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Optimum operating conditions for a combined power and cooling thermodynamic cycle
APPLIED ENERGY
Applied Energy 84 (2007) 254–265
www.elsevier.com/locate/apenergy
Optimum operating conditions for a combined power and cooling thermodynamic cycle S.M. Sadrameli a
a,*,1
, D.Y. Goswami
b
Department of Chemical Engineering, University of Florida, Gainesville, FL 32611-6300, United States b Department of Mechanical and Aerospace Engineering, University of Florida, Gainesville, FL 32611-6300, United States Received 2 May 2006; received in revised form 2 August 2006; accepted 5 August 2006
Abstract The combined production of thermal power and cooling with an ammonia–water based cycle proposed by Goswami is under intensive investigation. In the cycle under consideration, simultaneous cooling output is produced by expanding an ammonia-rich vapor in an expander to sub-ambient temperatures and subsequently heating the cool exhaust. When this mechanism for cooling production is considered in detail, it is apparent that the cooling comes at some expense to work production. To optimize this trade-off, a very specific coefficient-of-performance has been defined. In this paper, the simulation of the cycle was carried out in the process simulator ASPEN Plus. The optimum operating conditions have been found by using the Equation Oriented mode of the simulator and some of the results have been compared with the experimental data obtained from the cycle. The agreement between the two sets proves the accuracy of the optimization results. Published by Elsevier Ltd. Keywords: Ammonia cycle; Cooling; Power; Aspen simulation; Optimization
1. Introduction Multi-component working fluids such as a binary ammonia–water mixture in power cycles exhibit variable boiling temperatures during the boiling process which makes them *
1
Corresponding author. E-mail addresses: [email protected] (S.M. Sadrameli), [email protected]fl.edu (D.Y. Goswami). On his leave from Tarbiat Modarres University, Tehran, Iran.
0306-2619/$ - see front matter Published by Elsevier Ltd. doi:10.1016/j.apenergy.2006.08.003
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255
Nomenclature COP Ecool Qcool Qh Rc/w W
coefficient of performance exergy of cooling cooling heat transfer heat input ratio of cooling to work net work-output from cycle
Greek g efficiency g1st Law cycle’s first-law efficiency definition gII ref second-law cooling production efficiency Subscript work opt optimized for work output only w/cool dual output conditions suitable for a sensible heat source. Due to the low temperature-difference between the heat source and the working fluid, a good thermal match between the source and the working fluid is allowed, such that less irreversibility occurs during the heat-addition process. This work has stemmed from commonalities between Kalina-type power cycles and aquaammonia absorption cooling; there is now a small group of proposed configurations [1– 4]. The cited advantages of a combined operation include a reduction in capital equipment by sharing of components and the possibility of improved resource-utilization compared to separate power-and-cooling systems [1,2]. A novel ammonia–water binary mixture thermodynamic cycle capable of producing power and refrigeration has been proposed Goswami [3] and is intended primarily for power production while simultaneously producing a cooling output. Fig. 1 presents the schematic diagram of the proposed cycle, where it is shown that the combined cooling output is gained from a heat exchanger following the expander. The expander exhaust is cooled by expanding the vapor to sub-ambient temperatures. Application of low heatsource temperatures bellow 200 °C is one of the characteristics of the new cycle in comparison with other dual-output concepts reported in the literature. Due to the dual outputs of thermal power and cooling, evaluation of the cycle’s performance has not been straightforward and requires more consideration. To account for the quality of the cooling output, which dictates the expander’s exhaust temperature, it was typically weighted in efficiency definitions [5]. In this work, an analysis of cooling production has led to a new measure for the effectiveness of cooling production with this cycle. This parameter is used as an objective function to optimize the combined power-and-cooling. 2. Background Ammonia–water mixture as a working fluid for the absorption-cooling cycle applications, not only has excellent thermo-physical properties, but is an environmentally-friendly material. It is also the best substitute for the CFCs, for solving global-warming problems.
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Superheater Coolant
Rectifier Heat Source
Separator Expander
Generator Heat Source Boiler
Cooling Heat Exchanger
Cooled Fluid
Recovery Heat Exchanger Throttle Absorber
Coolant
Solution Pump Fig. 1. Schematic of the power and cooling cycle.
The thermo-chemical compression system of absorption refrigeration cycles is used as a heat-driven alternative to the mechanical vapor-compression cycle. It has also been incorporated into a power cycle with the ammonia–water working-fluid pair. One of the early studies of an absorption-based power cycle was performed by Maloney and Robertson [6] who concluded no significant advantage for the configuration over steam-cycle operation at the conditions considered. A cycle proposed by Kalina [7] several decades later, is a well-known absorption power-cycle which is a superior bottoming-cycle option over steam Rankine cycles. Independent studies have been performed, for example [8,9], that concede some advantages to the Kalina cycle under certain conditions. The key benefit of an ammonia–water working fluid is its boiling temperature glide, which allows for a better thermal match with sensible heat sources and therefore reduces heat-transfer irreversibilities. However, this temperature glide also exists during the condensation process and can limit expansion in the expander, especially at low resource temperatures. Rogdakis and Antonopoulos [10] proposed to take advantage of the chemical affinity of ammonia–water and replace condensation with absorption–condensation. This greatly improved performance with low heat-source temperatures because it increased the amount of expansion that could take place across the expander [10]. This modification requires that only partial vaporization of the working fluid takes place in the boiler so that the remaining liquid can be used to absorb vapor in the absorber.
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Given this configuration, Goswami [3] recognized that, under certain conditions, the vapor could be expanded to temperatures below those at which absorption–condensation is taking place. The absorption–condensation process is taking place in the absorber by using a coolant liquid as shown in Fig. 1. This is an obvious departure from pure working fluid Rankine-cycle operation, where the limiting expander’s exhaust-temperature is the vapor condensation temperature, sub-cooling effects aside. The proposal [3] takes advantage of the favorable boiling characteristics of ammonia–water, and capitalizes on the possible sub-ambient expander exhaust temperatures to form the basis of a combined powerand-cooling thermodynamic cycle. Since the original proposal [3], theoretical and experimental investigations have taken place. Initial investigations were performed theoretically and they focused on identifying operating trends [11]. Later studies concluded that the cycle could be optimized for work or cooling outputs and also efficiency. Optimization studies were performed, optimizing on the basis of the first law, second law, and energy-efficiency definitions [12]. Minimum cooling temperatures [13], working fluid combinations, and system configurations [14] were also studied and optimized. An initial experimental study was conducted, which generally verified the expected boiling and absorption processes [15]. 3. Combined cooling/power cycle Referring to Fig. 1, the relatively strong basic solution of 40 wt% ammonia leaves the absorber as a saturated liquid at the cycle’s low-pressure and pumped to a high pressure via the solution pump. Before entering the boiler, the basic solution recovers heat from the returning weak ammonia liquid solution in the recovery heat-exchanger. As the boiler operates between the bubble and dew-point temperatures of the mixture at the system pressure, the basic solution is partially boiled to produce a two-phase mixture: a liquid, which is relatively weak in ammonia, and a vapor with a high concentration of ammonia. This two-phase mixture is separated in the separator, and the weak liquid is throttled back to the absorber. The vapor’s ammonia-concentration is increased by cooling and condensate separation in the rectifier. Heat can be added in the super heater as the vapor proceeds to the expander. The expander extracts energy from the high-pressure vapor as it is throttled to the system’s low-pressure. The vapor rejoins the weak liquid in the absorber where, with heat rejection, the basic solution is regenerated. Several studies on the evaluation of ammonia–water mixture properties have been reported in the literature. A convenient semi-empirical scheme is used here that combines the Gibbs free-energy method for mixtures and bubble and dew-point temperature correlations for the phase equilibrium. A comparison between the calculated results and experimental mixture’s properties in the literature has been made by Xu and Goswami [11] shows a good agreement between two sets of data. By employing absorption–condensation, the vapor can be expanded to temperatures significantly below the temperature at which absorption is taking place. Cooling can thus be obtained by sensibly heating the expander’s exhaust. This behavior is due to the fact that the working fluid is a binary mixture, and at constant pressure the condensing temperature of an ammonia-rich vapor can be below the saturation temperature for a lower-concentration liquid. This is best illustrated with a binary mixture, phase equilibrium diagram, Fig. 2. The low concentration saturated liquid state with 40 wt% ammonia, approximates the basic solution exiting the absorber, while the high concentration vapor is typically at the expander’s exhaust conditions.
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Pressure = 0.203 MPa
Vapor
120
100
Temperature [C]
Two-Phase 80
Liquid
60
Basic solution in absorber
40
20
Expander exhaust
0
-20 0
0.1
0.2
0.3
0.4
0.5
0.6
0.7
0.8
0.9
1
Ammonia Mass Fraction
Fig. 2. ammonia–water phase equilibrium diagram [18].
4. Cooling To understand the compromise between work and cooling production with this cycle, a detailed look at the mechanisms of cooling production are now discussed. In the case of sensible cooling production, that is no phase change of the refrigerant is occurring, the exhaust temperature from the expander is the key to determining the amount of cooling that will be achieved. Also, because the basic solution is only partially vaporized in the boiler, the quantity of vapor produced is variable and is another factor in determining the quantity of cooling. The sensitivities of each are now discussed. Some comments regarding the expander’s exhaust-temperature were discussed in relation to the binary phase-diagram of Fig. 2. Further insight can be gathered by considering the entropy of the working fluid at the expander’s exit. Minimization of the exhaust temperature also implies a minimization of the vapor entropy at the expander’s exhaust, assuming constant exit pressure. From this consideration, an efficient expander is an obvious feature for low temperatures (low exhaust-vapor entropy of the turbine), but even an ideal device would only maintain the vapor entropy from inlet to exhaust. Therefore, inlet conditions should also be considered. For an ammonia–water vapor mixture, entropy decreases with increasing pressure, increasing ammonia concentration, and decreasing temperature. The limit of these conditions, while still maintaining vapor, would be saturated pure-ammonia. Considering these preferred expander’s inlet-conditions, the function of the rectifier, is immediately apparent. In the rectifier, the vapor’s ammonia concentration is increased by removing the small amounts of water that have also been vaporized. This is generally performed by taking advantage of the saturation properties of the mixture, which in its simplest form is cooling and condensate separation. The net change to the vapor is an increase in concentration and a decrease in temperature, accompanied by a minor drop in pressure and some reduction to the mass flow rate. These effects are mostly to the advantage of the cooler’s expander exhaust-temperatures.
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1
Vapor Concentration [kg/kg]
0.9
0.8
Saturated liquid conditions
0.7
Complete vaporization 0.6
0.5
0.4
0.3 0
0.1
0.2
0.3
0.4
0.5
0.6
0.7
0.8
0.9
1
Vapor Mass Fraction (kg/kg)
Fig. 3. Relationship between vapor mass fraction (kg/kg) and amount of partial vaporization.
The considerations for exhaust temperatures are now coupled with the mechanisms of vapor production. The working fluid was partially vaporized in the boiler and separated by phase in the separator. From initial to complete vaporization, the boiling process proceeds as indicated in Fig. 3, which is a plot of vapor concentration as a function of the vapor mass-flow fraction (the ratio of the vapor mass-flow to the basic solution massflow). As shown, with minimal vaporization, the concentration is highest, while at high amounts of vapor production, the concentration approaches the basic solution concentration (0.4 ammonia (kg/kg) in this case). From the previous discussion of expander inlet conditions, high concentrations are preferred, which implies low vaporization rates. These results imply that, for cooling production the partial-boiling operation should approximate to a distillation process separating ammonia from the working-fluid mixture. These are the same general requirements as of an aqua-ammonia absorption cooling cycle; however, they are in contrast to Rankine-based power production where the production of vapor, regardless of composition, is a critical consideration. 5. Cycle performance The fact that this cycle has dual outputs of power and cooling has raised questions regarding the evaluation of its performance. Central to these is the treatment of the quality of both the work and cooling outputs. Previous studies in this area are now discussed along with a description of the evaluation used in this study. 5.1. Prior work The question of appropriate efficiency expressions for the cycle was examined by Vijayaraghavan and Goswami [5]. It was noticed that the results obtained from an optimization
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of the cycle were influenced by the weight given to the cooling output in the objective function, which was typically an efficiency definition. In its simplest form, a first-law efficiency could be defined, where the work and cooling outputs of the cycle are added together as useful outputs, g1st Law ¼
ðW net þ Qcool Þ Qh
ð1Þ
However, by not accounting for the quality of the cooling output, Eq. (1) gives an overestimate of system’s performance [5]. The quality of the cooling output can be evaluated by using a COP to find a work equivalent for the cooling produced, g1st Law ¼
ðW net þ ðEcool =gII ref ÞÞ Qh
ð2Þ
The Carnot or Lorenz theoretical-efficiency, based on the appropriate temperature limits, could be used for this purpose. However, this would compute the theoretical minimum amount of work needed to produce the cooling, and thus greatly undervalues the cooling output [5]. A practically-obtainable COP value could be substituted; however, there is a subjective element to this approach. The authors concede that ultimately the value of work and cooling, that is the weighting, will be best decided by the end application [5]. 5.2. Effective COP In this study, a different approach is used to evaluate the effectiveness of combined cooling production. The discussion of cooling-production mechanisms revealed that some compromises to work production are needed in order to obtain a simultaneous cooling output. For instance, conditioning of the vapor to accommodate cooling typically reduces the amount of vapor entering the expander, which reduces the amount of work done. A new coefficient of performance, specific to this cycle, is proposed based on the idea that a compromise is needed for the dual outputs of power and cooling. This COP definition relates the cooling produced with this cycle to the theoretical amount of work production that was compromised in order to have combined-cooling production. The general formulation is provided as follows: COPeffective ¼
Cooling produced Potential-work lost
ð3Þ
This term is referred to as the effective COP since cooling and work are only indirectly related, in other words, there is no device directly producing cooling with the work that is not produced. Rather, the effective COP is a measure of the compromise made to accommodate a combined cooling output. Just as with conventional COP values, higher values of effective COP are desired since they maximize the cooling production for the penalty of less work production. The formulation with cycle parameters is presented as Eq. (4). In evaluating the lost potential-work, the difference between the work produced with cooling and the ideal work production is used. The ideal work production, Wwork opt, is the work output for a work-optimized (based on a first-law efficiency, but giving no value to cooling) power-cooling cycle operating between equivalent heat-source and rejection conditions.
S.M. Sadrameli, D.Y. Goswami / Applied Energy 84 (2007) 254–265
COPeffective ¼
Rc=w Qcool ¼ g work opt W work opt W w=cool 1
261
ð4Þ
gw=cool
The rightmost expression in Eq. (4) is an equivalent way to formulate the effective COP and it was used to evaluate the presented results. 6. Modeling The modeling for this work was based on the connection schematic of Fig. 1 and was performed using the Equation Oriented (EO) mode in ASPEN Plus, version 12.1 [16]. EO modeling is a different strategy for solving flow-sheet simulations in Aspen Plus, which is a very effective way of solving certain kinds of problems, such as process optimization. In this approach instead of solving each block of simulation in sequence (Sequential Modular, SM), EO gathers all the model equations together and solves them at the same time. The PSRK (Predictive Soave–Redlich–Kwong) equation-of-state (EQS) for the property algorithm within Aspen Plus was used to compute the thermodynamic properties of the ammonia–water working-fluid. This EQS formulation gave predictions that agreed with IGT (Institute of Gas Technology) experimentally-measured data [17]. Some of the simulation runs performed before optimization to for the verification of the developed model have been compared with the experimental data reported by Martin [18]. Averaged conditions for the experimental and simulated runs are as follows: Expander’s inlet-pressure = 0.0516 MPa Expander’s exit-pressure = 0.2080 MPa Vapor concentration = 0.9930 kg/kg Solution pump flow-rate = 0.0020 kg/s Noted procedures and constraints for the optimization were as follows: The boiling temperature and pressure were specified. The basic solution concentration was determined by specifying the pump’s inlet-temperature and pressure along with the assumption of a saturated liquid. The vaporization fraction (vapor quality at the boiler’s exit) was determined from saturation conditions. The absorber’s exit-temperature (pump inlet) was assumed to be approximately 5 °C higher than the assumed ambient temperature of 25 °C. Vapor rectification was limited by a specified minimum rectifier exit-temperature of 35 °C. The minimum amount of vapor leaving the rectifier was allowed to be 5% of the basic solution flow-rate. The quantity of cooling produced (if any) was calculated as the heat needed to raise the expander’s exhaust-temperature to the ambient temperature of 25 °C. The model used for this work has been simplified slightly when compared with the model used by Vijayaraghavan and Goswami [5]. In that work, the cooling needed in the rectifier was provided by a diverted portion of the basic solution stream from the absorber, thus providing some heat recovery. In this work, the rejected heat from the rectifier was judged to be small compared with the boiler’s input and has been rejected to an external cooling source.
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7. Results Fig. 4 shows the ASPEN Plus flow-sheet for the cycle simulation. Rectifier and absorber were considered as simple separation columns with heating and cooling utilities, to be identical to the distillation columns in the simulation. The following parameters and assumptions were used in the simulations: isentropic efficiencies of the turbine, and pump 80%, mechanical and electric efficiency 96%, a minimum temperature-difference for each of the heat exchangers of 10 °C, ambient temperature 25 °C, ambient pressure 1.0132 bar, and finally rectifier and separator temperatures of 39 °C and 83.5 °C, respectively. Table 1 shows one of the results for the turbine’s inlet-temperature of 39.1 °C. The comparison between the simulated and experimental results [18] for the turbine’s inlet-temperatures of 39.1 °C, 40.1 °C, 42.2 °C, and 47.9 °C are shown in Table 2. The agreement between two sets of the results validates the accuracy of the developed simulation model. An optimization of the cycle conditions was performed for a boiler’s exit-temperature of 122 °C and using the effective COP, Eq. (4), as the objective function. For a fixed value of the pump’s inlet-pressure, the values of Rc/w (parameter in Eq. (4) have been calculated for different values of the pump’s outlet-pressure. The calculations were repeated for different pumpinlet pressures from 2.5 to 5.0 bars until the optimum conditions were reached. The results are shown in the contour plot of Fig. 5. As can be seen from the figure, the maximum value of the effective COP was reached at 0.647, with a basic ammonia concentration of 0.554. The optimum conditions obtained are as follows: Boiling pressure = 2.3 MPa Absorber pressure = 0.42 MPa
Fig. 4. Major components in the Aspen Plus [16] process-model for the ammonia cycle.
Temperature (°C) Pressure (bar) Vapor frac (kg/kg) Mole flow (kmol/h) Mass flow (kg/h) Ammonia (mol%) Water (mol%) Ammonia (kmole/h) Water (kmole/h)
1
2
3
4
5
6
7
8
HE-O
Pump-O
Rec-O
Valv-O
Valv1-O
31.74 1.81 0.00 0.41 7.31 0.24 0.76 0.10 0.31
68.67 5.16 0.00 0.41 7.31 0.24 0.76 0.10 0.31
84.75 4.66 0.02 0.41 7.31 0.24 0.76 0.10 0.31
83.50 4.66 0.00 0.41 7.25 0.23 0.77 0.10 0.31
83.50 4.66 1.00 0.00 0.07 0.91 0.09 0.00 0.00
39.10 4.45 1.00 0.00 0.06 0.99 0.01 0.00 0.00
22.18 2.08 1.00 0.00 0.06 0.99 0.01 0.00 0.00
25.00 2.08 1.00 0.00 0.06 0.99 0.01 0.00 0.00
46.23 4.66 0.00 0.41 7.25 0.23 0.77 0.10 0.31
31.80 5.16 0.00 0.41 7.31 0.24 0.76 0.10 0.31
39.10 4.45 0.00 0.00 0.01 0.49 0.51 0.00 0.00
38.05 4.28 0.00 0.00 0.01 0.49 0.51 0.00 0.00
46.29 2.08 0.00 0.41 7.25 0.23 0.77 0.10 0.31
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Table 1 Simulation results by Aspen Plus [16] for one of the operating conditions
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Table 2 Comparison of the simulated and experimental operating-conditions Operating Condition
Turbine inlet (°C)
Absorber exit (°C)
Turbine exit (°C)
Boiler out (°C)
Boiler in (°C)
Separator liquid out (°C)
Separator vapor out (°C)
Experimental Modeling
39.1 39.1
31.74 31.7
21.95 22.2
84.75 84.8
68.67 68.7
83.18 83.5
84.24 83.5
Experimental Modeling
40.1 40.1
31.4 31.4
23.4 24.5
85.3 85.3
68.6 68.6
82.5 83
83.8 83
Experimental Modeling
42.2 42.1
31.76 31.8
24.5 24.1
83.7 83.7
65.0 65.1
81.9 82.5
83.2 82.5
Experimental Modeling
47.9 47.9
32.1 32.1
26.9 27.2
84.31 84.3
67.7 67.7
82.43 83
83.5 83
2.4 Effective COP
Boiling Pressure [MPa]
2.2
0.4 0.64
2 1.8
0.60
1.6 0.55 1.4
0.5 Eff COP optimum, 0.647 [5] 30 η II ref = 30% [5]1st optimum, [5] 50 η II ref = 50% [5]1st optimum,
1.2 0.40 1 0.345
0.395
0.445
0.495
0.545
Basic Solution Concentration [kg/kg]
Fig. 5. Operation map showing optimized region and comparison points (COPopt = 0.647 at a concentration of 0.554, expander’s outlet-temperature of 0.1 °C and pump pressure of 2.3 MPa.).
Basic concentration (ammonia wt%) = 0.554 Expander’s inlet-temperature = 69.0 °C Expander’s exit-temperature = 0.10 °C Boiler’s vapor fraction = 0.325 Ratio of cooling to work = 0.888 These results were obtained by considering a similar configuration (discussed in the modeling section) of the power and cooling cycle, simulating with equivalent heat-source and absorber conditions, but using a combined first-law efficiency definition, Eq. (2), as the objective function [5]. 8. Conclusions A thermodynamic analysis of a combined cooling-and-power cycle was carried out using an EO mode of simulation by ASPEN Plus 12.1 [16]. The results show that the con-
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ditions for the optimum cooling production are at odds with the power production. In this study, the trade-off of work to accommodate cooling is quantified by defining an effective COP. The defined parameter has been used as an objective function to optimize the cooling production. It appears that the effective COP optimum is biased towards higher expander exhaust-temperatures but with increased vapor flow. Some of the simulation runs have been performed before optimization for the verification of the developed model. The agreement between the simulation and experimental results proves the accuracy of the developed model. Finally, this analysis is not sufficient to determine the optimum parameters for the operating conditions. The optimum conditions have to be tested experimentally for the final verification of the results. Furthermore, it is not clear what conditions would favor a particular objective function. Exchanging the work-optimized cycle-efficiency in Eq. (4) with an analytic definition, such as Carnot efficiency, would aid in an analytical comparison of the effective COP with other performance evaluations. References [1] Erickson DC, Anand G, Kyung I. Heat-activated dual-function absorption cycle. ASHRAE Trans 2004;110(1):515–24. [2] Amano Y, Suzuki T, Hashizume T, Akiba M, Tanzawa Y, Usui A. A hybrid power-generation and refrigeration cycle with ammonia–water mixture, IJPGC2000-15058. In: Proceedings of AD 2000 joint power generation conference. ASME; 2000. [3] Goswami DY. Solar thermal power: status of technologies and opportunities for research, heat-and-mass transfer ’95. In: Proceedings of the 2nd ASME-ISHMT heat-and-mass transfer conference. New Delhi (India): Tata-McGraw Hill Publishers; 1995. p. 57–60. [4] Zheng D, Chen B, Qi Y. Thermodynamic analysis of a novel absorption power/cooling combined cycle. In: Proceedings of the 2002 international sorption heat-pump conference. Shanghai, China; 2002. p. 204–9. [5] Vijayaraghavan S, Goswami DY. On evaluating efficiency of a combined power-and-cooling cycle. In: Proceedings of IMECE 2002. New Orleans (LA): American Society of Mechanical Engineers; 2002. [6] Maloney JD, Robertson RC. Thermodynamic study of ammonia–water heat and power cycles. ORNL Report CF-53-8-43, Oak Ridge (TN); 1953. [7] Kalina AI. Combined-cycle system with novel bottoming cycle. ASME J Eng Gas Turb Power 1984;106:737–42. [8] Marston CH. Parametric analysis of the Kalina cycle. J Eng Gas Turb Power 1990;112:107–16. [9] Ibrahim OM, Klein SA. Absorption power cycles. Energy 1996;21(1):21–7. [10] Rogdakis ED, Antonopoulos KA. A high efficiency NH3/H2O absorption power-cycle. Heat Recov Syst CHP 1991;2:263–75. [11] Goswami DY, Xu F. Analysis of a new thermodynamic cycle for combined power-and-cooling using lowand mid-temperature solar collectors. J Solar Energ Eng 1999;121:91–7. [12] Lu S, Goswami DY. Optimization of a novel combined power/refrigeration thermodynamic cycle. J Solar Energ Eng 2003;125:212–7. [13] Lu S, Goswami DY. Theoretical analysis of ammonia-based combined power/refrigeration cycle at low refrigeration-temperatures. In: Proceedings of SOLAR 2002. Reno (NV): American Society of Mechanical Engineers; 2002. [14] Vijayaraghavan S. Thermodynamic studies on alternate binary working-fluid combinations and configurations for a combined power-and-cooling cycle. Ph.D. dissertation, University of Florida, Gainesville (FL); 2003. [15] Tamm G, Goswami DY. Novel combined power-and-cooling thermodynamic cycle for low-temperature heat sources, part II: Experimental investigation. J Solar Energ Eng 2003;125:223–9. [16] Aspen Plus, Version 12.1. Ten Canal Park, Cambridge (MA) 02141: Aspen Technology, Inc; 2004. [17] Macriss R, Eakin B, Ellington R, Huebler J. Physical and thermodynamic properties of ammonia–water mixtures. Chicago (IL): Institute of Gas Technology; 1964. [18] Martin CL. PhD thesis, Mechanical Engineering Department, University of Florida; 2004.
Applied Energy 84 (2007) 254–265
www.elsevier.com/locate/apenergy
Optimum operating conditions for a combined power and cooling thermodynamic cycle S.M. Sadrameli a
a,*,1
, D.Y. Goswami
b
Department of Chemical Engineering, University of Florida, Gainesville, FL 32611-6300, United States b Department of Mechanical and Aerospace Engineering, University of Florida, Gainesville, FL 32611-6300, United States Received 2 May 2006; received in revised form 2 August 2006; accepted 5 August 2006
Abstract The combined production of thermal power and cooling with an ammonia–water based cycle proposed by Goswami is under intensive investigation. In the cycle under consideration, simultaneous cooling output is produced by expanding an ammonia-rich vapor in an expander to sub-ambient temperatures and subsequently heating the cool exhaust. When this mechanism for cooling production is considered in detail, it is apparent that the cooling comes at some expense to work production. To optimize this trade-off, a very specific coefficient-of-performance has been defined. In this paper, the simulation of the cycle was carried out in the process simulator ASPEN Plus. The optimum operating conditions have been found by using the Equation Oriented mode of the simulator and some of the results have been compared with the experimental data obtained from the cycle. The agreement between the two sets proves the accuracy of the optimization results. Published by Elsevier Ltd. Keywords: Ammonia cycle; Cooling; Power; Aspen simulation; Optimization
1. Introduction Multi-component working fluids such as a binary ammonia–water mixture in power cycles exhibit variable boiling temperatures during the boiling process which makes them *
1
Corresponding author. E-mail addresses: [email protected] (S.M. Sadrameli), [email protected]fl.edu (D.Y. Goswami). On his leave from Tarbiat Modarres University, Tehran, Iran.
0306-2619/$ - see front matter Published by Elsevier Ltd. doi:10.1016/j.apenergy.2006.08.003
S.M. Sadrameli, D.Y. Goswami / Applied Energy 84 (2007) 254–265
255
Nomenclature COP Ecool Qcool Qh Rc/w W
coefficient of performance exergy of cooling cooling heat transfer heat input ratio of cooling to work net work-output from cycle
Greek g efficiency g1st Law cycle’s first-law efficiency definition gII ref second-law cooling production efficiency Subscript work opt optimized for work output only w/cool dual output conditions suitable for a sensible heat source. Due to the low temperature-difference between the heat source and the working fluid, a good thermal match between the source and the working fluid is allowed, such that less irreversibility occurs during the heat-addition process. This work has stemmed from commonalities between Kalina-type power cycles and aquaammonia absorption cooling; there is now a small group of proposed configurations [1– 4]. The cited advantages of a combined operation include a reduction in capital equipment by sharing of components and the possibility of improved resource-utilization compared to separate power-and-cooling systems [1,2]. A novel ammonia–water binary mixture thermodynamic cycle capable of producing power and refrigeration has been proposed Goswami [3] and is intended primarily for power production while simultaneously producing a cooling output. Fig. 1 presents the schematic diagram of the proposed cycle, where it is shown that the combined cooling output is gained from a heat exchanger following the expander. The expander exhaust is cooled by expanding the vapor to sub-ambient temperatures. Application of low heatsource temperatures bellow 200 °C is one of the characteristics of the new cycle in comparison with other dual-output concepts reported in the literature. Due to the dual outputs of thermal power and cooling, evaluation of the cycle’s performance has not been straightforward and requires more consideration. To account for the quality of the cooling output, which dictates the expander’s exhaust temperature, it was typically weighted in efficiency definitions [5]. In this work, an analysis of cooling production has led to a new measure for the effectiveness of cooling production with this cycle. This parameter is used as an objective function to optimize the combined power-and-cooling. 2. Background Ammonia–water mixture as a working fluid for the absorption-cooling cycle applications, not only has excellent thermo-physical properties, but is an environmentally-friendly material. It is also the best substitute for the CFCs, for solving global-warming problems.
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Superheater Coolant
Rectifier Heat Source
Separator Expander
Generator Heat Source Boiler
Cooling Heat Exchanger
Cooled Fluid
Recovery Heat Exchanger Throttle Absorber
Coolant
Solution Pump Fig. 1. Schematic of the power and cooling cycle.
The thermo-chemical compression system of absorption refrigeration cycles is used as a heat-driven alternative to the mechanical vapor-compression cycle. It has also been incorporated into a power cycle with the ammonia–water working-fluid pair. One of the early studies of an absorption-based power cycle was performed by Maloney and Robertson [6] who concluded no significant advantage for the configuration over steam-cycle operation at the conditions considered. A cycle proposed by Kalina [7] several decades later, is a well-known absorption power-cycle which is a superior bottoming-cycle option over steam Rankine cycles. Independent studies have been performed, for example [8,9], that concede some advantages to the Kalina cycle under certain conditions. The key benefit of an ammonia–water working fluid is its boiling temperature glide, which allows for a better thermal match with sensible heat sources and therefore reduces heat-transfer irreversibilities. However, this temperature glide also exists during the condensation process and can limit expansion in the expander, especially at low resource temperatures. Rogdakis and Antonopoulos [10] proposed to take advantage of the chemical affinity of ammonia–water and replace condensation with absorption–condensation. This greatly improved performance with low heat-source temperatures because it increased the amount of expansion that could take place across the expander [10]. This modification requires that only partial vaporization of the working fluid takes place in the boiler so that the remaining liquid can be used to absorb vapor in the absorber.
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Given this configuration, Goswami [3] recognized that, under certain conditions, the vapor could be expanded to temperatures below those at which absorption–condensation is taking place. The absorption–condensation process is taking place in the absorber by using a coolant liquid as shown in Fig. 1. This is an obvious departure from pure working fluid Rankine-cycle operation, where the limiting expander’s exhaust-temperature is the vapor condensation temperature, sub-cooling effects aside. The proposal [3] takes advantage of the favorable boiling characteristics of ammonia–water, and capitalizes on the possible sub-ambient expander exhaust temperatures to form the basis of a combined powerand-cooling thermodynamic cycle. Since the original proposal [3], theoretical and experimental investigations have taken place. Initial investigations were performed theoretically and they focused on identifying operating trends [11]. Later studies concluded that the cycle could be optimized for work or cooling outputs and also efficiency. Optimization studies were performed, optimizing on the basis of the first law, second law, and energy-efficiency definitions [12]. Minimum cooling temperatures [13], working fluid combinations, and system configurations [14] were also studied and optimized. An initial experimental study was conducted, which generally verified the expected boiling and absorption processes [15]. 3. Combined cooling/power cycle Referring to Fig. 1, the relatively strong basic solution of 40 wt% ammonia leaves the absorber as a saturated liquid at the cycle’s low-pressure and pumped to a high pressure via the solution pump. Before entering the boiler, the basic solution recovers heat from the returning weak ammonia liquid solution in the recovery heat-exchanger. As the boiler operates between the bubble and dew-point temperatures of the mixture at the system pressure, the basic solution is partially boiled to produce a two-phase mixture: a liquid, which is relatively weak in ammonia, and a vapor with a high concentration of ammonia. This two-phase mixture is separated in the separator, and the weak liquid is throttled back to the absorber. The vapor’s ammonia-concentration is increased by cooling and condensate separation in the rectifier. Heat can be added in the super heater as the vapor proceeds to the expander. The expander extracts energy from the high-pressure vapor as it is throttled to the system’s low-pressure. The vapor rejoins the weak liquid in the absorber where, with heat rejection, the basic solution is regenerated. Several studies on the evaluation of ammonia–water mixture properties have been reported in the literature. A convenient semi-empirical scheme is used here that combines the Gibbs free-energy method for mixtures and bubble and dew-point temperature correlations for the phase equilibrium. A comparison between the calculated results and experimental mixture’s properties in the literature has been made by Xu and Goswami [11] shows a good agreement between two sets of data. By employing absorption–condensation, the vapor can be expanded to temperatures significantly below the temperature at which absorption is taking place. Cooling can thus be obtained by sensibly heating the expander’s exhaust. This behavior is due to the fact that the working fluid is a binary mixture, and at constant pressure the condensing temperature of an ammonia-rich vapor can be below the saturation temperature for a lower-concentration liquid. This is best illustrated with a binary mixture, phase equilibrium diagram, Fig. 2. The low concentration saturated liquid state with 40 wt% ammonia, approximates the basic solution exiting the absorber, while the high concentration vapor is typically at the expander’s exhaust conditions.
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Pressure = 0.203 MPa
Vapor
120
100
Temperature [C]
Two-Phase 80
Liquid
60
Basic solution in absorber
40
20
Expander exhaust
0
-20 0
0.1
0.2
0.3
0.4
0.5
0.6
0.7
0.8
0.9
1
Ammonia Mass Fraction
Fig. 2. ammonia–water phase equilibrium diagram [18].
4. Cooling To understand the compromise between work and cooling production with this cycle, a detailed look at the mechanisms of cooling production are now discussed. In the case of sensible cooling production, that is no phase change of the refrigerant is occurring, the exhaust temperature from the expander is the key to determining the amount of cooling that will be achieved. Also, because the basic solution is only partially vaporized in the boiler, the quantity of vapor produced is variable and is another factor in determining the quantity of cooling. The sensitivities of each are now discussed. Some comments regarding the expander’s exhaust-temperature were discussed in relation to the binary phase-diagram of Fig. 2. Further insight can be gathered by considering the entropy of the working fluid at the expander’s exit. Minimization of the exhaust temperature also implies a minimization of the vapor entropy at the expander’s exhaust, assuming constant exit pressure. From this consideration, an efficient expander is an obvious feature for low temperatures (low exhaust-vapor entropy of the turbine), but even an ideal device would only maintain the vapor entropy from inlet to exhaust. Therefore, inlet conditions should also be considered. For an ammonia–water vapor mixture, entropy decreases with increasing pressure, increasing ammonia concentration, and decreasing temperature. The limit of these conditions, while still maintaining vapor, would be saturated pure-ammonia. Considering these preferred expander’s inlet-conditions, the function of the rectifier, is immediately apparent. In the rectifier, the vapor’s ammonia concentration is increased by removing the small amounts of water that have also been vaporized. This is generally performed by taking advantage of the saturation properties of the mixture, which in its simplest form is cooling and condensate separation. The net change to the vapor is an increase in concentration and a decrease in temperature, accompanied by a minor drop in pressure and some reduction to the mass flow rate. These effects are mostly to the advantage of the cooler’s expander exhaust-temperatures.
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1
Vapor Concentration [kg/kg]
0.9
0.8
Saturated liquid conditions
0.7
Complete vaporization 0.6
0.5
0.4
0.3 0
0.1
0.2
0.3
0.4
0.5
0.6
0.7
0.8
0.9
1
Vapor Mass Fraction (kg/kg)
Fig. 3. Relationship between vapor mass fraction (kg/kg) and amount of partial vaporization.
The considerations for exhaust temperatures are now coupled with the mechanisms of vapor production. The working fluid was partially vaporized in the boiler and separated by phase in the separator. From initial to complete vaporization, the boiling process proceeds as indicated in Fig. 3, which is a plot of vapor concentration as a function of the vapor mass-flow fraction (the ratio of the vapor mass-flow to the basic solution massflow). As shown, with minimal vaporization, the concentration is highest, while at high amounts of vapor production, the concentration approaches the basic solution concentration (0.4 ammonia (kg/kg) in this case). From the previous discussion of expander inlet conditions, high concentrations are preferred, which implies low vaporization rates. These results imply that, for cooling production the partial-boiling operation should approximate to a distillation process separating ammonia from the working-fluid mixture. These are the same general requirements as of an aqua-ammonia absorption cooling cycle; however, they are in contrast to Rankine-based power production where the production of vapor, regardless of composition, is a critical consideration. 5. Cycle performance The fact that this cycle has dual outputs of power and cooling has raised questions regarding the evaluation of its performance. Central to these is the treatment of the quality of both the work and cooling outputs. Previous studies in this area are now discussed along with a description of the evaluation used in this study. 5.1. Prior work The question of appropriate efficiency expressions for the cycle was examined by Vijayaraghavan and Goswami [5]. It was noticed that the results obtained from an optimization
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of the cycle were influenced by the weight given to the cooling output in the objective function, which was typically an efficiency definition. In its simplest form, a first-law efficiency could be defined, where the work and cooling outputs of the cycle are added together as useful outputs, g1st Law ¼
ðW net þ Qcool Þ Qh
ð1Þ
However, by not accounting for the quality of the cooling output, Eq. (1) gives an overestimate of system’s performance [5]. The quality of the cooling output can be evaluated by using a COP to find a work equivalent for the cooling produced, g1st Law ¼
ðW net þ ðEcool =gII ref ÞÞ Qh
ð2Þ
The Carnot or Lorenz theoretical-efficiency, based on the appropriate temperature limits, could be used for this purpose. However, this would compute the theoretical minimum amount of work needed to produce the cooling, and thus greatly undervalues the cooling output [5]. A practically-obtainable COP value could be substituted; however, there is a subjective element to this approach. The authors concede that ultimately the value of work and cooling, that is the weighting, will be best decided by the end application [5]. 5.2. Effective COP In this study, a different approach is used to evaluate the effectiveness of combined cooling production. The discussion of cooling-production mechanisms revealed that some compromises to work production are needed in order to obtain a simultaneous cooling output. For instance, conditioning of the vapor to accommodate cooling typically reduces the amount of vapor entering the expander, which reduces the amount of work done. A new coefficient of performance, specific to this cycle, is proposed based on the idea that a compromise is needed for the dual outputs of power and cooling. This COP definition relates the cooling produced with this cycle to the theoretical amount of work production that was compromised in order to have combined-cooling production. The general formulation is provided as follows: COPeffective ¼
Cooling produced Potential-work lost
ð3Þ
This term is referred to as the effective COP since cooling and work are only indirectly related, in other words, there is no device directly producing cooling with the work that is not produced. Rather, the effective COP is a measure of the compromise made to accommodate a combined cooling output. Just as with conventional COP values, higher values of effective COP are desired since they maximize the cooling production for the penalty of less work production. The formulation with cycle parameters is presented as Eq. (4). In evaluating the lost potential-work, the difference between the work produced with cooling and the ideal work production is used. The ideal work production, Wwork opt, is the work output for a work-optimized (based on a first-law efficiency, but giving no value to cooling) power-cooling cycle operating between equivalent heat-source and rejection conditions.
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COPeffective ¼
Rc=w Qcool ¼ g work opt W work opt W w=cool 1
261
ð4Þ
gw=cool
The rightmost expression in Eq. (4) is an equivalent way to formulate the effective COP and it was used to evaluate the presented results. 6. Modeling The modeling for this work was based on the connection schematic of Fig. 1 and was performed using the Equation Oriented (EO) mode in ASPEN Plus, version 12.1 [16]. EO modeling is a different strategy for solving flow-sheet simulations in Aspen Plus, which is a very effective way of solving certain kinds of problems, such as process optimization. In this approach instead of solving each block of simulation in sequence (Sequential Modular, SM), EO gathers all the model equations together and solves them at the same time. The PSRK (Predictive Soave–Redlich–Kwong) equation-of-state (EQS) for the property algorithm within Aspen Plus was used to compute the thermodynamic properties of the ammonia–water working-fluid. This EQS formulation gave predictions that agreed with IGT (Institute of Gas Technology) experimentally-measured data [17]. Some of the simulation runs performed before optimization to for the verification of the developed model have been compared with the experimental data reported by Martin [18]. Averaged conditions for the experimental and simulated runs are as follows: Expander’s inlet-pressure = 0.0516 MPa Expander’s exit-pressure = 0.2080 MPa Vapor concentration = 0.9930 kg/kg Solution pump flow-rate = 0.0020 kg/s Noted procedures and constraints for the optimization were as follows: The boiling temperature and pressure were specified. The basic solution concentration was determined by specifying the pump’s inlet-temperature and pressure along with the assumption of a saturated liquid. The vaporization fraction (vapor quality at the boiler’s exit) was determined from saturation conditions. The absorber’s exit-temperature (pump inlet) was assumed to be approximately 5 °C higher than the assumed ambient temperature of 25 °C. Vapor rectification was limited by a specified minimum rectifier exit-temperature of 35 °C. The minimum amount of vapor leaving the rectifier was allowed to be 5% of the basic solution flow-rate. The quantity of cooling produced (if any) was calculated as the heat needed to raise the expander’s exhaust-temperature to the ambient temperature of 25 °C. The model used for this work has been simplified slightly when compared with the model used by Vijayaraghavan and Goswami [5]. In that work, the cooling needed in the rectifier was provided by a diverted portion of the basic solution stream from the absorber, thus providing some heat recovery. In this work, the rejected heat from the rectifier was judged to be small compared with the boiler’s input and has been rejected to an external cooling source.
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7. Results Fig. 4 shows the ASPEN Plus flow-sheet for the cycle simulation. Rectifier and absorber were considered as simple separation columns with heating and cooling utilities, to be identical to the distillation columns in the simulation. The following parameters and assumptions were used in the simulations: isentropic efficiencies of the turbine, and pump 80%, mechanical and electric efficiency 96%, a minimum temperature-difference for each of the heat exchangers of 10 °C, ambient temperature 25 °C, ambient pressure 1.0132 bar, and finally rectifier and separator temperatures of 39 °C and 83.5 °C, respectively. Table 1 shows one of the results for the turbine’s inlet-temperature of 39.1 °C. The comparison between the simulated and experimental results [18] for the turbine’s inlet-temperatures of 39.1 °C, 40.1 °C, 42.2 °C, and 47.9 °C are shown in Table 2. The agreement between two sets of the results validates the accuracy of the developed simulation model. An optimization of the cycle conditions was performed for a boiler’s exit-temperature of 122 °C and using the effective COP, Eq. (4), as the objective function. For a fixed value of the pump’s inlet-pressure, the values of Rc/w (parameter in Eq. (4) have been calculated for different values of the pump’s outlet-pressure. The calculations were repeated for different pumpinlet pressures from 2.5 to 5.0 bars until the optimum conditions were reached. The results are shown in the contour plot of Fig. 5. As can be seen from the figure, the maximum value of the effective COP was reached at 0.647, with a basic ammonia concentration of 0.554. The optimum conditions obtained are as follows: Boiling pressure = 2.3 MPa Absorber pressure = 0.42 MPa
Fig. 4. Major components in the Aspen Plus [16] process-model for the ammonia cycle.
Temperature (°C) Pressure (bar) Vapor frac (kg/kg) Mole flow (kmol/h) Mass flow (kg/h) Ammonia (mol%) Water (mol%) Ammonia (kmole/h) Water (kmole/h)
1
2
3
4
5
6
7
8
HE-O
Pump-O
Rec-O
Valv-O
Valv1-O
31.74 1.81 0.00 0.41 7.31 0.24 0.76 0.10 0.31
68.67 5.16 0.00 0.41 7.31 0.24 0.76 0.10 0.31
84.75 4.66 0.02 0.41 7.31 0.24 0.76 0.10 0.31
83.50 4.66 0.00 0.41 7.25 0.23 0.77 0.10 0.31
83.50 4.66 1.00 0.00 0.07 0.91 0.09 0.00 0.00
39.10 4.45 1.00 0.00 0.06 0.99 0.01 0.00 0.00
22.18 2.08 1.00 0.00 0.06 0.99 0.01 0.00 0.00
25.00 2.08 1.00 0.00 0.06 0.99 0.01 0.00 0.00
46.23 4.66 0.00 0.41 7.25 0.23 0.77 0.10 0.31
31.80 5.16 0.00 0.41 7.31 0.24 0.76 0.10 0.31
39.10 4.45 0.00 0.00 0.01 0.49 0.51 0.00 0.00
38.05 4.28 0.00 0.00 0.01 0.49 0.51 0.00 0.00
46.29 2.08 0.00 0.41 7.25 0.23 0.77 0.10 0.31
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Table 1 Simulation results by Aspen Plus [16] for one of the operating conditions
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Table 2 Comparison of the simulated and experimental operating-conditions Operating Condition
Turbine inlet (°C)
Absorber exit (°C)
Turbine exit (°C)
Boiler out (°C)
Boiler in (°C)
Separator liquid out (°C)
Separator vapor out (°C)
Experimental Modeling
39.1 39.1
31.74 31.7
21.95 22.2
84.75 84.8
68.67 68.7
83.18 83.5
84.24 83.5
Experimental Modeling
40.1 40.1
31.4 31.4
23.4 24.5
85.3 85.3
68.6 68.6
82.5 83
83.8 83
Experimental Modeling
42.2 42.1
31.76 31.8
24.5 24.1
83.7 83.7
65.0 65.1
81.9 82.5
83.2 82.5
Experimental Modeling
47.9 47.9
32.1 32.1
26.9 27.2
84.31 84.3
67.7 67.7
82.43 83
83.5 83
2.4 Effective COP
Boiling Pressure [MPa]
2.2
0.4 0.64
2 1.8
0.60
1.6 0.55 1.4
0.5 Eff COP optimum, 0.647 [5] 30 η II ref = 30% [5]1st optimum, [5] 50 η II ref = 50% [5]1st optimum,
1.2 0.40 1 0.345
0.395
0.445
0.495
0.545
Basic Solution Concentration [kg/kg]
Fig. 5. Operation map showing optimized region and comparison points (COPopt = 0.647 at a concentration of 0.554, expander’s outlet-temperature of 0.1 °C and pump pressure of 2.3 MPa.).
Basic concentration (ammonia wt%) = 0.554 Expander’s inlet-temperature = 69.0 °C Expander’s exit-temperature = 0.10 °C Boiler’s vapor fraction = 0.325 Ratio of cooling to work = 0.888 These results were obtained by considering a similar configuration (discussed in the modeling section) of the power and cooling cycle, simulating with equivalent heat-source and absorber conditions, but using a combined first-law efficiency definition, Eq. (2), as the objective function [5]. 8. Conclusions A thermodynamic analysis of a combined cooling-and-power cycle was carried out using an EO mode of simulation by ASPEN Plus 12.1 [16]. The results show that the con-
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ditions for the optimum cooling production are at odds with the power production. In this study, the trade-off of work to accommodate cooling is quantified by defining an effective COP. The defined parameter has been used as an objective function to optimize the cooling production. It appears that the effective COP optimum is biased towards higher expander exhaust-temperatures but with increased vapor flow. Some of the simulation runs have been performed before optimization for the verification of the developed model. The agreement between the simulation and experimental results proves the accuracy of the developed model. Finally, this analysis is not sufficient to determine the optimum parameters for the operating conditions. The optimum conditions have to be tested experimentally for the final verification of the results. Furthermore, it is not clear what conditions would favor a particular objective function. Exchanging the work-optimized cycle-efficiency in Eq. (4) with an analytic definition, such as Carnot efficiency, would aid in an analytical comparison of the effective COP with other performance evaluations. References [1] Erickson DC, Anand G, Kyung I. Heat-activated dual-function absorption cycle. ASHRAE Trans 2004;110(1):515–24. [2] Amano Y, Suzuki T, Hashizume T, Akiba M, Tanzawa Y, Usui A. A hybrid power-generation and refrigeration cycle with ammonia–water mixture, IJPGC2000-15058. In: Proceedings of AD 2000 joint power generation conference. ASME; 2000. [3] Goswami DY. Solar thermal power: status of technologies and opportunities for research, heat-and-mass transfer ’95. In: Proceedings of the 2nd ASME-ISHMT heat-and-mass transfer conference. New Delhi (India): Tata-McGraw Hill Publishers; 1995. p. 57–60. [4] Zheng D, Chen B, Qi Y. Thermodynamic analysis of a novel absorption power/cooling combined cycle. In: Proceedings of the 2002 international sorption heat-pump conference. Shanghai, China; 2002. p. 204–9. [5] Vijayaraghavan S, Goswami DY. On evaluating efficiency of a combined power-and-cooling cycle. In: Proceedings of IMECE 2002. New Orleans (LA): American Society of Mechanical Engineers; 2002. [6] Maloney JD, Robertson RC. Thermodynamic study of ammonia–water heat and power cycles. ORNL Report CF-53-8-43, Oak Ridge (TN); 1953. [7] Kalina AI. Combined-cycle system with novel bottoming cycle. ASME J Eng Gas Turb Power 1984;106:737–42. [8] Marston CH. Parametric analysis of the Kalina cycle. J Eng Gas Turb Power 1990;112:107–16. [9] Ibrahim OM, Klein SA. Absorption power cycles. Energy 1996;21(1):21–7. [10] Rogdakis ED, Antonopoulos KA. A high efficiency NH3/H2O absorption power-cycle. Heat Recov Syst CHP 1991;2:263–75. [11] Goswami DY, Xu F. Analysis of a new thermodynamic cycle for combined power-and-cooling using lowand mid-temperature solar collectors. J Solar Energ Eng 1999;121:91–7. [12] Lu S, Goswami DY. Optimization of a novel combined power/refrigeration thermodynamic cycle. J Solar Energ Eng 2003;125:212–7. [13] Lu S, Goswami DY. Theoretical analysis of ammonia-based combined power/refrigeration cycle at low refrigeration-temperatures. In: Proceedings of SOLAR 2002. Reno (NV): American Society of Mechanical Engineers; 2002. [14] Vijayaraghavan S. Thermodynamic studies on alternate binary working-fluid combinations and configurations for a combined power-and-cooling cycle. Ph.D. dissertation, University of Florida, Gainesville (FL); 2003. [15] Tamm G, Goswami DY. Novel combined power-and-cooling thermodynamic cycle for low-temperature heat sources, part II: Experimental investigation. J Solar Energ Eng 2003;125:223–9. [16] Aspen Plus, Version 12.1. Ten Canal Park, Cambridge (MA) 02141: Aspen Technology, Inc; 2004. [17] Macriss R, Eakin B, Ellington R, Huebler J. Physical and thermodynamic properties of ammonia–water mixtures. Chicago (IL): Institute of Gas Technology; 1964. [18] Martin CL. PhD thesis, Mechanical Engineering Department, University of Florida; 2004.