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Energy 44 (2012) 1005e1016

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Energy

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Thermodynamic analysis of an absorption refrigeration system with ionic-liquid/refrigerant mixture as a working fluid

Yoon Jo Kim a,*, Sarah Kim b, Yogendra K. Joshi c, Andrei G. Fedorov c, Paul A. Kohl b

aDepartment of Mechanical Engineering, Washington State University Vancouver , 14204 NE Salmon Creek Avenue, Vancouver, WA 98686-9600, USAb School of Chemical and Biomolecular Engineering, Georgia Institute of Technology, Atlanta, GA 30332, USAc The George Woodruff School of Mechanical Engineering, Georgia Institute of Technology, Atlanta, GA 30332, USA

a r t i c l e i n f o

Article history:Received 6 December 2011Received in revised form17 April 2012Accepted 20 April 2012Available online 26 May 2012

Keywords:Ionic liquidAbsorption systemWaste-heatElectronics cooling

* Corresponding author. Tel.: þ1 360 546 9184; faxE-mail address: yoonjo.kim@wsu.edu (Y.J. Kim).

0360-5442/$ e see front matter � 2012 Elsevier Ltd.doi:10.1016/j.energy.2012.04.048

a b s t r a c t

Thermodynamics of an ionic-liquid (IL) based absorption refrigeration system has been numericallyanalyzed. It provides an alternative to the normally toxic working fluids, such as the ammonia inconventional absorption systems. The use of ILs also eliminates crystallization and metal-compatibilityproblems of the water/LiBr system. Mixtures of refrigerants and imidazolium-based ILs are theoreti-cally explored as the working fluid pairs in a miniature absorption refrigeration system, so as to utilizewaste-heat to power a refrigeration/heat pump system for electronics cooling. A non-random two-liquid(NRTL) model was built and used to predict the solubility of the mixtures. Saturation temperatures at theevaporator and condenser were set at 25 �C and 50 �C, respectively, with the power dissipation of 100 W.Water in combination with [emim][BF4] (1-ethyl-3-methylimidazolium tetrafluoroborate) gave thehighest coefficient of performance (COP) around 0.9. The refrigerant/IL compatibility indicated by thecirculation ratio, alkyl chain length of the IL, and thermodynamic properties of the refrigerants, such aslatent heat of evaporation were proven to be important factors in determining the performance of theabsorption system. The negative effect of high viscosity was mitigated by dilution of the IL with therefrigerant and the use of slightly larger microfluidic channel heat exchangers.

� 2012 Elsevier Ltd. All rights reserved.

1. Introduction

Recent advances in semiconductor technologies have led to anincrease in power density for high performance chips, such asmicroprocessors. According to the International Technology Road-map for Semiconductors (ITRS), these chips are expected to dissi-pate an average heat flux as high as 75 W/cm2, with the maximumjunction temperature not exceeding 85 �C, in 2012, while in 2024the numbers are more challenging, 120 W/cm2 and 70 �C, respec-tively [1]. Conventional chip packaging solutions, which use air-cooling, face difficulties in dissipating such high heat fluxes in thelimited space allocated to thermal management.

A variety of novel thermal solutions for electronic cooling havebeen reported, including thermosyphons [2], loop heat pipes [3],electro-osmotic pumping [4], stacked micro-channels [5],impinging jets [6], thermoelectric micro-coolers [7], vaporcompression refrigeration [8], and absorption based refrigerationsystems [9,10]. These cooling systems can be categorized into

: þ1 360 546 9438.

All rights reserved.

passive and active systems. Passive cooling systems utilize capillaryor gravitational buoyancy forces to circulate the workingfluid, while active cooling systems are driven by a pump orcompressor for higher cooling capacity and improved performance.Also, an active system may offer further increases in power dissi-pation by insertion of a negative thermal resistance into heat flowpath [8].

In this study, ionic-liquids (ILs), which are salts in a liquid stateusually with organic cations and inorganic anions, are used as anabsorbent fluid in a miniature absorption refrigeration systemdesigned for current electronic cooling requirements (i.e., bench-mark 100 W/cm2 power dissipation and 85 �C chip temperature).ILs have the character of molten salts, which are moisture and airstable at room temperature. Most ILs are thermally stable totemperatures well above those in vapor compression refrigerationsystems, >400 K [11e15]. However, in the case of a prolongedexposure to elevated temperatures, the effective decompositiontemperature could be lower. Blake et al. [15] reported that the half-life of [bmim][PF6] is only 138 days at 573 K, while at 423 K (thehighest operating temperature of an absorption system), it couldbe more than 10 years. Since the degradation products of ILs areoften volatile compounds [12], the operating temperature needs to

Table 1Saturation pressures and latent heats of various refrigerants at the evaporator andcondenser temperatures (Te¼ 25 �C, Tc¼ 50 �C) and their ozone depletion potentialsand global warming potentials [33].

Refrigerants Psat,e [kPa] Psat,c[kPa] Dhlv,e [kJ/kg] Dhlv,c [kJ/kg] ODPa GWPb

R114 214.4 446.9 128.07 117.53 1 3.9R124 382.7 775.8 146.56 130.56 0.02 620R125 1376.7 2533.2 110.4 76.4 0 3400R134a 665.3 1317.9 177.8 151.8 0 1300R143a 1262.3 2307.9 159.6 118.9 0 4300R152a 596.4 1177.4 279.4 245.4 0 120R32 1689.6 3141.2 270.9 209.6 0 620Water 3.170 12.35 2441.7 2382.0 0 e

a The ODP of all other refrigerants are compared to R11 [30].b GWP is a relative scale which compares the greenhouse gas to carbon dioxide

where GWP by definition is 1 [30].

Condenser

Expansion

device

Expansion

device

Liquid

pump

Evaporator

Solution HX

Absorber

Desorber

3

4

1

2

5

6

7

10

9

8

Weak-IL

Solution

Strong-IL

Solution

Refrigerant

Heat absorptionfrom source (Qe)

Heat rejectionto ambient (Qa)

Heat rejectionto ambient (Qc)

Heat absorptionfrom waste-heat(Qd)

Chemical

compressor

Fig. 1. Schematic diagram of an absorption refrigeration system using IL/refrigerantmixture as a working fluid pair.

Y.J. Kim et al. / Energy 44 (2012) 1005e10161006

be kept within acceptable limits. Due to the toxic and flammablenature of volatile organic compounds (VOCs) [16], ILs with thenegligible vapor pressure have been considered as a possiblereplacements of solvents for organic synthesis [17], biphasiccatalysis, separation and extraction processes [18], and dissolutionof biomaterials [19]. Note that although ILs are ‘non-flammable’,some of ILs could be ‘combustible’ at sufficiently high temperature(near their decomposition temperature) upon exposure to fire orcombustion [20]. Also, Jastorff et al. [21] found the toxicity of thevarious ILs spanned a range of about 1000 � (from least to mosttoxic) based on the toxicological tests of a series of ILs with imi-dazolium cations and [PF6�], [BF4�], [Cl�], [Br�] and tosylate anions.However, even the most toxic of the ILs has about the same toxicityas the least toxic of four common organic solvents (methanol,acetone, acetonitrile, MTBE) that were tested using the sameprotocol [13]. Thus, even though some ILs may not be intrinsically“green”, they can be designed to be environmentally benign, withlarge potential benefits for sustainable chemistry [22]. While thechemistry of ILs and their utilization for chemical processes havebeen reported in many publications, solubility data with othersolvents and thermodynamic properties of mixtures containing ILsand refrigerants are less common [11].

With the above mentioned features of ILs such as tunableproperties, zero vapor pressure, and high thermal stability, ILs arepromising absorbents [23]. In particular, the low volatility of the ILenables easy separation of the volatile working fluid from the IL bythermal stratification with the minimum harmful impacts onenvironment [24]. ILs can provide an alternative to the normallytoxic working fluids used in some absorption systems, such as theammonia/water system. Since many of the ILs have melting pointsbelow the lowest solution temperature in the absorption system(w300 K) [25e29], they also eliminate the crystallization andmetal-compatibility problems of the water/LiBr system. Among theILs considered in this study, the melting points of [emim][BF4] (1-ethyl-3-methylimidazolium tetrafluoroborate) and [bmim][PF6](1-butyl-3-methylimidazolium hexafluorophosphate) are 14.42 �C[28] and 10 �C [29], respectively, which are low enough to be usedin absorption system. Several theoretical works on the absorptionrefrigeration system using ILs as absorbent have been reported[23,24,30e32], and an experimental system was investigated byKim et al. [24].

In this study, the thermodynamic performance of a mini-ature absorption system using refrigerant/IL mixtures as theworking fluid pairs was theoretically investigated. Various mixturesof refrigerants and imidazolium ILs e [emim][Tf2N] (1-ethyl-3-methylimidazoliumbis(trifluoromethanesulfonyl)imide), [emim][BF4](1-ethyl-3-methylimidazoliumtetrafluoroborate), [bmim][BF4](1-butyl-3-methylimidazoliumtetrafluoroborate), [bmim][PF6] (1-butyl-3-methylimidazoliumhexafluorophosphate), [hmim][Tf2N] (1-hexyl-3-methylimidazoliumbis(trifluoromethanesulfonyl)imide), [hmim][BF4] (1-hexyl-3-methylimidazoliumtetrafluoroborate), and [hmim][PF6] (1-hexyl-3-methylimidazoliumhexafluorophosphate) e werecompared among each other in terms of system efficiencyand usability of low-grade waste heat as the heat source of thesystem. Several HFC refrigerants, R125 (pentafluoroethane),R134a (1,1,1,2-tetrafluoroethane), R143a (1,1,1-trifluoroethane),R152a (1,1-difluoroethane), and R32 (difluoromethane), wereinvestigated as the working fluid in the IL-based refrigerationsystem. Relevant characteristics of refrigerants, including ozonedepletionpotentials (ODP) andglobalwarmingpotentials (GWP)aregiven in Table 1. The performance of an IL with R114 (dichlorote-trafluoroethane) and R124 (chlorotetrafluoroethane) were evalu-ated and compared to hydrofluorocarbon refrigerants.Water, whichis attractive due to its high thermal conductivity and latent heat ofevaporation, was also explored as a refrigerant with IL absorbents.

In all calculations, the operating temperatures of the condenserand evaporator were set at 50 �C and 25 �C, respectively. Thecooling capacity of the evaporator was chosen to be 100 W. Thedegree of superheat at the evaporator outlet and subcooling atthe condenser outlet were set at 5 �C, with respect to saturationtemperatures at the evaporator and condenser pressures. Theoutlet temperatures of the desorber and absorber were treated asadjustable parameters, which can be tuned to find the optimumoperating conditions where the system efficiency reached itsmaximum value.

2. Waste-heat driven absorption system

Fig. 1 shows a schematic diagram of a thermally drivenabsorption refrigeration system using a refrigerant/IL mixture asa working fluid pair. The system consists of an evaporator, anabsorber, a desorber, a condenser, a liquid pump, and an expansiondevice. One of the major advantages of the absorption refrigerationsystem is the utilization of waste heat, which can have temperatureof 90 �C or lower. The use of waste-heat results in a significantreduction in operating cost and justifies the added balance-of-plantfor the absorption refrigeration system. Furthermore, the only

Table 3Adjustable parameters in Equation (4).

Y.J. Kim et al. / Energy 44 (2012) 1005e1016 1007

component of this system with moving mechanical parts is theliquid pump, so that relatively quiet operation is possible with nolubrication needed. In general, the key system performance metric,coefficient of performance (COP), is defined as the heat removed atthe evaporator per total power supplied to the system, Equation (1).

COP ¼ Qe�Wp þ Qd

� (1)

where Wp and Q are pumping work and the heat input/output,respectively. The subscripts “e” and “d” denote the evaporator anddesorber, respectively. Since the heat supply at the desorber is fromwaste heat (i.e., it is essentially free), a more relevant coefficient ofperformance (COPwaste) can be defined by eliminating Qd fromEquation (1), i.e., COPwaste ¼ Qe=Wp.

The cyclic process for the refrigerant loop is the same as that ofa vapor compression system, except the mechanical compressor isreplaced with a ‘chemical compressor’ which consists of anabsorber, liquid pump, solution heat exchanger, desorber andexpansion device. The pressurization process in the chemicalcompressor starts in the absorber, where the refrigerant vapor fromthe evaporator (state point 2) is exothermically absorbed into theweak (refrigerant) solution (state point 10), resulting in strong(refrigerant) solution at state point 5. Once the refrigerant isabsorbed, the solution is pressurized by the liquid pump. Thesolution heat exchanger preheats the strong solution of state point6 to state point 7 using the high temperature weak solution flowfrom the desorber. A high pressure and high temperature super-heated refrigerant vapor is generated in the desorber and therefrigerant is endothermically desorbed from the strong solution.The refrigerant vapor returns to the refrigerant loop. Meanwhile,the mixture solution becomes the weak solution and returns to theabsorber through solution heat exchanger and expansion device insequence, which completes the solution loop or chemicalcompression cycle. Table 2 summarizes the main components andtheir processes in the waste-heat driven absorption refrigerationsystem. The condensation/absorption process at the absorber andvaporization/desorption process at the desorber make it possible touse a low-power-input liquid pump to increase the pressurebetween the condenser and the evaporator. Although the presenceof the absorber and desorber increases the overall system volume,the displacement volume and power consumption for compressionof the liquid are much smaller than those for vapor compression.Also, the slight modification of the absorption refrigeration systemcan be used for power cycles, e.g., the Kalina cycle [34e36] anda heat transformer [37].

3. System analysis

The correlations of the refrigerant/IL mixture properties devel-oped based on non-random two-liquid (NRTL) activity coefficient

Table 2Main components and the processes of a waste-heat driven absorption refrigerationsystem for electronic cooling applications.

Components States Process

Evaporator 1 / 2 Heat absorption from anelectronic device

Absorber 2, 10 / 5 Refrigerant absorption/condensation into IL

Pump 5 / 6 Isentropic pressurizationSolution HX 6 / 7, 8 / 9 Regenerative pre-heatingDesorber 7 / 8, 3 Refrigerant desorption /

vaporization from ILCondenser 3 / 4 Heat rejection to ambientExpansion device 4 / 1, 9 / 10 Isenthalpic expansion

model by Shiflett and Yokozeki [38] are used in this study, Equa-tions (2) and (3).

lngm ¼ x2n

"snm

�Fnm

xm þ xnFnm

�2

þ snmFnmðxn þ xmFmnÞ2

#(2)

lngn ¼ x2m

"smn

�Fmn

xn þ xmFmn

�2þ smnFmn

ðxm þ xnFnmÞ2#

(3)

where x is the liquid-phase mole fraction and g represents theactivity coefficients. The subscripts “m” and “n” denote the refrig-erant and the IL, respectively. F is an adjustable binary interactionparameter and s are defined in Equations (4) and (5).

Fmnhexpð�usmnÞ and Fnmhexpð�usnmÞ (4)

smnhs0mn þ s1mnT

and snmhs0nm þ s1nmT

(5)

where T is absolute temperature, respectively, and u ¼ 0:2. Theparameters, s0mn, s

1mn, s

0nm, and s1nm, have been determined based on

literature vaporeliquid equilibrium (VLE) data, as shown in Table 3.It is assumed that upon mixing, the refrigerant/IL mixtures are onlyphysically interacting. To the best knowledge of the authors, nochemical reactions between the substances have been reported.The predicted mole fractions using the NRTL model are comparedwith the measured mole fraction data in references [38e42] inFig. 2, which shows good agreement between the measured andpredicted data. The correlations based on group contributionmethods were used to evaluate the viscosity [43], specific heat [44]and density [45] of the ILs. The REFPROP 6.0 software [46], devel-oped by the National Institute of Standards and Technology, hasbeen used to calculate the thermodynamic and transport propertiesof the refrigerants. It implements three models for the thermody-namic properties of pure fluids: (i) equations of state explicit in theHelmholtz energy, (ii) the modified BenedicteWebbeRubin equa-tion of state, and (iii) an extended corresponding states (ECS)model[46].

The specific enthalpy, hIL, and specific entropy, sIL, of ILs werecalculated using Equations (6) and (7) [47].

hIL ¼ZTT0

cp;ILdT þZPP0

vILdP þ h0 (6)

sIL ¼ZTT0

cp;ILT

dT þ s0 (7)

Working fluid pair (1)/(2) s0mn [-] s1mn [K] s0nm [-] s1nm [K]

R134a/[emim][Tf2N] [39] 6.6710 �716.04 �0.8502 �262.85R134a/[bmim][PF6] [39] 1.2510 411.45 0.57596 �406.43R134a/[hmim][Tf2N] [39] 13.186 �2904.5 �5.3330 1128.9R134a/[hmim][BF4] [39] 7.5975 �1176.7 �0.26344 �275.97R134a/[hmim][PF6] [39] 11.718 �2397.7 �3.5270 688.02R32/[bmim][BF4] [38] 0.93194 553.36 0.36807 �585.91R32/[bmim][PF6] [38] 6.1356 �1194.6 �1.9069 122.26R125/[bmim][PF6] [38] 2.4582 48.172 1.6394 �563.00R143a/[bmim][PF6] [38] 5.2848 �18.277 �0.11097 �68.435R152a/[bmim][PF6] [38] 6.5351 �871.79 �0.02731 �409.50R114/[emim][Tf2N] [40] 2.0631 1549.9 �1.0567 390.84R124/[emim][Tf2N] [40] 0.11312 1210.1 1.4565 �1040.3Water/[emim][Tf2N] [41] 1.7388 1074.7 �4.9829 1369.1Water/[emim][BF4] [42] 17.253 �6445.4 �12.650 4426.8

0

0.2

0.4

0.6

0.8

1

0 0.2 0.4 0.6 0.8 1

Mo

le

fra

ctio

n p

re

dic

te

d

Mole fraction measured

-5%

5%

Fig. 2. Comparison between the measured [38e42] and the predicted mole fractionsusing the NRTL model.

Y.J. Kim et al. / Energy 44 (2012) 1005e10161008

where cp,IL and vIL are the specific heat and specific volume of the IL,and P is the pressure. The IIR reference state was adopted;h0 ¼ 200 kJ=kg and s0 ¼ 1 kJ=kgK for the saturated liquid at T0 ¼0oC ðP0 ¼ PsatðT0ÞÞ. Then, the mixture enthalpy, H, and entropy, S,were calculated from Equation (8).

H ¼ Hid þ HE and S ¼ Sid þ SE (8)

where the ideal solution properties of enthalpy, Hid,and entropy,Sid, can be obtained from Equation (9).

Hid ¼ xmHm þ xnHn and Sid

¼ ðxmSm þ xnSnÞ � Rðxmlnxm þ xnlnxnÞ (9)

The excess enthalpy, HE, and excess entropy, SE, can be calcu-lated from Equations (10) and (11) [38].

HE ¼ �RT2�xm

�vlngmvT

�P;x

þ xn

�vlngnvT

�P;x

�(10)

TSE ¼ HE � GE (11)

where R is the universal gas constant and The excess Gibbs energy,GE, can be calculated from Equation (12).

GE

RT¼ xmlngm þ xnlngn (12)

Since the feasibility and the compatibility of ILs as an absorbentin an absorption system, in combination with refrigerants are theprincipal foci in this study, the performance of the system withrespect to the refrigerant/IL mixtures is thermodynamically eval-uated as a function of operating conditions.

Given the degree of subcooling at the condenser outlet, degreeof superheat at the evaporator outlet, and the saturation temper-atures for the condenser and evaporator, the condenser and evap-orator outlet states (states 4 and 2) can be determined. Isenthalpicthrottling through both the refrigerant and solution expansiondevices is assumed resulting in Equations (13) and (14).

h1 ¼ h4 (13)

h10 ¼ h9 (14)

The subscript numbers denote the states indicated in Fig. 1. Theenergy balance at the evaporator yields the refrigerant mass flowrate, Equation (15).

_mr ¼ ðh2 � h1ÞQe

(15)

The maximum desorber outlet temperature, T8, and theabsorber outlet temperature, T5, are adjustable parameters whichdetermine the refrigerant mass fractions of the strong ðxms ¼ xm5 Þand weak solutions ðxmw ¼ xm8 Þ. The solution mass flow rates can bedetermined from Equations (16) and (17).

_ms ¼ _mr

�1� xmw

��xms � xmw

� (16)

_mw ¼ _ms � _mr (17)

The subscripts “s” and “w” represent the strong- and weak-refrigerant solutions, respectively. Also, by assuming thermalequilibrium at the desorber outlet, T3 ¼ T8, the heat transfer rate atthe condenser can be calculated from Equation (18).

Qc ¼ _mrðh3 � h4Þ (18)

The pumping process is assumed to be isentropic, yieldingEquation (19).

s6 ¼ s5 (19)

The energy balances at the solution heat exchanger, assumingno heat losses to the environment, is expressed by Equation (20).

Qshx ¼ _msðh7 � h6Þ ¼ _mwðh8 � h9Þ (20)

where Qshx is the heat transfer from the high temperature weaksolution to the low temperature strong solution in the heatexchanger. The energy balances at the desorber and absorber isgiven by Equations (21) and (22).

Qd ¼ _mwh8 þ _mrh3 � _msh7 (21)

Qa ¼ _mwh10 þ _mrh2 � _msh5 (22)

where Qa is the heat removed from the refrigerant/IL mixture at theabsorber.

The viscosity of ILs are often relatively high; for example, for[bmim][PF6] at 294 K and atmospheric pressure it is 376 mPa s [48].Thus, the IL flow in micro-channels used in the absorber anddesorber may create large pressure drops, which can affect thesystem performance. The average pressure drops through themicrofluidic channel heat exchangers were evaluated using a two-phase pressure drop equation, Equation (23).

��dPdz

�¼

"2flG2

mð1�xvÞdhrl

#f2l þG2

mddz

"ðxvÞ2εrv

þð1�xvÞ2ð1� εÞrl

#(23)

where dh is the hydraulic diameter of channel, and f, G, xv, r, and ε

are the liquid-phase fanning friction factor, mass flux, vapor quality,density, and void fraction, respectively. z is the axial directioncoordinate along the channel length. Subscripts “l” and “v” repre-sent liquid and vapor phase, respectively. In the two-phase multi-plier correlation of Lockhart and Martinelli [49], fl is incorporatedwith the C value proposed by Lee and Mudawar [50], Equations(24)e(26).

f2l ¼ 1þ C

Xþ 1X2 (24)

C ¼ 2:16Re0:047lo We0:6lo

�laminar liquid� laminarvapor

(25)

0.5a

Y.J. Kim et al. / Energy 44 (2012) 1005e1016 1009

C ¼ 1:45Re0:25lo We0:23lo

�laminar liquid�turbulentvapor

(26)

0

0.1

0.2

0.3

0.4

60 70 80 90 100

CO

P(Q

e/(Q

d+

Wp))

Desorber outlet temperature [oC]

R32

R152a

T8,min ∼ 61.5oC

0.2

0.3

0.4

0.5

P(Q

e/(Q

d+

Wp))

T8,min ∼ 61.5oC

b

where Relo andWelo are liquid-only Reynolds andWeber numbers,respectively. The Martinelli parameter, X, and the single phaseempirical correlation of fanning friction factor for laminar flow ina rectangular channel by Shah and London [51] are expressed byEquations (27) and (28).

X ¼�mlmv

�0:5�1� xv

xv

�0:5�rvrl

�0:5

(27)

fRe ¼24�1� 1:3553bþ 1:9467b2 � 1:7012b3 þ 0:9564b4

� 0:2537b5

ð28Þ

where m is the viscosity and b is the aspect ratio of the channel.Also, the void fraction model of Zivi [52] is adopted in this study.Microchannel structures for the absorber and the desorber areadopted as a baseline, due to their ability to yield high heat andmass transfer rates and to minimize transport limitations on theperformance; yet, the extent of potentially negative effect of thehigh viscosity of ILs on the pressure drop in microchannel heat/mass exchangers and, in turn, the system performance is criticallyassessed. The dimensions (length � width) of the evaporator andcondenser are 2 � 2 cm, and 3 � 3 cm, respectively. The dimen-sions of the absorber and the desorber are 8 � 8 cm. The micro-fluidic channel cross-sectional area for the heat exchangers is1 � 1 mm.

0

0.1

60 70 80 90 100

CO

Desorber outlet temperature [o

C]

water

0

0.2

0.4

0.6

0.8

1

60 70 80 90 100

CO

P (Q

e/(Q

d+

Wp))

Desorber outlet temperature [o

C]

R32/[bmim][BF4]

T8,min ∼ 61.5oC

R134a/[hmim][BF4]

c

Fig. 3. COPs of the absorption system with respect to the desorber outlet temperatureusing different working fluid pairs: (a) refrigerant/[bmim][PF6] and R134a/[hmim][PF6]; (b) refrigerant/[emim][Tf2N] and R134a/[hmim][Tf2N]; (c) R134a/[hmim][BF4],R32/[bmim][BF4] and water/[emim][BF4].

4. Results & discussion

4.1. Effect of the desorber outlet temperature

Fig. 3 shows the COP variation of the absorption system withrespect to desorber outlet temperature for several working fluidpairs. Fig. 3 shows a relatively sharp rise in COP up to w80 �C fol-lowed by a more constant value as the desorber outlet temperaturerises beyond 80 �C. The variation of the desorber outlet tempera-ture affects the concentration of refrigerant in the weak solution,which can be approximated by Equation (29) [47].

xwwPc

gwPsatðT8Þ(29)

where Pc is the vapor pressure at the condenser and PsatðT8Þ is thesaturation vapor pressure at T8. Subscripts “c” and “sat” representthe condenser and saturation, respectively. As the temperatureincreases, the refrigerant mole fraction decreases. The flow rate ofthe solutionwill then vary correspondingly. The circulation ratio, a,is the amount of solution flow rate of the refrigerant/IL mixtureneeded to achieve compression for a given mass flow rate of therefrigerant, as given by Equation (30).

ah_ms_mr

¼�1� xmw

��xms � xmw

� (30)

The circulation ratio increases with mass fraction in the weaksolution, at fixed refrigerant mass fraction in the strong solution.Fig. 4 shows the circulation ratio variation with respect to thedesorber outlet temperature. The desorber outlet temperature risereduces the circulation ratio due to a decrease in the mass fractionof the weak solution. Equation (20) can be rewritten to yieldEquation (31).

Qdmr

¼ aðh8 � h7Þ � h8 þ h3 (31a)

Qdmr

¼ h8ða� 1Þ þ h3 � ah7 (31b)

Table 4Refrigerant mass fractions in the strong solution.

Working fluid pair (1)/(2) xms [e]

R143a/[bmim][PF6] 0.0787R125/[bmim][PF6] 0.2210R134a/[hmim][PF6] 0.2563R152a/[bmim][PF6] 0.2286R134a/[bmim][PF6] 0.3239R32/[bmim][PF6] 0.4362R114/[emim][Tf2N] 0.0599Water/[emim][Tf2N] 0.0091R134a/[hmim][Tf2N] 0.2910R134a/[emim][Tf2N] 0.2676R124/[emim][Tf2N] 0.3972R134a/[hmim][BF4] 0.2791R32/[bmim][BF4] 0.4487Water/[emim][BF4] 0.1542

1

10

100

1000

60 70 80 90 100

α(C

irc

ulatio

nratio

)

Desorber outlet temperature [o

C]

Desorber outlet temperature [o

C]

Desorber outlet temperature [o

C]

a

R134a

R152a

R125

1

10

100

1000

60 70 80 90 100

α(C

irc

ulatio

n ratio

)

b

1

10

100

1000

60 70 80 90 100

α(C

irc

ulatio

nratio

)

c

Fig. 4. Circulation ratios of the absorption system with respect to the desorber outlettemperature using different working fluid pairs: (a) refrigerant/[bmim][PF6] andR134a/[hmim][PF6]; (b) refrigerant/[emim][Tf2N] and R134a/[hmim][Tf2N]; (c) R134a/[hmim][BF4], R32/[bmim][BF4] and water/[emim][BF4].

Y.J. Kim et al. / Energy 44 (2012) 1005e10161010

Equation (31a) shows that a smaller circulation ratio results inless heat input to the desorber for the same amount of cooling atthe evaporator, whichwould increase COP. Thus, a lower circulationratio resulting from the temperature increase in Fig. 4 increasesCOP at low desorber outlet temperatures (<80 �C), as shown inFig. 3. However, as the temperature further increases, the change

rate of the circulation ratio ðva=vT8Þ is diminished (Fig. 4). Also, theincrease in enthalpies of the desorber outlet (h3 and h8) accom-panied by the increase of the temperature require larger heat inputto the desorber, Equation (31b). Consequently, at high desorberoutlet temperatures, the effect of the reduced circulation ratio andthe increased enthalpy on the heat input to the desorber conflictsand the effect on COP cancels each other out. This leads to a levelingoff of the COPs in Fig. 3.

The refrigerant mole fraction for the strong solution is approx-imated in Equation (32) [47].

xswPe

gsPsatðT5Þ(32)

Assuming gszgw, Equations (29) and (32) can be written toyield Equations (33) and (34) using the inequality, xs > xw.

PsatðT8Þ>PcPePsatðT5Þ (33)

T8>Tsat

�PcPePsatðT5Þ

�wT8;min (34)

From Equation (34), it can be seen that the lower bound on thedesorber outlet temperature, or maximum operating temperatureof the system, is a function of the operating conditions, i.e., vaporpressures at the evaporator and condenser as well as the absorberoutlet temperature. Fig. 3 also shows that the COP curves generallyconverge to zero at T8 w 61.5 �C (dashed line) for all refrigerants.Thus, the COP curves can be shifted to lower temperatures byadjusting the operating conditions so as to have optimum perfor-mance at a lower temperature, ca. 60 �C. Then, the utilization ofwaste-heat can be optimized.

4.2. Effect of refrigerant/IL compatibility

The R32/[bmim][PF6] pair showed the highest COPw0.42, of allthe refrigerant pairs in Fig. 3(a). The COP values for R134a andR152a with [bmim][PF6] were moderate, while for R125 and R143awere lower. Note that the COP trends for the HFC refrigerants inFig. 3(a) approximately correlate with the refrigerant mass fractionin the strong solution, xms , listed in Table 4.Equation (35) shows therelation between the mass fraction and mole fraction.

xm ¼ xMr

xMr þ ð1� xÞMIL(35)

where Mr and MIL are the molecular weights of refrigerant and IL,respectively. The effect of the strong solution solubility on the

a

R134a

R125

0.22

0.37

0.61

1

Refrig

eran

t m

as

s fra

ctio

n ratio

Desorber outlet temperature [oC]

57 67 78 88 99 111

R134a

R134a

/[hmim][Tf2N]

b

c

Desorber outlet temperature [oC]

57 67 78 88 99 1110.22

0.37

0.61

1

Refrig

eran

t m

as

s fra

ctio

n ratio

0.14

0.37

1

Refrig

eran

t m

as

s frac

tio

n ratio

0.61

0.22

Desorber outlet temperature [oC]

57 67 78 88 99 111

Fig. 5. Refrigerant mass fraction ratio in the weak solution relative to the strongsolution ðlnxmw=xms Þ of the absorption system as function of the desorber outlettemperature (lnT8), using different working fluid pairs: (a) refrigerant/[bmim][PF6] andR134a/[hmim][PF6]; (b) refrigerant/[emim][Tf2N] and R134a/[hmim][Tf2N]; (c) R134a/[hmim][BF4], R32/[bmim][BF4] and water/[emim][BF4].

Y.J. Kim et al. / Energy 44 (2012) 1005e1016 1011

performance is attributed to the smaller circulation ratio broughtabout by a larger value of refrigerant mass fraction in the strongsolution. Thus, it can be stated that high solubility of refrigerant(mass fraction solubility) in the IL solution improves the perfor-mance of the absorption system. However, the higher solubility ofR134a in [hmim][Tf2N] results in a lower COP value relative toR134a/[emim][Tf2N] in Fig. 3(b). Ren and Scurto [39] compared thesolubility (mole fraction solubility) of R134a in [emim][Tf2N] andR134a in [hmim][Tf2N]. They reported that a longer alkyl chainlength on the cation (i.e. hmimþ vs. emimþ) increases the solubilityof R134a. Shiflett and Yokozeki [38] and Aki et al. [53] measured thesolubility of R32 and carbon dioxide in the ILs, respectively. Theyboth also found that the solubility of the gas in the IL can beincreased by increasing the alkyl chain length on the organic cation,which is in agreement with the results in Fig. 5(b). On the otherhand, Kerle et al. [54] investigated the temperature dependence ofthe solubility of carbon dioxide in imidazolium-based ILs andshowed that as the alkyl chain length increases, the solubilitybecomes less sensitive to temperature. The temperature coefficientof solubility is directly related to the entropy of mixing in Equation(36) [55e58].�vlnx2vlnT

�P¼ Dsr

Rðvlna2=vlnx2ÞP;T(36)

where Dsr is the entropy of mixing (or relative partial molarentropy) of refrigerant in the solution, and a2 is the activity ofrefrigerant in the solution given by Equation (37).

a2 ¼ g2x2 (37)

In the Henry’s law regime (i.e., when g2 is independent of x2),the derivative in the right-hand side dominator in Equation (36)becomes unity. Therefore, the equation can be reduced to thevan’t Hoff form [57] as follows,

Dsr ¼ R�vlnx2vlnT

�P

(38)

Integrating Equation (38) yields [58]:

ðlnxÞP ¼ DsrR

lnT þ D (39)

where D is an integration constant. Since the relative partial molarentropy is reduced by increasing the alkyl chain length, Equation(39) supports the reduced dependence of the refrigerant massfraction on the desorber outlet temperature. In summary, longercation alkyl chain length cause a larger solubility but lowerdependence of the solubility on temperature. These effects of thecation result in larger circulation ratio for R134a/[hmim][Tf2N], asshown in Fig. 4(b), because the refrigerant mass fraction differencebetween the strong and weak solutions in Equation (30), ðxms � xmwÞ,is lowered due to the reduced dependence of the solubility ontemperature with the longer alkyl chain length of [hmimþ]. Thus,even with larger refrigerant solubility in the strong solution, theperformances of the system can be degraded due to the reduceddependence of the solubility on temperature. R143a/[bmim][PF6]and R134a/[hmim][PF6] are remarkably less sensitive to tempera-ture (Fig. 5(a)) and show relatively low COPs (Fig. 3(a)). Watersolubility in [emim][Tf2N] depends relatively more on the desorberoutlet temperature (Fig. 5(b)), but refrigerant mass fraction instrong solution is extremely low so that very large circulation ratioand thereby low COP are found in Fig. 4(b) and Fig. 3(b),respectively.

Ren and Scurto [39] investigated the effect of R134a solubility onanion type in the IL and found that the solubility scales with the

ionic radius of the anion or charge density. A lower charge densityallows the polar hydrofluorocarbons to more readily dissolve.Fig. 5(aec) shows that the solubility of R134a in [hmim][PF6] and[hmim][BF4] are similar while R134a is more soluble in [hmim][Tf2N] (Table 4). Therefore, the high solubility of R134a in [hmim]

0.001

0.01

0.1

1

10

250 300 350 400

Vis

co

sity

[P

a s

]

Temperature [K]

[hmim][PF6]

[bmim][PF6]

[hmim][BF4]

[bmim][BF4]

[emim][BF4]

[hmim][Tf2N]

[emim][Tf2N]

Fig. 6. Viscosities of ILs evaluated using group contribution method [43].

Y.J. Kim et al. / Energy 44 (2012) 1005e10161012

[Tf2N] and the high temperature dependence of the R134a solu-bility in [hmim][BF4] yield higher COP values than that of R134a/[hmim][PF6]. The differences in both the solubility (Fig. 5(a) and (c))and COP values (Fig. 3(a) and (c)) for R32 in [bmim][PF6] and[bmim][BF4], were not significant.

The COP values for the absorption system using the water/[emim][BF4] pair are notably higher than those with the otherworking fluid pairs (Fig. 3(c)). The maximum COP was 0.91 atTd ¼ 70 �C. The imidazolium-[BF4�] ILs have dramatically highersolubility for water (Fig. 4(c)), while ILs with [Tf2N�] showa decrease in water solubility [26] as shown in Fig. 4(b). The lowCOP values of water/[emim][Tf2N] in Fig. 3(a) are, therefore,attributed to the low solubility of water in [emim][Tf2N].However, the refrigerant mass fraction of water in [emim][BF4]isless than that of R32 (Fig. 5(c)). Moreover, the circulation ratio ofwater is the highest among the refrigerants for [emim][BF4] inFig. 4(c), by which the lower COP is expected than that of R32 in[emim][BF4]. This suggests that there should be another factoraffecting the performance of the absorption system, other thansolubility and circulation ratio. The latent heat of vaporization forwater is an order of magnitude higher than that of the otherrefrigerants as listed in Table 1. Then, the required mass flow rateof water to produce the target 100 W cooling capacity at theevaporator is significantly smaller (Table 5). It is evident fromEquation (31) that the low water flow rate contributes toreducing the heat input required at the desorber, Qd. Conse-quently, this leads to high COP values for the water/[emim][BF4]pair shown in Fig. 3(a). Similarly, the refrigerants R152a and R32also have large latent heats and small refrigerant mass flow rates.This improves the system performance when using theseworking fluid pairs.

4.3. Effect of IL viscosity

Higher viscosity ILs cause an increased pressure drop in thecompression loop, which would result in larger pumping poweror larger pipes and system volume. Fig. 6 shows the viscosity ofpure ILs vs. temperature evaluated using the group contributionmethod [43]. The viscosity increases with cation mass:emimþ < bmimþ < hmimþ. The viscosity is more dependent on theanionwith the following order: Tf2N� < BF4� < PF6�. The viscosity of[emim][Tf2N] is only 31.3 mPa s at 294 K, which is 10 times smallerthan that of [hmim][PF6].

Fig. 7 shows the total pumping power of the absorption systemfor various refrigerants/IL pairs. It is very encouraging that for mostof the working fluid pairs, except R125/[bmim][PF6], R143a/[bmim][PF6] and R114/[emim][Tf2N], the pumping power was less than10 W for the 100 W cooling capacity system. The total pumpingpower can be calculated using Equations (40) and (41).

Wp ¼ _ms

�Pc � Pe þ DPfriction

(40)

Table 5Refrigerant mass flow rates in the absorption system.

Refrigerants Refrigerant flow rate [g/s]

R114 0.897R124 0.785R125 1.155R134a 0.651R143a 0.761R152a 0.402R32 0.420Water 0.042

Wp_mr

¼ a�Pc � Pe þ DPfriction

(41)

where DPfriction is the pressure drop due to friction. The circu-lation ratio plays an important role in the pumping power, aswell. Since a lower solution flow rate is accompanied bya reduction in the circulation ratio at higher desorber tempera-tures, as shown in Equation (40), lower pumping power valuesresult from a higher desorber temperature. Also, COPwaste (ratioof the evaporator cooling capacity to the pumping power)improves as the pumping power decreases at higher desorbertemperatures, as shown in Fig. 8. Due to the small flow rate ofwater, _mr, the pumping power becomes negligible, resulting invery high COPwaste values for the water/[emim][Tf2N] and water/[emim][BF4] pairs. However, it should be noted that even thoughthe COPwaste values are large for water/[emim][Tf2N], the lowvalue of regular COP requires large desorber and absorber heatexchangers.

Fig. 9 shows the ratio of pumping power (due to friction) tothe total pumping power. For most working fluid pairs, less than10% of the total pumping power is attributed to friction. Most ofthe pumping power is used to pressurize the refrigerant/IL pairðPc � PeÞ, which is the reason that R32 shows the smallest fric-tional contribution while water has the largest. Since the pres-sure difference between the condenser and evaporator withwater is only 9.2 kPa, which is negligibly small and less than 1%of the pressure difference with R32, more than 90% ofthe pumping power for the water/[emim][Tf2N] and water/[emim][BF4] was used to overcome the frictional losses. For theR143a/[bmim][PF6] and R114/[emim][Tf2N] pairs, the high solu-tion flow rate, _ms, resulted from high circulation ratios and, inturn, caused large values of pumping power due to frictionallosses.

The relatively small contribution of friction to the total pumpingpower was mainly because of the low viscosity due to dilution ofthe refrigerants. However, it also should be noted that the heatexchanger microchannel cross-sectional area was slightly largerthan that of conventional microfluidic channels. As shown inEquation (28), the laminar flow friction factor is inversely propor-tional to the Reynolds number, flwRe�1wm=Gdh. Under fixed massflow rate, mass flux is inversely proportional to the channel cross-sectional area, GwA�1

cs w4=pd2h, where Acs is the channel cross-

1

10

100

1000

60 70 80 90 100

To

tal

pu

mp

ing

po

we

r ,

Wp[W

]

Desorber outlet temperature [oC]

a

R134aR32

R125

0.1

1

10

100

1000

60 70 80 90 100

To

tal p

um

pin

g p

ow

er,

Wp[W

]

Desorber outlet temperature [oC]

b

c

R114

0.01

0.1

1

10

100

60 70 80 90 100

To

tal p

um

pin

g p

ow

er,W

p[W

]

Desorber outlet temperature [oC]

Fig. 7. Total pumping power consumptions of the absorption system with respect tothe desorber outlet temperature using different working fluid pairs: (a) refrigerant/[bmim][PF6] and R134a/[hmim][PF6]; (b) refrigerant/[emim][Tf2N] and R134a/[hmim][Tf2N]; (c) R134a/[hmim][BF4], R32/[bmim][BF4] and water/[emim][BF4].

a

b

c

Fig. 8. COPwaste of the absorption system with respect to the desorber outlettemperature using different working fluid pairs: (a) refrigerant/[bmim][PF6] andR134a/[hmim][PF6]; (b) refrigerant/[emim][Tf2N] and R134a/[hmim][Tf2N]; (c) R134a/[hmim][BF4], R32/[bmim][BF4] and water/[emim][BF4].

Y.J. Kim et al. / Energy 44 (2012) 1005e1016 1013

sectional area. Assuming that the two-phase multiplier,f2l ¼ ðdP=dzÞTP=ðdP=dzÞl, is a weak function of diameter, Equation

(23) shows that the frictional pressure drop, which is the firstbracket of the right-hand side of the equation, can be given byEquation (42).

�dPdz

�friction

¼�2flG2ð1� xvÞ

dhrl

�f2l w

Gd2h

wd�4h (42)

Thus, small adjustments in the microfluidic channel dimensionscan have a large effect on the pumping power for systemswith highviscosity.

a

b

c

Fig. 9. Pumping power consumptions of the absorption system caused by friction withrespect to the desorber outlet temperature using different working fluid pairs: (a)refrigerant/[bmim][PF6] and R134a/[hmim][PF6]; (b) refrigerant/[emim][Tf2N] andR134a/[hmim][Tf2N]; (c) R134a/[hmim][BF4], R32/[bmim][BF4] and water/[emim][BF4].

a

b

c

Fig. 10. COPs of the absorption systemwith respect to the absorber outlet temperatureusing different working fluid pairs: (a) refrigerant/[bmim][PF6] and R134a/[hmim][PF6]; (b) refrigerant/[emim][Tf2N] and R134a/[hmim][Tf2N]; (c) R134a/[hmim][BF4],R32/[bmim][BF4] and water/[emim][BF4].

Y.J. Kim et al. / Energy 44 (2012) 1005e10161014

4.4. Effect of the absorber outlet temperature

The absorber outlet temperature is controlled by the ambientcooling temperature. A lower absorber temperature enhances the

COP of the system, as shown in Fig. 10. As the absorber outlettemperature rises, the solubility of the strong solution decreases, asdetermined by Equation (32). Thus, the difference in mole fractionbetween the strong and weak solutions is diminished, which leadsto a larger circulation ratio and lower COP. The increase in thedesorber inlet enthalpy, h7, driven by the absorber outlet temper-ature increase, also contributes to the enhancement of COP byreducing the required heat input to the desorber, Qd. However, the

Y.J. Kim et al. / Energy 44 (2012) 1005e1016 1015

effect of the circulation ratio change on COP dominates, so that themonotonic decrease in COP, with respect to the absorber temper-ature increase, is observed in Fig. 10.

5. Conclusion

The feasibility of several refrigerant/IL mixtures as working fluidpairs for the absorption based refrigeration systemwas numericallyevaluated. The refrigerant/IL compatibility indicated by the circu-lation ratio was one of the most important factors in determiningthe system performance with the new working fluid pairs. Thesolubility of the refrigerants in the ILs and their temperaturedependence affected the circulation ratio. R32 showed the lowestcirculation ratio with relatively high COP values for the ILshaving [PF6�] and [BF4�] anions. Although R124 showed the bestcompatibility and performance with [Tf2N�], due to its negativeenvironmental impact, R124 may not be the optimum choice ofrefrigerant. Cations with shorter alkyl chains are preferred([emimþ] > [bmimþ] > [hmimþ]) due to more sensitive depen-dence of the solubility on temperature. Overall, water/[emim][BF4]showed the highest COP value, ca. 0.91, where the better compat-ibility of water with [emim][BF4] and the superior properties ofwater as a heat transfer fluid, such as large latent heat of evapo-ration, followed by extremely small refrigerant (water) flow rateresulted in its high performance. Using water with [emim][Tf2N]has not been found promising due to the extremely low solubility ofIL in water (Table 4). Lafrate et al. [59], however, showed that diol-functionalized ILs with [Tf2N�] can be completely water-miscible,which could potentially deliver an improved performance of theabsorption system. However, the evaporator and condensertemperature and pressure values for water (as the refrigerant) arenot of interest for most electronic cooling applications. The effect ofthe higher viscosity of the ILs on the pressure drop and pumpingpower were mitigated by slightly increasing the microfluidicchannel cross-sectional area. The absorber outlet temperaturesignificantly affected the system performance. Lowering theabsorber outlet temperature is desirable due to the increasedsolubility of the refrigerant in the IL. The system level numericalinvestigation showed that refrigerant/IL pairs are promisingmaterials for absorption refrigeration utilizing low-grade wasteheat, as found in electronic systems. For practical implementationof ILs in the absorption systems the complimentary studies con-cerning the detailed effects of heat and mass transfer characteris-tics and geometries of heat/mass exchangers are needed. Also,optimal microfluidic absorber and desorber configurations need tobe developed to fully realize the promising potential of anabsorption refrigeration system based on the IL chemicalcompressor.

Acknowledgments

The authors acknowledge the support of the Interconnect FocusCenter, one of the five research centers funded under the FocusCenter Research Program, a Semiconductor Research Corporationand DARPA program.

Nomenclature

C coefficients in Lockhart and Martinelli correlationCOP coefficient of performanceCOPwaste coefficient of performance with waste-heatcp specific heat [J/kg K]dh hydraulic diameter [m]F adjustable binary interaction parameterf fanning friction factor

G Gibbs energy [J]Gm mass flux [kg/m2s]H enthalpy [J]h specific enthalpy [J/kg]M molecular weight [kg/mol]_m mass flow rate [kg/s]P pressure [Pa]Q heat transfer [W]R gas constant, 8.314 J/mol KRelo liquid-only Reynolds numberS entropy [J/K]s specific entropy [J/kg K]T temperature [K]v specific volume [m3/kg]Welo liquid-only Weber numberWp pumping work [W]X Martinelli parameterx liquid phase refrigerant mole fractionxm liquid phase refrigerant mole fractionxm liquid phase refrigerant mass fractionxn liquid phase ionic-liquid mole fractionxv vapor qualityz coordinate [m]

Greek symbolsb aspect ratioε void fraction4 two-phase multiplierg activity coefficientm viscosity [Pa s]r density [kg/m3]s adjustable parameters of Equation (5)

Subscripts0 reference state1, ., 10 state numbers indicated in Fig. 1a absorberc condenserd desorbere evaporatorIL ionic-liquidl liquidr refrigerants strong-refrigerant solutionsat saturationshx solution heat exchangerv vaporw weak-refrigerant solution

SuperscriptsE excess propertyid ideal solution

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