Capability of a falling film microstructured contactor for the separation of binary mixtures

Chemical Engineering Journal 167 (2011) 455–467

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Capability of a falling film microstructured contactor for the separation of binary mixtures Abdoulaye Kane, Hubert Monnier ∗ , Daniel Tondeur, Laurent Falk Nancy University, Laboratory of Reactions and Chemical Engineering, CNRS, 1 rue Grandville, 54 000 Nancy, France

a r t i c l e

i n f o

Article history: Received 14 April 2010 Received in revised form 22 September 2010 Accepted 23 September 2010 Keywords: Falling film Intensification Evaporation Boiling Separation Microprocess

a b s t r a c t The present study investigates evaporation and separation of a binary mixture of ethanol and n-propanol in a microstructured falling film contactor made from two electrically heated vertical plates grooved with straight microchannels and forming the sides of a common gas chamber. Experiments were performed at atmospheric pressure (101.3 kPa) and various wall temperatures above the saturation temperature of the inlet binary mixture. Electrical heating fluxes ranged from 2 to 4 kW/m2 and feed flowrate between 1 and 8 g/min. The feed concentration was fixed around 46 mol.% of ethanol. The results indicate that the microcontactor performances depend essentially on heat flux and feed flowrate and provide fundamental insight on critical values of these two factors. Analysis of the results point out that intermittent local dry-zone can occur at high heat flux or at low feed flow rate. This occurrence influences microcontactor performances which can be evaluated in terms of stripping efficiency and power of separation. © 2010 Elsevier B.V. All rights reserved.

1. Introduction Research in microprocess engineering is mainly oriented towards microstructured and intensified reactors, however there are also real needs for intensified and microstructured separators for sustainable development. Microstructured separators, associated with microstructured reactors would contribute to the development of mini-plants which could deliver moderate throughputs on site and on demand, for high value and/or dangerous chemical products. In this context, separation methods based on heat transfer (evaporation and distillation) are widely used in macro devices but application in micro technology is a difficult challenge mainly because capillary forces are often higher than gravity forces and thermal conductivity associated with compactness of equipments smooth the thermal gradients. Two important heat transfer modes, evaporation and flow boiling in small hydraulic diameter channels are becoming increasingly important in various applications. Micro and minichannels are characterized by high specific areas that permit to exchange heat and mass much faster than conventional technologies. It has been shown that miniaturization by the use of micro channels can result in higher system efficiency [1,2]. In this study, the developed devices use falling film technology, in which transfer intensification can be obtained through the thinness of the liquid layer. Falling film evaporation finds applications

∗ Corresponding author. E-mail address: [email protected] (H. Monnier). 1385-8947/$ – see front matter © 2010 Elsevier B.V. All rights reserved. doi:10.1016/j.cej.2010.09.090

in numerous industrial processes, e.g. for concentrating solutions. One particular aspect of falling film is that the gas liquid interface often displays waves [3], causing improvement of the heat and mass transfer rates, because of the presence of interfacial shear stress exerted by counter-current vapour flow. Due to a continuous vapour space, the associated pressure drop with falling film evaporation is very low compared to convective boiler. Falling liquid film systems are suitable for processing heat sensitive fluids [4–6] because of short residences times and are inherently efficient thermodynamically since high heat flux can be obtained with a low temperature driving force. In compact equipments, small dimensions may impose some difficulties: first, the driving force inducing liquid flow by gravity is very small. Second, the small size and the high thermal conductivity of the material induce thermal homogeneity in the system, which makes it difficult to create a longitudinal thermal gradient. Concerning phase-change heat transfer and mixtures separation in falling film, these difficulties are more pronounced because of the existence of high capillary forces against the phases separation, the possible formation of dry out zones, and the short residence time which is detrimental to the efficient contact between vapour and liquids. Moreover, the liquid falling film may sometimes break up because of surface tension and/or temperature gradient effect [7,8]. Numerous studies [9,10] have shown that structuration of the solid support allows generating and stabilizing thin liquid films. Classical falling film systems such as flat surfaces generate films with thicknesses of the order of millimetres. By contrast with micro channels machined on vertical plates, rivulet formation is avoided even at low flow-rates if the channel widths are sufficiently

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Nomenclature B cP d dhyd. D E F g h I Lv P Psep q Q R ReL S T0 U

x z

bottom flow rate (mol/s) calorific capacity (J/mol/K) diameter (m) hydraulic diameter (m) distillate flow rate (mol/s) separation efficiency (–) feed flow rate (mol/s) gravity (m/s) molar enthalpy (J/mol) Intensity (A) latent heat (J/kg) total pressure (Pa) power of separation (compositional exergy) (W) heat flux (W/m2 ) heat flux (J/S) molar gas constant liquid Reynolds number (–) surface (m2 ) temperature of environment, K current tension (V) average liquid velocity in a single channel(< uL >= Fvol /2nC ıL wC ) ethanol liquid mole fraction (–) axial position (m)

Greek letters ıL thickness of liquid film (m) L liquid dynamic viscosity (Pa s)  density (kg/m3 )  temperature (◦ C) Microcontactor geometrical parameters dC channel depth (m) number of channels (–) nC sC space between two channels (m) tV vapour vein thickness (space between two plates) (m) plate thickness tw wC channel width (m) wnc width of the vapour vein (equal to the overall width occupied by channels) (m) plate width (m) wP Z channel length (m) Subcripts B bottom C channel crit. critical D distillate elec electric F feed h hydraulic L liquid max. maximum sat. saturation vol. volumic m mass V vapour w wall

Fig. 1. Expanded view of the microstructured contactor.

small, and one can generate stable films less than 100 ␮m thick [11]. The aim of the present work is to study heat transfer phenomena and liquid hydrodynamics in a microstructured contactor based on falling liquid film technology. This study analysed experimental measurements of heat and mass transfer for a binary mixture evaporating in downflow in a heated vertical microstructured plate. Emphasis is placed on the effects of critical conditions (related to feed flow rate and heat flux) on heat and mass transfer, also on flow characteristics. It is therefore of great significance to understand the fundamental phenomena associated with falling films in micro channels, such as flow boiling heat transfer, critical heat flux and two-phase flow, in the perspective of developing related new technologies (e.g. micro separators). 2. Experimental procedure 2.1. Microstructured contactor The experimental apparatus is particularly designed for investigating the evaporation of a liquid film in a heated microstructured contactor. As shown in Fig. 1, the liquid film falls along a vertical plate from liquid distributors and is collected as it leaves the plate. The main element of the experimental apparatus is the microcontactor whose details are presented in Figs. 1 and 2 which represents a detailed view of the grooved channels in the contacting plate. In the inside of the microcontactor, there are two structured brass plates (18 cm long, 9 cm wide and 2 mm thick) symmetrically disposed. The geometric parameters are summarized in Table 1. The plates are grooved with micro channels in order to prevent the breakup of the liquid film at lowflow rates [12]. The separating walls of the channels play the role of fins that drastically improve the liquid heating. These two plates are separated by a hollow spacing plate (20 mm thickness) made of PTFE (Teflon® ) coated fiberglass. This spacing plate constitutes the vapour chamber. To minimize heat losses due to convection and conduction, the microcontactor and associated devices are thermally insulated with glass wool. The PTFE material used for the distributor and spacer plates also acts as good thermal insulator. The brass plates are heated by direct Joule effect using a resistive circuit located on the opposite side of the grooved surface and a direct current rectifier with a variable current output of 0–3 A over

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Fig. 2. Details of contacting plate with a cross-section view.

the unit. The contactor is equipped with a liquid trap in order to prevent vapour outlet at the bottom of the system. Vapour leaves at the top of the microcontactor to be then totally condensed in a water condenser. No additional gas is introduced in the device, so that the internal convection in the vapour chamber is very limited. By contrast with earlier designs [9], in this device, the vapour flowing in the spacing plate is in permanent and direct contact on both sides with the liquid phase flowing in the channels, minimizing wall condensation effects. Of course this configuration doesn’t allow the visualization of liquid and vapour flows. 2.2. Experimental set-up and conditions

Fig. 3. Arrangement and location of thermocouples on microstructured plates.

a voltage range of 0–30 V.The heating of the liquid phase is then ensured by conduction through the plate thickness. On each brass plate, four Chromel-Constantan thermocouples (0.5 mm wire) are attached to the outer surface, as described in Fig. 3, to measure the local wall temperature as a function of the z axial position. Before its enters at the top of the microcontactor, the liquid feed is preheated in a heat exchanger, and then flows over the microstructured plate to form a uniformly distributed film with a typical estimated thickness of about 60 ␮m. This estimation comes from observations on another microcontactor of the same type (IMM falling film microreactor) [10]. A more accurate estimation based on Kapitza’s equation is presented below. The falling film is partly vaporized in the microcontactor under a constant heat flux and the concentrated liquid phase is withdrawn at the bottom of

The experimental set-up is presented in Fig. 4, and the operating conditions are reported in Table 2. The liquid mixture stored in the feed reservoir is pumped into the liquid distributor. It is then distributed along the width of the entrance slot of the structured brass plate, and falls along the vertical plate as a falling film. Upon leaving the plate, the liquid is collected in a small container for chromatographic analysis and is then continuously pumped to prevent liquid accumulation in the contactor. The vapour is superheated (superheater 5) before being condensed to avoid condensation inside the device and a reflux effect which is not part of the study. The feed consists of a liquid alcohol mixture (0.46–0.48 ethanol mole fraction, and the complement as n-propanol). This liquid is preheated in the distribution chamber but arrives at the top of the film at a temperature below the mixture boiling point: 47 ± 2 ◦ C (Table 4). In these experiments, all chemicals were of analytical grade (ethanol 100%; n-propanol 99.6%) and obtained from commercial sources (Roth® ).Their physical properties are reported in Table 3, and the mixture is assumed to be ideal. The liquid flow rate must be chosen judiciously: on the one hand it should be higher than the minimum wetting flowrate in order to form a uniform

Table 1 Geometric parameters of microstructured brass plate. wp (mm)

wnc (␮m)

wC (␮m)

dC (␮m)

wC /dC (–)

sC (␮m)

nC (–)

tw (mm)

tV (mm)

Z (mm)

90

40

1000

200

5

100

36

2

20

150

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Table 2 Operating conditions. Fm (g/min)

F (mol/min) −2

1.81 × 10

1–8

−1

–1.5 × 10

Fvol (mL/min)

ıL (␮m) (Kapitza’s law)

ReL

xF (mol/mol)

qelec (W/m2 )

 sat. (◦ C)

0.6–5

60–145

1–6

0.46–0.48

2000–5000

87

Table 3 Physical properties of working fluids—P = 101.3 kPa. Fluid

 sat. (◦ C)

L (10−3 Pa s) at 20 ◦ C

L (kg/m3 ) at 20 ◦ C

Ethanol n-Propanol

78.4 97.1

1.08 1.95

789 805

falling film and to avoid the formation of dry zones on the contacting plate; on the other hand it should not exceed a certain value which would causes channels overflow. In this study, we used Kapitza’s equation to evaluate the film thickness in the micro channel and to predict the optimal feeding flow rate for pure ethanol [13]. In laminar flow regime, Kapitza’s relation writes:

 ıL =

2.42L ReL

1/3

L2 g

(1)

where ReL =

L < uL > dh L

(2)

ıL wC 2ıL + wC

(3)

and dh = 4

Eqs. (2) and (3) give ReL =

2L Fvol. L nC (2ıL + wC )

(4)

The flow rates (Table 2) have to be kept between 1 and 10 mL/min (for the two plates). These are the limiting conditions to avoid both the formation of dry zones and channel overflow. The

value of 10 mL/min flow rate corresponds to a Reynolds number of the liquid, ReL below 6. The contacting device presented in this study can be compared to a stripping section of a distillation column (often referred to as the lower section in a classical distillation column, located below the feed). The main difference is that there is no reflux nor reboiling: the vapour phase is generated in situ continuously along the falling film path. The bottom product (richer in the heavy component, npropanol) and the distillate (richer in the light component, ethanol) are obtained at the bottom and at the top of the microcontactor respectively. All experiments are carried out at a controlled heat flux density maintained by an electric resistance in contact with the brass plate. The total electric power dissipated in the test section was calculated as the product of voltage and current. The temperatures are monitored using a data acquisition system (Pico Technology Limited TC-08 data logger) associated with TestPoint® software. The accuracy of the readings was estimated to be within 1 ◦ C. Analogous experiments carried out with pure ethanol confirm this estimate. The wall temperature of the plate on the liquid film side was calculated by applying Fourier’s law accounting for the heat conducted through the wall and taking into account heat loss from the surface heating and the thermal conductivity of the plate. The collected liquid samples were analysed by gas chromatography, using a Varian 3900 gas chromatograph equipped with a DB-624 column (30 m × 0.53 mm i.d., 3 ␮m film thicknesses) and a flame ionisation detector (FID). The flow rates of both the feed and the vapour (after condensation) were measured by using the “bucket and stopwatch method”, and the flow rate of the bottom product was deduced

Fig. 4. Experimental setup for a microstructured falling film distillation (1—falling film contactor; 2—feed reservoir; 3—HPLC pump; 4—bottom reservoir; 5—electrical superheater with temperature controller; 6—electrical power supply; 7—cryostat with integrated pump (cooling); 8—peristaltic pump; 9a–9b—distillate and bottom sampling; 10a–10b—condensers).

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Fig. 5. Energy balance of the microcontactor distillation cell.

by difference. Data reproducibility was checked by replicating experiments.

represented schematically in Fig. 5, writes: Q = Qelec − Qloss = D · hV + B · hB − F · hF

(6)

2.3. Thermal characterization of the falling film device Evaporation of the binary film is induced by the electric power supplied by an electrical film heater located on the outer surface of the plate. The electrical heat flux dissipated by Joule effect through the two plates is defined as the ratio of the total power divided by the surface of the film heater Selec (Fig. 2). Considering the two plates of the cell, electrical heat flux is expressed as follows (Eq. (5)): qelec =

2UI UI = 2Selec Selec

(5)

with Selec = 2 × 7.42 × 10−3 m2 because there are two resistance plates (Fig. 2). The heat flux transferred to the liquid falling film is obtained through a heat balance over the distillation cell, to account for the heat losses of the system. In stationary conditions, the heat balance,

• Q is the heat power transmitted to the liquid film as sensible and latent heat. • Qelec is the overall electric power = 2UI • D, B and F are the distillate flow rate, the bottom flow rate and the feed flow rate respectively. • hV , hB , hF are the associated specific enthalpy of the different inlet and outlet flows [14].

The measurement of the composition and the temperature of the outlet flows allows to determine the different specific enthalpies. The accuracy of determination of Q is estimated from several experiments carried out under the same thermodynamic and hydrodynamic conditions. It is estimated as ±9%.The heat flux transferred to the liquid falling film is then calculated as the ratio of the heat power Q to the structured surface area, which is defined

Fig. 6. Longitudinal temperature profile at different electrical heat fluxes (F = 4.70 × 10−4 mol/s, xF = 0.46,  F = 47 ± 2 ◦ C,  sat. = 87 ◦ C).

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Fig. 7. Evolution of the temperature at the middle of the plate (z = Z/2) versus the electrical heat flux (F = 4.70 × 10−4 mol/s, xF = 0.46,  F = 47 ± 2 ◦ C,  sat. = 87 ◦ C).

3. Experimental results

as the total surface of the grooved channels bottom:

qliq =

Q Q = SL 2wnc Z

(7)

An important characteristic of the system, which is essential for the comprehension of the results presented hereafter in the paper, is the temperature profile of the plate. Fig. 6 shows the longitudinal temperature profile for different electrical heat fluxes, at a feed flow rate of 4.70 × 10−4 mol/s. One recognizes a slight longitudinal gradient between the inlet (top of the plate, z = 0) and the outlet (bottom) of the liquid falling film, which is induced by the evaporation of the binary mixture. This gradient is also due to the higher volatility of ethanol compared to n-propanol that induces the enrichment in n-propanol of the liquid falling film (saturation temperature of 97.1 ◦ C against 78.4 ◦ C for ethanol). Note that at high electrical power, the temperature of the plate can be higher than the mixture saturation temperature of the liquid. Fig. 7 represents the evolution of the temperature at the middle of the plate versus the electrical heat flux, for a feed flow rate of 4.70 × 10−4 mol/s. Below 3500 W/m2 , the temperature of the plate is lower than the saturation temperature of the inlet mixture. At higher electrical heat flux, there is a dramatic temperature increase leading to a high overheating of the plate above the inlet saturation of the liquid.

We characterize the overall performance of separation by an efficiency based on the molar concentrations of the more volatile component (ethanol) in the feed and in the bottom product, and defined as (Eq. (8)): E=

xF − xB xF

(8)

The experimental results of a typical run are shown in Table 4. 3.1. Boiling regimes in the falling film microcontactor Fig. 8 shows the evolution of the heat flux transferred to the fluid plotted against the difference between the plate wall temperature and the saturation temperature of the inlet mixture (87 ◦ C), obtained at different electrical heating powers. When increasing the heating power, the mean wall temperature of the plate rises towards an equilibrium temperature which depends on the inlet flow rate. In the present figure, one recognizes the typical behaviour identified by Nukiyama [15] of different boiling heat transfer regimes. At low temperature difference between the wall and the saturation temperature, the boiling heat transfer is characterized by nucleate boiling where the liquid phase is wetting the solid surface. Moderate increase of the wall temperature results in a stiff increase of the heat flux transferred to the fluid, resulting in an efficient operating regime. When the temperature

Table 4 Experimental data for a typical run—qelec = 2986 W/m2 , P = 101.3 kPa. Quantities xF (mol/mol) xB (mol/mol) xD (mol/mol) E (%)  feed (◦ C)  distillate (◦ C)  bottom (◦ C) Fm (g/min) Bm (g/min) Dm (g/min) Q (W)

Run 1 46.9% 37.8% 54% 19.4% 47.5 87.1 87.5 1.52 0.65 0.87 15.82

Run 2

Run 3

Average

Standard deviation

45.9% 38.3% 52.2% 16.6% 46.7 86.9 87.1 1.52 0.63 0.89 15.58

48.5% 39.9% 51.7% 17.7% 48.9 87.6 88.1 1.60 0.67 0.93 17.03

47.1.% 38.7% 52.6% 17.9% 47.9 87.2 87.6 ∼1.55 ∼0.65 ∼0.90 ∼16.14

0.013 0.011 0.012 0.014 1.114 0.361 0.503 0.046 0.020 0.031 0.78

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Fig. 8. Evolution of the heat flux transferred to the fluid versus the temperature difference between the wall and the inlet saturation temperature (F = 4.70 × 10−4 mol/s, xF = 0.46,  F = 47 ± 2 ◦ C,  sat. = 87 ◦ C, electrical heat flux: 2290 W/m2 , 2964 W/m2 , 3503 W/m2 , 3907 W/m2 , 4446 W/m2 ,  w (z = Z/2)).

difference is higher than 15–20 ◦ C, the thin layer of vapour is formed between the wall and the liquid, named calefaction phenomenon, resulting in the important diminution of the heat transfer to the liquid film. It is clearly seen that two regimes occur, materialized by the maximum of the heat flux transferred to the fluid, with a transition at some “critical” value of the electrical heating flux (here qcrit ∼3500 W/m2 ). elec The heat transfer coefficient in the nucleate boiling regime, given by the slope of the curve, is estimated to 200 W/m2 /C which is surprisingly low (4–10 times smaller) compared to experimental values obtained in macro falling systems by Palen et al. [4]. This is due to the value of the Reynolds number which is very low: the liquid flow is “smooth” laminar while it is usually laminar with waves or eddies. Therefore, in our case, the heat transfer in the liquid phase occurs mainly by diffusion. Calefaction involves also high overheating of the plate (40 ◦ C and 70 ◦ C above the inlet saturation temperature), which explains that, when the electric heating power is increased gradually with

constant power steps, the temperature of the plate does not rise regularly. In the literature, different interpretations are called upon to explain this phenomenon, depending on the situation: dry spots, dry-out, and calefaction. Usually [16] one calls dry-out the complete disappearance of the liquid phase, a phenomenon which can be ruled out a priori in our case, because of a measurable liquid flow rate at the bottom of the evaporator. Calefaction is the appearance of a continuous thin vapour film in contact with the wall, which acts as a thermal insulator for the liquid [17]. Dry zones may be expected by analogy with known thermocapillary flow phenomena [18] that occur in two-dimensional falling films. Notice that dry zones may occur as a spatial segregation (on part of the plate surface) and/or as an intermittent phenomenon at any given location, including in single channels.

Fig. 9. Wall temperature (at the middle of plate) as a function of time F = 4.70 × 10−4 mol/s, xF = 0.46,  F = 47 ± 2 ◦ C,  sat. = 87 ◦ C.

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Fig. 10. Bottom reduced concentration (xB /xF ), bottom reduced flow rate (B/F) and stripping efficiency (E) as a function of heat flux transferred to the film—F = 4.70 × 10−4 mol/s, xF = 0.46,  F = 47 ± 2 ◦ C,  sat. = 87 ◦ C; electrical heat flux: 2290 W/m2 , 2964 W/m2 , 3503 W/m2 , 3907 W/m2 , 4446 W/m2 ).

Fig. 9 gives a somewhat different illustration of the existence of different regimes. In this figure the wall temperature as a function of time is shown for two different electrical heat fluxes. At qelec = 2290 W/m2 , the wall temperature is practically constant and slightly higher than the boiling temperature of the initial mixture; this behaviour may be associated to a steady state quiescent boiling regime. At qelec = 3907 W/m2 , clear temperatures variations may be seen, characterized by periodical slow increase interrupted by sharp decrease. Also the average temperature level is much higher than in the previous case, thus than the boiling temperature. The fluctuations may be attributed to an unstable behaviour in the boiling regime, and suggest the occurrence of dry zones accompanied by wall superheating followed by rewetting of the wall surface. The higher average temperature level of the wall suggests a lower heat transfer efficiency towards the liquid film. Nevertheless, the usually interpretations of unstable boiling cannot be systematically transposed to the present situation because of the particularity of coupled heat and mass transfer in mini-channels, as discussed by Cheng and Mewes [19] and Thome [1]. The experimental results presented hereafter will not only help discriminating between these interpretations, but also highlight the performances and limits of such falling film contactors. 3.2. Evaporation and separation capability In order to analyse the microcontactor performances and limits, it is useful to represent, for a given feed rate (F) and ethanol concentration (xF ), the reduced values of bottom concentration (xB /xF ) and bottom flow rate (B/F), as well as separation efficiency E defined in Eq. (8), as a function of the heat flux transferred to the fluid. This is shown in Fig. 10. The bottom flow rate (B/F) decreases linearly with increasing the heat flux transferred to the liquid film. This expected result means that the heat flux transferred to the liquid film is almost proportional to the product of the flow rate of the vapour leaving the liquid film by the specific enthalpy of the vapour. The constant slope (straight line) also shows that the specific enthalpy of the vapour remains almost constant, i.e. the composition of the vapour is moderately changed whatever the heat flux transferred to the

liquid film, as illustrated by the evolution of the bottom reduced concentration (xB /xF ). As attested by Fig. 7, higher experimental values of the evaporated flow rate do not correspond to higher electric heating power. The maximal evaporation rate is obtained in mild boiling conditions at 1200 W/m2 , corresponding to an electric heat flux of 3503 W/m2 and a plate temperature of 15 ◦ C above the inlet saturation temperature. Higher electric heating leads to smaller evaporation rate. Fig. 10 is also reported the bottom reduced concentration (xB /xF ) and the separation efficiency (E) versus the transferred thermal flux, which exhibit a complex behaviour. For interpretation, let us consider in Fig. 11 the boiling point diagram with the equilibrium dew point curve and the boiling point curve of the mixture ethanol and n-propanol, near the composition of the inlet (xF = 0.46 mol fraction of ethanol). At low heating power, the temperature of the plate is relatively low near the boiling curve. In this case, the composition of the boiling liquid is equal to the inlet composition (point A, xB /xF = 1) and the composition of the vapor, given by the intersection of the horizontal line and the dew point curve, is higher than the inlet composition with xD /xF = 0.65/0.46 = 1.41. At higher temperature of the falling film equal to the limit given by the dew point curve (point B), the composition of the vapour is then equal to the inlet composition xD /xF ≈ 1 and the liquid at the bottom is xB /xF = 0.28/0.46 = 0.6. When the temperature of the falling film (point C) is progressively increased from the boiling point temperature up to the dew point temperature, the bottom and distillate compositions sweep the entire domain given by the boiling diagram. The composition curve in Fig. 11 of vapour and liquid at the outlet of the distillation cell is then strongly related to the temperature of the liquid falling film and also to the temperature of the plate. Since it was not possible in the device to measure the temperature of the falling film (because of dry out zone formation, calefaction, etc.), we have reported in Fig. 12 the evolution of the relative mole ratio of the distillate and the bottom liquid versus the temperature difference between the inlet saturation temperature and the plate temperature. One observes indeed the decrease of the relative mole ratio of the distillate from 1.34 (maximal theoretical value xD /xF = 1.41) towards 1, when increasing the temperature of

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463

Fig. 11. Boiling point diagram of the mixture ethanol–n-propanol. Enlargement around the inlet composition xF = 0.46,  F = 47 ± 2 ◦ C.

the plate and then the temperature of the liquid falling film. Similarly, the relative bottom mole ratio decreases from the value 1 down to 0.86 and then increases again to reach the value 1. At low heat flux values, the temperature of the plate (i.e. the film) is moderately higher than the saturation temperature and nucleate boiling is taking place in the falling film. In this way, essentially all the heat flux leaving the wall passes through the liquid film to supply the latent heat of vaporization at the interface, inducing the stripping of the more volatile component from the falling film all along the microcontactor. Thus xB /xF decreases,

and thereby stripping efficiency (E) increases, also B/F decreases gradually because of the evaporation of the more volatile (ethanol) component. At high heat flux values, the temperature of the plate is very high, above the boiling point curve which induces strong instabilities, calefaction and dry out zone formation. As a matter of fact, for the highest values of the heat flux, we visually observed some sputtering of the liquid leaving at the bottom as reported by Yoshioka and Hasegawa [17] suggesting a “wild calefaction” (Leidenfrost effect) which explains that liquid and vapour flow out of the distillation

Fig. 12. Bottom and top reduced concentration (xB /xF and xD /xF ) as a function of temperature difference—F = 4.70 × 10−4 mol/s, xF = 0.46,  F = 47 ± 2 ◦ C,  sat. = 87 ◦ C; electrical heat flux values: 2290 W/m2 , 2964 W/m2 , 3503 W/m2 , 3907 W/m2 , 4446 W/m2 ).

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Fig. 13. Heat transferred from the heated surface to the liquid film (Q), sensible heat (Qsensible ) and latent heat (Qlatent ) as a function of feed flow rate 2 Q = FcP (F − sat. ) + DLv − qelec = 2986 W/m , xF = 0.46,  F = 47 ± 2 ◦ C,  sat. = 87 ◦ C.





Qsensible

  Qlatent

cell without any separation of the binary mixture of methanol and n-propanol.

liquid from its inlet temperature to its boiling temperature, and a latent contribution, which is a measure of the flux evaporated (Eq. (9)).

3.3. Influence of feed flow rate

Q = FcP (F − sat. ) + DLv

Experiments were run with a fixed heat generation Q and at feed flow rate varying between 3 and 9 × 10−4 mol/s to investigate the influence of this parameter on the separation. Fig. 13 represents the heat transferred from the heated surface to the liquid film, as deduced from an overall enthalpy balance on the fluid (Eq. (7) and Fig. 5) as a function of the molar feed flowrate. This quantity Q comprises a sensible contribution to raise the entering





Qsensible





(9)

Qlatent

One can clearly distinguish two different regimes (Fig. 14) of increasing distillate flow rate at low feed flow rate (zone 1), and decreasing distillate at high feed rate (zone 2), with a transition roughly located at 5 × 10−4 mol/s. At first sight, this behaviour points toward the existence of an optimal regime for the separation, and this is strengthened when looking for example at the evolution

Fig. 14. Wall temperature (at the middle of plate) as a function of time qelec = 2986W/m2 , xF = 0.46,  F = 47 ± 2 ◦ C,  sat. = 87 ◦ C.

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Fig. 15. Heat flux transferred from the heated surface to the liquid film (Q) versus the temperature difference between the medium plate and the inlet saturation for different feed flow rates and constant electric heating power qelec = 2986W/m2 , xF = 0.46,  F = 47 ± 2 ◦ C,  sat. = 87 ◦ C.

of the separation efficiency as a function of the feed flow rate. However, some understanding of the mechanisms and causal relations must be brought in before drawing conclusions, calling for further studies. At very low feed flow rate (3 × 10−4 mol/s), the temperature of the plate is relatively high 116 ◦ C (Fig. 15) i.e. 29 ◦ C above the inlet saturation temperature, resulting in a moderate thermal flux transferred to the liquid film (650 W/m2 ), because of the calefaction phenomenon. Some partial dry-out takes place, meaning that the solid surface is not always and/or not everywhere wetted. This is substantiated by the local temperature measurements shown in Fig. 15, which at low flow rate shows a substantially higher temperature than at higher flow rate, and also periodic variations, both being hints to intermittent dry zones or spots, as already discussed in connection with Fig. 9.

At higher flow rate (5 × 10−4 mol/s), the temperature of the plate falls off at 98 ◦ C which is 11 ◦ C above the inlet saturation temperature. The resulting heat flux transferred to the liquid film is maximum at 1420 W/m2 . At still higher inlet flow rate (6.3 × 10−4 mol/s, 7.5 × 10−4 mol/s and 9.3 × 10−4 mol/s), the temperature of the plate stabilizes around 95 ◦ C, corresponding respectively to 8.3 ◦ C, 7.9 ◦ C and 7.6 ◦ C above the inlet saturation temperature. Fig. 16 represents the heat flux transferred from the heated surface to the liquid film at various temperature differences between the middle height of the plate and the inlet saturation, provides complementary information. Indeed, in this boiling regime, the heat flux transferred to the liquid film is very sensitive to the temperature difference which explains the strong impact of the feed flow rate upon the evaporation phenomenon.

Fig. 16. Amount of distillate produced as condensate at the top exit and stripping efficiency (E) as a function of feed flow rate—qelec = 2986 W/m2 , xF = 0.46,  F = 47 ± 2 ◦ C,  sat. = 87 ◦ C.

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Fig. 17. Power of separation (Psep ) as a function of the feed flow rate for different electrical heat fluxes xF = 0.46,  F = 47 ± 2 ◦ C,  sat. = 87 ◦ C.

3.4. Microcontactor performances in terms of power of separation To fully characterize the performance of the contactor, we need to distinguish between the degree of purification which involves only intensive quantities (we have used here the so-called separation efficiency), and production yield involving extensive quantities (for example, we have used here the bottoms flow). If we desire to characterize the performance with one single criterion, it appears that this criterion must combine mass balance, composition and the required energy for the separation. Thermodynamics provide a convenient way to do this since it considers the underlying ideas of quality of energy (required to achieve the separation). Sorin and Rheault [20] define a criterion of performance called “power of separation”. It is derived from exergy balance considerations and involves both intensive and extensive quantities. The “power of separation” represents the ideal compositional exergy change between input and output streams of the process (Eqs. (10) and (11)). According to Sorin and Rheault in many simple cases, “The maximum power of separation is achieved when half of the exergy spent in the process is destroyed and the other half is transformed into work of separation”.



Psep = RT0



Foutlet xoutlet ln xoutlet −

outlet





Finlet xinlet ln xinlet

inlet

(10) In our case: Psep = RT0 [(DxD ln xD + BxB ln xB ) − FxF ln xF ]

(11)

Fig. 17 shows for three heat fluxes, the evolution of the “power of separation” as a function of the feed flow rate for the operating conditions indicated in the legend. For each heat flux, the curves show a maximum value of the “power of separation”; this behaviour points toward the existence of an optimal regime for the separation. As discussed above, at low flow rate, dry-spots can occur, and this certainly influences the evolution of the heat transfer coefficient, and consequently affects the behaviour of the power of separation, Psep increases when feed flow rate increases.

For a given heat flux, once the maximum power of separation is reached, an increase in feed flow rate deteriorates the contactor performance (Psep decreases when feed flow rate increases). As depicted in Fig. 17, we can suggest operating with a feed flow rate around the “optimum point” of the process. For a given heat flux, this “optimum point” corresponds to the maximum of the power of separation; we may assume that under these conditions liquid flow equipartition per channel is sufficiently good and the wetting properties permit a good heat transfer between plate and falling film. This region corresponds to a thermodynamic compromise between productivity (rates of distillate), product quality (purity) and the amount of heat delivered (thermodynamic efficiency). This sets a bound, beyond which a further increase and/or decrease in feed flow rates will result in significant reduction of microcontactor performances. 4. Conclusions In the first part of this study a self-designed version of a microcontactor has been described. This microcontactor is based on a falling film technology and shown to be suitable for studying the effects of heat flux and feed flow rate. A better understanding of plate thermal management and falling films hydrodynamics in microcontactor implies understanding the phenomena which affect heat and mass transfer. At high heat flux, an average temperature well above the boiling point and clear temperature variations were observed. These fluctuations were attributed to an unstable behaviour in the boiling regime at high heat flux, suggesting the occurrence of dry zones accompanied by wall superheating followed by wall surface rewetting. On the other hand, at low heat fluxes, a steady state quiescent boiling regime occurs. Feed flowrate and heat flux were both shown to have a critical influence on the mixture flow boiling. It was found that there is a maximum value of the power of separation which is a good compromise between product purity, production yield and heat required for separation. The experimental observations emphasize the potential of this technology, but also indicate the inherent complexity of the process (e.g. dry spots occurrences, difficulty to impose a thermal gradient

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at small scales). The microcontactor design opens new horizons for distillation. In order to better characterize the microcontactor, forthcoming investigations are planned by using a sensitive thermal imaging technique to study the temperature field and flow patterns of falling liquid film. In the aim of heat and mass transfer intensification, further studies will focus on geometrical parameters effects by changing plate structuration and varying vapour vein thickness. Indeed, the device offers a certain flexibility by simply changing structured plates and the plate spacing (vapour vein thickness modulation). Acknowledgements The authors gratefully acknowledge the CNRS Energy program, the Lorraine Regional Council (LRC) for the financial support and guidance as well as Professor Mikhail SORIN (CAMNET Varennes – Canada) for our fruitful discussions. References [1] J. Thome, Boiling in micro channels: a review of experiment and theory, International Journal of Heat and Fluid Flow 25 (2004) 128–139. [2] S.G. Kandlikar, Heat Transfer and Fluid Flow in Minichannels and Micro Channels, Elsevier, 2006. [3] F.W. Pierson, S. Whitaker, Some theoretical and experimental observations of the wave structure of falling liquid films, Industrial and Engineering Chemistry Fundamentals 16 (4) (1977) 401–408. [4] J. Palen, Q. Wang, J. Chen, Falling film evaporation of binary mixtures, AIChE Journal 40 (2) (1994) 207–214. [5] U. Gropp, G. Schnabel, E.U. Schlunder, Influence of liquid-phase mass transfer resistance on the selectivity during partial evaporation of binary mixtures in a falling liquid film apparatus, Verfahrenstechnik 15 (10) (1981) 725–728. [6] U. Gropp, G. Schnabel, E.U. Schlunder, Effect of liquid-side mass transfer resistance on selectivity during partial evaporation of binary mixtures in a falling film, International Chemical Engineering 23 (1) (1983) 11–17.

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