006 0.008 0.01 0.012 0.014 Gas Injection Flow Rate [kg/s]
∆∆∆∆ pD F [P a ] 320°C 250°C 220°C 200°C
Figure 27. Driving force for enhanced circulation as function of the gas injection flow rate at different average temperatures of the system.
0 5000 10000 15000 20000 25000 30000 35000 40000 0 50 100 150 200 250
LBE Flow Rate [kg/s]
∆∆∆∆ pD F [P a ] 320°C 250°C 220°C 200°C
Figure 28. Driving force for enhanced circulation as function of the LBE flow rate at different average temperatures of the system
y = 1E-05x3.3309 0 100 200 300 400 500 600 700 800 900 0 50 100 150 200 250
LBE Flow Rate [kg/s]
Pp [W ] 320°C 250°C 220°C 200°C
Figure 29. Pumping power for enhanced circulation as function of the LBE flow rate at different average temperatures of the system
Similarly, figure 30 shows the trend of the pumping power as function of the injected gas flow rate; increasing the argon flow rate a higher pumping power is available for the circulation of the liquid metal.
0 100 200 300 400 500 600 700 800 900 0 0.002 0.004 0.006 0.008 0.01 0.012 0.014 Gas Injection Flow Rate [kg/s]
Pp [W ] 320°C 250°C 220°C 200°C
Figure 30. Pumping power for the enhanced circulation as function of the gas injection flow rate at different average temperatures of the system
In order to evaluate the efficiency of the gas lift technique, it is useful to define a performance index defined as:
p
comp P P
η
= (34)where P is the available pumping power, as already defined by equation (33), and p comp
P is the compression power supplied to the system, defined as:
comp g ad
P = ɺM l (35)
In the previous equation, l represents the specific adiabatic work required ad to compress and circulate the gas in the system (from the cover gas to the injection point). For a perfect gas, it is:
1 1 1 k k ad a k l RT k
β
− = − − (36)where:
• k Cp 1.67
Cv
= = ;
• R is the argon gas constant whose value is 208.21 J kg K; • T is the cover gas temperature; a
• inj
a p
p
β
= is the compression ratio, with p and a p cover gas and injinjection pressure respectively.
0 200 400 600 800 1000 1200 1400 1600 1800 0 50 100 150 200 250
LBE Flow Rate [kg/s]
Pc o m p [W ] 320°C 250°C 220°C 200°C
Figure 31. Compression power for the enhanced circulation as function of the LBE flow rate and at different average temperatures of the system
Figure 31 shows the results obtained by the application of the equation (35) on the available experimental data; the compression power increases with the liquid metal flow rate. Then, by equation (34), it is possible to obtain the trend of the performance index, shown in figure 32.
As it can be noted, the trend of the performance as function of the injected gas flow rate shows a maximum for gas flow rates in the range of 3÷6 Nl/s. In this condition, the LBE flow rate is equal to 200÷250 kg/s; so, for the nominal conditions,
the system has nearly the maximum ideal circulation performance, though it is less than the 60%. 0 0.1 0.2 0.3 0.4 0.5 0.6 0.7 0 0.002 0.004 0.006 0.008 0.01 0.012 0.014 Gas Injection Flow Rate [kg/s]
ηηηη
320°C 250°C
220°C 200°C
Figure 32. Ideal gas enhanced circulation performance as function of the gas injection flow rate at different average temperatures for the system.
3-6. REFERENCES
[1] G. Benamati, C. Foletti, N. Forgione, F. Oriolo, G. Scaddozzo, M. Tarantino, “Experimental study on gas-injection enhanced circulation performed with the CIRCE facility”, Nuclear Engineering and Design, vol. 237, pp. 768-777, Iss. 7, 2007
[2] Report ENEA HS-A-R-016, “Report on Gas Enhanced Circulation Experiments and Final Analysis (TECLA D41)”, G. Benamati, G. Bertacci, N. Elmi, G. Scaddozzo, January 2005;
[3] L. Cinotti., et al., “XADS cooled by Pb-Bi System Description”, Proceeding of International Workshop on P&T and ADS Development, Mol (Belgium), October 6-8, (2003), ISBN 90769971072, SCK-CEN (2003)
[4] Report ANSALDO CIRCE 8 SATX 0129, “Specifica tecnica di approvvigionamento: misuratori di portata sul circuito di ricircolo dell’argon”, February 2001 (in Italian);
[5] Report ANSALDO CIRCE 1 SREX 0004, “Specifica di Prova- Prove di Circolazione Assistita” August 2000 (in Italian);
[6] Report Ansaldo TRASCO 11 TNLX C016, “Dimensionamento degli orifizi che simulano le perdite di carico nelle prove di circolazione assistita in CIRCE”, October 2003 (in Italian);
[7] C. Rubbia, S. Buono, Y. Kadi and J.A. Rubio “Fast Neutron Incineration in the Energy Amplifier as an Alternative to Geologic Storage: the Case of Spain”, European Organisation for Nuclear Research, CERN/LHC/97-01 (EET), Geneva, 17th February 1997;
[8] C. Rubbia, J.A. Rubio, S. Buono, F. Carminati, N. Fiétier, J. Galvez, C. Gelès, Y.Kadi, R. Klapisch, P. Mandrillon, J.P. Revol and Ch. Roche “Conceptual Design of a Fast Neutron Operated High Power Energy Amplifier”, European
Organisation for Nuclear Research, CERN/AT/95-44 (ET), Geneva, September, 1995;
[9] M. Carta, G. Gherardi, S. Buono, L. Cinotti “The Italian R&D and Industrial Program for an Accelerator Driven System Experimental Plant”, IAEA-TCM on Core Physics and Engineering Aspects of Emerging Nuclear energy Systems for Energy Generation and Transmutation, ANL, Argonne (USA), 28 November -1 December 2000;
[10] W. Ambrosini, N. Forgione, F. Oriolo, M. Tarantino, "Analisi delle prove di circolazione assistita eseguite sull’apparecchiatura CIRCE", Università di Pisa, Dipartimento di Ingegneria Meccanica, Nucleare e delle Produzioni, DIMNP RL 1041 2004 (in Italian);
[11] N. E. Todreas, M. S. Kazimi, “Nuclear System I, Thermal Hydraulic Fundamentals” Taylor&Francis, New York, 1989;
[12] Y. Taitel, D. Bornea, A. E. Dukler, “Modelling flow pattern transitions for steady upward gas-liquid flow in vertical tubes”, AIChE J. 26:345, 1980;
[13] R. W. Lockart, R. C. Martinelli, “Proposed correlation of data for isothermal two-phase two-component flow in pipes”, Chem. Eng. Prog. 45:no. 39, 1949;
