
Research Article


Evaluation of Heat and Mass Transfer Coefficients for R134a/DMF Bubble Absorber 

M. Suresh
and
A. Mani



ABSTRACT

The Vapour Absorption Refrigeration System (VARS) has generated renewed interest and is being viewed as one of the alternatives for vapour compression refrigeration due to its potential for waste heat utilization. To improve the efficiency of these systems, it is necessary to study heat and mass transfer processes in absorption system components. The absorber, one of the crucial components in VARS is considered for study. Experimental investigation is carried out to study heat and mass transfer characteristics in a glass absorber. A new combination of R134a/DMF is used as the working fluid to overcome the limitations of well known working pairs, ammoniawater and lithium bromidewater. The effects of parameters viz., gas flow rate, solution initial concentration, solution pressure and solution temperature on absorber performance are analyzed. Heat and mass transfer coefficients evaluated from the experiments are compared with the numerical model and it is found that agreement is good. Heat and mass transfer coefficients increase as the gas flow rate, solution initial concentration and solution temperature increase whereas they decrease as the solution pressure increases. Sherwood number and Nusselt number evaluated from the experimental data are compared with those obtained from the numerical correlations developed earlier by the authors.





Received: October 22, 2010;
Accepted: November 01, 2010;
Published: April 18, 2011


INTRODUCTION
Fulfillment of refrigeration requirements through lowgrade waste heat recovery
is part of the drive for the reduction of electrical energy consumption and
optimal usage of resources. Intensive research has been focused in the absorption
refrigeration technology, since it uses waste heat as energy source. Many environment
friendly fluid combinations have been suggested by number of investigators in
order to overcome some of the limitations of well known working pairs viz.,
ammoniawater and lithium bromidewater for the VARS. Though HCFC refrigerant
R22organic solvent based absorption refrigeration systems have been extensively
studied by Fatouh and Srinivasa Murthy (Fatouh and Murthy,
1995; Fatouh and Murthy, 1996ac),
HCFCs along with CFCs, are also covered by Montreal and other International
Protocols and are being phased out. So environment friendly HFC refrigerant
R134a based VARS are being investigated. Nezu et al.
(2002) examined the possibility of testing R134a as a refrigerant in VARS
with various organic solvents and showed that R134aDMA and R134aDMF systems
are considered attractive as the workingfluid pairs for the absorption refrigeration
system than other R134a/absorbent systems. Yokozeki (2005)
studied the theoretical performance of various refrigerantabsorbent pairs in
a VARS cycle by the use of equations of state. Of these, R134aDMF and DMA systems
exhibit better performance, compared to other R134aabsorbent systems. Also
circulation ratio is less and COP is more for the R134aDMF system compared
to R134aDMA system. Mani (2009) carried out experimental
studies on R134a/DMF based compact vapour absorption refrigeration system with
plate heat exchangers. He reported that this system could be very competitive
for applications ranging from 10 to 10°C, with heat source temperature
in the range of 80 to 90°C and with cooling water as coolant for the absorber
and condenser in the temperature range of 20 to 35°C.
The absorber is considered as one of the crucial components in vapour absorption
refrigeration system. Kang et al. (2000) carried
out an analytical investigation of falling film and bubbles type absorbers and
found that the absorption rate of bubble type absorber is found to be always
higher than that of the falling film mode. Bubble type absorber provides better
heat and mass transfer coefficients, also good wettability and mixing between
the liquid and vapour.
Absorption process is characterized by simultaneous heat and mass transfer
phenomena. These mechanisms, though complicated, influence the system performance
significantly. Elperin and Fominykh (2003) studied the
combined heat and mass transfer mechanisms at all stages of bubble growth and
rise in a bubble absorber, which can be useful in the design calculations of
gasliquid absorbers. Lee et al. (2003) performed
both numerical and experimental analyses in the absorption process of a bubble
absorber. Numerical model in these studies can be used for the optimum design
of absorber. Merrill and PerezBlanco (1997) developed
an analytical model to predict the bubble dynamics in binary subcooled solutions.
This model improves the understanding of bubble absorption dynamics.
Sujatha et al. (1997a, b)
carried out numerical analysis in a vertical tubular bubble absorber working
with R22 as refrigerant and five organic fluids namely DMF, DMA, DMETEG, DMEDEG
and NMP as absorbents. The model is validated by comparing with the results
available in literature. Based on these results, a correlation for mass transfer
coefficient has been suggested for the vertical tubular bubble absorber. Sujatha
et al. (1999) also carried out experimental studies on a vertical
tubular bubble absorber working with R22DMF. The experimental pressure drop,
heat transfer coefficient and mass transfer coefficient are compared with the
results obtained from the numerical model.
Kang et al. (1998) developed a model for bubble
absorber with a plate type heat exchanger by considering the combined heat and
mass transfers analysis in both liquid and vapour regions. All geometric variables
such as distance between the two plates, number of plates and width of the plates
could be selected optimally for the given thermal conditions by the developed
design model for ammoniawater combination.
Staicovici (2000ac) used nonequilibrium
phenomenological theory to evaluate the gasliquid interaction. The design of
bubble absorber, based on nonequilibrium thermodynamics could be suited to
a modern compact plate type construction and offer better absorption efficiency
and minimum pressure loss on the gas side. Suresh and Mani
(2010) developed a numerical model on bubble dynamics, heat and mass characteristics
of R134a/DMF based bubble absorber using phenomenological theory and validated
the model by comparing with the results available in the literature.
Kang et al. (2002a) developed an experimental
correlation of mass transfer coefficient for ammoniawater bubble absorption.
They (Kang et al., 2002b) also developed a correlation for initial bubble
diameter, which can be used to calculate the interfacial area in the design
of ammoniawater bubble absorber. Cerezo et al. (2009)
carried out experimental studies using a plate heat exchanger as absorber
and ammonia water as working fluid. They concluded that increase in pressure,
solution and cooling flow rates positively affects the absorber performance
and increase in the concentration, cooling and solution temperature negatively
affects the absorber performance.
The present experimental work is carried out to evaluate the heat and mass transfer coefficients and study the effect of parameters viz., gas flow rate, solution initial concentration, solution pressure and solution temperature on the performance of R134a/DMF bubble absorber. MATERIALS AND METHODS Schematic diagram of the experimental setup has been illustrated in Fig. 1. The setup consists of a glass bubble absorber, strong and weak solution tanks, solution pump, cooling water thermostat, instrumentation and valves. The glass absorber consists of two concentric tubes. DMF solution is pumped from weak solution tank, through the bottom of inner tube by a solution pump. R134a gas is supplied from a high pressure cylinder through a mass flow controller unit and injected through a nozzle installed at the bottom of inner tube. Strong DMF solution is collected in the strong solution tank at the top of the absorber. Cooling water is supplied by cooling water simulator, through absorber annulus counter flow to the solution and gas. Cooling water simulator consists of a R22 based Vapour Compression Refrigeration (VCR) circuit of 3.4 TR capacity, a cooling water tank insulated with Expanded Polyethylene (EPE) sheets, electric heaters, pump, flow meter, PT100 sensor, PID temperature controller, contactor, piping and valves. VCR circuit consists of a hermetically sealed reciprocating compressor, an air cooled condenser, a thermostatic expansion valve and cooling coil.
The location of various temperature sensors, pressure sensors, flow meters
and valves are indicated in Fig. 1. All these measuring instruments
are precalibrated. 12 numbers of copperconstant thermocouples are used as
temperature sensors with a measurement uncertainty up to±0.5°C. 3
numbers of piezoelectric type pressure transducers are used as pressure sensors
with a measurement uncertainty up to ±1.2%. Glass rotameters are used
to measure the flow of solution and cooling water with a measurement uncertainty
up to±2.5%. Mass flow controller unit is used to measure the volume flow
rate of gas with a measurement uncertainty of±1%. An online density meter
is used to measure the density of strong and weak solutions with a measurement
uncertainty of ±0.1%. Concentrations of strong and weak solutions are
evaluated from the measured density values using HBT (HankinsonBrobstThomson)
equation used by Reid et al. (1989).
 Fig. 1: 
Schematic diagram of glass bubble absorber experimental setup 
Readings from all these instruments and sensors are monitored continuously
by connecting them to a data acquisition system and a computer.
Initially DMF solution is charged into the weak solution tank. Then it is pumped through the inner tube of glass absorber by solution pump. Solution is collected in the strong solution tank at the top of the absorber. Solution is returned to the weak solution tank through a needle valve, after completion of one run of experiment. Initial concentration of the solution at the absorber inlet is measured by online density meter. Solution inlet pressure and temperature are monitored continuously and kept constant. Cooling water is allowed at a constant flow rate, through the outer tube of the glass absorber, counter flow to the solution. Then R134a gas is injected from a high pressure cylinder through the nozzle at the bottom of the absorber. Gas flow rate is accurately measured by the mass flow controller unit. Gas temperature is maintained constant. During this process, by keeping solution flow rate constant, gas flow rate is varied and measured using DC power supply unit connected to mass flow controller. All the parameters viz., solution inlet and outlet pressure, temperature and concentration, solution flow rate, gas flow rate, pressure and temperature, cooling water flow rate, inlet and outlet temperature are monitored and recorded in the computer using data acquisition unit. During next run of experiment, by keeping solution, gas and water flow rates constant, solution initial concentration is increased by injecting R134a gas through a charging line in the weak solution tank and monitored by the online density meter. All readings are monitored and noted. This process is repeated for various solution pressures and solution initial concentrations. In another run, the solution inlet temperature is varied by varying the cooling water temperature and keeping all other parameters constant. All readings are monitored and noted. RESULTS AND DISCUSSION
Experimentation was conducted by varying the operating parameters viz., gas
flow rate from 0.5 to 2.5 lpm, solution pressure from 120 kPa to 400 kPa, solution
initial concentration from 0.01 to 0.2 kg^{1} and solution temperature
from 20 to 30°C. Solution flow rate and cooling water flow rate was maintained
at 50 lph. Experimental results are compared with the numerical model for the
bubble absorber developed earlier by Suresh and Mani (2010)
to study heat and mass transfer characteristics using phenomenological theory,
which was validated with the literature results. Experimental values of volumetric
mass transfer coefficient, heat transfer coefficient, Sherwood number and Nusselt
number are calculated using the equations given in Appendix A.
Figure 2 compares the experimental volumetric mass transfer
coefficient with that of the numerical model for various gas flow rates. The
agreement is good within ±14% deviation. Mass transfer coefficient increases
as the gas flow rate increases due to increase in absorption rate at high gas
flow rates. Figure 3 compares the experimental heat transfer
coefficient with that of the numerical model for various gas flow rates. The
agreement is good within ±10% deviation. Heat transfer coefficient increases
as the gas flow rate increases due to increase in coupled heat transfer rate
at high absorption rates.
 Fig. 2: 
Effect of gas flow rate on volumetric mass transfer coefficient,
gas inlet pressure = 650 kPa, gas inlet temperature = 32°C, solution
flow rate = 50 lph, solution inelt pressure = 120 kPa, solution inlet temperature
= 30°C, solution inlet concentration = 0.01 kg kg^{1}, Cooling
water flow rate = 50 lph 
 Fig. 3: 
Effect of gas flow rate on heat transfer coefficient, gas
inlet pressure = 650 kPa, gas inlet temperature = 32°C, solution flow
rate = 50 lph, solution inlet pressure = 120 kPa, solution inlet temperature
30°C, solution inlet concentration = 0.01 kgkg^{1}, cooling
water flow rate = 50 lph 
Figure 4 compares the experimental volumetric mass transfer
coefficient with that of the numerical model for various solution pressures.
The agreement is good within ±15% deviation. Mass transfer coefficient
decreases as the solution pressure increases. The reason is that though the
absorption rate is almost constant with respect to increase in solution pressure,
the Log Mean Concentration Difference (LMCD) increases as solution pressure
increases resulting in lower mass transfer coefficients.
 Fig. 4: 
Effect of solution pressure on volumetric mass transfer coefficient
, gas flow rate = 2 lpm, gas inlet pressure = 650 kPa, gas inlet temperature
= 32°C, solution flow rate = 50 lph, solution inlet temperature = 0.01
kg kg^{1}, cooling water flow rate = 50 lph 
 Fig. 5: 
Effect of solution pressure on heat transfer coefficient,
gas flow rate = 2 lpm, gas inlet pressure = 650 kPa, gas inlet temperature
= 32°C, solution flow rate = 50 lph, solution inlet temperature = 30
°C, solution inlet concentration = 0.01 kg kg^{1}, cooling
water flow rate = 50 lph 
Figure 5 compares the experimental heat transfer coefficient
with that of the numerical model for various solution pressures. The agreement
is good within ±10% deviation. Heat transfer coefficient decreases as
the solution pressure increases. The reason is that though the heat transfer
rate is almost constant with respect to increase in solution pressure, the Log
Mean Temperature Difference (LMTD) increases as solution pressure increases
resulting in lower heat transfer coefficients.
 Fig. 6: 
Effect of solution initial concentration on volumetric mass
transfer coefficient, gas flow rate = 2 lpm, gas inlet pressure = 650 kPa,
gas inlet temperature = 32°C, solution flow rate = 50 lph, solution
inlet pressure = 120 kPa, solution inlet temperature = 30°C, cooling
water flow rate = 50 lph 
 Fig. 7: 
Effect of solution initial concentration on heat transfer
coefficient, gas flow rate = 2 lpm, gas inlet pressure = 650 kPa, gas inlet
temperature = 32°C, solution flow rate = 50 lph, solution inlet pressure
= 120 kPa, solution inlet temperature = 30°C, coolingwater flow rate
= 50 lph 
Figure 6 compares the experimental volumetric mass transfer
coefficient with that of the numerical model for various solution initial concentrations.
The agreement is good within ±10% deviation. Mass transfer coefficient
increases as the solution initial concentration increases. Though the absorption
rate is almost constant with respect to increase in solution initial concentration,
the LMCD decreases as solution inlet concentration increases resulting in higher
mass transfer coefficients.
Figure 7 compares the experimental heat transfer coefficient
with that of the numerical model for various solution initial concentrations.
 Fig. 8: 
Effect of solution inlet temperature on volumetric mass transfer
coefficient, gas flow rate = 2 lpm, gas inlet pressure = 650 kPa, gas inlet
temperature = 32°C, solution flow rate = 50 lph, solution inlet pressure
= 120 kPa, solution inlet concentration = 0.01 kg kg^{1}, coolingwater
flow rate = 50 lph 
 Fig. 9: 
Effect of solution inlet temperature on heat transfer coefficient,
gas flow rate = 2 lpm, gas inlet pressure = 650 kPa, gas inlet temperature
= 32°C, solution flow rate = 50 lph, solution inlet pressure = 120 kPa,
solution inlet concentration = 0.01 kgkg^{1}, coolingwater flow
rate = 50 lph 
The agreement is good within ±10% deviation. Heat transfer coefficient
increases as the solution initial concentration increases. Though the heat transfer
rate is almost constant with respect to increase in solution initial concentration,
the LMTD decreases as solution inlet concentration increases resulting in higher
heat transfer coefficients.
Figure 8 compares the experimental volumetric mass transfer
coefficient with that of the numerical model for various solution inlet temperatures.
The agreement is good within ±10% deviation. Mass transfer coefficient
increases as the solution inlet temperature increases.
 Fig. 10: 
Variation of Sherwood number based on experiments with Sherwood
number based on numerical correlation 
 Fig. 11: 
Variation of Nusselt number based on experiments with Nusselt
number based on numerical correlation 
Though the absorption rate decreases with respect to increase in solution inlet
temperature, the LMCD also decreases as the solution inlet temperature increases
resulting in higher mass transfer coefficients.
Figure 9 compares the experimental heat transfer coefficient with that of the numerical model for various solution inlet temperatures. The agreement is good within±7% deviation. Heat transfer coefficient increases as the solution inlet temperature increases. Though the heat transfer rate decreases with respect to increase in solution inlet temperature, the LMTD also decreases as the solution inlet temperature increases resulting in higher heat transfer coefficients.
Experimental values of Sherwood number and Nusselt number are evaluated and
plotted against those obtained from the following correlations developed by
the authors from their numerical model (Suresh and Mani,
2010).
Figure 10 illustrates the comparison between experimental
and predicted Sherwood numbers and Fig. 11 illustrates the
comparison between experimental and predicted Nusselt numbers.
The agreement is fair within ±25% deviation. The deviation could be due to the combined effect of assumptions made in the theoretical model and also the inaccuracies in the measurements. CONCLUSIONS Experimental investigations have been carried out on a glass bubble absorber to study heat and mass transfer characteristics of R134a in Dimethyl Formamide (DMF) and the effect of parameters viz., gas flow rate, solution initial concentration, solution pressure and solution temperature on absorber performance. Heat and mass transfer coefficients determined from the experiments are compared with the numerical model and it is found that the agreement is good. Experimental values of Sherwood number and Nusselt number are evaluated and compared with those obtained from the correlations developed by the authors from their numerical model.
The following conclusions are drawn from the present study.
• 
Volumetric mass transfer coefficient and heat transfer coefficient
determined from the experiments are compared with the numerical model for
various gas flow rates, solution pressures, solution initial concentrations
and solution inlet temperatures and the agreement is generally good with
maximum deviation of ±15% 
• 
Volumetric mass transfer coefficient and heat transfer coefficients increase
as the gas flow rate, solution initial concentration and solution inlet
temperature increase 
• 
Volumetric mass transfer coefficient and heat transfer coefficient decrease
as the solution pressure increases 
• 
Experimental values of Sherwood number and Nusselt number are evaluated
and plotted against those obtained from the correlations developed by the
authors from their numerical model. The agreement is generally fair within
±25% deviation 
NOTATIONS
Units, abbreviations and symbols
A 
= 
Absorber crosssection area ( m^{2}) 
Cp 
= 
Specific heat capacity (kJ kg^{1}K^{1}) 
D 
= 
Absorber diameter (m) 
Dc 
= 
Diffusion coefficient (m^{2 }sec^{1}) 
h 
= 
Heat transfer coefficient (W m^{2}K^{1}) 
K 
= 
Thermal conductivity (W m^{1}K^{1}) 
L 
= 
Absorber length (m) 
M 
= 
Volumetric mass transfer coefficient (kg m^{3 }sec^{1}) 
m 
= 
Mass flow rate (kg sec^{1}) 
NU 
= 
Nusselt number 
Pr 
= 
Prandtl number 
p 
= 
Solution pressure (bar) (kPa) 
p_{0} 
= 
Atmospheric pressure (bar) (kPa) 
Q 
= 
Heat transfer rate (W) (kW) 
Re 
= 
Reynolds number 
Sc 
= 
Schmidt number 
Sh 
= 
Sherwood number 
X 
= 
Liquid mass fraction (kg kg^{1}) 
T 
= 
Solution temperature (K) 
U 
= 
Overall heat transfer coefficient (Wm ^{2}K^{1}) 
V 
= 
Volumetric flow rate (m^{3 }sec^{1}) 
Subscripts
1 
= 
Inner 
2 
= 
Outer 
eq 
= 
Equilibrium 
g 
= 
Gas 
1 
= 
Liquid 
in 
= 
Inlet 
out 
= 
Outlet 
v 
= 
Volumetric 
w 
= 
Water 
ws 
= 
Weak solution 
Greek symbols
μ 
= 
Dynamic viscosity (Pas) 
γ 
= 
Kinematic viscosity (m^{2}s^{1}) 
ρ 
= 
Density, kgm^{3} 
α 
= 
Thermal diffusivity, m^{2}sec^{1} 
θ 
= 
nondimensional temperature 
APPENDIXA
θ = nondimensional solution temperature =

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