A 1-D analysis of ejector performance. Analyse unidimensionnelle de la performance d un éjecteur

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1 International Journal of Refrigeration (1999) A 1-D analysis of ejector performance B.J. Huang*, J.M. Cang, C.P. Wang, V.A. Petrenko Department of Mecanical Engineering, National Taiwan University, Taipei 106, Taiwan Received 4 May 1998; received in revised form 5 November 1998; accepted 3 December 1998 Abstract A 1-D analysis for te prediction of ejector performance at critical-mode operation is carried out in te present study. Constant-pressure mixing is assumed to occur inside te constant-area section of te ejector and te entrained flow at coking condition is analyzed. We also carried out an experiment using 11 ejectors and R141b as te working fluid to verify te analytical results. Te test results are used to determine te coefficients, p, s, f p and f m defined in te 1-D model by matcing te test data wit te analytical results. It is sown tat te1-d analysis using te empirical coefficients can accurately predict te performance of te ejectors Elsevier Science Ltd and IIR. All rigts reserved. Keywords: Refrigerating system; Ejector system; Refrigerant; R141b; Ejector; Performance Analyse unidimensionnelle de la performance d un éjecteur Résumé Dans cette étude, on a effectué une analyse unidimensionnelle pour prédire la performance d un éjecteur fonctionnant en mode critique. Les auteurs sont partis du principe que le mélange s effectue à pression constante dans la partie de l éjecteur dont la section est constante et ont analysé le flux entraîné au niveau de l onde de coc. Onze éjecteurs utilisant le R141b comme fluide actif ont été utilisés par les auteurs afin de vérifier les résultats analytiques. Les résultats expérimentaux sont utilisés pour déterminer les coefficients p, s, f p et f m définis dans le modèle unidimensionnel. L étude a montré que l analyse unidimensionnelle utilisant les coefficients empiriques peut prédire la performance des éjecteurs de façon précise Elsevier Science Ltd and IIR. All rigts reserved. Mots clés: Système frigorifique; Système à éjecteur; Frigorigène; R141b; Éjecteur; Performance Nomenclature A area, m a sonic velocity, m/s C p specific eat of gas at constant pressure, kj kg 1 K 1 C v specific eat of gas at constant volume, kj kg 1 K 1 * Corresponding autor. Tel.: ; fax: address: bjuang@tpts6.seed.net.tw (B.J. Huang) d _m M * P c P e P g R T T c * /99/$ Elsevier Science Ltd and IIR. All rigts reserved. PII: S (99) diameter, m entalpy, kj kg 1 mass flowrate, kg s 1 Mac number critical back pressure of te ejector, MPa vapor pressure at te suction port of te ejector, MPa vapor pressure at te nozzle inlet of te ejector, MPa gas constant, kj kg 1 K 1 temperature, K saturated-vapor temperature corresponding to te critical back pressure P c *, K

2 B.J. Huang et al. / International Journal of Refrigeration (1999) T e vapor temperature at te suction port of te ejector, K T g vapor temperature at te nozzle inlet of te ejector, K T gs saturated-vapor temperature corresponding to P g, K V gas velocity, m s 1 x nozzle position, m y position of te ypotetical troat g ˆ C p /C V Superscripts * critical mode operation of ejector Subscripts c exit of ejector; condenser co limiting condition of ejector operational mode e inlet port of te entrained flow; ypotetical troat g nozzle inlet m mixed flow p primary flow p1 nozzle exit py primary flow at te location of coking for te entrained flow s suction or entrained flow sy entrained flow at te location of coking for te entrained flow t nozzle troat y location of coking for te entrained flow 1 nozzle exit entrance of te constant-area section 3 exit of te constant-area section 1. Introduction Ejector air-conditioning or refrigeration system powered by low-grade energy as been studied since te mid-1950s. For utilizing solar or waste eat energy as te eat source, many researcers used refrigerant suc as R11, R1, R13, R, R113, R114, R14, or R14b as te working fluid [1,]. Te performance of refrigerant ejector cooling system is owever relatively low as compared to te conventional system. Te key problem is in te ejector design. Te ejector design can be classified into two categories according to te position of te nozzle []. For te nozzle wit its exit located witin te constant-area section of an ejector, te mixing of te primary and te entrained flows occurs inside te constant-area section and te ejector is known as constant-area mixing ejector. For te nozzle wit its exit located witin te suction camber wic is in front of te constant-area section, te ejector is referred as constant-pressure mixing ejector. For tis kind of ejector, it was assumed tat te mixing of te primary and te entrained streams occurs in te suction camber wit a uniform or constant pressure. It is known tat te constant-pressure ejector as a better performance tan te constant-area ejector and is tus widely used. Terefore, we focus on te design of a constant-pressure ejector in te present study, but wit a new concept tat te mixing occurs witin te constant-area section. Te constant-pressure mixing teory of ejector developed by Keenan et al. [3] was frequently used in te analysis of constant-pressure ejector [1,4,5]. Keenan et al. [3] assumed tat te pressures of te primary and te entrained flows at te exit of te nozzle ave an identical pressure. Mixing of te two streams begins tere wit a uniform pressure, i.e. constant pressure, until te inlet of te constant-area section. In practice, two coking penomena exist in te ejector performance [6,7]: one in te primary flow troug te nozzle and te oter in te entrained flow. In addition to te coking in te nozzle, te second coking of an ejector results from te acceleration of te entrained flow from a stagnant state at te suction port to a supersonic flow in te constant-area section. Fig. 1 sows te variation of entrainment ratio v wit te discarge or back pressure P c at fixed suction pressure P e and fixed primary flow pressure P g. Te ejector performance can ten be divided into tree operational modes, according to te back pressure P c : 1. double-coking or critical mode as P c P c *, wile te primary and te entrained flows are bot coking and te entrainment ratio is constant, i.e. v ˆ constant;. single-coking or subcritical mode as P c * P c P co, wile only te primary flow is coked and v canges wit te back pressure P c ; and 3. back-flow or malfunction mode as P c P co, wile bot te primary and te secondary flow are not coked and te entrained flow is reversed (malfunction), i.e.v 0. Te ejector ad better perform at critical mode in order to obtain a better efficiency. Te 1-D constant-pressure mixing teory of Keenan et al. [3] is owever unable to analyze te coking of te entrained flow at critical operation mode. In te present study, we developed an 1-D model for te analysis of ejector performance at te critical-mode operation. Te constant- Fig. 1. Operational modes of ejector. Fig. 1. Modes de fonctionnement de l éjecteur.

3 356 B.J. Huang et al. / International Journal of Refrigeration (1999) Fig.. Scematic diagram of ejector performance. Fig.. Scéma du fonctionnement de l éjecteur. pressure mixing is assumed to occur inside te constant-area section and te coking of te entrained flow is predicted. We ten carried out an experiment to compare te analytical and test results using various ejectors and R141b as te working fluid.. Ejector performance analysis Keenan et al. [3] assumed tat mixing of te two streams takes place inside te suction camber wit a constant or uniform pressure from te exit of te nozzle to te inlet of te constant-area section. Munday and Bagster [6] postulated tat after exausting from te nozzle, te primary flow fans out witout mixing wit te entrained flow and induces a converging duct for te entrained flow. Tis duct acts as a converging nozzle suc tat te entrained flow is accelerated to a sonic velocity at some place, i.e. ypotetical troat. After tat, mixing of te two streams starts wit a uniform pressure. A ypotetical troat area, or effective area A e [6,7], was defined for te entrained flow at critical operation mode. Huang et al. [7] furter determined experimentally te ypotetical troat area A e for R113 ejector. In te present study, we assume tat te ypotetical troat occurs inside te constant-area section of te ejector. Tus, te mixing of two streams occurs inside te constantarea section wit a uniform pressure. Fig. is a scematic diagram sowing te mixing process of te two streams in te ejector. Te following assumptions are made for te analysis: 1. Te working fluid is an ideal gas wit constant properties C p and g.. Te flow inside te ejector is steady and one dimension. 3. Te kinetic energy at te inlets of primary and suction ports and te exit of diffuser are negligible. 4. For simplicity in deriving te 1-D model, te isentropic relations are used as an approximation. But to account for non-ideal process, te effects of frictional and mixing losses are taken into account by using some coefficients introduced in te isentropic relations. Tese coefficients are related to te isentropic efficiency and needs to be determined experimentally. 5. After exausting from te nozzle, te primary flow fans out witout mixing wit te entrained flow until at some cross section y y (ypotetical troat) wic is inside te constant-area section. 6. Te two streams starts to mix at te cross section y y (ypotetical troat) wit an uniform pressure, i.e. P py ˆ P sy, before te sock wic is at te cross section s s. 7. Te entrained flow is coked at te cross section y y (ypotetical troat). 8. Te inner wall of te ejector is adiabatic..1. Governing equations.1.1. Primary flow troug nozzle For a given inlet stagnant pressure P g and temperature T g, te mass flow troug te nozzle at coking condition follows te gas dynamic equation: s _m p ˆ PgA t g g 1 = g 1 p p T g R g 1 p; 1 were p is a coefficient relating to te isentropic efficiency of te compressible flow in te nozzle. Te gas dynamic relations between te Mac number at te exit of nozzle M p1 and te exit cross section area A p1 and pressure P p1 are, using isentropic relations as an

4 B.J. Huang et al. / International Journal of Refrigeration (1999) approximation, A p1 1 g 1 1 A t g 1 M p1 M p1 g 1 = g 1 ; P g g 1 g= g 1 1 Mp1 : 3 P p1.1.. Primary-flow core (from section 1 1 to section y y) Te Mac number M py of te primary flow at te y y section follows te isentropic relations as an approximation: g= g 1 P py 1 g 1 = Mp1 P g= g 1 : 4 p1 1 g 1 = Mpy For te calculation of te area of te primary flow core at te y y section, we use te following isentropic relation, but an arbitrary coefficientf p is included to account for te loss of te primary flow from section 1 1 to y y: A py A p1 ˆ f p=m py = g 1 1 g 1 = Mpy i g 1 = g 1 : 1=M p1 = g 1 1 g 1 = Mp1 5 i g 1 = g 1 Te loss may result from te slipping or viscous effect of te primary and te entrained flows at te boundary. Te loss actually reflects in te reduction of troat area A py at y y section troug te introduction of te coefficient f p in Eq. (5) Entrained flow from inlet to section y y From assumption (6), te entrained flow reaces coking condition at te y y section, i.e. M sy ˆ 1. For a given inlet stagnant pressure P e, we ave P e 1 g 1 g= g 1 Msy : 6 P sy Te entrained flow rate at coking condition follows s _m s ˆ Pe A sy g g 1 = g 1 p p T e R g 1 s; 7 were s is te coefficient related to te isentropic efficiency of te entrained flow Cross-sectional area at section y y Te geometrical cross-sectional area at section y y is A 3 tat is te sum of te areas for te primary flow A py and for te entrained flow A sy. Tat is, A py A sy ˆ A 3 : Temperature and Mac number at section y y Te temperature and te Mac number of te two stream at section y y follows T g T py ˆ 1 g 1 T e ˆ 1 g 1 T sy M py; M sy: Mixed flow at section m m before te sock Two streams starts to mix from section y y. A sock ten takes place wit a sarp pressure rise at section s s. A momentum balance relation tus can be derived as i f m _m p V py _m s V sy ˆ _m p _m s V m ; 11 were V m is te velocity of te mixed flow and f m is te coefficient accounting for te frictional loss [8]. Similarly, an energy balance relation can be derived as!! _m p C p T py V py _m s C p T sy V sy! ˆ _m p _m s C p T m V m ; 1 were V py and V sy are te gas velocities of te primary and entrained flow at te section y y: q V py ˆ M py a py ; a py ˆ grt py ; 13 V sy ˆ M sy a sy ; q a sy ˆ grt sy : 14 Te Mac number of te mixed flow can be evaluated using te following relation: M m ˆ Vm p ; a a m ˆ grt m : 15 m.1.7. Mixed flow across te sock from section m m to section 3 3 A supersonic sock will take place at section s s wit a sarp pressure rise. Assuming tat te mixed flow after te sock undergoing an isentropic process, te mixed flow between section m m and section 3 3 inside te constantarea section as a uniform pressure P 3. Terefore, te following gas dynamic relations exist: P 3 ˆ 1 g P m g 1 M m 1 ; 16 M 3 ˆ 1 g 1 = M m gm m g 1 = : Mixed flow troug diffuser Te pressure at te exit of te diffuser follows te

5 358 B.J. Huang et al. / International Journal of Refrigeration (1999) section A 3. For a given nozzle troat area A t (or diameter d t ) and nozzle exit area A p1 (or diameter d p1 ), te performance of an ejector is caracterized by te stagnant temperature and pressure at te nozzle inlet T g ; P g and at te suction inlet port T e ; P e, te critical back pressure P c *. Tat is, tere are 5 independent variables T g ; P g ; T e ; P e ; P c * in te ejector performance analysis. Te analysis procedure follows te flowcart sown in Fig. 3. Te output of te analysis includes te primary flow _m p, te entrained flow _m s, te entrainment ratio v, te cross sectional area of te constant-area section A 3 and te area ratio A 3 /A t. 3. Experimental verification 3.1. Experimental setup Fig. 4 sows te scematic diagram of te test facility. R141b is selected as te working fluid in te present study as R141b as a positive-slope saturated-vapor line in te termodynamic T s diagram [9]. Tis will not produce a condensation for te vapor expansion in te ejector and tus reduces losses. Te design and operation of te test facility is te same as tat described in te previous article [10]. It takes about 1 to warm up te test facility and about 30 min for eac steady-state test run. 3.. Ejector specifications Fig. 3. Simulation flowcart in te ejector performance analysis. Fig. 3. Diagramme d analyse pour la simulation de la performance de l éjecteur. relation, assuming isentropic process P c ˆ 1 g 1 g= g 1 M3 : 18 P 3.. Ejector performance analysis procedure Using te above 1-D model of ejector, we can carry out te performance analysis to determine te entrainment ratio v and te required cross sectional area of te constant-area To verify te teoretical analysis using te present model, we ave tested 11 different ejectors. Te ejector is designed in tree major parts: nozzle, suction camber body, and constant-area section (including diffuser). Te standard connections between te different parts are used so tat all te parts are intercangeable. Two nozzles (A, E) are designed and fabricated in te experiment. Te specifications of te nozzles are listed in Table 1. We designed 8 different constant-area sections (including diffuser) as listed in Table. Eleven ejectors are used in te present experiment. Te area ratio A 3 /A t of te tested ejectors ranges from 6.44 to Comparison of analysis wit test results Te 11 ejectors were tested under various operating conditions. For easy understanding, we also present te saturated vapor temperatures in te parentesis for te pressures Table 1 Nozzle design Tableau 1 Caractéristiques des tuyères Nozzle Troat diameter, d t (mm) Exit diameter, d pl (mm) A pl /A t A E

6 B.J. Huang et al. / International Journal of Refrigeration (1999) Fig. 4. Scematic diagram of te ejector test facility. Fig. 4. Scéma du dispositif utilisé pour tester les éjecteurs. sown in te figures. Tat is, T c * represents te saturated vapor temperature of te critical back pressure P c *. T gs represents te saturated vapor temperature of te primary flow pressure at te inlet of te nozzle P g. For abbreviation, ejector AB represents te ejector assembly of nozzle A and constant-area section B, for example. Sown in Tables 3 and 4 are te teoretical calculations of te required ejector area ratio A 3 =A t and te critical-mode entrainment ratio v at various operating conditions. For te ejectors aving a A 3 =A t consistent wit te calculated (or required) value to witin ^ 10% deviation (as sown in Figs. 5 and 6), te measured v coincide fairly well wit te calculated results using te present 1-D analysis, mostly witin ^ 15% error, as can be seen from Figs. 7 and 8. As all te ejectors were performed at critical mode, only te critical-point performance is sown and discussed in te present study. In te 1-D analysis, te coefficients accounting for te loss in te primary flow in te nozzle and in te suction flow before mixing are taken as p ˆ 0.95 and s ˆ 0.85, respectively. Tey are not very sensitive to te analytical results as te values adopted approximate tat for isentropic process. Te coefficient of te primary flow leaving te nozzle is taken as f p ˆ It was found tat te loss coefficient f m in Eq.(11) is more sensitive tan te oter coefficients and sould be taken to vary sligtly wit te ejector area ratio A 3 =A t in order to fit te test results. An empirical relation is found 8 0:80; for A 3 =A t 8:3; >< f m ˆ 0:8; for A 3 =A t 8:3; 19 >: 0:84; for A 3 =A t 6:9: Fig. 9 sows te variations of te ejector area ratio wit Table Constant-area section design and ejector specification ((XX): ejector model) Tableau Caractéristiques de la partie à section constante et de l éjecteur [(XX): modèle d éjecteur] Constant-area section Ejector specification Serial No. d 3 (mm) Inlet converging angle, ( ) A 3 /A t (wit Nozzle A) A 3 /A t (wit Nozzle E) A (AA) B (AB) G (AG) 6.77 (EG) C (AC) 7.6 (EC) D (AD) 8.5 (ED) E (EE) F (EF) H (EH)

7 360 B.J. Huang et al. / International Journal of Refrigeration (1999) Table 3 Comparisons of test and analytical results at P e ˆ MPa (8 C) ((XX): ejector specification; error ˆ (teory experiment)/experiment) Tableau 3 Comparaison des résultats expérimentaux et analytiques où P e ˆ 0,040 MPa (8 C) [(XX): modèle d éjecteur; erreur (téorique expérimentale)/(expérimentale)] P g, Mpa (T gs, C) T c *( C) A 3 /A t v Teory Experiment Error, % Teory Experiment Error, % (EH) (95) (EF) (AD) (EE) (AC) (ED) (EC) (AG) (EG) (AA) (AD) (90) (AC) (AG) (AB) (AA) (AD) (84) (AC) (AG) (AB) (AA) (AD) (78) (AC) (AG) (AB) (AA) Table 4 Comparisons of test and analytical results at P e ˆ MPa (1 C) ((XX): ejector specification; error ˆ (teory experiment)/experiment) Tableau 4 Comparaison des résultats expérimentaux et analytiques où P e ˆ 0,047 MPa (1 C) [(XX): modèle d éjecteur; erreur ˆ (téorique expérimentale)/(expérimentale)] P g, Mpa (T gs, C) T c *( C) A 3 /A t v Teory Experiment Error, % Teory Experiment Error, % (EF) (95) (EE) (AD) (AG) (EC) (AA) (AD) (90) (AG) (AA) (AD) (84) (AG) (AA) (AD) (78) (AG)

8 B.J. Huang et al. / International Journal of Refrigeration (1999) Fig. 5. Comparison between te teoretical calculation and te experimental design of te ejector area ratio at P e ˆ MPa (8 C). Fig. 5. Comparaison entre les valeurs calculée et expérimentale du rapport des surfaces de l éjecteur pour P e ˆ 0,040 MPa (8 C). Fig. 6. Comparison between te teoretical calculation and te experimental design of te ejector area ratio at P e ˆ MPa (1 C). Fig. 6. Comparaison entre les valeurs calculée et expérimentale du rapport de la superficie de l éjecteur pour P e ˆ 0,047 MPa (1 C).

9 36 B.J. Huang et al. / International Journal of Refrigeration (1999) Fig. 7. Comparison between te teoretical calculation and te experimental design of te ejector entrainment ratio at P e ˆ MPa (8 C). Fig. 7. Comparaison entre les valeurs calculée et expérimentale du rapport de l entraînement de l éjecteur pour P e ˆ 0,040 MPa (8 C). Fig. 8. Comparison between te teoretical calculation and te experimental design of te ejector entrainment ratio at P e ˆ MPa (1 C). Fig. 8. Comparaison entre les valeurs calculée et expérimentale du rapport de l entraînement de l éjecteur pour P e ˆ 0,047 MPa (1 C).

10 B.J. Huang et al. / International Journal of Refrigeration (1999) Fig. 9. Variation of ejector area ratio wit te vapor pressure at te nozzle inlet. Fig. 9. Variation du rapport des surfaces de l éjecteur en fonction de la pression à l entrée de la tuyère. te primary flow pressure and its saturated temperature at te inlet of te nozzle. Te teoretical or required A 3 =A t is seen to increase wit increasing primary flow inlet pressure P g for a fixed ejector critical back pressure P c *. A 3 =A t of an ejector also increases wit decreasing P c * for a fixed P g. Fig. 10 sows tat te measured v coincide wit te analysis. Te teoretical calculations also sow tat te entrainment ratio can be furter improved by raising te primary flow pressure wit a matcing ejector aving iger A 3 =A t. For R141b, it is seen tat v can reac 0.7 Fig. 10. Variation of ejector entrainment ratio wit te vapor pressure at te nozzle inlet. Fig. 10. Variation du rapport d entraînement de l éjecteur en fonction de la pression à l entrée de la tuyère.

11 364 B.J. Huang et al. / International Journal of Refrigeration (1999) at P g ˆ MPa (T gs ˆ 100 C) for an ejector aving a A 3 =A t 16:0. 4. Discussion In te 1-D analysis, we need to know te coefficients p ; s ; f p, and f m wic account for various losses in ejector. Tese values are related to te ejector design and manufacturing tecnique. Tey depend on macining, center-line alignment, interior surface polising, material used and suction port configuration etc. Te determination of tese coefficients relies on experiences. Te test results are used to determine te coefficients p ; s ; f p, and f m defined in te 1-D model by matcing te test data wit te analytical results. Te present 1-D analysis tus can be treated as semi-empirical from tis viewpoint. Bases on te experimental results obtained from te 11 ejectors wit good quality in macining, it is satisfactory to take p ˆ 0:95; s ˆ 0:85 and f p ˆ 0:88. Te loss coefficientf m was found to vary sligtly wit te ejector area ratio A 3 =A t and follows Eq. (19). However, a more convenient but roug relation can also be used: f m ˆ 1:037 0:0857 A 3 : 0 A t Te nozzle position is also an important factor affecting te ejector performance [11]. In te present study, te distance of te nozzle exit measured from te inlet of te constant-area section X is adjusted suc tat te ejector obtains te best performance at eac operating condition. It is found experimentally tat te ratio X=d 3 is around 1.50 for te best ejector performance and te test results obtained are used to compare wit te analysis. Te degree of supereating of te primary and entrained flows before entering te ejector is anoter factor tat may affect te performance. In te present study, we used no supereating in te primary flow. Te supereating of te entrained flow is in te range of 5 0 K in te present experiment. We found tat te ejector performance will not vary wit te degree of supereating of te entrained flow in tis range. Tis means tat te present results are valid for te ejector wit te degree of supereating exceeding 5 K in te entrained flow. Te need in te supereating depends on te particular termodynamic properties of te working fluid. For working fluid wit a negative-slope saturated-vapor line in te termodynamic T s diagram (i.e. bell-saped saturation line), an isentropic expansion of te vapor would possibly induce a condensation tat will seriously affect te gas dynamic process in te ejector and te performance of te ejector as well. Te working fluid R141b used in te present study as a positive-slope saturated-vapor line in te termodynamic T s diagram [9]. Terefore, supereating is not as important as oter working fluids. 5. Conclusion In te present study, we carried out a 1-D analysis for te prediction of te ejector performance at criticalmode operation. Te constant-pressure mixing is assumed to occur inside te constant-area section of te ejector and te entrained flow at coking condition is analyzed. We ave also carried out an experiment to verify te analytical results using 11 ejectors and R141b as te working fluid. Te test results are used to determine te coefficients p ; s ; f p,andf m defined in te 1-D model by matcing te test data wit te analytical results. It is sown tat te1-d analysis using te empirical coefficients can accurately predict te performance of te ejectors. Acknowledgements Te present study was supported by te National Science Council, ROC, Taiwan, troug Grant No.NSC E R. References [1] Sun D-W, Emes IW. Performance caracteristics of HCFC-13 ejector refrigeration cycles. Int. J. Energy Res. 1996;0: [] Sun D-W. Recent developments in te design teories and applications of ejectors a review. J. Inst. Energy 1995;68: [3] Keenan H, Neumann EP, Lustwerk F. An investigation of ejector design by analysis and experiment. J. Appl. Mec., Trans. ASME 1950;7: [4] Sun D-W. Variable geometry ejectors, teir applications in ejector refrigeration systems. Energy 1996;1: [5] Sun D-W. Experimental investigation of te performance caracteristics of a steam jet refrigeration system. Energy Sources 1997;19: [6] Munday, Jon T, Bagster DF. A new ejector teory applied to steam jet refrigeration. Ind. Engng Cem., Process Des. Dev. 1977;16: [7] Huang BJ, Jiang CB, Fu FL. Ejector performance caracteristics and design analysis of jet refrigeration system. ASME J. Engng Gas Turbines and Power 1985;107: [8] Eames IW, Apornratana S, Haider H. A teoretical and experimental study of a small-scale steam jet refrigerator. Int. J. Refrig. 1995;18: [9] Morrison G. Te sape of te temperature entropy saturation boundary. Int. J. Refrig. 1994;18:1 30. [10] Huang BJ, Cang JM. Empirical correlation for ejector design. Int. J. Refrig. 1999;: [11] Apornratana S, Eames IW. A small capacity steam-ejector refrigerator: experimental investigation of a system using ejector wit movable primary nozzle. Int. J. Refrig. 1997;0: