Theoretical and Experimental Investigation of an Absorption

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Theoretical and Experimental investigation of an absorption refrigeration system using R134/[bmim][PF6] working fluid Sarah Kim, and Paul A Kohl Ind. Eng. Chem. Res., Just Accepted Manuscript • DOI: 10.1021/ie400985c • Publication Date (Web): 22 Aug 2013 Downloaded from http://pubs.acs.org on August 26, 2013

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Industrial & Engineering Chemistry Research

Theoretical and Experimental Investigation of an Absorption Refrigeration System Using R134/[bmim][PF6] Working Fluid

Sarah Kim and Paul A. Kohl*

School of Chemical and Biomolecular Engineering Georgia Institute of Technology Atlanta, GA 30332

* Corresponding author: [email protected]

ACS Paragon Plus Environment


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Abstract 1,1,2,2-tetrafluoroethane (R134) and ionic liquid (IL), 1-butyl-3-methylimidazolium hexafluorophosphate ([bmim][PF6]), were investigated as a new working fluid pair in an absorption refrigeration system. The R134/[bmim][PF6] pair was compared to previous studies using 1,1,1,2-tetrafluoroethane (R134a) and the same IL. The R134/[bmim][PF6] fluid pair had up to 92.3% greater cooling-to-total-energy efficiency than the R134a/[bmim][PF6] fluid pair even though R134 and R134a are isomers and have nearly identical physical properties. The coefficient of performance of the R134/[bmim][PF6] fluid pair was up to three times larger than that of R134a/[bmim][PF6] fluid pair when only waste heat was used at the desorber. A working refrigeration system with R134/[bmim][PF6] was constructed and the measurements of its performance showed that R134/[bmim][PF6] had 1.9 times larger cooling capability than R134a/[bmim][PF6] at a desorber temperature as low as 63°C.


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1. Introduction The United States Department of Energy (DOE) has established energy efficiency standards that apply to various types of appliances and equipment in an effort to reduce the exponentially growing electricity consumption. Manufacturers of residential central air conditioners and heat pumps have been required to comply with the DOE energy conservation standards since 19921. The mandatory standards drive the need for more efficient heat pumps. Toward this end, the absorption refrigeration cycle based on an ionic liquid (IL) working fluid is of interest as a means of utilizing low-quality waste heat. Absorption chillers offer an opportunity to recycle large amounts of industrial waste heat. Also, renewable energy sources which generate heat can be directly used for refrigeration without the need of conversion to electrical power2. However, the use of the absorption refrigeration cycle has been somewhat limited due to technological challenges, health hazards, and environmental concerns of existing systems3. Commonly used absorption refrigeration working fluid pairs, ammonia/water and water/LiBr, have drawbacks including toxicity (ammonia), negative effects of crystallization (LiBr), incompatibility with metal components (corrosion), and the need for a rectifier for postdesorption separation of the two fluid streams4. Further, the pressure and temperature of the evaporator in the ammonia/water and water/LiBr systems is significantly different from that of Freon-based vapor-compression systems so that many refrigeration applications are not within reach. ILs are a liquid salt at ambient temperature and are being considered as absorbents in a variety of applications because of their tunable properties, zero vapor pressure, high thermal stability, and environmental safety5. In particular, the near-zero volatility of the IL enables easy separation of the volatile working fluid by thermal stratification. There have been several



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theoretical studies that have evaluated the performance of IL based working fluids including 1ethyl-3-methylimidazolium

ethylsulfate

([emim][EtSO4])6,

1,3-dimethylimidazolium

dimethylphosphate ([emim][DMP])7 , and various imidazolium, pyridinium, and pyrrolidinium type ILs8,9. In a previous report, an experimental demonstration of a cooler/heat pump using a Freon/IL working fluid pair was reported for the first time10. The application was targeted for cooling in high power electronics using waste heat. An IL, [bmim][PF6], was selected as the absorbent because of its high molar uptake of freons, which is essential to deliver maximum cooling. The analysis showed 1,1,1,2-tetrafluoroethane (R134a) to be a promising refrigerant when paired with 1-butyl-3-methylimidazolium hexafluorophosphate ([bmim][PF6]) as the working fluid for a waste heat recycling absorption refrigeration system. 1,1,2,2-tetrafluoroethane (R134), an isomer of R134a (Figure 1), was expected to have similar results with that of R134a. However, the shift in fluorine substitution (1,1,1,2-tetrafluoroethane to 1,1,2,2-tetrafluoroethane) has resulted in a significant change in the absorption of the fluorocarbon in [bmim][PF6] and the resulting performance of the absorption refrigeration system. In this study, the performance of the R134/[bmim][PF6] working fluid pair was modeled using a two-phase pressure drop Equationof-State (EOS) model11. The effect of a counter-flowing solution heat exchanger, desorber temperature, and absorber temperature on the system performance was evaluated. The fluid was also tested in the bench-top absorption cooling cycle for an experimental demonstration. Both computational and experimental results of R134/[bmim][PF6] were compared with that of R134a/[bmim][PF6].

2. Computational Model


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2.1 Thermodynamic system analysis The principal features of the absorption refrigeration cycle are shown in Figure 2. The cycle resembles that of the vapor compression refrigeration (heat pump) system, except the vapor compressor is replaced with a thermochemical process. Pressurization of the thermochemical process starts in the absorber, where the refrigerant vapor from the evaporator (state point 2) is exothermically absorbed into the strong-IL solution (state point 10), resulting in a weak IL solution at state point 5. The IL solution is pressurized by the liquid pump after the absorber. Then, the solution (regenerative) heat exchanger preheats the weak-IL solution at state point 6 creating state point 7 using heat from the strong-IL solution flowing back to the absorber (from the desorber). In the desorber, superheated refrigerant vapor is released at high temperature and pressure from the IL via desorption from the weak-IL solution by the addition of heat (preferably waste heat). The strong-IL returns to the absorber through the solution heat exchanger and an expansion

device.

The

condensation/absorption

process

at

the

absorber

and

vaporization/desorption process at the desorber both occur in the liquid phase. This allows use of a liquid pump to create the pressure difference between the condenser and evaporator. System-level simulations have been carried out with refrigerant/ionic liquid combinations. In all the calculations, the operating temperature of the condenser and evaporator were set at 50°C and 25°C, respectively. The cooling capacity at the evaporator was set at 100 W. The energy and mass conservation equations for all components in the system were simultaneously solved to determine the heat and workloads. The overall energy balance for the system is given in Equation 1. Qd + Qe + W p = Qc + Qa


(1)


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Where Wp is the liquid pump work and Q is the heat input/output. The subscripts d, e, c, and a correspond to the desorber, evaporator, condenser, and absorber, respectively. All values of heat are expressed as positive (magnitude) values regardless of the direction (in or out) of heat flow. For the calculations, the heat exchanger efficiency, 35% was used. The estimated heat transfer is thus given by Equation 211. = Qactual 0.35 = Qmax 0.35Cmin (Thot ,in − Tcold ,in )

(2)

Where Cmin is the heat capacity coefficient of the stream that limits the amount of heat transfer. This estimate of the heat transfer at the counter-flow heat exchanger can then be used to determine the temperature and enthalpy of the exit streams of the solution heat exchanger. The energy conservation for a sub-system consisting of a regenerative heat exchanger and a pump is given by Equation 3.

( h9 − h8 )( mw − mr ) =( h7 − h5 ) mw − Wp

(3)

Where h is enthalpy and the subscripts correspond to the locations shown in the system diagram, Figure 2. Also, mw and mr are the mass flow rates of the weak-IL solution and refrigerant leaving the desorber, respectively. Energy conservation for the desorber is given by Equation 4.

Qd= h8 ( mw − mr ) + h3 mr − h7 mw

(4)

Similarly, the heat rejected at the absorber is given by Equation 5.

Qa = h5 mw − h10 ( mw − mr ) − h2 mr

(5)

Energy conservation for the condenser and the evaporator yield the respective heat loads, Equations 6 and 7.


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Q= c

( h4 − h3 ) mr

(6)

Q = e

( h2 − h1 ) mr

(7)

The cooling-to-total-energy (CE) is defined as the heat removed at the evaporator divided by the power supplied to the desorber and the pump, Equation 8.

CE =

Qe Q ≈ e Qd + Wp Qd

(8)

Since waste heat is intended to be used to heat the desorber, the practical coefficient of performance (COP), η, is defined by Equation 9.

η=

Qe Wp

(9)

Equation 9 is the usual figure of merit for absorption refrigeration/heat pump systems where waste heat is used. In this analysis, the Redlich-Kwong (RK) type EOS was used to calculate the thermodynamic properties of the fluids at each point in the cycle. The RK-EOS was chosen because it has a temperature dependent attractive term, a(T), and fits the data very well. Binary interaction parameters (BIPs) were introduced to improve the accuracy of the model. Several assumptions were made for convenience in the calculations: (i) the expansion process is isenthalpic; (ii) the compression process is isentropic; (iii) state 4 is saturated liquid refrigerant; (iv) state 2 is saturated vapor refrigerant; and (v) the vapor quality at state 5 is zero. The general RK-EOS can be written in the following form, Equations 10 to 1212.

= P

RT a (T ) − V − b V (V + b)


(10)


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a (T ) = 0.42748

R 2Tc2 α (T ) Pc

b = 0.08664

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(11)

RTc Pc

(12)

The subscript c represents the critical properties of the substance. P is the pressure, T is the temperature, V is the molar volume, R is the gas constant, and a and b are constants. The temperature dependent function of the α parameter is expressed by Equation 13.

α (T= )

≤3

∑β k =0

k

(1/ Tr − Tr ) k , Tr ≡ T / Tc

(13)

The parameter βk was determined so as to yield the vapor pressure of each pure compound, the refrigerant and the ionic liquid11,13. The critical properties along with the β values are summarized in Table 1. Three binary interaction parameters (BIPs), τ, l, and k were introduced in the a and b parameters for N component mixtures15.

a =

N

∑

i , j =1

0.42748 ai a j fij (T )(1 − lij ) xi x= j , ai

R 2Tci2 α (T ) Pci

fij (T ) = 1 + τ ij / T , where τ ij = τ ji and τ ii = 0

= b

RT 1 N bi + b j ) (1 − lij )(1 − kij ) xi x j , = bi 0.08664 ci ( ∑ 2 i , j =1 Pci

(14)

(15)

(16)

Where, lij = l ji ; lii = 0 ; kij = k ji ; kii = 0 . A detailed description of finding the BIP can be found elsewhere11. The optimum BIPs for R134 and [bmim][PF6] mixture are listed in Table 2.


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The EOS was then used to find the enthalpy values at the point of interest, Equation 17. T N

H= H R + ∫ ∑ xi C pi0 dT + H 0 T0

(17)

i =1

Where HR is the residual enthalpy, T0 is the reference temperature which was set at 273.15 K, and H0 is an arbitrary constant (an enthalpy at the reference state). The ideal-gas heat capacity of the i -th species, C pio , was modeled in Equation 18. C pi0 =C0 + C1T + C2T 2 + C3T 3

(18)

The coefficients in Equation 18 are given in Table 3. Using the heat capacity for each species in the mixture, the enthalpy for the RK EOS is given by Equation 19. T N V  a T da  H=  − RT Z xi C pi0 dT + H 0 + − + ln ( 1) ∑  ∫ T 0  b b dT  V + b i =1

(19)

2.2 Pressure Drop The viscosity of ILs is often relatively high. For example, [bmim][PF6] at 294 K and atmospheric pressure has a viscosity of 376 mPa⋅s16. Thus, the IL flow in the microchannels used at the absorber and desorber may create a significant pressure drop, which can affect the system performance. The average pressure drop through the microfluidic channel heat exchangers was evaluated using a two-phase pressure drop equation, Equation 2017.

 dP  = −   dz 

 x v )2 (1 − x v )2   2 f lGm2 (1 − x v )  2 2 d (    ϕl + Gm + d h ρl dz  ερ v (1 − ε ) ρ l     

(20)

Where d h is the hydraulic diameter of the channel, and f, G, xv, ρ, and ε are the liquid-phase fanning friction factor, mass flux, vapor quality, density, and void fraction, respectively. z is the axial direction coordinate along the channel length. Subscripts “l” and “v” stand for liquid and



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vapor phase, respectively. In the two-phase multiplier correlation of Lockhart and Martinelli18, ϕl is incorporated with the C value proposed by Lee and Mudawar19, Equations 21 to 23.

ϕl2 =+ 1

C 1 + 2 X X

(21)

C = 2.16 Re0.047 Welo0.6 (laminar liquid-laminar vapor) lo

(22)

0.23 C = 1.45 Re0.25 (laminar liquid-turbulent vapor) lo We lo

(23)

Where Relo and Welo are liquid-only Reynolds and Weber numbers, respectively. The Martinelli parameter, X, and the single phase empirical correlation of the fanning friction factor for laminar flow in a rectangular channel by Shah and London20 are expressed by Equations 24 and 25. 0.5

0.5

 µl   1 − x v   ρ v  X =   v     µv   x   ρl 

0.5

f Re = 24(1 − 1.3553β + 1.9467 β 2 − 1.7012 β 3 + 0.9564 β 4 − 0.2537 β 5 )

(24)

(25)

Where µ is the viscosity and β is the aspect ratio of the channel. Also, the void fraction model of Zivi21 was adopted in this study. Microchannel structures are used in the absorber and the desorber in this study due to their high heat and mass transfer rates. However, the magnitude of the negative effects of the IL high viscosity on the pressure drop within the microchannel heat exchangers and, in turn, the system performance is assessed in this study. The dimensions (length × width) of the evaporator and condenser are 2×2 cm, and 3×3 cm, respectively. The dimensions of the absorber and the desorber are 8×8 cm. The microfluidic channel crosssectional area for the heat exchangers is 1×1 mm.

3. Experimental Methods


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The ionic liquid [bmim][PF6] (iolitec, 99%), R134 (SynQuest Laboratory, Inc., 98%), and R134a (Airgas, 99.9%) were used as received. An experimental setup for a laboratory scale absorption refrigeration system using the IL based working fluid has been built and operated, the details of which are described elsewhere10. To investigate the effect of desorber power (waste heat) input on the evaporator cooling capacity, the desorber power was increased up to ∼100 W to keep the evaporator temperature constant at 42oC with the condenser and absorber coolant (secondary fluid) inlet temperatures of 22oC. When the refrigerant was changed from R134a to R134, the R134a was discharged from the bench top system using a refrigerant recovery unit. A port on the refrigerant loop prevented drainage of the ionic liquid. The presence of a check valve and vertical cylinder between the solution loop and refrigerant loop also worked as a barrier for ionic liquid overflow. The [bmim][PF6] in the system remained under vacuum for several days until there was no further change in the pressure. R134 was then loaded into the system to the same level as R134a as determined by the saturated pressure. The measured parameters were temperature, absolute pressure, and electrical power input. The uncertainty in the temperature reading was 0.1 K for the calibrated thermocouples relative to each other. The uncertainty in absolute pressure measurement was 0.25% of the maximum value of 2068.43 kPa (300 psi). The uncertainty in the output value of the electrical power transducer was 0.14% of the measured value.

4. Results and Discussion 4.1 Computational results It was previously shown that the R134a/[bmim][PF6] working fluid pair could be used in a waste-heat recycling absorption refrigeration system10. The change in fluorine position in



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tetrafluoroethane from the first carbon to the second carbon (R134a is 1,1,1,2-tetrafluoroethane and R134 is 1,1,2,2-tetrafluoroethane) has little effect on its physical properties, such as density, and boiling point. However, this difference in chemical structure has a significant effect on its interaction with [bmim][PF6], which affects the performance of the absorption refrigeration cycle. The solubility of R134 in the IL absorbent was predicted using the two phase pressure drop EOS model, Figure 3. The calculated values are in good agreement with the experimentally measured data reported by Shiflett and Yokozeki 22. In Fig. 3, the experimentally determined values are the data points and the solid lines correspond to the fitted values. Surprisingly, R134 showed significantly higher solubility at the same temperature and pressure in comparison to R134a at the same temperatures, Figure 4. This is most likely due to the higher probablity for hydrogen bonding between the symmetrical R134 and [bmim][PF6]24. The operating pressure of the system is determined by the saturated pressure of the refrigerant. Therefore, the solubility difference between the absorber and desorber is important in determining the overall solubility of the refrigerant. The vapor pressure of R134 and R134a at the condenser and evaporator temperature of 50 °C and 25 °C, respectvely, are listed in Table 4. Effect of absorber/desorber outlet temperature on CE: The effect of changing several key operating conditions was evaluated using the model developed in this work. The effect of lowering the absorber temperature from 309.65K to 300.65K was evaluated. Figure 5 shows the value of CE for the absorber at 309.6K and Figure 6 shows the resulting value of CE for the absorber at 300.65K. The CE value was generally observed to increase when the absorber temperature was lowered from 309.65K to 300.65K due to higher refrigerant solubility in the absorber. R134 had a higher CE value than R134a at the same temperature, with an average improvement of 32.3% and 92.3% for the absorber operated at 300.65 K and 309.65 K,


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respectively. This larger difference in CE values between the refrigerants at 309.65 K is due to the higher solubility of R134 compared to R134a. The effect of solubility on the system efficiency becomes increasingly important for cases where the solubility difference between the absorber and desorber is small (i.e. high absorber temperature). Effect of solution heat exchanger on CE: The introduction of a solution heat exchanger between the inlet and outlet of the desorber improves the performance of the cycle. The solution heat exchanger transfers heat from the strong IL solution at system point 8 to the weak IL solution at system point 6, Figure 2. The heat exchanger improves the system efficiency because it lowers the amount of waste heat needed to increase the temperature of the solution mixture to the desorber temperature and less heat is discharged at the absorber. A 35% efficient heat exchanger improved the CE values by 16.4% and 32.6% for R134 and R134a (see Figure 7), respectively, compared to the case without the heat exchanger, as shown in Figure 5. The system performance was enhanced to a greater extent for the refrigerant with lower CE values, R134a. Effect of absorber/desorber outlet temperature on η: The coefficient of performance η is plotted with respect to desorber outlet temperature in Figures 8 and 9 for absorber temperatures of 309.65 K and 300.65K, respectively. At higher desorber outlet temperatures, the pumping work was reduced due to an increase in the refrigerant-to-absorbent circulation ratio. This means that the IL absorbent can carry more refrigerant on each pass which leads to an efficiency increase. The liquid pumping work is the product of the liquid volumetric flow rate and pressure difference between the absorber and desorber: W p= Vl × ∆P . Therefore, when free, waste-heat is used at the desorber to drive the system, the operating pressure range and solubility difference between the absorber and desorber are both important because the goal is to minimize the



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pumping work by reducing the amount of IL pumped. Both refrigerants had higher COP values at lower absorber temperature due to the higher refrigerant-to-absorbent ratio resulting in a lower liquid volumetric flow rate. The R134/IL pair had twice (Figure 9) or three times (Figure 8) the efficiency of the R134a/IL pair depending on the absorber temperature. The improvement factor by replacing R134a/IL with R134/IL is greater for the COP values than it is for CE because the pumping work, Wp, which dominates COP is two orders of magnitude smaller than the heat required at the desorber (Qd) and absorber (Qa). Wp is the only term in the denominator for computing COP. The R134/IL pair had less pumping work than the R134a/IL pair due to several thermophysical properties including: (i) larger molar enthalpy of vaporization (Table 5); (ii) smaller operating pressure range (Table 4); (iii) larger solubility and solubility difference in [bmim][PF6] (smaller liquid volumetric flow); and (iv) larger liquid density (Table 5). 4.2 Experimental Results The evaporator junction temperature, which reflects the operating temperature of an electronic device or other component requiring cooling, was maintained to be at 42 °C by controlling the heater power at the evaporator.

Data were collected at different desorber

temperatures by adjusting the heater power attached to the desorber. The desorber outlet temperature was measured prior to the separation of the refrigerant vapor produced. That is, it is an IL-vapor mixture combining states 3 and 8 in Figure 2. The evaporator junction temperature, evaporator outlet fluid temperature, separator outlet fluid temperature (state 8 in Figure 2), and desorber outlet fluid temperature at specific desorber power values are shown in Figure 10. The temperature of the solution leaving the separator and desorber increased, while the evaporator outlet temperature decreased with respect to the desorber heater power. This shows that more vapor refrigerant was generated at the desorber as a result of the higher desorber fluid


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temperature. The evaporator outlet temperature tends to converge to a specific temperature at high desorber temperature (higher desorber power). The CE had a parabolic shape, Figure 11, as predicted by the computational analysis due to the tradeoff between the desorber heater power and the cooling occurring at the evaporator. The experimental R134/[bmim][PF6] efficiency values were smaller than the theoretical values, as shown in Figure 5 due to experimental losses in heat at various locations. The experimental bench-top system was insulated but not optimized. The temperature and pressure sensors and excess piping added to accommodate their installation made the system larger and less insulated than an optimized system. In addition, the performance can be improved by adding a solution heat exchanger and improving the design of the microchannel absorber and desorber. Nevertheless, the experimental system results proved that the performance trends predicted in the model were correct and that a fluorocarbon/IL refrigerant pair can be made into a working system. The cooling capacity of the R134/IL and R134a/IL working pairs were compared, as shown in Figure 12. The R134/IL pair reached nearly the maximum cooling capacity at a relatively low separator temperature, 63.26°C. This is 1.9 times larger than that of the R134a/IL working pair. Therefore, the R134/IL pair forms a highly effective refrigerant pair for recycling low grade waste heat. The results are in good agreement with the EOS model when the measured temperature and pressure values are given as input parameters. The EOS model calculations show that R134/IL is predicted to have 1.82 times larger cooling capacity which is very close to the measured values.

5. Conclusion



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The R134/[bmim][PF6] fluid pair was evaluated in an energy recycling absorption refrigeration system. The EOS model showed that R134 had higher solubility difference in [bmim][PF6] than R134a in the same IL at the cooling cycle operating conditions, which resulted in higher CE values. The larger molar enthalpy of vaporization, smaller operating pressure range, and larger liquid density of R134 compared to R134a also contributed to the improvement in COP values. Experimental results confirmed the larger cooling capacity of the R134/[bmim][PF6] working fluid pair than the previously reported R134a /[bmim][PF6] pair. The maximum cooling capacity of the R134/[bmim][PF6] working fluid pair was reached at a relatively low separator fluid temperature (63.26°C), which would allow more effective waste heat utilization.

Acknowledgement: The financial support of the Interconnect Focus Center, one of six focus centers of the Semiconductor Research Corporation, is gratefully acknowledged. The authors would also like to thank Mark Evans and Nishith Patel for assistance in operating the bench-top absorption system and modeling of new fluids.


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Zhang, X.; Hu, D. Performance simulation of the absorption chiller using water and ionic

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Martin, A.; Bermejo, M. D. Thermodynamic analysis of absorption refrigeration cycles

using ionic liquid + supercritical CO2 pairs. J. Supercrit. Fluids 2010, 55 (2), 852-859.



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Kim, Y. J.; Kim, S.; Joshi, Y. K.; Fedorov, A. G.; Kohl, P. A. Thermodynamic analysis

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pump/refrigeration system utilizing ionic liquid and hydrofluorocarbon refrigerants. J. Electron. Packag. 2012, 134 (3) 031009-1-031009-9. 11.

Kim, S. Patel, N.; Kohl, P. A. Performance Simulation of Ionic Liquid and

Hydrofluorocarbon Working Fluids for an Absorption Refrigeration System. Ind. Eng. Chem. Res. 2013, 52 (19), 6329-6335. 12.

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absorption refrigeration cycle by the use of equations of state. Appl. Energy 2005, 80 (4), 383399. 14.

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Figure Captions 1. Structures of materials used. (a) 1-butyl-3-methylimidazolium hexafluorophosphate ([bmim][PF6]), (b) 1,1,2,2-tetrafluoroethane (R134), and (c) 1,1,1,2-tetrafluoroethane (R134a). 2. Schematic diagram of an absorption refrigeration system using IL/refrigerant mixture as a working fluid. 3. R134 solubility in [bmim][PF6] as a function of temperature (K) and pressure (MPa). Symbols: experimental solubility data22; lines: computed EOS model. 4. R134a solubility in [bmim][PF6] as a function of temperature (K) and pressure (MPa)10. Symbols: experimental solubility data23; lines: computed EOS model. 5. Effect of desorber outlet temperature on CE (without solution heat exchanger). Ta= 309.65 K. 6. Effect of desorber outlet temperature on CE (without solution heat exchanger). Ta= 300.65 K. 7. Effect of desorber outlet temperature on CE (with solution heat exchanger ε=0.35). Ta=309.65 K. 8. Effect of desorber outlet temperature on η (without solution heat exchanger). Ta= 309.65 K. 9. Effect of desorber outlet temperature on η (without solution heat exchanger). Ta= 300.65 K. 10. Experimentally measured desorber outlet fluid temperature, Separator outlet fluid temperature, evaporator junction temperature, and evaporator outlet temperature with respect to desorber power input for R134/[bmim][PF6] working fluid.


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11. Experimentally measured CE for R134/[bmim][PF6] working fluid with respect to desorber outlet fluid temperature. 12. Comparison of cooling capacity between R134/[bmim][PF6] and R134a/[bmim][PF6] working fluid pairs.



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Table 1. EOS constants of [bmim][PF6] and HFC refrigerants Pure compound [bmim][PF6]14 R13413

Tc (K) 860.5 391.97

Pc (kPa) 2645 4641

β0 1.0 1.0012

β1

β2

β3

0.62627 0.48291

0 -0.05070

0 0


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Table 2. Binary interaction parameters for R134 and [bmim][PF6]. Fluid pair a

l12 = l21

k12

τ12 (K)

R134/[bmim][PF6] 0.0418 -0.0451 24.4896 Standard deviations of non-linear least squares in pressures.


ΔP(MPa)a 0.002


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Table 3. Coefficients for ideal gas heat capacity of pure compounds [J⋅mol-1]. Pure compound [bmim][PF6]5 R13413

-1 C0 (J⋅mol )

-2.214 15.58

-1 -1 C1 (J⋅mol K )

-1 -2 C2 (J⋅mol K )

0.57685 0.28694

-3.854×10-4 -2.028×10-4


-1 -3 C3 (J⋅mol K )

9.785×10-8 5.39633×10-8

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Table 4. Vapor pressure of R13422 and R134a [kPa]25. Temperature [K]

R134

R134a

298.15

526.0

664.9

323.15

1062.9

1318.6



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Table 5. Enthalpy of vaporization (ΔHvap) and saturated liquid density (ρ) values of R134 and R134a at 298.15K. ΔHvap [kJ/mol]

ρ25,26 [g/cm3]

R134

16.33

1.290

R134a

14.64

1.207


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Figure 1. 165x73mm (150 x 150 DPI)



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Figure 2. 110x111mm (150 x 150 DPI)


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Figure 3. 147x109mm (150 x 135 DPI)



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Figure 4. 149x109mm (150 x 129 DPI)


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Figure 5. 149x98mm (133 x 124 DPI)



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Figure 6. 164x102mm (121 x 121 DPI)


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Figure 7. 150x85mm (150 x 150 DPI)



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Figure 8. 164x95mm (121 x 121 DPI)


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Figure 9. 164x95mm (121 x 121 DPI)



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Figure 10. 129x102mm (150 x 150 DPI)


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Figure 11. 122x90mm (150 x 150 DPI)



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Figure 12. 114x106mm (150 x 150 DPI)


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For Table of Contents Only 254x190mm (300 x 300 DPI)


Theoretical and Experimental Investigation of an Absorption

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