Transient volumetric heat transfer coefficient ...

Transient volumetric heat transfer coefficient prediction of a three-phase direct contact condenser Hameed B. Mahood, Adel O. Sharif & Rex B. Thorpe

Heat and Mass Transfer Wärme- und Stoffübertragung ISSN 0947-7411 Heat Mass Transfer DOI 10.1007/s00231-014-1403-4

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Author's personal copy Heat Mass Transfer DOI 10.1007/s00231-014-1403-4

ORIGINAL

Transient volumetric heat transfer coefficient prediction of a three-phase direct contact condenser Hameed B. Mahood • Adel O. Sharif Rex B. Thorpe

•

Received: 24 September 2013 / Accepted: 26 June 2014 Springer-Verlag Berlin Heidelberg 2014

Abstract An experimental investigation for the time dependent volumetric heat transfer coefficient of the bubbles type, three-phase direct contact condenser has been carried out utilising a short column (70 cm in total height and 4 cm inner diameter). A 47 cm active height was chosen with five different mass flow rate ratios and three different initial dispersed phase temperatures. Vapour pentane and constant temperature tap water as dispersed and continuous phases were implemented. The results showed that the volumetric heat transfer coefficient decreases with increased time until it almost reaches its steady state conditions. A sharp decrease in the volumetric heat transfer coefficient was found at the beginning of the operation and, diminished over a short time interval. Furthermore, a positive effect of the mass flow rate ratios on the volumetric heat transfer coefficient was noted and this was more pronounced at the beginning of the operation. On the other hand, the volumetric heat transfer coefficient decreased with an increase in the continuous phase mass flow rate and there was no considerable effect of the initial dispersed phase temperatures, which confirms that latent heat transfer is dominant in the process.

H. B. Mahood A. O. Sharif R. B. Thorpe Centre for Osmosis Research and Applications (CORA), Chemical and Process Engineering Department, Faculty of Engineering and Physical Sciences, University of Surrey, Guildford GU2 7XH, UK e-mail: [email protected] H. B. Mahood (&) University of Misan, Misan, Iraq e-mail: [email protected]; [email protected] A. O. Sharif The Qatar Foundation, Qatar Energy and Environment Research Institute, Doha, Qatar

List A Cpc hfg m_ Q t T u V Z

of symbols Cross-section area (m2) Specific heat of continuous phase (kJ/kg C) Latent heat of condensation (kJ/kg) Mass flow rate (kg/s) Heat transfer rate (kW) Time (s) Temperature (C) Velocity (m/s) Column volume (m3) Axial height (m)

Greek symbols q Density (kg/m3) D Difference (–) Subscripts c Continuous phase (–) d Dispersed phase i Inlet LM Log-mean temperature o Outlet

1 Introduction Direct contact heat condensers have many advantages over the surface type, such as a shell and tube condenser. These befits are exhibited in a vast area of applications such as water desalination, geothermal power generation, solar energy applications and emergency cooling of nuclear reactor systems. However, the main features of the direct contact heat exchangers are: high heat transfer rate, capability to work with a very low temperature driving force, very low fouling and corrosion, low cost and simple design and scale up.

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Author's personal copy Heat Mass Transfer

Practically, direct contact condensers can be classified into two types; one-component and two-component condensers. The main difference between them is the nature of the contacting fluids which is the same substance in the former but completely immiscible in the second. Much attention has been directed to the studying the one component direct contact condenser, which enable an increase in their industrial applications. On the other hand, very modest attempts can be found elsewhere in the literature regarding the two-component direct contact condenser, which is defiantly reflected in the lack of applications and the ambiguity in its mathematical design formulas. Nevertheless, the direct contact condensation phenomenon of two immiscible fluids, as a basis of the direct contact condenser, has been researched widely since 50 years ago. Most of these studies were concentrated on the heat transfer and hydrodynamics of a single bubble [1–13] and only very few addressed the multibubbles and bubble in train [14, 15]. The increasing attention on a new economic and sustaining energy necessitates the development of efficient energy recovery systems. However, the recovering energy from such processes and other thermal downstream processes might be enhanced by utilising the direct contact condenser. Accordingly, author and others have carried out a series of investigations regarding the three-phase direct contact condenser [16–18]. The characteristics of the threephase direct contact condenser as well as the effect of different parameters have been investigated and discussed. In this paper, the volumetric heat transfer coefficient of a bubbles type, three-phase direct contact condenser is developed. The analysis concerns the transient operation period of the condenser. In addition, the effect of mass flow rate ratios and initial dispersed phase temperatures are investigated experimentally.

2 Experimental setup and procedure A schematic layout of the experimental test rig is illustrated in Fig. 1. It consists of three main parts; the test column (DCC), the vaporising vessel and the water storage tank with auxiliary equipment. The test section, as a direct contact condenser column, is a Perspex cylindrical column of 70 cm height and 4 cm internal diameter. The heating (vaporising) vessel is a Perspex cylinder of large diameter with ID about 22 cm and height about 13 cm, with two covers fixed tightly on both ends and set on a digital scale to take the initial and the final vessel weights. The vessel contains a copper coil with an internal diameter of about 6 mm and a length of about 7 m, used for heating purposes by carrying hot water as a heating medium from a constant temperature water bath. In addition, the heating vessel is half filled with water. The liquid pentane as a dispersed phase is injected inside the

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heating vessel using a small feeding tube. This water has two main functions; it helps to make the heat distribution uniform inside the heating vessel and it can be used as a heating medium to maintain an almost constant temperature during dispersed phase vapour injection. The volume and weight of liquid pentane are measured before injection into the heating vessel. A K-type calibrated thermocouple is fixed to the upper part of the heating vessel to measure the dispersed phase (vapour) temperature. In addition, a venting pipe and pressure gauge are fixed on the upper vessel cover. At the same time, about 160 litres of water is pumped at a constant temperature from its storage tank as a continuous cooling phase. A calibrated rota meter is used for measuring the continuous phase (water) mass flow rate. A short isolated pipe, about 3 cm length, is used for connecting the heating vessel with the test section (direct contact column). The vapour injected through the short tube is controlled by a check valve in a similar technique to that used by [19–21]. Six calibrated thermocouples are fixed on the direct contact column at positions from the bottom of the column: TC2 @ 0 cm, TC3 @ 13.5 cm, TC4 @ 25 cm, and TC5 @ 37. One is for the cooling water inlet (@ 47 cm) and the other for condensing outlet temperature (nearly @ 52 cm). Another rota meter is fixed at the water outlet stream from the test column. The experiment procedure starts by preparing the cooling water by maintaining it at the desired temperature 19 C. Due to the large size of the tank, there is no considerable change in water temperature during the experiment, which takes about 3 min. A known volume and weight of liquid pentane are injected inside the heating vessel utilising a special inlet or small sized tube. Due to the immiscibility and the density difference of pentane and water, a layer of liquid pentane forms over the water. Hot water for a heating medium is pumped from a constant temperature water bath through the inlet coil into the heating vessel and returns to the water bath again through another coil end. The heating temperature is controlled and increased gradually until the desired temperature inside the heating vessel has been achieved. A specific level of cooling water or continuous phase (47 cm) in the direct contact column is chosen by adjusting the inlet and outlet water valves. A data logger of eight channels is implemented to measure the vapour temperature in the heating vessel and the water in the direct contact column and this reads them directly onto a PC computer. When liquid pentane starts boiling and vapour forms (the boiling point of pentane is 37 C), all the heating vessel valves are closed until the desired temperature is reached. A calibrated size of the open check valve (vapour injection line) is determined before the beginning of the real test by making calibration runs. Hence, the real test begins when the pentane vapour reaches its desirable value. Then the check valve is opened at the calibrated limit. The


Fig. 1 Schematic diagram of the experimental rig

Table 1 The initial conditions of the experiments Tv (C)

Tco (C)

Ho (m)

_v R¼m m_ c 100 %

37.6

19

0.47

43.69

19

0.47

22.97

19

0.47

12.23

19

0.47

8.61

38.4

41.7

19

0.47

6.69

19 19

0.47 0.47

43.69 22.97

19

0.47

12.23

19

0.47

8.61

19

0.47

6.69

19

0.47

43.69

19

0.47

22.97

19

0.47

12.23

19

0.47

8.61

19

0.47

6.64

pentane) injection into the column. This is useful to make the temperature uniform within the column. At the moment of vapour injection, the temperatures along the column start to be recorded on the PC. The vapour temperature is controlled to try to maintain it at its injection value. The fluctuation in the vapour temperature should be within ;0.5 C. The dispersed phase flow rate is determined by both weighting the remaining pentane in the heating vessel and taking away from the initial value, and by measuring the volume of condensate and dividing both results by the experiment time. The condensate is recovered completely from the water that sometimes drains with it using a conical separator flask. The associated water is thus drained out while the condensate is returned back to the heating vessel or collected and used again in another run. The initial conditions of the experiments can be shown in Table 1.

3 Results and discussion cooling water flow rate is determined and controlled using both inlet and outlet rota meters and recycled throughout the test column before the dispersed phase (vapour

The volumetric heat transfer coefficient of a bubbles type, three-phase direct contact condenser is measured

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experimentally employing five different mass flow rate ratios and three initial vapour temperatures. Based on the following assumption, the calculation procedure is made utilising energy balance: •

•

•

The sensible heat is small in comparison with the latent heat of condensing vapour; therefore, it is negligible in the calculations. Both continuous phase and dispersed phase mass flow rates are assumed constant along the column. This can be made reasonably acceptable by means of a constant holdup ratio along the column height which has been demonstrated for both direct contact evaporator [22– 24] and direct contact condenser [17, 18]. The heat losses from the direct contact column to the environment are ignored.

Fig. 2 Volumetric heat transfer coefficient versus time

The calculations are obtained via Eq. (1) below: Uv ¼

Q VDTLM

ð1Þ

where Q, V and DTLM denote the total heat transfer rate, direct contact column volume and log-mean temperature difference respectively. A direct contact heat transfer process with change of phase, generally, is associated with a dominant latent heat transfer; therefore the simple heat balance leads to: Q ¼ m_ c Cpc DTc ¼ m_ d hfg

ð2Þ


where DTc represents the continuous phase temperature’s difference ðTco Tci Þ. Exploiting assumption 2 above, the continuous velocity can be found by: uc ¼

m_ c ð1 aÞqc A

ð3Þ

where m_ c , A, qc and a denote the continuous phase mass flow rate, the column area, the continuous phase density and the holdup ratio respectively. Using log-mean temperature difference formula and exploiting the velocity definition u ¼ Zt , the column height which appears implicitly in the volume in Eq. (1) could be written as: Z¼

m_ c t ð1 aÞqc A

ð4Þ

Equations (1) to (4) can be implemented to obtain the volumetric heat transfer coefficient expression as: 3 2 ðTdi Tco Þ þ mm__ dc Chfgpc Cpc ð1 aÞqc 5 Uv ¼ ð5Þ ln4 t ðTdi Tco Þ In accordance with Eq. (5), the volumetric heat transfer coefficient calculations, in addition to the other parameters,

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rely on the initial vapour temperature and the outlet continuous phase temperature which are practically available. Figures 2, 3 and 4 show the variation of the volumetric heat transfer coefficient, with time, for different mass flow rate ratios and at a constant dispersed phase initial temperature. A sharp decline of Uv is noted after a short time from the start of the column operation, where the


Fig. 5 Volumetric heat transfer coefficient versus mass flow rate ratio

Fig. 7 Volumetric heat transfer coefficient versus continuous phase mass flow rate

observed which confirms our previous conclusion and is clearly shown in Fig. 6, regarding the latent heat transfer dominating process. On the contrary, and irrespective of the initial dispersed phase temperature, an increase of the continuous phase mass flow rate results in a decrease in Uv (Fig. 7) due to the simple energy balance, where a high continuous phase amount normally requires a high energy source for heating up.

4 Conclusions Fig. 6 Volumetric heat transfer coefficient versus initial dispersed phase temperature

maximum Uv value is recorded and this then dropped to a nearly constant value at t = 100 s. This might result from the intensive condensation during the beginning of vapour injection, where the driving temperature force is at its maximum; consequently a lot of heat is absorbed by the continuous phase, which has been confirmed previously [17, 18]. In addition, the Uv is noted to be increase with an increase in the mass flow rate ratio as shown in (2–4) and this effect appreciably appears at the start of a direct contact condensation process and diminishes with time. It is apparent that a high heating media (vapour) mass flow rate or a low cooling media (continuous phase) mass flow rate, entails the provision of more heat during the direct contact process and in consequence more heat is absorbed by the cooling phase. This, of course, leads to increase the volumetric heat transfer coefficient as shown by Eq. (1). However, the mass flow rate ratio impact on Uv is confirmed clearly in Fig. 5, for different initial dispersed phase temperatures. No significant effect of the initial dispersed temperature on Uv value can be

An experimental investigation for the prediction of the time dependent volumetric heat transfer coefficient of a bubbles type, three phase direct contact condenser has been carried out. In accordance with the experimental results, the following conclusions can be made: •

•

•

The volumetric heat transfer coefficient decreases with time, and a sharp decrease is observed at the beginning of the operation. Contrary to the continuous phase, the mass flow rate ratio is positively affected by the volumetric heat transfer coefficient. No considerable effect of the dispersed phase initial temperatures on the volumetric heat transfer coefficient is shown.

References 1. Sideman S, Hirsch G (1965) Direct contact heat transfer with change of phase: condensation of single vapour bubbles in an immiscible liquid medium. Preliminary studies. AIChE J 11:1019–1025 2. Isenberg J, Sideman S (1970) Direct contact heat transfer with change of phase: bubble condensation in immiscible liquids. Int J Heat Mass Transf 13:997–1011

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Author's personal copy Heat Mass Transfer 3. Moalem D, Sideman S (1973) The effect of motion on bubble collapse. Int J Heat Mass Transf 16:2321–2329 4. Higeta K, Mori YH, Komotori K (1979) Condensation of a single vapor bubble rising in another immiscible liquid. In: Heat transfer-San Diego, AI ChE symposiums series, vol 75, pp 256–265 5. Jacobs HR, Major BM (1982) The effect of noncondensable gases on bubble condensation in an immiscible liquid. J Heat Transfer 104:487–492 6. Raina G, Wanchoo R, Grover P (1984) Direct contact heat transfer with phase change: motion of evaporating droplets. AIChE J 30:835–837 7. Lerner Y, Kalman H, Letan R (1987) Condensation of an accelerating-decelerating bubble: experimental and phenomenological analysis. J Heat Transfer 109:509–517 8. Lerner Y, Letan R (1990) Dynamics of condensing bubbles at intermediate frequencies of injection. J Heat Transfer 112:825–829 9. Wanchoo RK (1993) Forced convection heat transfer around large two-phase bubbles condensing in an immiscible liquid. Heat Recover Syst CHP 13:441–449 10. Wanchoo RK, Sharma SK, Raina GK (1997) Drag coefficient and velocity of rise of a single collapsing two-phase bubbler. AIChE J 43:1955–1963 11. Kalman H, Mori YH (2002) Experimental analysis of a single vapour bubble condensing in subcooled liquid. Chem Eng J 85:197–206 12. Kalman H (2003) Condensation of bubbles miscible liquids. Int J Heat Mass Transf 46:3451–3463 13. Kalman H (2006) Condensation of a bubble train in immiscible liquids. Int J Heat Mass Transf 49:2391–2395 14. Moalem D, Sideman S, Orell A, Hetsroni G (1973) Direct contact heat transfer with change of phase: condensation of a bubble train. Int J Heat Mass Transf 16:2305–2319

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15. Sideman S, Moalem D (1974) Direct contact heat exchangers: comparison of counter and co-current condensers. Int J Multiph Flow 1:555–572 16. Mahood HB, Sharif AO, Al-Aibi S, Thorpe R (2013) Heat transfer modelling of two-phase bubbles swarm condensing in three-phase direct-contact condenser. J Therm Sci. doi:10.2298/ TSCI130219015M 17. Mahood HB, Sharif AO, Alaibi S, Hwakis D, Thorpe R (2013) Analytical solution and experimental measurements for temperatures distribution prediction of three-phase direct contact condenser. J Energy 67:538–547 18. Mahood HB, Thorpe RB, Sharif AO (2014) Experimental measurements and theoretical prediction for the transient characteristic of a three-phase direct contact condenser. Sub J Applied Energy 19. Oh S, Revankar ShT (2005) Effect of noncondensable gas in a vertical tube condenser. Nucl Eng Desi 235:1699–1712 20. Monning C, Numrich R (1999) Condensation of vapours of immiscible liquids in the presence of a non-condensable gas. Int J Therm Sci 38:684–694 21. Bontozoglou V, Karabelas AJ (1995) Direct-contact stream condensation with simultaneous noncondensable gas absorption. AIChE J 41:241–250 22. Coban T, Boehm R (1989) Performance of a three-phase, spray column, direct-contact heat exchanger. J Heat Transfer 111:166–172 23. Jacobs HR, Golafshani M (1989) Heuristic evaluation of the governing mode of heat transfer in a liquid–liquid spray column. J Heat Transfer 111:773–779 24. Mahood HB, Sharif AO, Hossini SA, Thorpe R (2013) Analytical modelling of a spray column three-phase direct contact heat exchanger. ISRN Chem Eng. doi:10.1155/2013/457805