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Effectiveness of cooling production with a combined power and cooling thermodynamic cycle
Applied Thermal Engineering 26 (2006) 576–582 www.elsevier.com/locate/apthermeng
Effectiveness of cooling production with a combined power and cooling thermodynamic cycle C. Martin, D.Y. Goswami
*
Solar Energy and Energy Conversion Laboratory, Department of Mechanical and Aerospace Engineering, University of Florida, P.O. Box 116300, Gainesville, FL 32611-6300, USA Received 23 November 2004; accepted 15 July 2005 Available online 1 September 2005
Abstract The combined production of power and cooling with an ammonia–water based cycle is under investigation. Cooling is produced by expanding an ammonia-rich vapor in an expander to sub-ambient temperatures and it is shown that a compromise exists between cooling and work production. A new parameter, termed the effective COP, is used to relate the gain in cooling to the compromise in work production. When the parameter is used to optimize conditions for the rectifier, the effective COP values are good, having values of up to 5. However, when combined operation is compared to work-optimized results, the maximum effective COP values are near 1.1. This implies that per unit of cooling production, nearly equal amounts of work are compromised for combined operation. 2005 Elsevier Ltd. All rights reserved. Keywords: Absorption power cycle; Cooling; Ammonia–water; Coefficient of performance
1. Introduction A topic of recent interest is the idea of combined power and cooling cycles that use an ammonia–water working fluid. This work has stemmed from commonalities between Kalina-type power cycles and aqua-ammonia absorption cooling, and there is now a small group of proposed configurations [1–4]. The cited advantages of 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]. The particular cycle under investigation in this work was originally proposed by Goswami [3] and is intended primarily for power production while simultaneously producing a cooling output. Prior studies have shown that the *
Corresponding author. Tel.: +1 352 392 0812; fax: +1 352 392 1071. E-mail address: [email protected]fl.edu (D.Y. Goswami). 1359-4311/$ - see front matter 2005 Elsevier Ltd. All rights reserved. doi:10.1016/j.applthermaleng.2005.07.007
concept has the potential to work with low temperature (100–200 C), sensible heat sources, which if provided by non-concentrating solar thermal collectors could reduce the capital cost of solar thermal energy conversion [5]. Fig. 1 presents the configuration 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. This cycle is distinguished from other dual-output concepts by the method it uses to produce cooling and the low heat source temperatures (<200 C) it is intended to interface with. Evaluation of the cycleÕs performance has not been straightforward because of the dual outputs of power and cooling. To account for the quality of the cooling output, it was typically weighted in efficiency definitions [6]. 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 to optimize combined power and cooling production.
C. Martin, D.Y. Goswami / Applied Thermal Engineering 26 (2006) 576–582
577
Nomenclature COP Q Qh
coefficient of performance quantity of heat transfer quantity of heat transfer to cycle from heat source
2. Cycle derivation 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. The first study of an absorption-based power cycle was performed by Maloney and Robertson [7] who concluded no significant advantage to the configuration over steam cycle operation at the conditions considered. Several decades later, Kalina [8], reintroduced the idea of an ammonia– water power cycle as a superior bottoming cycle option over steam Rankine cycles. Independent studies have been performed, for example [9,10], that concede some advantage of 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 pro-
Rc/w W g
ratio of cooling produced to the net amount of work produced quantity of work produced cycle first law efficiency
cess and can limit expansion in the expander, especially at low resource temperatures. Rogdakis and Antonopoulos [11] 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 [11]. This modification requires that only partial vaporization of the working fluid take place in the boiler so that the remaining liquid can be used to absorb vapor in the absorber. 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. This is an obvious departure from pure working fluid Rankine cycle operation, where the limiting expander 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 power and cooling thermodynamic cycle.
Superheater Coolant
Rectifier
Heat Source
3. Cycle operation
Separator
Expander 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, used to model performance.
Fig. 1 is the flow schematic for the combined power and cooling cycle. Referring to the figure, basic-concentration fluid is drawn from the absorber and pumped to high pressure via the solution pump. Before entering the boiler, the basic solution recovers heat from the returning weak-in-ammonia liquid solution in the recovery heat exchanger. In the boiler, 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 vaporÕs ammonia concentration is increased by cooling and condensate separation in the rectifier. Heat can be added in the superheater as the vapor proceeds to the expander. The expander extracts energy from the high-pressure vapor as it is throttled to the system low-pressure. The vapor rejoins the weak liquid in the absorber where, with heat rejection, the basic solution is regenerated.
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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 highlighting the source of cooling temperatures.
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 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 approximates the basic solution exiting the absorber, while the high concentration vapor is typical of the expander exhaust conditions.
4. Previous work Since the original proposal [3] theoretical and experimental investigations have taken place. Initial investigations were performed theoretically and they focused on identifying operating trends [12]. 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 first law, second law, and energy efficiency definitions [13]. Minimum cooling temperatures [14], working fluid combinations, and system configurations [15] were also studied and optimized. Also, an initial experimental study was conducted, which generally verified the expected boiling and absorption processes [16].
used to determine the relative value of each output so that they can be combined into an efficiency definition. Previous work in this area is now discussed along with a description of the new evaluation used for this work. The question of appropriate efficiency expressions for the cycle was examined by Vijayaraghavan and Goswami [6]. It had been noticed that the results obtained from an optimization of the cycle were heavily 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, Eq. (1): 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 performance [6]. The quality of the cooling output can be evaluated by using a COP to find a work equivalent for the cooling produced (Eq. (2)): g1st Law ¼
ðW net þ ðQcool =COPÞÞ Qh
ð2Þ
The Carnot COP, 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 [6]. 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 [6]. In this work, a different approach is used to evaluate the effectiveness of cooling production. Analysis of cooling production with the combined power and cooling cycle (which is discussed in more detail in the next section) reveals that some compromises to work production have to be made in order to obtain the simultaneous cooling output. That is a work-optimized system will not produce the conditions necessary for simultaneous cooling production. A new COP definition 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 and the theoretical amount of work production that was compromised in order to have combined cooling production. The simple formulation is provided as Eq. (3): Cooling Produced Potential Work Lost
5. Performance measures
COPeffective ¼
The fact that this cycle has dual outputs of power and cooling raises some unique questions regarding the evaluation of its performance. For instance, the method
This term has been called the effective COP since cooling and work are only indirectly related, in other words there is no device directly producing cooling with the
ð3Þ
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6. Cooling production To understand the compromise between work and cooling production with this cycle, a detailed look at the mechanisms of cooling production are now discussed. Since the cooling produced with this method is sensible, that is no phase change of the refrigerant is occurring, the exhaust temperature from the expander is key in 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 exhaust temperature were discussed in relation to the binary phase diagram of Fig. 2. A simple reiteration is that the separation between the basic solution concentration and the rectified vapor concentration is critical to achieving a significant, sub-ambient temperature difference. Further insight can be gathered by considering the entropy of the working fluid at the expander exit. Minimization of the exhaust temperature also implies a minimization of the vapor entropy at expander exhaust, assuming constant exit pressure. From this consideration an efficient expander is an obvious feature for low temperatures, 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 inlet conditions, the function of the rectifier is immediately apparent. In the rectifier the vaporÕs 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 mass flow rate. These effects are mostly to the advantage of cooler expander exhaust temperatures. It may be noted that the same rectification process is used in aqua-ammonia absorption systems, but for
1 0.9
Vapor Concentration [kg/kg]
work that is given up. Rather, the effective COP is a measure of the compromise made to accommodate a combined cooling output. Just as with other workdriven COP values higher values of effective COP are desired since they maximize cooling production for the penalty to work production.
579
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 Flow Fraction
Fig. 3. Relationship between vapor concentration and amount of partial vaporization.
slightly different reasons. In an absorption cooling system, since the vapor is condensed to liquid, trace amounts of water do not necessarily prevent evaporator temperatures that are close to the saturation temperature of pure ammonia (at least not in the dramatic fashion that it effects vapor saturation temperatures, refer again to Fig. 2). However, excess water in the evaporator can create an unwanted temperature glide or change system chemistry so that design point operation cannot be maintained [17]. The considerations for exhaust temperatures are now coupled with the mechanisms of vapor production. It was mentioned that 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. Fig. 3 is a plot of vapor concentration as a function of the vapor mass flow fraction, which is the ratio of the vapor mass flow to the basic solution mass flow. 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 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 a distillation process separating ammonia from the working fluid mixture. These are the same general requirements 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.
7. Modeling The modeling is based on thermodynamic simulation of the schematic in Fig. 1. Straightforward conservation
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models were used for each component of the cycle. Noted procedures and constraints were as follows.
5
The correlations used to compute ammonia–water property data are based on a combination of the Gibbs free energy method for mixture properties and empirical equations for bubble and dew point temperatures to determine phase equilibrium. Details of this formulation are covered in Ref. [18].
Eff. COP
4.5
• The boiling conditions were completely specified, that is boiling temperature, pressure, and basic solution concentration were provided as inputs. The vaporization fraction (vapor quality at boiler exit) was determined from saturation conditions. • The basic solution ammonia concentration and absorber exit temperature (held constant at 35 C) were both specified. The system low pressure was determined by assuming saturated conditions at absorber exit. • Vapor rectification was limited by either the specified rectifier exit temperature or an ammonia mass fraction of 0.999, whichever was encountered first. The minimum rectification temperature considered was 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 turbine exhaust temperature to 15 C.
4 90 %
3.5 Expander Efficiency =
60 %
3
2.5 60
65
70
75
80
85
90
95
100
Boiling Temp. [°C]
Fig. 4. Maximum effective COP values where the work component is the amount of work lost due to operation with rectification vs. equivalent conditions with no rectification.
objective function and maximized. The variable parameters were the basic solution concentration, boiler exit pressure, and rectifier exit temperature. The maximum effective COP values are presented in Fig. 4. The results of the maximum effective COP appear quite good. In the best cases, the work that is lost due to rectification penalties is effectively traded for 4–5 times the quantity of cooling. This is comparable performance to other work-driven cooling systems, however, it is not quite the complete picture of energy efficiency since work production can also be optimized. For the case with expander efficiency of 60%, the decline at low temperatures is due to the minimum limit placed on rectifier exit temperature.
8. Optimum rectification From the discussion of cooling production, the expander inlet conditions were identified as key in obtaining cool exhaust temperatures. These conditions are determined by the amount of cooling that takes place in the rectifier. While more cooling is generally beneficial for cool exhaust temperatures, rectification penalizes work production in two ways. First it reduces the available energy of the vapor stream by lowering the vapor temperature. Second, the mass flow rate is reduced due to partial condensation that occurs in the rectifier. This scenario implies a that there could be a gain in cooling but at a cost to work production. To quantify the balance between cooling and work production, the effective COP concept given in Eq. (3) is used to determine the optimum amount of rectification. Adapting the general definition of Eq. (3) to an evaluation of rectifier operation results in Eq. (4): COPeffective ¼
Qcool w=rect ðW no rect W with rect Þ
ð4Þ
To determine the most effective balance between gains and penalties of rectification, Eq. (4) was used as an
9. Work optimization For a given heat source temperature, work optimization of the combined cycle results in the system configuration evolving toward a pure component working fluid Rankine cycle. General optimum characteristics for the power-cooling cycle have been discussed and a sense of these differences can be gained from the typical conditions for a work optimized and power-cooling optimized systems (Table 1). In the cooling-optimized case the basic solution concentration is kept to a moderate level so that absorption pressures are low enough for cooling to take place, and
Table 1 Typical operating characteristics for cooling and work optimized cycles Optimization
Vapor mass flow fraction (%)
Basic solution concentration
Cooling Work
5–10 +90
0.45 0.95
C. Martin, D.Y. Goswami / Applied Thermal Engineering 26 (2006) 576–582 1.2
581
Vapor Mass Flow Fraction = 0 %
1.7
5% 10 %
1.1
Boiling Pressure [MPa]
15 %
1
Eff. COP
90 %
0.9
0.8 60 %
= Expander Efficiency
1.0
1.4 -4.5° -7°
1.1 -10°
0.7
0.8 0°
0.7 0.5 0.4
0.6
Overall Effective COP = 0.9
1.08
0.44
4.5°
0.48
9° = Exhaust Temp. [°C]
0.52
0.56
0.6
Basic Solution Concentration 0.5 60
65
70
75
80
85
90
95
100
Boiling Temp. [°C]
Fig. 5. Maximum overall effective COP values as defined by Eq. (5).
the amount of vaporization is low to maintain a high ammonia content in the vapor. Work-optimization, on the other hand, leads to a nearly pure, ammonia Rankine cycle. While the pressure ratio is sufficient for power production, the high absorption–condensation pressure precludes any cooling temperatures. For a more stringent evaluation, cooling production can be assessed by the amount of lost work that could be obtained from a work-optimized system. When considering this scenario the overall effective COP has the following formulation. COPoverall ¼
Rc=w Qcool ¼ g work opt W work opt W w=cool 1
ð5Þ
gw=cool
where the subscripts work opt and w/cool refer to parameters for the work optimized case and those with cooling production, respectively. The rightmost formulation represents the actual implementation with dimensionless parameters. The term Rc/w is the ratio of cooling to net work production and g is the simple first law efficiency formulation (not accounting for cooling production), which is given as Eq. (6): g1st Law ¼
ðW expander W pump Þ ðQboiler þ Qsuperheat Þ
ð6Þ
Fig. 5 presents the maximum values of the overall, effective COP as a function of boiling temperature. Based solely on energy considerations and the assumptions inherent to Fig. 5, at the best conditions nearly equal amounts of work must be forfeited for the amount of cooling gained.
Fig. 6. Operation map showing the relative sensitivity of overall effective COP to vapor mass flow fraction and exhaust temperature. Sensitivity to mass flow is high, while with a mild penalty to effective COP a wide range of temperatures could be expected.
perature from the turbine is key in determining the amount of cooling that can be achieved. Furthermore, the temperature of the exhaust will dictate the suitable cooling applications. Consideration of the relationship between overall effective COP and other system parameters provides a good indication of achievable exhaust temperatures. Fig. 6 attempts to show this relationship by plotting results of constant temperature operation in a boiling pressure-basic solution concentration plane. Three parameters are plotted in Fig. 6: the overall effective COP value, the expander exhaust temperature, and the vapor mass flow fraction (the percentage of the basic solution mass flow that is vaporized and passes through the expander). The maximum effective COP value is centered within the 1.08 contour line at approximately a pressure of 1.03 MPa and a basic solution concentration of roughly 0.46. The corresponding exhaust temperature and vapor flow fraction is 4.5 C and 7.8%, respectively. Fig. 6 shows that the effective COP is much more sensitive to the vapor flow fraction than the exhaust temperature. Therefore, with a mild penalty to effective COP values, a large range of exhaust temperatures could be accessed. For example, while operating within the 1.08 effective COP contour line, the exhaust temperatures could vary between approximately 7 C and 2.3 C. Furthermore, if the effective COP were diminished by 7.5% (the 1.0 contour line), the range of possible exhaust temperatures would be between roughly 12 C and 5.6 C.
11. Conclusions 10. Exhaust temperatures Since the cooling produced with this method is sensible, i.e. no phase change is occurring, the exhaust tem-
With the present combined power and cooling cycle, the conditions for optimal cooling production are at odds with power production. The need for efficient
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separation of ammonia and water in the generated vapor, and the preference for low expander inlet temperatures detract from the conditions needed for work production alone. In this study, the trade-off of work to accommodate cooling is quantified by defining an effective COP. By using this effective COP as an objective function, the gain in cooling was optimized and evaluated. In the first such exercise, the value of the rectifier was made apparent by producing effective COP values of up to 5. However, when combined operation is compared to a work-optimized system, the maximum overall effective COP is approximately 1.1. Compared to other work-driven cooling cycles, a COP of 1.1 is somewhat unimpressive. This can be viewed as the cost of combined power and cooling production with this particular cycle. The previously cited advantages of combined power and cooling production would have to outweigh this seeming deficit in efficiency. References [1] D.C. Erickson, G. Anand, I. Kyung, Heat-activated dual-function absorption cycle, ASHRAE Transactions 110 (1) (2004) 515–524. [2] Y. Amano, T. Suzuki, T. Hashizume, M. Akiba, Y. Tanzawa, A. Usui, A hybrid power generation and refrigeration cycle with ammonia–water mixture, IJPGC2000-15058, in: Proceedings of 2000 Joint Power Generation Conference, ASME, 2000. [3] D.Y. Goswami, Solar thermal power: status of technologies and opportunities for research, heat and mass transfer Õ95, in: Proceedings of the 2nd ASME-ISHMT Heat and Mass Transfer Conference, Tata-McGraw Hill Publishers, New Delhi, India, 1995, pp. 57–60. [4] D. Zheng, B. Chen, Y. Qi, Thermodynamic analysis of a novel absorption power/cooling combined cycle, in: Proceedings of the 2002 International Sorption Heat Pump Conference, Shanghai, China, 2002, pp. 204–209. [5] D.Y. Goswami, F. Xu, Analysis of a new thermodynamic cycle for combined power and cooling using low and mid temperature solar
[6]
[7]
[8]
[9] [10] [11]
[12]
[13]
[14]
[15]
[16]
[17] [18]
collectors, ASME Journal of Solar Energy Engineering 121 (1999) 91–97. S. Vijayaraghavan, D.Y. Goswami, On evaluating efficiency of a combined power and cooling cycle, in: Proceedings of IMECE 2002, American Society of Mechanical Engineers, New Orleans, LA, 2002. J.D. Maloney, R.C. Robertson, Thermodynamic study of ammonia–water heat power cycles, ORNL Report CF-53-8-43, Oak Ridge, TN, 1953. A.I. Kalina, Combined cycle system with novel bottoming cycle, ASME Journal of Engineering for Gas Turbines and Power 106 (1984) 737–742. C.H. Marston, Parametric analysis of the Kalina cycle, Journal of Engineering for Gas Turbines and Power 112 (1990) 107–116. O.M. Ibrahim, S.A. Klein, Absorption power cycles, Energy 21 (1) (1996) 21–27. E.D. Rogdakis, K.A. Antonopoulos, A high efficiency NH3/H2O absorption power cycle, Heat Recovery Systems and CHP 2 (1991) 263–275. D.Y. Goswami, F. Xu, Analysis of a new thermodynamic cycle for combined power and cooling using low and mid temperature solar collectors, Journal of Solar Energy Engineering 121 (1999) 91– 97. S. Lu, D.Y. Goswami, Optimization of a novel combined power/ refrigeration thermodynamic cycle, Journal of Solar Energy Engineering 125 (2003) 212–217. S. Lu, D.Y. Goswami, Theoretical analysis of ammonia-based combined power/refrigeration cycle at low refrigeration temperatures, in: Proceedings of SOLAR 2002, American Society of Mechanical Engineers, Reno, NV, 2002. S. Vijayaraghavan, 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. G. Tamm, D.Y. Goswami, Novel combined power and cooling thermodynamic cycle for low temperature heat sources. Part II: experimental investigation, Journal of Solar Energy Engineering 125 (2003) 223–229. K.E. Herold, R. Radermacher, S.A. Klein, Absorption Chillers and Heat Pumps, CRC Press, Boca Raton, FL, 1996. F. Xu, D.Y. Goswami, Thermodynamic properties of ammonia– water mixtures for power-cycle applications, Energy 24 (1999) 525–536.
Effectiveness of cooling production with a combined power and cooling thermodynamic cycle C. Martin, D.Y. Goswami
*
Solar Energy and Energy Conversion Laboratory, Department of Mechanical and Aerospace Engineering, University of Florida, P.O. Box 116300, Gainesville, FL 32611-6300, USA Received 23 November 2004; accepted 15 July 2005 Available online 1 September 2005
Abstract The combined production of power and cooling with an ammonia–water based cycle is under investigation. Cooling is produced by expanding an ammonia-rich vapor in an expander to sub-ambient temperatures and it is shown that a compromise exists between cooling and work production. A new parameter, termed the effective COP, is used to relate the gain in cooling to the compromise in work production. When the parameter is used to optimize conditions for the rectifier, the effective COP values are good, having values of up to 5. However, when combined operation is compared to work-optimized results, the maximum effective COP values are near 1.1. This implies that per unit of cooling production, nearly equal amounts of work are compromised for combined operation. 2005 Elsevier Ltd. All rights reserved. Keywords: Absorption power cycle; Cooling; Ammonia–water; Coefficient of performance
1. Introduction A topic of recent interest is the idea of combined power and cooling cycles that use an ammonia–water working fluid. This work has stemmed from commonalities between Kalina-type power cycles and aqua-ammonia absorption cooling, and there is now a small group of proposed configurations [1–4]. The cited advantages of 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]. The particular cycle under investigation in this work was originally proposed by Goswami [3] and is intended primarily for power production while simultaneously producing a cooling output. Prior studies have shown that the *
Corresponding author. Tel.: +1 352 392 0812; fax: +1 352 392 1071. E-mail address: [email protected]fl.edu (D.Y. Goswami). 1359-4311/$ - see front matter 2005 Elsevier Ltd. All rights reserved. doi:10.1016/j.applthermaleng.2005.07.007
concept has the potential to work with low temperature (100–200 C), sensible heat sources, which if provided by non-concentrating solar thermal collectors could reduce the capital cost of solar thermal energy conversion [5]. Fig. 1 presents the configuration 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. This cycle is distinguished from other dual-output concepts by the method it uses to produce cooling and the low heat source temperatures (<200 C) it is intended to interface with. Evaluation of the cycleÕs performance has not been straightforward because of the dual outputs of power and cooling. To account for the quality of the cooling output, it was typically weighted in efficiency definitions [6]. 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 to optimize combined power and cooling production.
C. Martin, D.Y. Goswami / Applied Thermal Engineering 26 (2006) 576–582
577
Nomenclature COP Q Qh
coefficient of performance quantity of heat transfer quantity of heat transfer to cycle from heat source
2. Cycle derivation 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. The first study of an absorption-based power cycle was performed by Maloney and Robertson [7] who concluded no significant advantage to the configuration over steam cycle operation at the conditions considered. Several decades later, Kalina [8], reintroduced the idea of an ammonia– water power cycle as a superior bottoming cycle option over steam Rankine cycles. Independent studies have been performed, for example [9,10], that concede some advantage of 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 pro-
Rc/w W g
ratio of cooling produced to the net amount of work produced quantity of work produced cycle first law efficiency
cess and can limit expansion in the expander, especially at low resource temperatures. Rogdakis and Antonopoulos [11] 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 [11]. This modification requires that only partial vaporization of the working fluid take place in the boiler so that the remaining liquid can be used to absorb vapor in the absorber. 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. This is an obvious departure from pure working fluid Rankine cycle operation, where the limiting expander 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 power and cooling thermodynamic cycle.
Superheater Coolant
Rectifier
Heat Source
3. Cycle operation
Separator
Expander 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, used to model performance.
Fig. 1 is the flow schematic for the combined power and cooling cycle. Referring to the figure, basic-concentration fluid is drawn from the absorber and pumped to high pressure via the solution pump. Before entering the boiler, the basic solution recovers heat from the returning weak-in-ammonia liquid solution in the recovery heat exchanger. In the boiler, 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 vaporÕs ammonia concentration is increased by cooling and condensate separation in the rectifier. Heat can be added in the superheater as the vapor proceeds to the expander. The expander extracts energy from the high-pressure vapor as it is throttled to the system low-pressure. The vapor rejoins the weak liquid in the absorber where, with heat rejection, the basic solution is regenerated.
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C. Martin, D.Y. Goswami / Applied Thermal Engineering 26 (2006) 576–582
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 highlighting the source of cooling temperatures.
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 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 approximates the basic solution exiting the absorber, while the high concentration vapor is typical of the expander exhaust conditions.
4. Previous work Since the original proposal [3] theoretical and experimental investigations have taken place. Initial investigations were performed theoretically and they focused on identifying operating trends [12]. 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 first law, second law, and energy efficiency definitions [13]. Minimum cooling temperatures [14], working fluid combinations, and system configurations [15] were also studied and optimized. Also, an initial experimental study was conducted, which generally verified the expected boiling and absorption processes [16].
used to determine the relative value of each output so that they can be combined into an efficiency definition. Previous work in this area is now discussed along with a description of the new evaluation used for this work. The question of appropriate efficiency expressions for the cycle was examined by Vijayaraghavan and Goswami [6]. It had been noticed that the results obtained from an optimization of the cycle were heavily 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, Eq. (1): 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 performance [6]. The quality of the cooling output can be evaluated by using a COP to find a work equivalent for the cooling produced (Eq. (2)): g1st Law ¼
ðW net þ ðQcool =COPÞÞ Qh
ð2Þ
The Carnot COP, 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 [6]. 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 [6]. In this work, a different approach is used to evaluate the effectiveness of cooling production. Analysis of cooling production with the combined power and cooling cycle (which is discussed in more detail in the next section) reveals that some compromises to work production have to be made in order to obtain the simultaneous cooling output. That is a work-optimized system will not produce the conditions necessary for simultaneous cooling production. A new COP definition 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 and the theoretical amount of work production that was compromised in order to have combined cooling production. The simple formulation is provided as Eq. (3): Cooling Produced Potential Work Lost
5. Performance measures
COPeffective ¼
The fact that this cycle has dual outputs of power and cooling raises some unique questions regarding the evaluation of its performance. For instance, the method
This term has been called the effective COP since cooling and work are only indirectly related, in other words there is no device directly producing cooling with the
ð3Þ
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6. Cooling production To understand the compromise between work and cooling production with this cycle, a detailed look at the mechanisms of cooling production are now discussed. Since the cooling produced with this method is sensible, that is no phase change of the refrigerant is occurring, the exhaust temperature from the expander is key in 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 exhaust temperature were discussed in relation to the binary phase diagram of Fig. 2. A simple reiteration is that the separation between the basic solution concentration and the rectified vapor concentration is critical to achieving a significant, sub-ambient temperature difference. Further insight can be gathered by considering the entropy of the working fluid at the expander exit. Minimization of the exhaust temperature also implies a minimization of the vapor entropy at expander exhaust, assuming constant exit pressure. From this consideration an efficient expander is an obvious feature for low temperatures, 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 inlet conditions, the function of the rectifier is immediately apparent. In the rectifier the vaporÕs 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 mass flow rate. These effects are mostly to the advantage of cooler expander exhaust temperatures. It may be noted that the same rectification process is used in aqua-ammonia absorption systems, but for
1 0.9
Vapor Concentration [kg/kg]
work that is given up. Rather, the effective COP is a measure of the compromise made to accommodate a combined cooling output. Just as with other workdriven COP values higher values of effective COP are desired since they maximize cooling production for the penalty to work production.
579
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 Flow Fraction
Fig. 3. Relationship between vapor concentration and amount of partial vaporization.
slightly different reasons. In an absorption cooling system, since the vapor is condensed to liquid, trace amounts of water do not necessarily prevent evaporator temperatures that are close to the saturation temperature of pure ammonia (at least not in the dramatic fashion that it effects vapor saturation temperatures, refer again to Fig. 2). However, excess water in the evaporator can create an unwanted temperature glide or change system chemistry so that design point operation cannot be maintained [17]. The considerations for exhaust temperatures are now coupled with the mechanisms of vapor production. It was mentioned that 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. Fig. 3 is a plot of vapor concentration as a function of the vapor mass flow fraction, which is the ratio of the vapor mass flow to the basic solution mass flow. 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 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 a distillation process separating ammonia from the working fluid mixture. These are the same general requirements 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.
7. Modeling The modeling is based on thermodynamic simulation of the schematic in Fig. 1. Straightforward conservation
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models were used for each component of the cycle. Noted procedures and constraints were as follows.
5
The correlations used to compute ammonia–water property data are based on a combination of the Gibbs free energy method for mixture properties and empirical equations for bubble and dew point temperatures to determine phase equilibrium. Details of this formulation are covered in Ref. [18].
Eff. COP
4.5
• The boiling conditions were completely specified, that is boiling temperature, pressure, and basic solution concentration were provided as inputs. The vaporization fraction (vapor quality at boiler exit) was determined from saturation conditions. • The basic solution ammonia concentration and absorber exit temperature (held constant at 35 C) were both specified. The system low pressure was determined by assuming saturated conditions at absorber exit. • Vapor rectification was limited by either the specified rectifier exit temperature or an ammonia mass fraction of 0.999, whichever was encountered first. The minimum rectification temperature considered was 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 turbine exhaust temperature to 15 C.
4 90 %
3.5 Expander Efficiency =
60 %
3
2.5 60
65
70
75
80
85
90
95
100
Boiling Temp. [°C]
Fig. 4. Maximum effective COP values where the work component is the amount of work lost due to operation with rectification vs. equivalent conditions with no rectification.
objective function and maximized. The variable parameters were the basic solution concentration, boiler exit pressure, and rectifier exit temperature. The maximum effective COP values are presented in Fig. 4. The results of the maximum effective COP appear quite good. In the best cases, the work that is lost due to rectification penalties is effectively traded for 4–5 times the quantity of cooling. This is comparable performance to other work-driven cooling systems, however, it is not quite the complete picture of energy efficiency since work production can also be optimized. For the case with expander efficiency of 60%, the decline at low temperatures is due to the minimum limit placed on rectifier exit temperature.
8. Optimum rectification From the discussion of cooling production, the expander inlet conditions were identified as key in obtaining cool exhaust temperatures. These conditions are determined by the amount of cooling that takes place in the rectifier. While more cooling is generally beneficial for cool exhaust temperatures, rectification penalizes work production in two ways. First it reduces the available energy of the vapor stream by lowering the vapor temperature. Second, the mass flow rate is reduced due to partial condensation that occurs in the rectifier. This scenario implies a that there could be a gain in cooling but at a cost to work production. To quantify the balance between cooling and work production, the effective COP concept given in Eq. (3) is used to determine the optimum amount of rectification. Adapting the general definition of Eq. (3) to an evaluation of rectifier operation results in Eq. (4): COPeffective ¼
Qcool w=rect ðW no rect W with rect Þ
ð4Þ
To determine the most effective balance between gains and penalties of rectification, Eq. (4) was used as an
9. Work optimization For a given heat source temperature, work optimization of the combined cycle results in the system configuration evolving toward a pure component working fluid Rankine cycle. General optimum characteristics for the power-cooling cycle have been discussed and a sense of these differences can be gained from the typical conditions for a work optimized and power-cooling optimized systems (Table 1). In the cooling-optimized case the basic solution concentration is kept to a moderate level so that absorption pressures are low enough for cooling to take place, and
Table 1 Typical operating characteristics for cooling and work optimized cycles Optimization
Vapor mass flow fraction (%)
Basic solution concentration
Cooling Work
5–10 +90
0.45 0.95
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581
Vapor Mass Flow Fraction = 0 %
1.7
5% 10 %
1.1
Boiling Pressure [MPa]
15 %
1
Eff. COP
90 %
0.9
0.8 60 %
= Expander Efficiency
1.0
1.4 -4.5° -7°
1.1 -10°
0.7
0.8 0°
0.7 0.5 0.4
0.6
Overall Effective COP = 0.9
1.08
0.44
4.5°
0.48
9° = Exhaust Temp. [°C]
0.52
0.56
0.6
Basic Solution Concentration 0.5 60
65
70
75
80
85
90
95
100
Boiling Temp. [°C]
Fig. 5. Maximum overall effective COP values as defined by Eq. (5).
the amount of vaporization is low to maintain a high ammonia content in the vapor. Work-optimization, on the other hand, leads to a nearly pure, ammonia Rankine cycle. While the pressure ratio is sufficient for power production, the high absorption–condensation pressure precludes any cooling temperatures. For a more stringent evaluation, cooling production can be assessed by the amount of lost work that could be obtained from a work-optimized system. When considering this scenario the overall effective COP has the following formulation. COPoverall ¼
Rc=w Qcool ¼ g work opt W work opt W w=cool 1
ð5Þ
gw=cool
where the subscripts work opt and w/cool refer to parameters for the work optimized case and those with cooling production, respectively. The rightmost formulation represents the actual implementation with dimensionless parameters. The term Rc/w is the ratio of cooling to net work production and g is the simple first law efficiency formulation (not accounting for cooling production), which is given as Eq. (6): g1st Law ¼
ðW expander W pump Þ ðQboiler þ Qsuperheat Þ
ð6Þ
Fig. 5 presents the maximum values of the overall, effective COP as a function of boiling temperature. Based solely on energy considerations and the assumptions inherent to Fig. 5, at the best conditions nearly equal amounts of work must be forfeited for the amount of cooling gained.
Fig. 6. Operation map showing the relative sensitivity of overall effective COP to vapor mass flow fraction and exhaust temperature. Sensitivity to mass flow is high, while with a mild penalty to effective COP a wide range of temperatures could be expected.
perature from the turbine is key in determining the amount of cooling that can be achieved. Furthermore, the temperature of the exhaust will dictate the suitable cooling applications. Consideration of the relationship between overall effective COP and other system parameters provides a good indication of achievable exhaust temperatures. Fig. 6 attempts to show this relationship by plotting results of constant temperature operation in a boiling pressure-basic solution concentration plane. Three parameters are plotted in Fig. 6: the overall effective COP value, the expander exhaust temperature, and the vapor mass flow fraction (the percentage of the basic solution mass flow that is vaporized and passes through the expander). The maximum effective COP value is centered within the 1.08 contour line at approximately a pressure of 1.03 MPa and a basic solution concentration of roughly 0.46. The corresponding exhaust temperature and vapor flow fraction is 4.5 C and 7.8%, respectively. Fig. 6 shows that the effective COP is much more sensitive to the vapor flow fraction than the exhaust temperature. Therefore, with a mild penalty to effective COP values, a large range of exhaust temperatures could be accessed. For example, while operating within the 1.08 effective COP contour line, the exhaust temperatures could vary between approximately 7 C and 2.3 C. Furthermore, if the effective COP were diminished by 7.5% (the 1.0 contour line), the range of possible exhaust temperatures would be between roughly 12 C and 5.6 C.
11. Conclusions 10. Exhaust temperatures Since the cooling produced with this method is sensible, i.e. no phase change is occurring, the exhaust tem-
With the present combined power and cooling cycle, the conditions for optimal cooling production are at odds with power production. The need for efficient
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separation of ammonia and water in the generated vapor, and the preference for low expander inlet temperatures detract from the conditions needed for work production alone. In this study, the trade-off of work to accommodate cooling is quantified by defining an effective COP. By using this effective COP as an objective function, the gain in cooling was optimized and evaluated. In the first such exercise, the value of the rectifier was made apparent by producing effective COP values of up to 5. However, when combined operation is compared to a work-optimized system, the maximum overall effective COP is approximately 1.1. Compared to other work-driven cooling cycles, a COP of 1.1 is somewhat unimpressive. This can be viewed as the cost of combined power and cooling production with this particular cycle. The previously cited advantages of combined power and cooling production would have to outweigh this seeming deficit in efficiency. References [1] D.C. Erickson, G. Anand, I. Kyung, Heat-activated dual-function absorption cycle, ASHRAE Transactions 110 (1) (2004) 515–524. [2] Y. Amano, T. Suzuki, T. Hashizume, M. Akiba, Y. Tanzawa, A. Usui, A hybrid power generation and refrigeration cycle with ammonia–water mixture, IJPGC2000-15058, in: Proceedings of 2000 Joint Power Generation Conference, ASME, 2000. [3] D.Y. Goswami, Solar thermal power: status of technologies and opportunities for research, heat and mass transfer Õ95, in: Proceedings of the 2nd ASME-ISHMT Heat and Mass Transfer Conference, Tata-McGraw Hill Publishers, New Delhi, India, 1995, pp. 57–60. [4] D. Zheng, B. Chen, Y. Qi, Thermodynamic analysis of a novel absorption power/cooling combined cycle, in: Proceedings of the 2002 International Sorption Heat Pump Conference, Shanghai, China, 2002, pp. 204–209. [5] D.Y. Goswami, F. Xu, Analysis of a new thermodynamic cycle for combined power and cooling using low and mid temperature solar
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