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propane mixture
International Journal of Refrigeration 22 (1999) 313–318
Thermostatic valve control using a non-azeotropic refrigerant, isobutane/propane mixture T.P. Castle a, R.N. Richardson a,*, T.J. Ritter b a
Institute of Cryogenics, University of Southampton, Southampton SO17 1BJ, UK b Calor Gas Refrigeration, Leamington Spa CV43 6RL, UK
Received 22 April 1998; received in revised form 5 October 1998; accepted 8 October 1998
Abstract This article describes the evaluation and comparison of a conventional R12 cross-charged thermostatic valve and an electronic expansion valve using a non-azeotropic refrigerant mixture (NARM); isobutane/propane mixture (CARE30). The superheat temperature setting on an expansion valve needs to compensate for the temperature glide associated with a nonazeotropic refrigerant as these can be of similar magnitude. It is also advisable to increase the superheat setting to make allowance for change in refrigerant composition as a result of preferential refrigerant/oil solubility. The majority of refrigeration systems operate at fixed evaporating temperatures, hence, once superheat setting is trimmed during commissioning, then there should be no further problems associated with evaporation of a non-azeotropic refrigerant provided the system is leaktight. An R12 expansion valve with a factory superheat setting of 5⬚C tested over a wide range of evaporating temperatures proved satisfactory in operation with CARE30 after increasing the superheat temperature screw setting equivalent to 5⬚C. 䉷 1999 Elsevier Science Ltd and IIR. All rights reserved. Keywords: Thermostatic valve; Refrigerant; Isobutane; Propane
Re´gulation d’un de´tendeur thermostatique utilisant un me´lange d’isobutane et de propane comme frigorige`ne non aze´otropique Resume´ Dans cet article, on e´value et on compare un de´tendeur thermostatique classique a` chargement en R12 et un de´tendeur e´lectronique qui utilise un me´lange non aze´otropique constitue´ d’isobutane et de propane (CARE30). La valeur de la surchauffe de tempe´rature fixe´e sur le de´tendeur doit compenser le glissement de tempe´rature lie´ a` un frigorige`ne non aze´otropique, puisqu’ils peuvent eˆtre du meˆme ordre. Il est e´galement recommande´ d’augmenter la valeur de re´glage de la surchauffe afin de permettre des changements dans la composition du frigorige`ne, compte tenu de la solubilite´ de l’huile par rapport au frigorige`ne pre´fe´rentiel. La majorite´ des syste`mes frigorifiques fonctionnent a` des tempe´ratures d’e´vaporation fixes; aussi, une fois que la valeur de re´glage de la surchauffe aura e´te´ e´tablie au cours de la mise en e´tat de fonctionnement, n’y aura-t-il plus aucun proble`me lie´ a` l’e´vaporation d’un me´lange non aze´otropique puisque le syste`me est e´tanche. Un de´tendeur au R12 avec une valeur de re´glage de la surchauffe de 5⬚C, teste´ pour une large plage de tempe´ratures d’e´vaporation, a fonctionne´ de fac¸on satisfaisante avec du CARE30, apre`s que l’on a eu augmente´ la surchauffe de 5⬚C. 䉷 1999 Elsevier Science Ltd and IIR. All rights reserved. Mots cle´s: De´tendeur thermostatique; Frigorige`ne; Isobutane; Propane
0140-7007/99/$20.00 䉷 1999 Elsevier Science Ltd and IIR. All rights reserved. PII: S0140-700 7(98)00059-0
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1. Introduction The role of chlorofluorocarbons in the process of ozone depletion and their impact on global warming is now widely accepted [1,2]. Both hydrofluorocarbons, in particular R134a, and hydrocarbon refrigerants are being used as substitutes for R12. In an optimized system the efficiencies of R134a are comparable to those of R12, in spite of this it is still not clear whether R134a will become the industry standard replacement for R12. Previous studies at low evaporating temperatures have shown inferior performance and there remain doubts over its environmental acceptability [3–6]. Hydrocarbons, however, meet all the requirements of an environmentally safe refrigerant and also offer the potential for high efficiency [7–10]. CARE30 [11,12] is a nonazeotropic refrigerant mixture (NARM) of isobutane and propane designed as an R12 ‘drop-in’ replacement refrigerant (Fig. 1). However, unlike R12, CARE30 exhibits a temperature glide of approximately 8⬚C during evaporation. This affects the expansion valve refrigerant control which relies on the ‘measured’ temperature rise across the evaporator. An experimental system was designed (Fig. 2) using standard R12 components to evaluate the performance of two types of R12 expansion valve using CARE30. A calorimeter evaporator arrangement comprising of a coil immersed in methanol contained in a Dewar flask is used and the methanol is electrically heated and stirred. The thermostatic expansion valve (TEV) is of the R12 cross-charged type which is the industry norm for an R12 system with an external pressure equalization port [13]. It is charged with methyl chloride (R40), a certain percentage of an inert gas and a small amount of air. The design gives a virtually constant superheat value over the evaporating
range. The test system also includes an electronic expansion valve (EEV). For reasons of experimental convenience this is manually modulated; the valve position being controlled to maintain a constant temperature difference between the evaporator outlet and inlet during the calorimeter tests. No oil separator was used at the compressor discharge, hence a certain amount of oil is entrained in the refrigerant vapour. The behaviour of the circuit is, therefore, representative of a refrigeration system in which the expansion valve must control a refrigerant/oil mixture [14].
2. Testing of TEV and EEV operation with CARE30 Table 1 shows the temperature of the refrigerant entering (TIN) and leaving (TOUT) the evaporator in relation to the CARE30 bubble (TB) and dew point (TD) temperatures during calorimeter tests for the TEV and EEV valves. The accuracy of the temperature and power measurements are ^0.1⬚C and ^1 W respectively. The EEV was controlled manually, such that the refrigerant temperature difference between the outlet and inlet of the evaporator, (TOUT ⫺ TIN), remained constant at 12⬚C by adjusting the position of the valve needle. The refrigerant superheat, (TOUT ⫺ TD), which the TEV controls was maintained at 8⬚C, by adjusting the TEV superheat screw setting during calorimeter tests. The condensing temperature was determined by the evaporator load and ambient temperature (air cooled condenser). All the experiments were conducted at an ambient temperature of 21⬚C ^ 1⬚C which resulted in condensing temperatures between 26⬚C and 35⬚C depending on the evaporator load. This is, of course, representative of the behaviour of a practical refrigeration system which employs
Fig. 1. Comparison of saturation curves for R12 [15] and CARE30. Fig. 1. Comparaison des courbes de saturation pour le R12 et le CARE30.
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315
Fig. 2. Schematic diagram of the experimental system. Fig. 2. Sche´ma du syste`me expe´rimental.
an air cooled condenser and must operate at variable load. The coefficient of performance, compressor power and evaporator duty are plotted against CARE30 bubble point temperature (Figs. 3–5). For both the valves, the inlet vapour fraction reduces as the evaporator saturation temperature increases. The reduction in vapour fraction reflects the smaller temperature differential between the vaporizing temperature and the temperature of the liquid approaching the refrigerant control
device. Consequently, at higher vaporizing temperatures a smaller fraction of the refrigerant vaporizes in the control device. These changes to vapour fraction with evaporation temperature are not specific to NARMs. The refrigerant vapour fraction at the evaporator inlet for the TEV was lower than that for the EEV. As (TOUT ⫺ TIN) is constant for the EEV, and the vapour fraction reduces as the saturation temperature increases, the refrigerant superheat, (TOUT ⫺ TD) reduces slightly. The
Table 1 Comparison of the inlet and outlet temperatures of the evaporator for the EEV and TEV control device Tableau 1 Comparaison des tempe´ratures d’aspiration et de refoulement de l’e´vaporateur, pour le re´glage du de´tendeur thermostatique et du de´tendeur e´lectronique Duty W
100 150 200 250 300 350 400 450 a
EEV TB a
TD b
TIN c
TOUT d
Inlet e %
Superheat
⫺ 24.9 ⫺ 20.4 ⫺ 14.7 ⫺ 9.8 ⫺ 6.3 ⫺ 1.2
⫺ 16.2 ⫺ 11.8 ⫺ 6.2 ⫺ 1.4 2.1 7.1
⫺ 22.0 ⫺ 17.8 ⫺ 12.1 ⫺ 7.3 ⫺ 3.9 1.2
⫺ 9.9 ⫺ 5.7 ⫺ 0.1 4.7 8.1 12.9
33 30 31 25 29 29
6.3 6.1 6.1 6.1 6.0 5.8
4.2
12.3
5.9
17.9
21
5.6
TEV TB a ⫺ 20.7 ⫺ 15.1 ⫺ 10.8 ⫺ 7.2 ⫺ 3.5 ⫺ 0.6 2.6
TD b
TIN c
TOUT d
Inlet e %
Superheat
⫺ 12.1 ⫺ 6.6 ⫺ 2.4 1.2 4.8 7.5 10.8
⫺ 18.8 ⫺ 13.4 ⫺ 9.1 ⫺ 5.4 ⫺ 2.2 0.7 3.8
⫺ 4.1 1.4 5.6 9.2 12.8 15.5 18.8
22 20 20 21 16 16 15
8 8 8 8 8 8 8
Bubble point temperature (⬚C) based on measured evaporating pressure. Dew point temperature (⬚C) based on measured evaporating pressure. c Evaporator inlet temperature (⬚C). d Evaporator outlet temperature (⬚C). e Evaporator inlet molar vapour fraction (%). The refrigerant vapour fraction was determined from measured evaporator pressure and evaporator inlet temperature. b
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Fig. 3. Cooling coefficient of performance for different expansion devices as a function of evaporation temperature. Fig. 3. Coefficient de performance du refroidissement pour plusieurs de´tendeurs, en fonction de la tempe´rature d’e´vaporation.
reduction in superheat would be greater but for the fact that the temperature glide also reduces as the saturation temperature of CARE30 increases. For instance, at TB equal to ⫺20⬚C and 0⬚C, the corresponding glide temperatures for CARE30 are 8.6⬚C and 8.2⬚C respectively.
Experiments were conducted using the TEV to investigate how the measured evaporator superheat varied with evaporator temperature for a particular setting of the static superheat at the valve. The TEV superheat characteristic over the vaporizing temperature range depends on the
Fig. 4. Compressor power for different expansion devices as a function of evaporation temperature. Fig. 4. Puissance du compresseur pour plusieurs de´tendeurs en fonction de la tempe´rature d’e´vaporation.
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317
Fig. 5. Evaporator duty for different expansion devices as a function of evaporation temperature. Fig. 5. Travail de l’e´vaporateur pour plusieurs de´tendeurs en fonction de la tempe´rature d’e´vaporation.
slope of the pressure–temperature relationship of the thermostatic liquid cross-charge element and the refrigerant used. The TEV used in the experiment was designed for nearly constant superheat over the evaporator temperature range when used with R12. As a result of the similarity in their saturation characteristics, no significant deviation in superheat would therefore be expected when using CARE30. Only the reduction in temperature glide at higher evaporating temperatures for CARE30 may have an effect on the superheat value for a constant superheat screw setting. Table 2 shows the results of this test. The ‘static’ superheat setting at the valve is 8⬚C at an initial evaporating temperature of ⫺ 20⬚C. As the evaporating temperature rises, the measured superheat reduces in a non-linear manner. For instance, the superheat reduced by 1.5⬚C between the evaporating temperatures ⫺ 20⬚C and ⫺ 10⬚C, whilst between ⫺ 10⬚C and 0⬚C, the superheat reduced by 1⬚C. As a result of this reduction in superheat at higher evaporation temperatures the superheat setting must be carefully considered so as to ensure that no
Table 2 Change in TEV evaporator superheat for a fixed superheat setting of 8⬚C set at an evaporator saturation temperature of ⫺ 20⬚C Tableau 2 Changement de la surchauffe de l’e´vaporateur avec le de´tendeur thermostatique pour une valeur de re´glage de la surchauffe fixe´e a` 8⬚C a` une tempe´rature de saturation de l’e´vaporateur de ⫺ 20⬚C Evaporation temperature TB (⬚C)
Superheat (TOUT ⫺ TB) (⬚C)
⫺ 20 ⫺ 10 0
8 6.5 5.5
liquid carry over occurs during system pulldown before reaching the steady state. The drop in evaporator superheat with increase in vaporizing temperature does not alter significantly the refrigerating effect per unit mass of refrigerant circulated or the valve capacity. For example, the refrigerant enthalpy change in the evaporator for a 35⬚C condensing temperature, 5⬚C subcooling, evaporation at 0⬚C and superheats of 8⬚C and 5.5⬚C are 328.8 kJ/kg and 331 kJ/kg respectively.
3. Adjustment to expansion valves The number of degrees of superheat in the propane/isobutane mixture leaving the evaporator can only be approximated because the temperature of the two-phase refrigerant entering the evaporator is somewhere between the bubble and dew point temperatures; the precise composition depending on the quantity of the flash gas. To avoid evaporator liquid through-flow, the set temperature difference of the EEV must compensate for the temperature glide of the two-phase flow. This also applies to the superheat screw setting for the TEV. A typical factory superheat setting of 5⬚C will be of approximately the same magnitude as the CARE30 temperature glide between the evaporator inlet temperature and the dew point temperature (Table 1). Therefore, the temperature difference which the EEV controls will have to be at least twice that to achieve 5⬚C superheat. As for the TEV, the superheat screw setting will have to be trimmed to achieve 5⬚C with the plant in operation. The EEV superheat value over the evaporating range could be kept a constant if a pressure transducer was used to measure the evaporator pressure, converting pressure to the corresponding dew point temperature according to a
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table or formula. As conversion of pressure to temperature is based on the saturation curve, there is no glide problem as this can be incorporated in the formula. 4. Potential control problems associated with preferential leakage With any refrigerant mixture exhibiting temperature glide, that is a mixture that can be fully or partly separated into individual components, there is a possibility of a change in composition compared to the original charge in the event of a leak. Practical experience with CARE30 indicates that the loss of a small fraction of charge has a minimal effect on the mixture composition and consequently the effect on system performance is negligible. Obviously, any loss of refrigerant is undesirable and good practice should ensure leak-tight systems. Nevertheless, measures can be taken to minimise the problems of system stability in the event of loss of refrigerant. Operation of the TEV will be compromised if the refrigerant composition, and hence the saturation characteristics, change significantly as these will no longer correspond to the pressure/temperature characteristic of the valve. For example, following a loss of refrigerant, a higher evaporation temperature for a given pressure reduces the valve’s static superheat (point at which the valve needle moves away from the seat – typically 4⬚C–5⬚C). If it is required to allow this eventuality, the TEV can be adjusted for higher superheat to compensate for any possible change in composition without compromising on the performance of the system. Our experiments indicate that an adequate margin of safety to ensure no liquid carry-over is achieved with an increase in the static superheat setting of approximately 2⬚C–3⬚C. Such measures are applicable to other refrigerant mixtures with more than 5⬚C temperature glide. In practice there have been no reports of refrigerant leakage which have caused control problems when using refrigerant blends. It is likely that any loss of refrigerant serious enough to result in a significant change in the mixture composition will cause the system to fail, or at least under perform, as a result of lack of charge and it can be investigated accordingly. 5. Conclusions As the temperature glide of a non-azeotropic refrigerant such as CARE30 (temperature glide approximately 8⬚C) is of similar magnitude to that of the static superheat setting of a TEV, it must be taken into account when initially adjusting the valve and balancing the system. To calculate the exact dew point temeperature and therefore the correct superheat, the evaporator pressure must be accurately measured so as to determine the saturation temperature.
It was demonstrated that a standard R12 thermostatic expansion and electronic expansion valve can be used with CARE30 after the superheat setting is trimmed during commisioning because the saturation characterisitcs and volumetric refrigerating effect of R12 and CARE30 are very similar. The superheat reduces as the evaporation temperature increases if no adjustment is made to the static superheat setting, especially for the TEV. However, this does not have a noticeable effect on the utilisation of the evaporator or system performance. It is appreciated that whilst there may be a change in the condensing temperarture, majority of the users operate the plant at fixed evaporating temperatures. Therefore, once the correct superheat setting is established, no further problems should be encountered in the normal operation when using a NARM.
References [1] Wuebbles DJ. The role of refrigerants in climate change. Int J Refrig 1994;17:1. [2] World Meteorlogical Organisation. WMO Rep.20, vol. 1 and 2, 1990. [3] Petterson B, Thorsell H. Comparison of the refrigeerants HFC 134a and CFC12. Int J Refrig 1990;13(5):176–180. [4] Tickell O. Up in the air. New Scientist, 128 (No.1739; Oct 20), 1990:41–43. [5] Wallington TJ, Hurley MD, Ball JC, Kalser EW. Atmospheric Chemistry of Hydrofluorocarbon 134a Fate of the Alkoxy radical CF3CFHO. Environ Sci Technol 1992;26(7):1318– 1324. [6] Banks RE. Sceptism about R134a justified — Prof, Refrigeration and Air Conditioning, 1993:16. [7] Richardson RN, Butterworth JS. The performance of propane/ isobutane mixtures in a vapour-compression refrigeration system. Int J Refrig 1995;18(1):58–62. [8] Richardson RN, Ritter TJ. The performance of hydrocarbon alternatives to R12 and R22. AIRAH Journal, 1996;21–28. [9] Stene J. Country reports — International status, workshop proceedings, compression systems with natural working fluids, Trondheim, 1995. [10] Lystad T. Testing of a heat pump with propane as working fluid, workshop proceeding, compression systems with natural working fluids, Trondheim, 1995. [11] Hydrocarbons take off. Refrigeration and air conditioning, October 1995. [12] Calor Gas Limited. CARE refrigerant property data and KMKREIS computer simulation program. [13] Refrigerant control devices, in: ASHRAE Fundamentals Handbook, ch. 19, 1989. [14] Gosney WB. Principles of refrigeration. Cambridge: Cambridge University Press, 1982. [15] ASHRAE thermodynamic properties of refrigerants, 1986.
Thermostatic valve control using a non-azeotropic refrigerant, isobutane/propane mixture T.P. Castle a, R.N. Richardson a,*, T.J. Ritter b a
Institute of Cryogenics, University of Southampton, Southampton SO17 1BJ, UK b Calor Gas Refrigeration, Leamington Spa CV43 6RL, UK
Received 22 April 1998; received in revised form 5 October 1998; accepted 8 October 1998
Abstract This article describes the evaluation and comparison of a conventional R12 cross-charged thermostatic valve and an electronic expansion valve using a non-azeotropic refrigerant mixture (NARM); isobutane/propane mixture (CARE30). The superheat temperature setting on an expansion valve needs to compensate for the temperature glide associated with a nonazeotropic refrigerant as these can be of similar magnitude. It is also advisable to increase the superheat setting to make allowance for change in refrigerant composition as a result of preferential refrigerant/oil solubility. The majority of refrigeration systems operate at fixed evaporating temperatures, hence, once superheat setting is trimmed during commissioning, then there should be no further problems associated with evaporation of a non-azeotropic refrigerant provided the system is leaktight. An R12 expansion valve with a factory superheat setting of 5⬚C tested over a wide range of evaporating temperatures proved satisfactory in operation with CARE30 after increasing the superheat temperature screw setting equivalent to 5⬚C. 䉷 1999 Elsevier Science Ltd and IIR. All rights reserved. Keywords: Thermostatic valve; Refrigerant; Isobutane; Propane
Re´gulation d’un de´tendeur thermostatique utilisant un me´lange d’isobutane et de propane comme frigorige`ne non aze´otropique Resume´ Dans cet article, on e´value et on compare un de´tendeur thermostatique classique a` chargement en R12 et un de´tendeur e´lectronique qui utilise un me´lange non aze´otropique constitue´ d’isobutane et de propane (CARE30). La valeur de la surchauffe de tempe´rature fixe´e sur le de´tendeur doit compenser le glissement de tempe´rature lie´ a` un frigorige`ne non aze´otropique, puisqu’ils peuvent eˆtre du meˆme ordre. Il est e´galement recommande´ d’augmenter la valeur de re´glage de la surchauffe afin de permettre des changements dans la composition du frigorige`ne, compte tenu de la solubilite´ de l’huile par rapport au frigorige`ne pre´fe´rentiel. La majorite´ des syste`mes frigorifiques fonctionnent a` des tempe´ratures d’e´vaporation fixes; aussi, une fois que la valeur de re´glage de la surchauffe aura e´te´ e´tablie au cours de la mise en e´tat de fonctionnement, n’y aura-t-il plus aucun proble`me lie´ a` l’e´vaporation d’un me´lange non aze´otropique puisque le syste`me est e´tanche. Un de´tendeur au R12 avec une valeur de re´glage de la surchauffe de 5⬚C, teste´ pour une large plage de tempe´ratures d’e´vaporation, a fonctionne´ de fac¸on satisfaisante avec du CARE30, apre`s que l’on a eu augmente´ la surchauffe de 5⬚C. 䉷 1999 Elsevier Science Ltd and IIR. All rights reserved. Mots cle´s: De´tendeur thermostatique; Frigorige`ne; Isobutane; Propane
0140-7007/99/$20.00 䉷 1999 Elsevier Science Ltd and IIR. All rights reserved. PII: S0140-700 7(98)00059-0
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1. Introduction The role of chlorofluorocarbons in the process of ozone depletion and their impact on global warming is now widely accepted [1,2]. Both hydrofluorocarbons, in particular R134a, and hydrocarbon refrigerants are being used as substitutes for R12. In an optimized system the efficiencies of R134a are comparable to those of R12, in spite of this it is still not clear whether R134a will become the industry standard replacement for R12. Previous studies at low evaporating temperatures have shown inferior performance and there remain doubts over its environmental acceptability [3–6]. Hydrocarbons, however, meet all the requirements of an environmentally safe refrigerant and also offer the potential for high efficiency [7–10]. CARE30 [11,12] is a nonazeotropic refrigerant mixture (NARM) of isobutane and propane designed as an R12 ‘drop-in’ replacement refrigerant (Fig. 1). However, unlike R12, CARE30 exhibits a temperature glide of approximately 8⬚C during evaporation. This affects the expansion valve refrigerant control which relies on the ‘measured’ temperature rise across the evaporator. An experimental system was designed (Fig. 2) using standard R12 components to evaluate the performance of two types of R12 expansion valve using CARE30. A calorimeter evaporator arrangement comprising of a coil immersed in methanol contained in a Dewar flask is used and the methanol is electrically heated and stirred. The thermostatic expansion valve (TEV) is of the R12 cross-charged type which is the industry norm for an R12 system with an external pressure equalization port [13]. It is charged with methyl chloride (R40), a certain percentage of an inert gas and a small amount of air. The design gives a virtually constant superheat value over the evaporating
range. The test system also includes an electronic expansion valve (EEV). For reasons of experimental convenience this is manually modulated; the valve position being controlled to maintain a constant temperature difference between the evaporator outlet and inlet during the calorimeter tests. No oil separator was used at the compressor discharge, hence a certain amount of oil is entrained in the refrigerant vapour. The behaviour of the circuit is, therefore, representative of a refrigeration system in which the expansion valve must control a refrigerant/oil mixture [14].
2. Testing of TEV and EEV operation with CARE30 Table 1 shows the temperature of the refrigerant entering (TIN) and leaving (TOUT) the evaporator in relation to the CARE30 bubble (TB) and dew point (TD) temperatures during calorimeter tests for the TEV and EEV valves. The accuracy of the temperature and power measurements are ^0.1⬚C and ^1 W respectively. The EEV was controlled manually, such that the refrigerant temperature difference between the outlet and inlet of the evaporator, (TOUT ⫺ TIN), remained constant at 12⬚C by adjusting the position of the valve needle. The refrigerant superheat, (TOUT ⫺ TD), which the TEV controls was maintained at 8⬚C, by adjusting the TEV superheat screw setting during calorimeter tests. The condensing temperature was determined by the evaporator load and ambient temperature (air cooled condenser). All the experiments were conducted at an ambient temperature of 21⬚C ^ 1⬚C which resulted in condensing temperatures between 26⬚C and 35⬚C depending on the evaporator load. This is, of course, representative of the behaviour of a practical refrigeration system which employs
Fig. 1. Comparison of saturation curves for R12 [15] and CARE30. Fig. 1. Comparaison des courbes de saturation pour le R12 et le CARE30.
T.P. Castle et al. / International Journal of Refrigeration 22 (1999) 313–318
315
Fig. 2. Schematic diagram of the experimental system. Fig. 2. Sche´ma du syste`me expe´rimental.
an air cooled condenser and must operate at variable load. The coefficient of performance, compressor power and evaporator duty are plotted against CARE30 bubble point temperature (Figs. 3–5). For both the valves, the inlet vapour fraction reduces as the evaporator saturation temperature increases. The reduction in vapour fraction reflects the smaller temperature differential between the vaporizing temperature and the temperature of the liquid approaching the refrigerant control
device. Consequently, at higher vaporizing temperatures a smaller fraction of the refrigerant vaporizes in the control device. These changes to vapour fraction with evaporation temperature are not specific to NARMs. The refrigerant vapour fraction at the evaporator inlet for the TEV was lower than that for the EEV. As (TOUT ⫺ TIN) is constant for the EEV, and the vapour fraction reduces as the saturation temperature increases, the refrigerant superheat, (TOUT ⫺ TD) reduces slightly. The
Table 1 Comparison of the inlet and outlet temperatures of the evaporator for the EEV and TEV control device Tableau 1 Comparaison des tempe´ratures d’aspiration et de refoulement de l’e´vaporateur, pour le re´glage du de´tendeur thermostatique et du de´tendeur e´lectronique Duty W
100 150 200 250 300 350 400 450 a
EEV TB a
TD b
TIN c
TOUT d
Inlet e %
Superheat
⫺ 24.9 ⫺ 20.4 ⫺ 14.7 ⫺ 9.8 ⫺ 6.3 ⫺ 1.2
⫺ 16.2 ⫺ 11.8 ⫺ 6.2 ⫺ 1.4 2.1 7.1
⫺ 22.0 ⫺ 17.8 ⫺ 12.1 ⫺ 7.3 ⫺ 3.9 1.2
⫺ 9.9 ⫺ 5.7 ⫺ 0.1 4.7 8.1 12.9
33 30 31 25 29 29
6.3 6.1 6.1 6.1 6.0 5.8
4.2
12.3
5.9
17.9
21
5.6
TEV TB a ⫺ 20.7 ⫺ 15.1 ⫺ 10.8 ⫺ 7.2 ⫺ 3.5 ⫺ 0.6 2.6
TD b
TIN c
TOUT d
Inlet e %
Superheat
⫺ 12.1 ⫺ 6.6 ⫺ 2.4 1.2 4.8 7.5 10.8
⫺ 18.8 ⫺ 13.4 ⫺ 9.1 ⫺ 5.4 ⫺ 2.2 0.7 3.8
⫺ 4.1 1.4 5.6 9.2 12.8 15.5 18.8
22 20 20 21 16 16 15
8 8 8 8 8 8 8
Bubble point temperature (⬚C) based on measured evaporating pressure. Dew point temperature (⬚C) based on measured evaporating pressure. c Evaporator inlet temperature (⬚C). d Evaporator outlet temperature (⬚C). e Evaporator inlet molar vapour fraction (%). The refrigerant vapour fraction was determined from measured evaporator pressure and evaporator inlet temperature. b
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Fig. 3. Cooling coefficient of performance for different expansion devices as a function of evaporation temperature. Fig. 3. Coefficient de performance du refroidissement pour plusieurs de´tendeurs, en fonction de la tempe´rature d’e´vaporation.
reduction in superheat would be greater but for the fact that the temperature glide also reduces as the saturation temperature of CARE30 increases. For instance, at TB equal to ⫺20⬚C and 0⬚C, the corresponding glide temperatures for CARE30 are 8.6⬚C and 8.2⬚C respectively.
Experiments were conducted using the TEV to investigate how the measured evaporator superheat varied with evaporator temperature for a particular setting of the static superheat at the valve. The TEV superheat characteristic over the vaporizing temperature range depends on the
Fig. 4. Compressor power for different expansion devices as a function of evaporation temperature. Fig. 4. Puissance du compresseur pour plusieurs de´tendeurs en fonction de la tempe´rature d’e´vaporation.
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Fig. 5. Evaporator duty for different expansion devices as a function of evaporation temperature. Fig. 5. Travail de l’e´vaporateur pour plusieurs de´tendeurs en fonction de la tempe´rature d’e´vaporation.
slope of the pressure–temperature relationship of the thermostatic liquid cross-charge element and the refrigerant used. The TEV used in the experiment was designed for nearly constant superheat over the evaporator temperature range when used with R12. As a result of the similarity in their saturation characteristics, no significant deviation in superheat would therefore be expected when using CARE30. Only the reduction in temperature glide at higher evaporating temperatures for CARE30 may have an effect on the superheat value for a constant superheat screw setting. Table 2 shows the results of this test. The ‘static’ superheat setting at the valve is 8⬚C at an initial evaporating temperature of ⫺ 20⬚C. As the evaporating temperature rises, the measured superheat reduces in a non-linear manner. For instance, the superheat reduced by 1.5⬚C between the evaporating temperatures ⫺ 20⬚C and ⫺ 10⬚C, whilst between ⫺ 10⬚C and 0⬚C, the superheat reduced by 1⬚C. As a result of this reduction in superheat at higher evaporation temperatures the superheat setting must be carefully considered so as to ensure that no
Table 2 Change in TEV evaporator superheat for a fixed superheat setting of 8⬚C set at an evaporator saturation temperature of ⫺ 20⬚C Tableau 2 Changement de la surchauffe de l’e´vaporateur avec le de´tendeur thermostatique pour une valeur de re´glage de la surchauffe fixe´e a` 8⬚C a` une tempe´rature de saturation de l’e´vaporateur de ⫺ 20⬚C Evaporation temperature TB (⬚C)
Superheat (TOUT ⫺ TB) (⬚C)
⫺ 20 ⫺ 10 0
8 6.5 5.5
liquid carry over occurs during system pulldown before reaching the steady state. The drop in evaporator superheat with increase in vaporizing temperature does not alter significantly the refrigerating effect per unit mass of refrigerant circulated or the valve capacity. For example, the refrigerant enthalpy change in the evaporator for a 35⬚C condensing temperature, 5⬚C subcooling, evaporation at 0⬚C and superheats of 8⬚C and 5.5⬚C are 328.8 kJ/kg and 331 kJ/kg respectively.
3. Adjustment to expansion valves The number of degrees of superheat in the propane/isobutane mixture leaving the evaporator can only be approximated because the temperature of the two-phase refrigerant entering the evaporator is somewhere between the bubble and dew point temperatures; the precise composition depending on the quantity of the flash gas. To avoid evaporator liquid through-flow, the set temperature difference of the EEV must compensate for the temperature glide of the two-phase flow. This also applies to the superheat screw setting for the TEV. A typical factory superheat setting of 5⬚C will be of approximately the same magnitude as the CARE30 temperature glide between the evaporator inlet temperature and the dew point temperature (Table 1). Therefore, the temperature difference which the EEV controls will have to be at least twice that to achieve 5⬚C superheat. As for the TEV, the superheat screw setting will have to be trimmed to achieve 5⬚C with the plant in operation. The EEV superheat value over the evaporating range could be kept a constant if a pressure transducer was used to measure the evaporator pressure, converting pressure to the corresponding dew point temperature according to a
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table or formula. As conversion of pressure to temperature is based on the saturation curve, there is no glide problem as this can be incorporated in the formula. 4. Potential control problems associated with preferential leakage With any refrigerant mixture exhibiting temperature glide, that is a mixture that can be fully or partly separated into individual components, there is a possibility of a change in composition compared to the original charge in the event of a leak. Practical experience with CARE30 indicates that the loss of a small fraction of charge has a minimal effect on the mixture composition and consequently the effect on system performance is negligible. Obviously, any loss of refrigerant is undesirable and good practice should ensure leak-tight systems. Nevertheless, measures can be taken to minimise the problems of system stability in the event of loss of refrigerant. Operation of the TEV will be compromised if the refrigerant composition, and hence the saturation characteristics, change significantly as these will no longer correspond to the pressure/temperature characteristic of the valve. For example, following a loss of refrigerant, a higher evaporation temperature for a given pressure reduces the valve’s static superheat (point at which the valve needle moves away from the seat – typically 4⬚C–5⬚C). If it is required to allow this eventuality, the TEV can be adjusted for higher superheat to compensate for any possible change in composition without compromising on the performance of the system. Our experiments indicate that an adequate margin of safety to ensure no liquid carry-over is achieved with an increase in the static superheat setting of approximately 2⬚C–3⬚C. Such measures are applicable to other refrigerant mixtures with more than 5⬚C temperature glide. In practice there have been no reports of refrigerant leakage which have caused control problems when using refrigerant blends. It is likely that any loss of refrigerant serious enough to result in a significant change in the mixture composition will cause the system to fail, or at least under perform, as a result of lack of charge and it can be investigated accordingly. 5. Conclusions As the temperature glide of a non-azeotropic refrigerant such as CARE30 (temperature glide approximately 8⬚C) is of similar magnitude to that of the static superheat setting of a TEV, it must be taken into account when initially adjusting the valve and balancing the system. To calculate the exact dew point temeperature and therefore the correct superheat, the evaporator pressure must be accurately measured so as to determine the saturation temperature.
It was demonstrated that a standard R12 thermostatic expansion and electronic expansion valve can be used with CARE30 after the superheat setting is trimmed during commisioning because the saturation characterisitcs and volumetric refrigerating effect of R12 and CARE30 are very similar. The superheat reduces as the evaporation temperature increases if no adjustment is made to the static superheat setting, especially for the TEV. However, this does not have a noticeable effect on the utilisation of the evaporator or system performance. It is appreciated that whilst there may be a change in the condensing temperarture, majority of the users operate the plant at fixed evaporating temperatures. Therefore, once the correct superheat setting is established, no further problems should be encountered in the normal operation when using a NARM.
References [1] Wuebbles DJ. The role of refrigerants in climate change. Int J Refrig 1994;17:1. [2] World Meteorlogical Organisation. WMO Rep.20, vol. 1 and 2, 1990. [3] Petterson B, Thorsell H. Comparison of the refrigeerants HFC 134a and CFC12. Int J Refrig 1990;13(5):176–180. [4] Tickell O. Up in the air. New Scientist, 128 (No.1739; Oct 20), 1990:41–43. [5] Wallington TJ, Hurley MD, Ball JC, Kalser EW. Atmospheric Chemistry of Hydrofluorocarbon 134a Fate of the Alkoxy radical CF3CFHO. Environ Sci Technol 1992;26(7):1318– 1324. [6] Banks RE. Sceptism about R134a justified — Prof, Refrigeration and Air Conditioning, 1993:16. [7] Richardson RN, Butterworth JS. The performance of propane/ isobutane mixtures in a vapour-compression refrigeration system. Int J Refrig 1995;18(1):58–62. [8] Richardson RN, Ritter TJ. The performance of hydrocarbon alternatives to R12 and R22. AIRAH Journal, 1996;21–28. [9] Stene J. Country reports — International status, workshop proceedings, compression systems with natural working fluids, Trondheim, 1995. [10] Lystad T. Testing of a heat pump with propane as working fluid, workshop proceeding, compression systems with natural working fluids, Trondheim, 1995. [11] Hydrocarbons take off. Refrigeration and air conditioning, October 1995. [12] Calor Gas Limited. CARE refrigerant property data and KMKREIS computer simulation program. [13] Refrigerant control devices, in: ASHRAE Fundamentals Handbook, ch. 19, 1989. [14] Gosney WB. Principles of refrigeration. Cambridge: Cambridge University Press, 1982. [15] ASHRAE thermodynamic properties of refrigerants, 1986.