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A study of a bubble absorber using a plate heat exchanger with NH3–H2O, NH3–LiNO3 and NH3–NaSCN
Applied Thermal Engineering 31 (2011) 1869e1876
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A study of a bubble absorber using a plate heat exchanger with NH3eH2O, NH3eLiNO3 and NH3eNaSCN J. Cerezo a, *, R. Best b, R.J. Romero c a
Centro de Estudio de las Energías Renovables, Instituto de Ingeniería, UABC, Calle de la Normal s/n, Col. Insurgentes Este, 21280, B.C., Mexico Centro de Investigación en Energía, UNAM, Privada Xochicalco s/n, 62580 Morelos, Mexico c CIICAp e UAEM, Av. Universidad 1001, Chamilpa, 62209 Morelos, Mexico b
a r t i c l e i n f o
a b s t r a c t
Article history: Received 30 November 2010 Accepted 18 February 2011 Available online 12 March 2011
There is a continuous research effort being carried out worldwide on the development of absorption cooling systems with the objective of increasing their performance, one important area is the search for more efficient heat exchangers and alternative working fluids that can improve the performance of NH3eH2O and H2OeLiBr that are commonly used in commercial absorption refrigeration machines. In this work the study of a plate heat exchangers used as bubble absorbers with NH3eLiNO3 and NH3eNaSCN as alternative working fluids is carried out. A mathematical model was developed in order to analyze the absorption process in a bubble absorber with NH3eH2O, NH3eLiNO3 and NH3eNaSCN as working fluids using a plate heat exchanger at refrigeration conditions and low generator temperatures. The results show that NH3eH2O and NH3eNaSCN working fluids obtained higher absorber thermal loads and absorbed vapor mass values than NH3eLiNO3, the lower values for the latter were caused mainly by the high solution viscosity that decreases the efficiency of the absorption process. On the other hand, NH3eLiNO3 obtained the highest COP value from a single effect absorption refrigeration system simulation, however, NH3eNaSCN obtained a higher COP than NH3eH2O; therefore NH3eNaSCN seems to be a good working fluid to be tested in an absorption machine. Ó 2011 Elsevier Ltd. All rights reserved.
Keywords: Advanced cycle Plate heat exchanger Ammoniaewater Ammoniaelithium nitrate Ammoniaesodium thiocyanate heat pumps
1. Introduction Absorption refrigeration absorption machines are a very good option to significantly reduce electricity consumption as they use thermal energy to be activated (e.g. industrial waste heat, solar collectors, etc.). However, the absorption systems have lower coefficient of performance values than conventional vapor compression systems, caused mainly by inefficiencies in the absorber component [1]. Moreover, thermodynamically, the conventional working fluid NH3eH2O has a low performance in a single-stage absorption refrigeration system at low generator temperature and high absorber and condenser temperature conditions [2]. The absorption refrigeration system could be improved substituting the conventional NH3eH2O working fluid for other mixtures that have better thermodynamic and thermophysical properties for higher efficiency and heat and mass transfer processes. Sun [3] compared a refrigeration single-stage absorption
* Corresponding author. E-mail address: [email protected] (J. Cerezo). 1359-4311/$ e see front matter Ó 2011 Elsevier Ltd. All rights reserved. doi:10.1016/j.applthermaleng.2011.02.032
system with NH3eLiNO3, NH3eNaSCN and NH3eH2O at absorber temperature of 25 C. The author found that NH3eLiNO3 and NH3eNaSCN cycles were suitable alternatives to NH3eH2O absorption systems. Antonopoulos and Rogdakis [4] studied the performance of solar-driven NH3eLiNO3 and NH3-NaSCN absorption systems operating as coolers or heat pumps in the Athens area. For heating purposes, the NH3eLiNO3 system was superior to NH3eNaSCN, since it provided a higher heat-gain factor and useful thermal power. For cooling, the choice depended on the special requirements of each application, for the reason that the NH3eLiNO3 system provided a higher cooling power, whilst the NH3eNaSCN system achieved a higher coefficient of performance. The NH3eLiNO3 and NH3eNaSCN mixtures are thermodynamically attractive as they allow lower activation temperatures than NH3eH2O and do not require a rectifier. However, the study of the heat and mass transfer processes for these working mixtures in the absorber component is required as their higher viscosity could reduce the effectiveness of the absorption process [5]. The bubble absorber has been recommended because it has several advantages, due their very high heat and mass transfer coefficients [6]. Some researchers have studied the NH3eH2O
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absorbers in bubble mode with different configurations [6e10], however, Infante Ferreira [11] and Venegas et al. [12,13] are the only authors that have studied other absorbers with NH3eLiNO3. Infante Ferreira [11] studied a vertical tubular bubble absorber with the NH3eLiNO3 and NH3eNaSCN mixtures and concluded that the application of the Penetration and Nusselt theories to solve the mass and heat transfer equations respectively resulted in a low accuracy when compared with experimental results. Venegas et al. [12] applied the Newman model to predict the mass transfer at different zones: liquid jet, drop deceleration and uniform movement and concluded that the drop deceleration zone absorbed around 60 percent of the total vapor mass. The following year Venegas et al. [13] developed a mathematical model to study the absorption of droplets in a spray absorber using NH3eLiNO3 and concluded that the droplets reached an equilibrium state in less than 1 s, also an increase in subcooling causes the droplets to absorb more refrigerant vapor. There is scarce information regarding the absorption process with NH3eLiNO3 and NH3eNaSCN solutions; in particular the latter could be of interest because it has lower solution viscosity than NH3eLiNO3. The objective of this work is to compare the absorber performance in bubble mode between NH3eH2O, NH3eLiNO3 and NH3eNaSCN working fluids using a plate heat exchanger at commercial refrigeration conditions, and high condenser and absorber temperatures to allow for wet or dry cooling tower heat rejection conditions.
The Coefficient of performance (COP) for this system is defined as the ratio between evaporator load and generator load. The lowest generator temperature for each mixture was selected when stable values of the COP were reached, that is to say when a generator temperature does not significantly affect the COP. Fig. 2 shows the effect of generator temperature on COP. A minimum generator temperature value of 115 C was selected for the NH3eH2O and NH3eNaSCN working fluids, with a COP of 0.48 and 0.55, respectively. For NH3eLiNO3 a generator temperature of 105 C was selected with a COP value of 0.57. As it can be seen, NH3eLiNO3 mixture has the best COP, and it can be activated at lower generator temperatures than the other working fluids. NH3eH2O obtained the lowest COP value caused mainly by the penalty of the rectifier. Table 2a shows the input operation conditions for the absorber component obtained for a refrigeration capacity of 1 kW. The system was simulated using water as a cooling medium at an inlet temperature of 30 C and a mass flow rate of 110 kg/h to extract the absorber thermal load. The inlet solution temperature (TS,IN) of 55 C, was very similar for the three working fluids, however the solution concentration, solution flow rate and absorber capacity were different. NH3eH2O demanded the highest absorber heat load, QAB at 1.8 kW, whilst NH3eLiNO3 required the highest input ammonia concentration at 40.39% wt, which could help to reduce the problems of high solution viscosity. 3. Mathematical model of the bubble absorber
2. Absorption refrigeration cycle working conditions The main objective and advantage for absorption refrigeration cycles utilizing ammonia as refrigerant is to work at evaporator temperatures below the temperature of freezing water. Therefore, a temperature of 5 C was a predetermined condition, as well as ambient temperatures above 30 C for heat rejection. A simulation program of a single effect absorption refrigeration system was developed in order to select the working conditions for the system, specially the required generator temperatures. The model was based on mass and energy balances for each component [2]. Fig. 1 shows the components of a single effect absorption cycle for NH3eH2O used for this work. A single effect absorption refrigeration system for NH3eLiNO3 and NH3eNaSCN is very similar to NH3eH2O, the only difference being that a rectifier is not required. The transport and thermodynamic properties for the working fluids were obtained mainly from correlations reported by Infante [14], Conde [15] and Ibrahim [16]. A sensitivity analysis was developed fixing the evaporator temperature at 5 C and the condenser and absorber temperature at 40 C. The generator temperature was varied from 85 to 150 C.
Condenser
Rectifier
The mathematical model was based on constant absorber control volume increments as can be seen in Fig. 3 and explained in [10]. A plate heat exchanger is used as absorber due to the higher mass and heat transfer processes it can achieve [6,8]. Solution (mS) and vapor (mV) streams enter at the lower end and exit at the upper end of the control volume, whilst the cooling water flows in counter-current with the solution. Table 1 shows some geometrical details of the plate heat exchanger. 3.1. Model assumptions The model is based in the following assumptions: The absorption process takes place at steady state conditions. The vapor and liquid phases are in equilibrium at the interface zone.
Generator
Economizer
Evaporator
Absorber
Fig. 1. Single-stage absorption cycle for NH3eH2O.
Fig. 2. Coefficient of performance as a function of TGE in a single effect absorption system.
J. Cerezo et al. / Applied Thermal Engineering 31 (2011) 1869e1876
mS(i+1) x(i+1) hS(i+1)
mC(i) TC(i)
mV(i+1) y(i+1) hV(i+1)
Interface zone hREF,INT,V hABSOR,INT,V
hREF,INT,S hABSOR,INT,S QSEN,S
QAB mC(i) TC(i+1)
mS(i) x(i) hS(i)
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QSEN,V
mV(i) y(i) hV(i)
b
a Fig. 3. Control volume of bubble absorber.
The bubble velocity is constant. The vapor bubble has a spherical shape. There is no break-up and no interaction or coalescence amongst bubbles.
Q SEN;V ¼ cteV ðTINT TV Þ
(4)
The heat transferred from the interface to the liquid bulk is
Q SEN;S ¼ cteS ðTINT TS Þ The total heat transferred is then:
3.2. Mass and energy balance at the interface zone The mass transfer in the control volume was calculated by simultaneously solving the diffusion equation for the liquid and vapor phases. Equations (1) and (2) calculate the mass absorption flux (FAB) from the bulk vapor to the interface zone and from the interface to the bulk liquid phase, respectively, and both equations must be equal [17]:
z yINT FAB;REF þ FAB;ABSOR ¼ kmV In zy
(1)
zx FAB;REF þ FAB;ABSOR ¼ kmS In z xINT
(2)
Equations (1) and (2) are used to calculate the vapor absorbed for the NH3eH2O working fluid as the vapor and liquid phases contain both components (water and ammonia); however Equation (1) is eliminated for the NH3eLiNO3 and the NH3eNaSCN mixtures as the vapor phase contains only ammonia. The mass transfer correlations (kmV and kmS) are shown in Appendix A. The energy transferred between the vapor and liquid bulk is calculated as follows [17]:
cteS;V
mREF CpS;V;ABSOR þ mREF CpS;V;ABSOR hINT;S;V ¼ dA h mREF CpS;V;REF þ mABSOR CpS;V;ABSOR INT;S;V M hINT;SV 1e (3)
The heat transferred from the interface to the vapor bulk is defined in Fig. 3 as: Table 1 Plate heat exchanger geometry. Geometry
Value
Model Type No. of plates Plate area, m2 Plate length, m Pinch, m
NB51 L 4 0.05 0.5 0.002
(5)
Q l ¼ Q SEN;S þ Q SEN;V
(6)
The outlet vapor enthalpy is calculated from an energy balance on its respective phases with the following equation [10]:
mV ði þ 1Þ hV ði þ 1Þ ¼ Q SEN;V þ mV ðiÞ hV ðiÞ mREF hREF;INT;V mABSOR hABSOR;INT;V ð7Þ The absorber thermal load (QAB) and the outlet solution enthalpy (hS) are calculated solving simultaneously Equations (8), (9) and (12). The energy balance for Fig. 2 can be written in the following way: Table 2 Input data for bubble absorber operation, a) minimum activation temperature, b) similar flow rates, c) similar generator temperature. Parameter
NH3eH2O
NH3eLiNO3
NH3eNaSCN
a) Minimum activation temperature (115 C for NH3eH2O and NH3eNaSCN and 105 C for NH3eLiNO3) TS,IN, C 55.3 55.0 55.5 33.4 40.4 37 xIN, % weight 99 100 100 yIN, % weight Pressure, kPa 320 355 355 23.5 30.1 33.9 mS,IN, kg/h 1.82 1.60 1.62 QAB, kW COP 0.48 0.57 0.55 b) Similar flow rates (23.5 kg/h). 55.3 TS,IN, C xIN, % weight 33.4 99 yIN, % weight Pressure, kPa 320 23.5 mS,IN, kg/h 1.82 QAB, kW COP 0.48
55.0 40.4 100 355 23.5 1.24 0.57
55.5 37.0 100 355 23.5 1.13 0.55
c) Similar generator temperature (120 C) 55.8 TS,IN, C xIN, % weight 31.2 99 yIN, % weight Pressure, kPa 320 18.6 mS,IN, kg/h 1.76 QAB, kW COP 0.49
55.6 35.6 100 355 16.8 1.48 0.61
55.5 35.6 100 355 27.4 1.62 0.55
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Fig. 6. Temperature profiles along bubble absorber length for NH3eNaSCN.
Fig. 4. Temperature profiles along bubble absorber length for NH3eH2O.
ms ði þ 1Þ hs ði þ 1Þ ¼ Q SEN;S Q AB þ ms ðiÞ hs ðiÞ þ mREF hREF;INT;S þ mABSOR hABSOR;INT;S
ð8Þ
where hREF,INT,V, hREF,INT,S, hABSOR,INT,V, hABSOR,INT,S are the specific enthalpies for each component in the vapor and solution phases at the interfacial conditions. The total absorber heat must be calculated by UA method:
Q AB ¼ UADTML
(9)
where A is the heat transfer area in m2, and the logarithmic mean temperature for this configuration is defined as:
DTML ¼
TS;IN TC;OUT TS;OUT TC;IN TS;IN TC;OUT Ln TS;OUT TC;IN
(10)
The global heat transfer coefficient considers three sections:
U ¼
1 Dw 1 1 þ þ hS K hC
(11)
where the solution (hS) and cooling water (hC) heat transfer coefficients are calculated from Herbine and Perez-Blanco [9] and Cerezo [2], respectively as shown in Appendix A. For a constant pressure, the heat transferred to the cooling water is calculated as follows:
Q AB ¼ mC CpC ½TC ði þ 1Þ TðiÞ
(12)
Fig. 5. Temperature profiles along bubble absorber length for NH3eLiNO3.
The bubble absorber simulation was previously validated with experimental data for the NH3eH2O working fluid with acceptable values [10]. 4. Results The absorber analysis was realized comparing the different working mixtures at three different conditions: minimum activation temperature, similar flow rates and similar generator temperatures. Table 2 shows a summary of the input data for the three analyzed conditions. 4.1. Results for minimum activation temperature conditions The minimum activation temperature refers to the minimum generator temperature at which the performance of the absorption refrigeration system is stable as it was explained on Section 2 and Table 2a. Figs. 4e6 show temperature profiles as a function of absorber length for the NH3eH2O, NH3eLiNO3 and NH3eNaSCN mixtures, respectively. The vapor temperature increases exponentially up to 0.2 m length from the bottoms of the absorber, after that the vapor temperature increases slightly for each mixture. This temperature increment is caused by the large vapor sensible heat caused by the temperature difference between the bulk vapor and the interface (5 and 55 C, respectively). The solution temperature decreases almost linearly with a temperature drop of 17.9, 18.3 and 20.6 C for NH3eH2O, NH3eLiNO3 and NH3eNaSCN, respectively. The solution
Fig. 7. Total absorber heat load and vapor flow rate along bubble absorber length.
J. Cerezo et al. / Applied Thermal Engineering 31 (2011) 1869e1876
900
1873
0.05
800 700 600 NH3-H2O NH3-LiNO3 NH3-NaSCN
Re
500
0.03
400 0.02 300 200
Velocity, m/s
0.04
0.01
100 0
0 0
0.1
0.2
0.3
0.4
0.5
Length, m
Fig. 10. Absorber thermal load along bubble absorber length at similar solution and vapor flow rates.
Fig. 8. Reynolds number and velocity of solution along bubble absorber length.
As Fig. 9 shows, for the NH3eH2O, NH3eNaSCN and NH3eLiNO3 mixtures, the viscosity range goes from the values of 0.50 to 0.67, 1.85 to 1.83 and 2.00 to 2.77 cp, respectively, at the solution temperatures shown in Figs. 4e6. The viscosity of the NH3eLiNO3 and NH3eNaSCN solutions is 4.1 and 2.7 times higher than for NH3eH2O at the end absorber. Appendix A shows the viscosity and hS and hc correlations used for this study.
and interface temperatures have a similar tendency and values. The cooling water temperature increases from 30.0 to 43.5 C for NH3eH2O, from 30.0 to 40.4 C for NH3eLiNO3 and from 30.0 to 43.8 C for NH3eNaSCN. Fig. 7 shows the absorber thermal load and vapor flow rate as a function of absorber length for the different working mixtures. The thermal load value was around 1.68 kW for NH3eH2O and NH3eNaSCN, but for NH3eLiNO3 a lower value of 1.30 kW was obtained. The absorber heat load for NH3eH2O and NH3eLiNO3 resulted in a value 7.7% and 18.7% lower than the value obtained from the simulation of the refrigeration system, respectively, as shown in Table 2a, however, the absorber heat load for NH3eNaSCN resulted in a 3.6% higher value. In the same figure, the vapor flow rate decreases from 0.00092 to 0.000065, 0.00093 to 0.00035 and 0.00093 to 0.000081 kg/s for NH3eH2O, NH3eLiNO3 and NH3eNaSCN, respectively. In terms of absorption flux (FAB), the following values were obtained 0.0086, 0.0058 and 0.0085 kg/m2 s, for NH3eH2O, NH3eLiNO3 and NH3eNaSCN respectively. It is clear that NH3eLiNO3 has the higher COP than the other mixtures, although it has difficulties in the absorption process. As it can be seen, the vapor flow rate is not totally absorbed in the absorber, so extra heat and mass transfer area is required. Fig. 8 shows that NH3eNaSCN achieves the highest solution velocity values (w0.05 m/s) caused by the high solution flow rate (see Table 2a), although NH3eH2O resulted with the highest turbulence values (Re ¼ 250e200) and NH3eLiNO3 the lowest values (Re ¼ 90e110) due to its higher viscosity.
In this case the inlet vapor and solution flow rates for each working fluid in the bubble absorber have the same values than those obtained for the NH3eH2O mixture at the simulated conditions shown in Table 2b. Figs. 10e13 show the absorber thermal load and vapor flow rate, the overall heat transfer coefficient, the hold-up and the Reynolds number and velocity, respectively, for each working fluid as a function of absorber length. The absorber thermal load obtained was 1.68 kW for NH3eH2O, 1.47 kW for NH3eNaSCN and the lowest value corresponded to NH3eLiNO3 at 1.35 kW, as shown in Fig. 10. As Fig. 11 shows, the NH3eH2O and NH3eNaSCN mixtures obtained higher U values than NH3eLiNO3 due to their higher hold-up and thermal conductivity values, respectively, that improved the individual heat transfer coefficients. The FAB values were 0.0082 and 0.0056 kg/m2 s for NH3eNaSCN and NH3eLiNO3 respectively. The FAB and thermal load decrease as expected, due to the lower value of the solution flow rate.
Fig. 9. Solution viscosity along bubble absorber length.
Fig. 11. Overall heat transfer coefficient profile along bubble absorber length at similar solution and vapor flow rates.
4.2. Results for similar flow rate conditions
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Fig. 12. Gas hold-up profile along bubble absorber length at similar solution and vapor flow rates.
The NH3eH2O mixture obtained larger U values than NH3eNaSCN at the beginning of the absorber length, but the U value relationship is inverted for lengths higher than 0.20 m as it is shown in Fig. 11. The low value of the NH3eLiNO3 overall heat transfer coefficient is caused mainly by the high viscosity (similar viscosity values from Fig. 9). The U value depends mainly on the solution side, as the cooling water heat transfer coefficient is kept almost constant (6 kW/m2 K) along the absorber length. Fig. 12 shows that NH3eH2O achieves the highest hold-up value (defined as the volume ratio of gas phase in the mixture) and it then decreases linearly from 0.95 to 0.30. For NH3eLiNO3 and NH3eNaSCN the hold-up values decreases exponentially from 0.80 to 0.30 and 0.80 to 0.15, respectively. The solution velocity is different for each working fluid, 0.041, 0.036 and 0.032 m/s for NH3eH2O, NH3eNaSCN and NH3eLiNO3, respectively, as shown in Fig. 13. Moreover, the Reynolds number increases from 69 to 90 for NH3eLiNO3, but it decreases from 256 to 211 and from 69 to 57 for NH3eH2O and NH3eNaSCN, respectively, this decrement is caused by the lower solution temperature at the absorber outlet. The NH3eH2O mixture obtained the maximum turbulence, whilst the minimum turbulence value was obtained for NH3-LiNO3 mixture. 4.3. Results for similar activation temperature conditions For similar activation temperatures the input data were obtained from the absorption refrigeration system simulation at similar generator temperatures. The generator temperature was
Fig. 13. Solution Reynolds number and velocity profile along bubble absorber length at similar solution and vapor flow rates.
Fig. 14. Absorber thermal load and vapor flow rate along bubble absorber length at TGE ¼ 120 C.
fixed at 120 C, TCO and TAB was fixed at 40 C and TEV at 5 C, for a 1 kW of cooling capacity as shown in Table 2c. The NH3eLiNO3 mixture obtains the highest COP value of 0.61 whilst NH3eH2O obtained the lowest value of 0.49. Fig. 14 shows the absorber thermal load and vapor flow rate as a function of absorber length for each working fluid. The absorber thermal load follows a very similar tendency and value for NH3eH2O and NH3eNaSCN at 1.66 kW, however NH3eLiNO3 obtained a value of 1.27 kW. The FAB obtained similar values for NH3eH2O and NH3eNaSCN (0.0087 kg/m2 s), but NH3eLiNO3 obtained the lowest value (0.0067 kg/m2 s), caused by the high viscosity that affects the heat and mass transfer process. 5. Conclusions A bubble absorber was simulated with three different working fluids: NH3eH2O, NH3eLiNO3 and NH3eNaSCN in order to analyze the absorber performance using a plate exchanger. The bubble absorber working conditions were obtained from an absorption refrigeration system simulation of 1 kW cooling capacity at three different conditions: minimum generator temperature, similar flow rate and similar generator temperature. The following conclusions were drawn from the absorber analysis. At minimum generator temperature conditions the NH3eH2O and NH3eNaSCN working fluids had an absorber thermal load value of 1.68 kW, while NH3eLiNO3 obtained a value of 1.30 kW, also the mass absorption flux was 0.0086, 0.0058 and 0.0085 kg/m2 s for NH3eH2O, NH3eLiNO3 and NH3eNaSCN, respectively. The NH3eLiNO3 mixture obtained the lowest values for both parameters. At similar flow rates, NH3eH2O, NH3eNaSCN and NH3eLiNO3 obtained 1.69, 1.47 and 1.35 kW of absorber thermal load respectively. NH3eH2O and NH3eNaSCN obtained a higher heat transfer coefficient than NH3eLiNO3 caused by higher values of gas hold-up and thermal conductivity. At similar generator temperatures, the NH3eH2O and NH3eNaSCN working fluids obtained higher absorber thermal load values than NH3eLiNO3. Also, NH3eLiNO3 obtained the lowest mass absorption flux value (0.0067 kg/m2 s) caused by its higher viscosity. The NH3eLiNO3 working fluid obtained the lower absorption mass transfer and absorber thermal load values than NH3eH2O and NH3eNaSCN in the three cases studied, due mainly to its higher viscosity that causes problems with the absorption of NH3 vapor in the NH3eLiNO3 solution; despite that it, it obtained the highest COP in the absorption refrigeration system simulation. On the other
J. Cerezo et al. / Applied Thermal Engineering 31 (2011) 1869e1876
hand NH3eNaSCN seems to be a better alternative as it attained better results for the absorber performance than NH3eLiNO3 and a higher COP than NH3eH2O. Acknowledgements The authors would like to thank UNAM for partially funding the work through project PAPIIT IN109009. This document has been produced with the financial support of the European Union and Mexico through project FONCICYT 94256. The contents of this document are the sole responsibility of the Instituto de Ingeniería de la Universidad Autónoma de Baja California and the Centro de Investigación en Energía de la Universidad Nacional Autónoma de México and can under no circumstance be regarded as reflecting the position of the European Union and Mexico. Dr. Jesus Cerezo would like to thank UNAM for the postdoctoral scholarship at the Centro de Investigación en Energía. The authors would like to thank Carmen Huerta for her support on the use of EES software.
Appendix A This section shows correlation used for different mixtures used to calculate the viscosity (m). They were obtained from Infante [14] and Conde [15]. NH3eH2O
m ¼ exp x ln mNH3 ;S þ ð1 xÞln mNH3 ;S
(A.1)
T þ 273:15 fs 0:00001 þ 0:52 0:815 TNH3 ;CRIT þ 273:15 NH3eLiNO3
m ¼ ð5:1835T þ 992:337Þ ð1 xÞð0:08333Tþ6:8333Þ þ expð0:01147T 1:744Þ 0:001
hS ¼ hLIQ ð1 holdupÞ0:8
(A.5)
1= PrC 3 kC d1 hc ¼ 0:99 Re0:529 C H
(A.6)
hLIQ is the solution heat transfer coefficient at liquid phase, but there is no values for this type of plate heat exchanger, then hLIQ was supposed to be obtained from experimental values for water from Equation A.6 The solution and vapor mass transfer coefficients were calculated from Incropera [18] and Sherwood [19].
0:666 0:116 dB g0:333 bS ShINT;S ¼ 2 þ 0:0187Re0:779 Sc0:546 B s 1=3
ShINT;V ¼ 0:664Re0:5 B ScV
A COP cp dH FAB h K km m Q Re Sh T U x
z
(A.2)
m
x < 50% wt
m ¼ 5:79278:9xþ390x2 176x3
þ 2:692þ11:95x17:24x2 þ7:974x3 T þ 0:058850:2701xþ0:3994x2 0:1882x3 T2 þ 0:0004977þ0:002339x0:003532x2 þ0:001691x3 T3 0:001
ðA:3Þ
x > 50% wt
m ¼ ð0:5289T þ 29695Þ ð1 xÞ3:9365 þ 2:19 11
10
4:636
ð120 TÞ
þ0:1 0:001
(A.7) (A.8)
Nomenclature
y
NH3eNaSCN
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area (m2) coefficient of operation heat capacity (kJ/kg K) hydraulic diameter (m) Mass absorption flux per heat transfer area (kg/m2 s) Specific enthalpy (kJ/kg), or heat transfer coefficient (kW/ m2 K) Thermal conductivity (kW/m K) Mass transfer coefficient (kg/m2 s) Mass flow rate (kg/s) Thermal load (kW) Reynolds number Sherwood number temperature ( C) overall heat transfer coefficient (kW/m2 K) weight concentration in solution phase (kg refrigerant/kg solution) weight concentration in vapor phase (kg refrigerant/kg solution) ratio of ammonia mass flux to the total mass flux absorbed dynamic viscosity (cp)
Subscripts AB absorber ABSOR absorbent C cooling water water H2O IN input INT interface ammonia NH3 OUT outlet REF refrigerant S solution SEN sensible V vapor phase w wall References
(A.4)
The solution (hS) and cooling (hc) heat transfer coefficients were obtained by Herbine and Perez-Blanco [9] and Cerezo [10], respectively
[1] A.M. Selim, M.M. Elsayed, Performance of a packed bed absorber for water ammonia absorption refrigeration system, International Journal of Refrigeration 22 (1999) 283e292. [2] Cerezo, J. Estudio del proceso de absorción con amoniaco-agua en intercambiadores de placas para equipos de refrigeración por absorción, PhD Thesis, Universidad Rovira i Virgili, Spain, 2006.
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[3] D.W. Sun, Comparison of the performances of NH3-H2O, NH3-LiNO3 and NH3-NaSCN absorption refrigeration systems, Energy Conversion and Management 39 (5/6) (1998) 357e368. [4] K.A. Antonopoulos, E.D. Rogdakis, Performance of solar-driven ammonialithium nitrate and ammonia-sodium thiocyanate absorption systems operating as coolers or heat pumps in Athens, Applied Thermal Engineering 16 (2) (1996) 127e147. [5] Heard, C.L. On the viscosity of the absorption heat pump working pair: ammonia/lithium nitrate and improved heat transfer correlation for boiling in a vertical tube, Instituto Mexicano del Petroleo, Mexico, personal communication, 2000. [6] Y.T. Kang, R.N. Christensen, T. Kashiwagi, Ammonia-water bubble absorber with a plate heat exchanger, ASHRAE Transaction 204 (1) (1998). [7] J. Fernández-Seara, J. Sieres, C. Rodríguez, M. Vázquez, Ammoniaewater absorption in vertical tubular absorbers, International Journal of Thermal Sciences 44 (2005) 277e288. [8] K.B. Lee, B.H. Chung, J.C. Lee, C.H. Lee, S.H. Kim, Experimental analysis bubble mode in a plate-type absorber, Chemical Enginnering Science 57 (2002) 1923e1929. [9] G.S. Herbine, H. Perez-Blanco, Model of an ammonia-water bubble absorber, Ashrae Transactions 101 (1) (1995) 1324e1331. [10] J. Cerezo, R. Best, M. Bourouis, A. Coronas, Comparison of numerical and experimental performance criteria of an ammonia-water bubble absorber
[11] [12]
[13]
[14]
[15]
[16] [17] [18] [19]
using plate heat exchanger, International Journal of Heat and Mass Transfer 53 (17e18) (2010) 3379e3386. C.A. Infante Ferreira, Combined momentum, heat and mass transfer in vertical slug flow absorbers, International Journal of Refrigeration 8 (6) (1985) 326e334. M. Venegas, M. Izquierdo, P. Rodriguez, A. Lacuona, Heat and mass transfer during absorption of ammonia vapour by LiNO3-NH3 solution droplets, International Journal of Heat and Mass Transfer 47 (2004) 2653e2667. M. Venegas, P. Rodriguez, A. Lecuona, M. Izquierdo, Spray absorbers in absorption systems using lithium nitrate-ammonia solution, International Journal of Refrigeration 28 (4) (2005) 554e564. C.A. Infante Ferreira, Thermodynamic and physical property data equations for ammonia-lithium nitrate and ammonia-sodium thiocyanate solutions, Solar Energy 32 (1984) 231e236. M.R. Conde Petit, Thermophysical Properties of NH3-H2O Solutions for Industrial Design for Absorption Refrigeration Equipment (2004) M. Conde Engineering, Zurich, Switzerland. O.M. Ibrahim, S.A. Klein, Thermodynamic properties of ammonia-water mixtures, ASHRAE Transactions: Symposia 21 (2) (1993). R.E. Treybal, Mass Transfer Operations, third ed.. McGraw Hill, Tokyo, 1980. F.P. Incropera, D.P. De Witt, Fundamental of Heat and Mass Transfer, third ed. John Wiley & Sons Cop., New York, 1990. T.K. Sherwood, R.L. Pigford, C.R. Wlke, Mass Transfer. Mc Graw-Hill, New York, 1975.
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A study of a bubble absorber using a plate heat exchanger with NH3eH2O, NH3eLiNO3 and NH3eNaSCN J. Cerezo a, *, R. Best b, R.J. Romero c a
Centro de Estudio de las Energías Renovables, Instituto de Ingeniería, UABC, Calle de la Normal s/n, Col. Insurgentes Este, 21280, B.C., Mexico Centro de Investigación en Energía, UNAM, Privada Xochicalco s/n, 62580 Morelos, Mexico c CIICAp e UAEM, Av. Universidad 1001, Chamilpa, 62209 Morelos, Mexico b
a r t i c l e i n f o
a b s t r a c t
Article history: Received 30 November 2010 Accepted 18 February 2011 Available online 12 March 2011
There is a continuous research effort being carried out worldwide on the development of absorption cooling systems with the objective of increasing their performance, one important area is the search for more efficient heat exchangers and alternative working fluids that can improve the performance of NH3eH2O and H2OeLiBr that are commonly used in commercial absorption refrigeration machines. In this work the study of a plate heat exchangers used as bubble absorbers with NH3eLiNO3 and NH3eNaSCN as alternative working fluids is carried out. A mathematical model was developed in order to analyze the absorption process in a bubble absorber with NH3eH2O, NH3eLiNO3 and NH3eNaSCN as working fluids using a plate heat exchanger at refrigeration conditions and low generator temperatures. The results show that NH3eH2O and NH3eNaSCN working fluids obtained higher absorber thermal loads and absorbed vapor mass values than NH3eLiNO3, the lower values for the latter were caused mainly by the high solution viscosity that decreases the efficiency of the absorption process. On the other hand, NH3eLiNO3 obtained the highest COP value from a single effect absorption refrigeration system simulation, however, NH3eNaSCN obtained a higher COP than NH3eH2O; therefore NH3eNaSCN seems to be a good working fluid to be tested in an absorption machine. Ó 2011 Elsevier Ltd. All rights reserved.
Keywords: Advanced cycle Plate heat exchanger Ammoniaewater Ammoniaelithium nitrate Ammoniaesodium thiocyanate heat pumps
1. Introduction Absorption refrigeration absorption machines are a very good option to significantly reduce electricity consumption as they use thermal energy to be activated (e.g. industrial waste heat, solar collectors, etc.). However, the absorption systems have lower coefficient of performance values than conventional vapor compression systems, caused mainly by inefficiencies in the absorber component [1]. Moreover, thermodynamically, the conventional working fluid NH3eH2O has a low performance in a single-stage absorption refrigeration system at low generator temperature and high absorber and condenser temperature conditions [2]. The absorption refrigeration system could be improved substituting the conventional NH3eH2O working fluid for other mixtures that have better thermodynamic and thermophysical properties for higher efficiency and heat and mass transfer processes. Sun [3] compared a refrigeration single-stage absorption
* Corresponding author. E-mail address: [email protected] (J. Cerezo). 1359-4311/$ e see front matter Ó 2011 Elsevier Ltd. All rights reserved. doi:10.1016/j.applthermaleng.2011.02.032
system with NH3eLiNO3, NH3eNaSCN and NH3eH2O at absorber temperature of 25 C. The author found that NH3eLiNO3 and NH3eNaSCN cycles were suitable alternatives to NH3eH2O absorption systems. Antonopoulos and Rogdakis [4] studied the performance of solar-driven NH3eLiNO3 and NH3-NaSCN absorption systems operating as coolers or heat pumps in the Athens area. For heating purposes, the NH3eLiNO3 system was superior to NH3eNaSCN, since it provided a higher heat-gain factor and useful thermal power. For cooling, the choice depended on the special requirements of each application, for the reason that the NH3eLiNO3 system provided a higher cooling power, whilst the NH3eNaSCN system achieved a higher coefficient of performance. The NH3eLiNO3 and NH3eNaSCN mixtures are thermodynamically attractive as they allow lower activation temperatures than NH3eH2O and do not require a rectifier. However, the study of the heat and mass transfer processes for these working mixtures in the absorber component is required as their higher viscosity could reduce the effectiveness of the absorption process [5]. The bubble absorber has been recommended because it has several advantages, due their very high heat and mass transfer coefficients [6]. Some researchers have studied the NH3eH2O
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absorbers in bubble mode with different configurations [6e10], however, Infante Ferreira [11] and Venegas et al. [12,13] are the only authors that have studied other absorbers with NH3eLiNO3. Infante Ferreira [11] studied a vertical tubular bubble absorber with the NH3eLiNO3 and NH3eNaSCN mixtures and concluded that the application of the Penetration and Nusselt theories to solve the mass and heat transfer equations respectively resulted in a low accuracy when compared with experimental results. Venegas et al. [12] applied the Newman model to predict the mass transfer at different zones: liquid jet, drop deceleration and uniform movement and concluded that the drop deceleration zone absorbed around 60 percent of the total vapor mass. The following year Venegas et al. [13] developed a mathematical model to study the absorption of droplets in a spray absorber using NH3eLiNO3 and concluded that the droplets reached an equilibrium state in less than 1 s, also an increase in subcooling causes the droplets to absorb more refrigerant vapor. There is scarce information regarding the absorption process with NH3eLiNO3 and NH3eNaSCN solutions; in particular the latter could be of interest because it has lower solution viscosity than NH3eLiNO3. The objective of this work is to compare the absorber performance in bubble mode between NH3eH2O, NH3eLiNO3 and NH3eNaSCN working fluids using a plate heat exchanger at commercial refrigeration conditions, and high condenser and absorber temperatures to allow for wet or dry cooling tower heat rejection conditions.
The Coefficient of performance (COP) for this system is defined as the ratio between evaporator load and generator load. The lowest generator temperature for each mixture was selected when stable values of the COP were reached, that is to say when a generator temperature does not significantly affect the COP. Fig. 2 shows the effect of generator temperature on COP. A minimum generator temperature value of 115 C was selected for the NH3eH2O and NH3eNaSCN working fluids, with a COP of 0.48 and 0.55, respectively. For NH3eLiNO3 a generator temperature of 105 C was selected with a COP value of 0.57. As it can be seen, NH3eLiNO3 mixture has the best COP, and it can be activated at lower generator temperatures than the other working fluids. NH3eH2O obtained the lowest COP value caused mainly by the penalty of the rectifier. Table 2a shows the input operation conditions for the absorber component obtained for a refrigeration capacity of 1 kW. The system was simulated using water as a cooling medium at an inlet temperature of 30 C and a mass flow rate of 110 kg/h to extract the absorber thermal load. The inlet solution temperature (TS,IN) of 55 C, was very similar for the three working fluids, however the solution concentration, solution flow rate and absorber capacity were different. NH3eH2O demanded the highest absorber heat load, QAB at 1.8 kW, whilst NH3eLiNO3 required the highest input ammonia concentration at 40.39% wt, which could help to reduce the problems of high solution viscosity. 3. Mathematical model of the bubble absorber
2. Absorption refrigeration cycle working conditions The main objective and advantage for absorption refrigeration cycles utilizing ammonia as refrigerant is to work at evaporator temperatures below the temperature of freezing water. Therefore, a temperature of 5 C was a predetermined condition, as well as ambient temperatures above 30 C for heat rejection. A simulation program of a single effect absorption refrigeration system was developed in order to select the working conditions for the system, specially the required generator temperatures. The model was based on mass and energy balances for each component [2]. Fig. 1 shows the components of a single effect absorption cycle for NH3eH2O used for this work. A single effect absorption refrigeration system for NH3eLiNO3 and NH3eNaSCN is very similar to NH3eH2O, the only difference being that a rectifier is not required. The transport and thermodynamic properties for the working fluids were obtained mainly from correlations reported by Infante [14], Conde [15] and Ibrahim [16]. A sensitivity analysis was developed fixing the evaporator temperature at 5 C and the condenser and absorber temperature at 40 C. The generator temperature was varied from 85 to 150 C.
Condenser
Rectifier
The mathematical model was based on constant absorber control volume increments as can be seen in Fig. 3 and explained in [10]. A plate heat exchanger is used as absorber due to the higher mass and heat transfer processes it can achieve [6,8]. Solution (mS) and vapor (mV) streams enter at the lower end and exit at the upper end of the control volume, whilst the cooling water flows in counter-current with the solution. Table 1 shows some geometrical details of the plate heat exchanger. 3.1. Model assumptions The model is based in the following assumptions: The absorption process takes place at steady state conditions. The vapor and liquid phases are in equilibrium at the interface zone.
Generator
Economizer
Evaporator
Absorber
Fig. 1. Single-stage absorption cycle for NH3eH2O.
Fig. 2. Coefficient of performance as a function of TGE in a single effect absorption system.
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mS(i+1) x(i+1) hS(i+1)
mC(i) TC(i)
mV(i+1) y(i+1) hV(i+1)
Interface zone hREF,INT,V hABSOR,INT,V
hREF,INT,S hABSOR,INT,S QSEN,S
QAB mC(i) TC(i+1)
mS(i) x(i) hS(i)
1871
QSEN,V
mV(i) y(i) hV(i)
b
a Fig. 3. Control volume of bubble absorber.
The bubble velocity is constant. The vapor bubble has a spherical shape. There is no break-up and no interaction or coalescence amongst bubbles.
Q SEN;V ¼ cteV ðTINT TV Þ
(4)
The heat transferred from the interface to the liquid bulk is
Q SEN;S ¼ cteS ðTINT TS Þ The total heat transferred is then:
3.2. Mass and energy balance at the interface zone The mass transfer in the control volume was calculated by simultaneously solving the diffusion equation for the liquid and vapor phases. Equations (1) and (2) calculate the mass absorption flux (FAB) from the bulk vapor to the interface zone and from the interface to the bulk liquid phase, respectively, and both equations must be equal [17]:
z yINT FAB;REF þ FAB;ABSOR ¼ kmV In zy
(1)
zx FAB;REF þ FAB;ABSOR ¼ kmS In z xINT
(2)
Equations (1) and (2) are used to calculate the vapor absorbed for the NH3eH2O working fluid as the vapor and liquid phases contain both components (water and ammonia); however Equation (1) is eliminated for the NH3eLiNO3 and the NH3eNaSCN mixtures as the vapor phase contains only ammonia. The mass transfer correlations (kmV and kmS) are shown in Appendix A. The energy transferred between the vapor and liquid bulk is calculated as follows [17]:
cteS;V
mREF CpS;V;ABSOR þ mREF CpS;V;ABSOR hINT;S;V ¼ dA h mREF CpS;V;REF þ mABSOR CpS;V;ABSOR INT;S;V M hINT;SV 1e (3)
The heat transferred from the interface to the vapor bulk is defined in Fig. 3 as: Table 1 Plate heat exchanger geometry. Geometry
Value
Model Type No. of plates Plate area, m2 Plate length, m Pinch, m
NB51 L 4 0.05 0.5 0.002
(5)
Q l ¼ Q SEN;S þ Q SEN;V
(6)
The outlet vapor enthalpy is calculated from an energy balance on its respective phases with the following equation [10]:
mV ði þ 1Þ hV ði þ 1Þ ¼ Q SEN;V þ mV ðiÞ hV ðiÞ mREF hREF;INT;V mABSOR hABSOR;INT;V ð7Þ The absorber thermal load (QAB) and the outlet solution enthalpy (hS) are calculated solving simultaneously Equations (8), (9) and (12). The energy balance for Fig. 2 can be written in the following way: Table 2 Input data for bubble absorber operation, a) minimum activation temperature, b) similar flow rates, c) similar generator temperature. Parameter
NH3eH2O
NH3eLiNO3
NH3eNaSCN
a) Minimum activation temperature (115 C for NH3eH2O and NH3eNaSCN and 105 C for NH3eLiNO3) TS,IN, C 55.3 55.0 55.5 33.4 40.4 37 xIN, % weight 99 100 100 yIN, % weight Pressure, kPa 320 355 355 23.5 30.1 33.9 mS,IN, kg/h 1.82 1.60 1.62 QAB, kW COP 0.48 0.57 0.55 b) Similar flow rates (23.5 kg/h). 55.3 TS,IN, C xIN, % weight 33.4 99 yIN, % weight Pressure, kPa 320 23.5 mS,IN, kg/h 1.82 QAB, kW COP 0.48
55.0 40.4 100 355 23.5 1.24 0.57
55.5 37.0 100 355 23.5 1.13 0.55
c) Similar generator temperature (120 C) 55.8 TS,IN, C xIN, % weight 31.2 99 yIN, % weight Pressure, kPa 320 18.6 mS,IN, kg/h 1.76 QAB, kW COP 0.49
55.6 35.6 100 355 16.8 1.48 0.61
55.5 35.6 100 355 27.4 1.62 0.55
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Fig. 6. Temperature profiles along bubble absorber length for NH3eNaSCN.
Fig. 4. Temperature profiles along bubble absorber length for NH3eH2O.
ms ði þ 1Þ hs ði þ 1Þ ¼ Q SEN;S Q AB þ ms ðiÞ hs ðiÞ þ mREF hREF;INT;S þ mABSOR hABSOR;INT;S
ð8Þ
where hREF,INT,V, hREF,INT,S, hABSOR,INT,V, hABSOR,INT,S are the specific enthalpies for each component in the vapor and solution phases at the interfacial conditions. The total absorber heat must be calculated by UA method:
Q AB ¼ UADTML
(9)
where A is the heat transfer area in m2, and the logarithmic mean temperature for this configuration is defined as:
DTML ¼
TS;IN TC;OUT TS;OUT TC;IN TS;IN TC;OUT Ln TS;OUT TC;IN
(10)
The global heat transfer coefficient considers three sections:
U ¼
1 Dw 1 1 þ þ hS K hC
(11)
where the solution (hS) and cooling water (hC) heat transfer coefficients are calculated from Herbine and Perez-Blanco [9] and Cerezo [2], respectively as shown in Appendix A. For a constant pressure, the heat transferred to the cooling water is calculated as follows:
Q AB ¼ mC CpC ½TC ði þ 1Þ TðiÞ
(12)
Fig. 5. Temperature profiles along bubble absorber length for NH3eLiNO3.
The bubble absorber simulation was previously validated with experimental data for the NH3eH2O working fluid with acceptable values [10]. 4. Results The absorber analysis was realized comparing the different working mixtures at three different conditions: minimum activation temperature, similar flow rates and similar generator temperatures. Table 2 shows a summary of the input data for the three analyzed conditions. 4.1. Results for minimum activation temperature conditions The minimum activation temperature refers to the minimum generator temperature at which the performance of the absorption refrigeration system is stable as it was explained on Section 2 and Table 2a. Figs. 4e6 show temperature profiles as a function of absorber length for the NH3eH2O, NH3eLiNO3 and NH3eNaSCN mixtures, respectively. The vapor temperature increases exponentially up to 0.2 m length from the bottoms of the absorber, after that the vapor temperature increases slightly for each mixture. This temperature increment is caused by the large vapor sensible heat caused by the temperature difference between the bulk vapor and the interface (5 and 55 C, respectively). The solution temperature decreases almost linearly with a temperature drop of 17.9, 18.3 and 20.6 C for NH3eH2O, NH3eLiNO3 and NH3eNaSCN, respectively. The solution
Fig. 7. Total absorber heat load and vapor flow rate along bubble absorber length.
J. Cerezo et al. / Applied Thermal Engineering 31 (2011) 1869e1876
900
1873
0.05
800 700 600 NH3-H2O NH3-LiNO3 NH3-NaSCN
Re
500
0.03
400 0.02 300 200
Velocity, m/s
0.04
0.01
100 0
0 0
0.1
0.2
0.3
0.4
0.5
Length, m
Fig. 10. Absorber thermal load along bubble absorber length at similar solution and vapor flow rates.
Fig. 8. Reynolds number and velocity of solution along bubble absorber length.
As Fig. 9 shows, for the NH3eH2O, NH3eNaSCN and NH3eLiNO3 mixtures, the viscosity range goes from the values of 0.50 to 0.67, 1.85 to 1.83 and 2.00 to 2.77 cp, respectively, at the solution temperatures shown in Figs. 4e6. The viscosity of the NH3eLiNO3 and NH3eNaSCN solutions is 4.1 and 2.7 times higher than for NH3eH2O at the end absorber. Appendix A shows the viscosity and hS and hc correlations used for this study.
and interface temperatures have a similar tendency and values. The cooling water temperature increases from 30.0 to 43.5 C for NH3eH2O, from 30.0 to 40.4 C for NH3eLiNO3 and from 30.0 to 43.8 C for NH3eNaSCN. Fig. 7 shows the absorber thermal load and vapor flow rate as a function of absorber length for the different working mixtures. The thermal load value was around 1.68 kW for NH3eH2O and NH3eNaSCN, but for NH3eLiNO3 a lower value of 1.30 kW was obtained. The absorber heat load for NH3eH2O and NH3eLiNO3 resulted in a value 7.7% and 18.7% lower than the value obtained from the simulation of the refrigeration system, respectively, as shown in Table 2a, however, the absorber heat load for NH3eNaSCN resulted in a 3.6% higher value. In the same figure, the vapor flow rate decreases from 0.00092 to 0.000065, 0.00093 to 0.00035 and 0.00093 to 0.000081 kg/s for NH3eH2O, NH3eLiNO3 and NH3eNaSCN, respectively. In terms of absorption flux (FAB), the following values were obtained 0.0086, 0.0058 and 0.0085 kg/m2 s, for NH3eH2O, NH3eLiNO3 and NH3eNaSCN respectively. It is clear that NH3eLiNO3 has the higher COP than the other mixtures, although it has difficulties in the absorption process. As it can be seen, the vapor flow rate is not totally absorbed in the absorber, so extra heat and mass transfer area is required. Fig. 8 shows that NH3eNaSCN achieves the highest solution velocity values (w0.05 m/s) caused by the high solution flow rate (see Table 2a), although NH3eH2O resulted with the highest turbulence values (Re ¼ 250e200) and NH3eLiNO3 the lowest values (Re ¼ 90e110) due to its higher viscosity.
In this case the inlet vapor and solution flow rates for each working fluid in the bubble absorber have the same values than those obtained for the NH3eH2O mixture at the simulated conditions shown in Table 2b. Figs. 10e13 show the absorber thermal load and vapor flow rate, the overall heat transfer coefficient, the hold-up and the Reynolds number and velocity, respectively, for each working fluid as a function of absorber length. The absorber thermal load obtained was 1.68 kW for NH3eH2O, 1.47 kW for NH3eNaSCN and the lowest value corresponded to NH3eLiNO3 at 1.35 kW, as shown in Fig. 10. As Fig. 11 shows, the NH3eH2O and NH3eNaSCN mixtures obtained higher U values than NH3eLiNO3 due to their higher hold-up and thermal conductivity values, respectively, that improved the individual heat transfer coefficients. The FAB values were 0.0082 and 0.0056 kg/m2 s for NH3eNaSCN and NH3eLiNO3 respectively. The FAB and thermal load decrease as expected, due to the lower value of the solution flow rate.
Fig. 9. Solution viscosity along bubble absorber length.
Fig. 11. Overall heat transfer coefficient profile along bubble absorber length at similar solution and vapor flow rates.
4.2. Results for similar flow rate conditions
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Fig. 12. Gas hold-up profile along bubble absorber length at similar solution and vapor flow rates.
The NH3eH2O mixture obtained larger U values than NH3eNaSCN at the beginning of the absorber length, but the U value relationship is inverted for lengths higher than 0.20 m as it is shown in Fig. 11. The low value of the NH3eLiNO3 overall heat transfer coefficient is caused mainly by the high viscosity (similar viscosity values from Fig. 9). The U value depends mainly on the solution side, as the cooling water heat transfer coefficient is kept almost constant (6 kW/m2 K) along the absorber length. Fig. 12 shows that NH3eH2O achieves the highest hold-up value (defined as the volume ratio of gas phase in the mixture) and it then decreases linearly from 0.95 to 0.30. For NH3eLiNO3 and NH3eNaSCN the hold-up values decreases exponentially from 0.80 to 0.30 and 0.80 to 0.15, respectively. The solution velocity is different for each working fluid, 0.041, 0.036 and 0.032 m/s for NH3eH2O, NH3eNaSCN and NH3eLiNO3, respectively, as shown in Fig. 13. Moreover, the Reynolds number increases from 69 to 90 for NH3eLiNO3, but it decreases from 256 to 211 and from 69 to 57 for NH3eH2O and NH3eNaSCN, respectively, this decrement is caused by the lower solution temperature at the absorber outlet. The NH3eH2O mixture obtained the maximum turbulence, whilst the minimum turbulence value was obtained for NH3-LiNO3 mixture. 4.3. Results for similar activation temperature conditions For similar activation temperatures the input data were obtained from the absorption refrigeration system simulation at similar generator temperatures. The generator temperature was
Fig. 13. Solution Reynolds number and velocity profile along bubble absorber length at similar solution and vapor flow rates.
Fig. 14. Absorber thermal load and vapor flow rate along bubble absorber length at TGE ¼ 120 C.
fixed at 120 C, TCO and TAB was fixed at 40 C and TEV at 5 C, for a 1 kW of cooling capacity as shown in Table 2c. The NH3eLiNO3 mixture obtains the highest COP value of 0.61 whilst NH3eH2O obtained the lowest value of 0.49. Fig. 14 shows the absorber thermal load and vapor flow rate as a function of absorber length for each working fluid. The absorber thermal load follows a very similar tendency and value for NH3eH2O and NH3eNaSCN at 1.66 kW, however NH3eLiNO3 obtained a value of 1.27 kW. The FAB obtained similar values for NH3eH2O and NH3eNaSCN (0.0087 kg/m2 s), but NH3eLiNO3 obtained the lowest value (0.0067 kg/m2 s), caused by the high viscosity that affects the heat and mass transfer process. 5. Conclusions A bubble absorber was simulated with three different working fluids: NH3eH2O, NH3eLiNO3 and NH3eNaSCN in order to analyze the absorber performance using a plate exchanger. The bubble absorber working conditions were obtained from an absorption refrigeration system simulation of 1 kW cooling capacity at three different conditions: minimum generator temperature, similar flow rate and similar generator temperature. The following conclusions were drawn from the absorber analysis. At minimum generator temperature conditions the NH3eH2O and NH3eNaSCN working fluids had an absorber thermal load value of 1.68 kW, while NH3eLiNO3 obtained a value of 1.30 kW, also the mass absorption flux was 0.0086, 0.0058 and 0.0085 kg/m2 s for NH3eH2O, NH3eLiNO3 and NH3eNaSCN, respectively. The NH3eLiNO3 mixture obtained the lowest values for both parameters. At similar flow rates, NH3eH2O, NH3eNaSCN and NH3eLiNO3 obtained 1.69, 1.47 and 1.35 kW of absorber thermal load respectively. NH3eH2O and NH3eNaSCN obtained a higher heat transfer coefficient than NH3eLiNO3 caused by higher values of gas hold-up and thermal conductivity. At similar generator temperatures, the NH3eH2O and NH3eNaSCN working fluids obtained higher absorber thermal load values than NH3eLiNO3. Also, NH3eLiNO3 obtained the lowest mass absorption flux value (0.0067 kg/m2 s) caused by its higher viscosity. The NH3eLiNO3 working fluid obtained the lower absorption mass transfer and absorber thermal load values than NH3eH2O and NH3eNaSCN in the three cases studied, due mainly to its higher viscosity that causes problems with the absorption of NH3 vapor in the NH3eLiNO3 solution; despite that it, it obtained the highest COP in the absorption refrigeration system simulation. On the other
J. Cerezo et al. / Applied Thermal Engineering 31 (2011) 1869e1876
hand NH3eNaSCN seems to be a better alternative as it attained better results for the absorber performance than NH3eLiNO3 and a higher COP than NH3eH2O. Acknowledgements The authors would like to thank UNAM for partially funding the work through project PAPIIT IN109009. This document has been produced with the financial support of the European Union and Mexico through project FONCICYT 94256. The contents of this document are the sole responsibility of the Instituto de Ingeniería de la Universidad Autónoma de Baja California and the Centro de Investigación en Energía de la Universidad Nacional Autónoma de México and can under no circumstance be regarded as reflecting the position of the European Union and Mexico. Dr. Jesus Cerezo would like to thank UNAM for the postdoctoral scholarship at the Centro de Investigación en Energía. The authors would like to thank Carmen Huerta for her support on the use of EES software.
Appendix A This section shows correlation used for different mixtures used to calculate the viscosity (m). They were obtained from Infante [14] and Conde [15]. NH3eH2O
m ¼ exp x ln mNH3 ;S þ ð1 xÞln mNH3 ;S
(A.1)
T þ 273:15 fs 0:00001 þ 0:52 0:815 TNH3 ;CRIT þ 273:15 NH3eLiNO3
m ¼ ð5:1835T þ 992:337Þ ð1 xÞð0:08333Tþ6:8333Þ þ expð0:01147T 1:744Þ 0:001
hS ¼ hLIQ ð1 holdupÞ0:8
(A.5)
1= PrC 3 kC d1 hc ¼ 0:99 Re0:529 C H
(A.6)
hLIQ is the solution heat transfer coefficient at liquid phase, but there is no values for this type of plate heat exchanger, then hLIQ was supposed to be obtained from experimental values for water from Equation A.6 The solution and vapor mass transfer coefficients were calculated from Incropera [18] and Sherwood [19].
0:666 0:116 dB g0:333 bS ShINT;S ¼ 2 þ 0:0187Re0:779 Sc0:546 B s 1=3
ShINT;V ¼ 0:664Re0:5 B ScV
A COP cp dH FAB h K km m Q Re Sh T U x
z
(A.2)
m
x < 50% wt
m ¼ 5:79278:9xþ390x2 176x3
þ 2:692þ11:95x17:24x2 þ7:974x3 T þ 0:058850:2701xþ0:3994x2 0:1882x3 T2 þ 0:0004977þ0:002339x0:003532x2 þ0:001691x3 T3 0:001
ðA:3Þ
x > 50% wt
m ¼ ð0:5289T þ 29695Þ ð1 xÞ3:9365 þ 2:19 11
10
4:636
ð120 TÞ
þ0:1 0:001
(A.7) (A.8)
Nomenclature
y
NH3eNaSCN
1875
area (m2) coefficient of operation heat capacity (kJ/kg K) hydraulic diameter (m) Mass absorption flux per heat transfer area (kg/m2 s) Specific enthalpy (kJ/kg), or heat transfer coefficient (kW/ m2 K) Thermal conductivity (kW/m K) Mass transfer coefficient (kg/m2 s) Mass flow rate (kg/s) Thermal load (kW) Reynolds number Sherwood number temperature ( C) overall heat transfer coefficient (kW/m2 K) weight concentration in solution phase (kg refrigerant/kg solution) weight concentration in vapor phase (kg refrigerant/kg solution) ratio of ammonia mass flux to the total mass flux absorbed dynamic viscosity (cp)
Subscripts AB absorber ABSOR absorbent C cooling water water H2O IN input INT interface ammonia NH3 OUT outlet REF refrigerant S solution SEN sensible V vapor phase w wall References
(A.4)
The solution (hS) and cooling (hc) heat transfer coefficients were obtained by Herbine and Perez-Blanco [9] and Cerezo [10], respectively
[1] A.M. Selim, M.M. Elsayed, Performance of a packed bed absorber for water ammonia absorption refrigeration system, International Journal of Refrigeration 22 (1999) 283e292. [2] Cerezo, J. Estudio del proceso de absorción con amoniaco-agua en intercambiadores de placas para equipos de refrigeración por absorción, PhD Thesis, Universidad Rovira i Virgili, Spain, 2006.
1876
J. Cerezo et al. / Applied Thermal Engineering 31 (2011) 1869e1876
[3] D.W. Sun, Comparison of the performances of NH3-H2O, NH3-LiNO3 and NH3-NaSCN absorption refrigeration systems, Energy Conversion and Management 39 (5/6) (1998) 357e368. [4] K.A. Antonopoulos, E.D. Rogdakis, Performance of solar-driven ammonialithium nitrate and ammonia-sodium thiocyanate absorption systems operating as coolers or heat pumps in Athens, Applied Thermal Engineering 16 (2) (1996) 127e147. [5] Heard, C.L. On the viscosity of the absorption heat pump working pair: ammonia/lithium nitrate and improved heat transfer correlation for boiling in a vertical tube, Instituto Mexicano del Petroleo, Mexico, personal communication, 2000. [6] Y.T. Kang, R.N. Christensen, T. Kashiwagi, Ammonia-water bubble absorber with a plate heat exchanger, ASHRAE Transaction 204 (1) (1998). [7] J. Fernández-Seara, J. Sieres, C. Rodríguez, M. Vázquez, Ammoniaewater absorption in vertical tubular absorbers, International Journal of Thermal Sciences 44 (2005) 277e288. [8] K.B. Lee, B.H. Chung, J.C. Lee, C.H. Lee, S.H. Kim, Experimental analysis bubble mode in a plate-type absorber, Chemical Enginnering Science 57 (2002) 1923e1929. [9] G.S. Herbine, H. Perez-Blanco, Model of an ammonia-water bubble absorber, Ashrae Transactions 101 (1) (1995) 1324e1331. [10] J. Cerezo, R. Best, M. Bourouis, A. Coronas, Comparison of numerical and experimental performance criteria of an ammonia-water bubble absorber
[11] [12]
[13]
[14]
[15]
[16] [17] [18] [19]
using plate heat exchanger, International Journal of Heat and Mass Transfer 53 (17e18) (2010) 3379e3386. C.A. Infante Ferreira, Combined momentum, heat and mass transfer in vertical slug flow absorbers, International Journal of Refrigeration 8 (6) (1985) 326e334. M. Venegas, M. Izquierdo, P. Rodriguez, A. Lacuona, Heat and mass transfer during absorption of ammonia vapour by LiNO3-NH3 solution droplets, International Journal of Heat and Mass Transfer 47 (2004) 2653e2667. M. Venegas, P. Rodriguez, A. Lecuona, M. Izquierdo, Spray absorbers in absorption systems using lithium nitrate-ammonia solution, International Journal of Refrigeration 28 (4) (2005) 554e564. C.A. Infante Ferreira, Thermodynamic and physical property data equations for ammonia-lithium nitrate and ammonia-sodium thiocyanate solutions, Solar Energy 32 (1984) 231e236. M.R. Conde Petit, Thermophysical Properties of NH3-H2O Solutions for Industrial Design for Absorption Refrigeration Equipment (2004) M. Conde Engineering, Zurich, Switzerland. O.M. Ibrahim, S.A. Klein, Thermodynamic properties of ammonia-water mixtures, ASHRAE Transactions: Symposia 21 (2) (1993). R.E. Treybal, Mass Transfer Operations, third ed.. McGraw Hill, Tokyo, 1980. F.P. Incropera, D.P. De Witt, Fundamental of Heat and Mass Transfer, third ed. John Wiley & Sons Cop., New York, 1990. T.K. Sherwood, R.L. Pigford, C.R. Wlke, Mass Transfer. Mc Graw-Hill, New York, 1975.