TY - JOUR
AB - Abstract In this article, a hybrid photovoltaic–thermal (PV/T) module generating both electrical and thermal energy simultaneously has been used in a closed-cycle system to provide regeneration heat via a dynamic solar radiation model as well as electrical power relative to location, time of the day and day of the month. Electrical power generated drives the air fan, water and solution pumps, while the thermal component is used for desiccant’s regeneration. This combination enhances energy efficiency of the air conditioning system. A simplified analytical model of the complex occurrence of coupled heat and mass transmission phenomenon in liquid desiccant regeneration system powered by a hybrid PV/T module is developed. The interfacial air–desiccant interaction in a structured packing vertical column using lithium bromide solution and Mellapak was analysed. The resulting differential equations are solved simultaneously using separative evaluation and step-by-step iterative procedure. The system’s performance was projected with regeneration effectiveness, subject to varying temperatures of air and desiccant solution, moisture content and mass flow rate. It was established that subject to the prevailing local weather conditions, the PV/T module significantly raised desiccant temperature to a high of 67.22°C good enough for the regeneration process. The regeneration rate and effectiveness improved with upsurge in mass flow rate but reduces with a rise in humidity ratio. The optimum flow mix for effective regeneration was therefore established to be 0.847 and 0.00331 kg/min for air and desiccant solution, respectively, for maximum effectiveness of 69.3%. The liquid desiccant solution concentration increased by 30% during when solar radiation peak hours. The obtained theoretical outcomes matched with experimental results from the available literature show a permissible discrepancy of within ±20%, largely due to the fact that the simulation parameters were not the same as the prevailing experimental conditions. 1 INTRODUCTION The urgent need for energy efficiency coupled with environmental consciousness in air conditioning systems has in the recent decade, drawn attention towards liquid desiccant dehumidification and regeneration. As an innovative substitute technology to conventional vapour-compression systems, liquid desiccant air conditioning systems (LDACS) have currently gained prominence in both domestic and industrial applications due to its ability to use low-grade industrial waste heat as well as renewable sources like solar energy. The inception of solar-powered LDAC system dates back to 1980s, pioneered by Lof et al. [1]. Since then, the popularity of solar-powered LDACS has stepped up 3-fold with numerous advances made with emphasis in feasibility, use, energy intake and economic analysis [2]. Solar-powered desiccant schemes are categorized as closed- or open-loop configurations. A chilling unit is built in a closed-loop system utilizing water as a heat transfer medium to the desiccant. On the other hand, water is supplied from external source in an open cycle and is used as the refrigeration agent, hence replacing the need for energy-intensive condenser [3]. The advances of the LDAC technology was demonstrated in the experimental study by Gommed and Grossman [4] on the use of solar energy in liquid desiccant dehumidification and cooling. On the feasibility and economic analysis of solar application in LDAC systems, Li and Yang [5] and Halliday et al. [6] presented an open-cycle dehumidification in Hong Kong and UK environments, respectively. More recently, Aly et al. [3] used Matlab to simulate an open-cycle solar-powered, two-desiccant system. Closed-cycle systems of solar-powered desiccant regeneration are not so common and have very limited literature. In addition, the application of a hybrid photovoltaic and thermal (PV/T) module brings a new dimension to this technology. Direct solution regeneration through a solar collector in which the desiccant solution is the heat collecting medium has been demonstrated to be more effective compared to indirect systems [7]. This article presents a simplified theoretical breakdown of interfacial heat and mass transfer characteristics of an autonomous and self-sustainable liquid desiccant regenerator powered by PV/T collector in a closed loop through computer modelling. The theoretical models for solar radiation, solution side mass transfer, heat transfer as well as the pressure drop on the airside are developed and used to establish the influence of inlet properties of both air and desiccant solution on the regenerator effectiveness (e.g. temperatures, humidity ratio and mass flow rate) for near-zero carry-overs. 2 LDAC SYSTEM A graphic illustration of a typical hybrid solar PV/T-powered LDAC system considered for this analysis is presented in Figure 1. The arrangement is made up of three major units; the dehumidifier within which strong desiccant solution is spewed from the top of a packed vertical column, while air blown from the bottom mixes with the desiccant crystals in a counterflow pattern. Strong/concentrated desiccant solution absorbs water vapour from humid air as they come into contact. Figure 1. View largeDownload slide Graphical illustration of a solar-powered direct liquid desiccant regeneration and dehumidification scheme. Figure 1. View largeDownload slide Graphical illustration of a solar-powered direct liquid desiccant regeneration and dehumidification scheme. The regenerator functions in a similar manner as the dehumidifier except for heated dilute solution crystals flowing from the topside of a packed-bed vertical column where a counterflow interaction is initiated with stream of air from the bottom. The interaction results in evaporation of water molecules leaving in a strong desiccant liquid at near-initial concentration ready for recirculation in the dehumidifier. The third unit is the hybrid PV/T collector in which both electrical power and thermal energy are generated subject to prevailing solar radiation. The PV unit converts solar radiation into electrical energy used to drive the air fans and solution pumps, while residual heat energy is absorbed and transferred to the weak/dilute desiccant solution circulating through the tubes attached to the absorber plate. The solution leaves the collector at raised temperatures. At this state, it becomes difficult to hold water vapour and is pumped to the regeneration column. Theoretical analysis was only limited to the solar radiation and interfacial interaction amongst the liquid desiccant and air in the regeneration process. The interactions in dehumidification process were not considered in this study but provided merely for easy understanding and completeness of the concept. 3 PERFORMANCE ANALYSIS The performance index used for regenerator analysis was the specific effectiveness ( εr) defined as: εr=ωi−ωoωi−ωe (1) where ωi, ωo and ωe are air humidity ratios at inlet, outlet and equilibrium settings, respectively. Equilibrium air humidity ratio was evaluated in terms of atmospheric, pa and partial vapour pd pressures of air and bulk fluid, respectively, using the following relationship. ωe=0.622pdpa−pd (2) According to Montieth and Unsworth [8], partial vapour pressure was further given by Teten’s equation as follows: pd=0.61078exp(17.27TT+237.3) (3) where T is the temperature of desiccant solution. 3.1 Solar energy model The surface total incident solar radiation GT was evaluated with respect to time and day of the year and location as a function of direct, diffuse and reflected beam radiation. The relationship adopted from the study of Aly et al. [3] is applied, thus, GT=RbGbn+CGd(1+cosφ)2+(Gbn+Gd)(ϒg)(1+cosφ)2 (4) where Gb is the horizontal surface beam radiation, Rb is the beam radiation tilt factor, Gd is the diffuse sky radiation, φ is the surface tilt angle taken to be 35°, while Υg is the ground reflectance whose value depends on the surface texture. C is the diffuse radiation factor determined by Becker [9] as follows: C=1+0.11×cos(n−15)×2π/365 (5) where n is the number of ith days during the year under consideration. The expression for beam radiation on horizontal surface Gbwas established as Gb=GbncosΦ (6) where Φ is the solar altitude angle and Gbn is the incident beam radiation normal to the module surface on a clear day computed as a function of the solar altitude angle using the following expression: Gbn=1366.1{1+0.033cos(360ι365)}exp(−31sinΦ) (7) Additionally, diffuse solar radiation Gd was estimated as: Gd=0.5(1+cosφ)CGbn (8) The altitude angle Φ was evaluated in terms of the latitude L, hour and declination angles all in a simplified expression as follows: sinΦ=sinLcosA+cosLcosAcos(±0.25Γ) (9) where A = (23.45sin{360/365(248+ι)}) The ± implies negative and positive values of the hour angle in the morning and afternoon, respectively, while Γ is the quantity of time in minutes, before or after confined solar noon at a particular time in consideration. As defined in [3], the beam radiation tilt factor was formulated as the ratio of sunbeam heat on slanted and flat surfaces respectively. This relationship was further simplified and expressed as: Rb=sin(L−φ)cosA+cos(L−φ)cosAcos(±0.25Γ)sinLcosA+cosLcosAcos(±0.25Γ) (10) Since the main focus of this analysis was on the thermal energy component generated by the PV/T collector, the electrical energy element has not been detailed. Hence, from the total solar radiation on the collector surface, the overall heat energy available to the solution in the collector was expressed as follows: Gc=GT{(1−ρ)τα−(τα−τpvηel)} (11) where ρ,τ, and α are the reflectance, transmittance and absorptance of the collector glazing and PV cells attached directly to the absorber plate, respectively. ηel represents the PV/T module electrical efficiency. Therefore, taking heat energy balance for the hybrid PV/T collector, an expression was then derived as follows: Gcdx=ṁadha+ṁddhd+UL(Td−Tamb)+mhfg (12) where ṁa and ṁd are air and desiccant mass flow rates, respectively, in kg/s; ha and hd are air- and desiccant-specific enthalpies, respectively, in J/kg; m is the quantity of water evaporated in kg; hfg is the hidden heat of vaporization of water in J/kg, UL is the overall heat loss coefficient in W/m2K, Td and Tamb are the desiccant and ambient temperatures, respectively, in °C. Likewise, the air stream energy equilibrium entering the regenerator was also derived as follows: ṁadha=ϑa(Td−Ta)dx−ϑd(Ta−Tamb)dx (13) where ϑa and ϑd are the heat transfer coefficients of air and desiccant, respectively. 3.2 Mass transfer Considering a unit volume and cross section of a counterflow packed column differential segment presented in Figure 2, heat and mass exchange process occurring at the air–desiccant solution interface can theoretically be analysed based on the derivations of [2], [10], and [11]. Figure 2. View largeDownload slide Differential section of air–desiccant interface. Figure 2. View largeDownload slide Differential section of air–desiccant interface. The regeneration process was considered adiabatic and hence, negligible liquid-phase heat flow resistance was experienced. Taking both heat and mass transfer to happen in crosswise directions of air–desiccant flow on a uniform area throughout the interface, correlations and differential equations were derived and solved by separative evaluation and step-by-step iterative methods. The desiccant-specific enthalpy was expressed as the sum of solution’s weakening heat ΔQsol at set point Tref and heat gain as follows: hd=Cp,d(T−Tref)+ΔQsol (14) where Cp,d represents desiccant-specific heat capacity at T °C. On the other hand, specific enthalpy of air was formulated as: ha=Cp,a(T−Tref)+ω{Cp,w(Ta−Tref)+Qw,ref} (15) where ω is the humidity ratio, Cp,a and Cp,w are specific of dry air and water, respectively, while Qw,ref is the latent heat of water at reference temperature Tref. From the principle of conservation of mass, the correlation for water vapour content was derived as a function of air- and liquid-phase-specific mass flow rates ṁa and ṁd, respectively, expressed as: dṁd=ṁadω (16) 3.2.1 Airside The interfacial mass transfer is expressed as a function of molecular mass Mw and specific molar flow rate Nw of water formulated by Babakhani and Soleymani [12] as: −ṁadω=Mwa˜dHϑaln[1−βi1−β] (17) The airside humidity ratio was formulated by further simplifying (17) to yield a rudimentary differential equation for as follows: dωdH=−a˜Mwϑaṁaln[1−βi1−β] (18) where β is the moisture content in air and subscript i denotes initial conditions, ã is the exact active interfacial contact area in m2/m3 of packed-bed capacity which is dependent on the stuffing arrangement and operating environments given for structured packing as: a˜=an(dpsinθ4ξ)ϕt1.5(ρdg3μdνd)0.5 (19) where an is the packing material's exact surface area for each single element volume, while dp is the equivalent diameter of channel, θ is the angle of inclination of packing channels, ξ represents the hollow part of dry packing and ϕt is the total liquid hold up. The air mass transfer coefficient ϑa is in direct proportionality to the air mean partial pressure Pm,a, but inversely proportional to the product of temperature and universal gas constant RT. Considering this relationship, a mathematical expression of mass transfer coefficient was developed, thus, ϑa=τaPm,aRT (20) The constant of proportionality is the gas-phase mass transfer coefficient τa which depends on Reynold’s Re and Prandlt Pr numbers for air as was formulated by Fair and Bravo [13] for structured packing and thus, τa=0.0338(∂a/dp)Rea0.8Pra1/3 (21) where dp is the equivalent packing channel diameter and ∂a is the airside molecular diffusivity. Reynoldsnumber is given by: Rea=ρadpμa{νaξ(1−ϕt)sinθ+νdξϕtsinθ} (22) where νa and νd are the superficial velocities of air and desiccant, respectively, ϕt is the total liquid residence on packing surface, while θ denotes the inclination angle of packing channel from the vertical axis. From [14], we obtain an expression for total liquid hold up ϕt for Mellapack structured packing as follows: ϕt=8.66an0.83v′1.39(μdμw)0.25 (23) where μd and μw are viscosities for desiccant fluid and water, respectively. 3.2.2 Solution side Considering the desiccant solution side, the interfacial mass transfer was found by the following correlation: Nd=ϑdln[1−ψ1−ψi] (24) where ψ and ψi are the solution’s bulk and interfacial preoccupation in water, while ϑd is the desiccant solution’s mass transfer factor given as: ϑd=ϱd(ρdM˜d)ψ˜ (25) where ψ̃ is the mean solution’s molar salt preoccupation, M̃d is the desiccant’s mean molar mass and ϱd is mass transfer factor in liquid phase in structured packing derived from the study of Fair and Bravo [13] as a function of molecular diffusivity of the desiccant fluid ∂d given by: ϱd=2{∂dvd/ξϕtsinθ}1/πdp (26) For equilibrium conditions, the exact airside interfacial mass transfers must be equal to those of desiccant side, therefore, taking the mass balance at the interface, the airside interfacial molar mass strength in water βi was found by: βi=[1−(1−β)]{1−ψ1−ψi}ϑd/ϑa (27) This Equation (27) was then solved concurrently with air–desiccant equilibrium equation by an iterative procedure while taking the molar mass concentration β as dependent on the humidity ratio given by the expression: β=ωω+Mw/Ma (28) 3.3 Heat transfer The coupled nature of heat and mass transfer processes confirms they happen simultaneously. However, for ease of understanding, analysis of heat transfer was performed separately on the respective phases of air and liquid. The simultaneous airside interfacial sensible heat flow qa was formulated as a function of bulk air and interfacial temperatures Ta and Ti, respectively, by taking the thermal energy balance on the airside: qaa˜dz=σ′aa˜(Ta−Ti)dz (29) where σ́aã is the simultaneous heat transfer factor for air obtained from Ackermann correction in [15] and given as: σ′aa˜=−ṁaCpwdω/dH{1−exp(ṁaCpwdω/dHαaa˜)} (30) where αa denotes the airside heat transfer factor which was formulated for structured packing as: αa=0.0338(ka/dd)Rea0.8Pra1/3 (31) where ka is the airside thermal conductivity. Therefore, taking the thermal energy balance on the airside, we obtained ṁaha−ṁa(ha+dha)+ṁaω{Cp,w(Ta−Tref)+Qw,ref}=σ′aa˜(Ta−Ti)dH (32) But the enthalpy of air was taken to vary across the differential element according to the following expression. dha=Cp,adTa+ωCp,wdTa+ṁaω{{Cp,w(Ta−Tref)+Qw,ref} (33) Considering (34) and (35), a differential equation for air temperature was generated, thus, dTadH=−σ′aa˜(Ta−Ti)ṁa(Cp,a+ωCp,w) (34) And the overall heat stability of the differential control element within the packed unit was then formulated as: ṁadha=ṁddhd+ṁadωhd (35) Substituting (14) and (35) into (37), we get: Cp,adTa+ωCp,wdTa+ṁaω{Cp,w(Ta−Tref)+Qw,ref}ṁa=ṁd{Cp,ddTd+d(ΔQsol)}+ṁadω{Cp,d(Td−Tref)+ΔQsol} (36) However, since the variation in thermal energy for dilution d(∆Qsol) is negligibly small, it was ignored and a new basic differential equation for desiccant temperature emerged as follows: dTddH=ṁaṁd(Cp,a+ωCp,w)dTadH+{Cp,w(Ta−Tref)+Qw,ref}dωdH−{Cp,d(Td−Tref)+ΔQsol}dωdH (37) Upholding the principle of conservation of mass in the differential volume, the expression for salt content was developed as: ṁdχ=(ṁd+dṁd)(χ+dχ) (38) where χ denotes the desiccant content in salty solution. Consequently, a basic differential equation was obtained by substituting Eqn. (16) into (38) to give; dχdH=−χ(ṁdṁa)dωdH (39) Again Eqn. (39) was solved iteratively to obtain the concentration of the desiccant solution leaving the regenerator. 4 RESULTS AND DISCUSSION Apart from design parameters that were investigated by Aly et al. [3] and Gandhidasan [16], the PV/T module performance in desiccant regeneration was affected by the ambient properties of air such as temperature and vapour pressure. 4.1 The air and desiccant solution temperatures variations The change in solution temperature during the sunshine hours of the day under consideration is shown in Figure 3. The solution conditions at the inlet and outlet followed an expected trend, i.e. the solution left the module at a higher temperature than at entry point. The variation of temperature with respect to various flow rate levels was demonstrated. It was observed that the lower desiccant flow rates corresponded to higher temperatures due to increased residence time in the module. Figure 3. View largeDownload slide The temperature variations of desiccant at various mass flow rates over time of day. Figure 3. View largeDownload slide The temperature variations of desiccant at various mass flow rates over time of day. From the previous study of Oyieke and Inambao [17], it was shown that the electrical efficiency of the PV/T module increased when a working fluid flow through and carried away residual heat. Based on this previous knowledge, apart from heating the desiccant solution, the PV/T efficiency was also kept to near maximum. Different desiccant flow rates were considered in regeneration performance evaluation. For this analysis to happen, the air flow rate ma was kept at an optimum constant value of 5.082 kg/min while varying the desiccant flow rate md. Since the mass transfer ability of desiccant solution is in direct proportionality to the water evaporation rate along the regenerator height, the mass transfer coefficient was unchanged during the iteration. The air and desiccant temperature profiles at the regenerator entry and exit were plotted as shown in Figure 4 over a 24-h span. The desiccant solution entered the regenerator at same exit temperatures of the PV/T module. The highest inlet temperature Ti,d exhibited was 67.22°C with a corresponding outlet temperature To,d of 36.14°C at 12:30 h. However, air entered the regenerator at low temperature Ti,a and left at elevated temperature To,a signifying a gain due to the heat transfer occurrence. During 24-h day, regeneration process only occurred between 6:00 h and 18:00 h, which corresponded to the sunrise and sunset times. Outside the range of these hours, the vessel operated as a dehumidifier whose performance is not included in this work. Figure 4. View largeDownload slide The temperature variations of air and desiccant at entry and exit of the regenerator over time of day. Figure 4. View largeDownload slide The temperature variations of air and desiccant at entry and exit of the regenerator over time of day. 4.2 Influence of mass flow rates on regeneration effectiveness To evaluate the influence of air mass flow rates on regenerator effectiveness, a plot is presented in Figure 5 The regeneration process occurred with the air flow rates ranging from 0.065 to 0.095 kg/min. A general observation was that effectiveness improved with the upsurge of mass flow rate. However, maximum effectiveness of 69.3% was achieved at a corresponding air flow rate of 0.0847 kg/min, beyond this value, the effectiveness begun to drop due the system’s reliance on the solar energy. The maximum effectiveness was noted to occur at solar noon which on this day was seen to be at 12:30 h. Figure 5. View largeDownload slide Effect of mass flow rates of air on regeneration effectiveness. Figure 5. View largeDownload slide Effect of mass flow rates of air on regeneration effectiveness. Similarly, the variation of desiccant solution’s flow rate against regeneration effectiveness was plotted as shown in Figure 6. The effectiveness improved proportionally with desiccant flow rate like in previous case, but the maximum effectiveness was attained at a solution flow rate of 0.00331 kg/min which was below that of air. The regeneration process was observed to occur between flow rate ranges of 0.1665–0.2262 kg/min. Figure 6. View largeDownload slide Effect of mass flow rates of desiccant on regeneration effectiveness. Figure 6. View largeDownload slide Effect of mass flow rates of desiccant on regeneration effectiveness. From these two scenarios, it was observed that maximum effectiveness was achieved at air and desiccant flow rates of 0.847 and 0.00331 kg/min, respectively. These values were taken as the optimum flow mix for effective regeneration performance of the unit. Any other values outside this combination demonstrated desiccant fluid carryover to the process air. The outlet solution concentration at this point is 82%, which is near the initial concentration, hence the confirmation of a regeneration process. The relationship between regenerator effectiveness and mass flow rate ratio mḋ/mȧ is shown in Figure 7. The effectiveness reduced with an increase in mass flow rate ratio as depicted by the negative gradient of up to 20.64. Low mḋ and high mȧ resulted in a low flow rate ratio, which in turn gave higher regenerator effectiveness. On the other hand, higher mḋ and low mȧ gave a higher flow rate ratio, which results in reduced effectiveness. Since the simulation was conducted at varying flow rates of desiccant and air at an alternating pattern, i.e. reducing one and increasing the other and vice versa, the individual instantaneous flow rate ratios appeared to be scattered but suggested a reducing trend. Therefore, it was logically concluded that for effective desiccant regeneration, the solution flow rate must be lower than the air flow rate at any instant. Figure 7. View largeDownload slide Effect of mass flow rate ratio on regeneration effectiveness. Figure 7. View largeDownload slide Effect of mass flow rate ratio on regeneration effectiveness. 4.3 Effect of mass flow rates on the mass transfer coefficient The mass transfer coefficient values were calculated by an iterative procedure and the results provided in this section. As can be observed in Figure 8, the variation of overall mass transfer factor with regards to alteration of air mass flow rate was plotted. An improvement of air the mass flow rate triggered an exponential growth in overall mass transfer coefficient over the regeneration period. When air mass flow rate increased over a span of 2–6 kg/min, a corresponding increase of between 2.4 and 9.8 kg/m2 in mass transfer coefficient was observed. Figure 8. View largeDownload slide Effect of mass flow rates of air on the overall mass transfer coefficient. Figure 8. View largeDownload slide Effect of mass flow rates of air on the overall mass transfer coefficient. Similar exponential increment of mass transfer coeffcient was realized as a result of the rise in desiccant flow rate as presented in Figure 9. When desiccant solution’s flow rate varies between the range of 0.1 and 0.25 kg/min, mass transfer factor improves from 2.18 to 9.0 kg/m2. This steep alteration in mass transfer factor is because of rapid rate desiccant crystallization, thus increasing the air–desiccant interfacial surface area for mass transfer. Figure 9. View largeDownload slide Effect of mass flow rates of desiccant on the overall mass transfer coefficient. Figure 9. View largeDownload slide Effect of mass flow rates of desiccant on the overall mass transfer coefficient. The variation of solution concentration during the active regeneration period is shown in Figure 10. The liquid desiccant sollution at the initial concentration level of 50% was monitored during the solar peak hours from 10.00 to 15.00 h when solar radiation was at its highest. The final concentration achieved was 82%, this showed an increase of 30%. A corresponding escalation of the relative mass of water vapour per kilogram of solution of the same proportion was also realized. Figure 10. View largeDownload slide The variation of solution concentration during regeneration period. Figure 10. View largeDownload slide The variation of solution concentration during regeneration period. However, with respect to the mass flow rate ratio, it was observed that there was an uphazard distribution with a slight reduction in concentration with an increase in the mass flow rate ratio as depicted in Figure 11. Generally, for the individual flow rates of air and desiccant, there was a negligible change in concentration with higher desiccant flow rates, while significant increase was seen with an increase in air flow rates. Figure 11. View largeDownload slide The variation of solution concentration (a) during regeneration period and (b) with change in mass flow rate ratio. Figure 11. View largeDownload slide The variation of solution concentration (a) during regeneration period and (b) with change in mass flow rate ratio. 4.4 Assessment of predicted and experimental outcomes The average relative deviation between investigational and hypothetical outcomes are evaluated based on average deviation, Equation (40). It was realized from the existing literature that the application of PV/T in air conditioning was not a common phenomenon and had virtually not been documented prior to this study. Hence, to check the trend of some selected parameters, existing experimental results by Zhang et al. [18] were used for validation. The comparison of heat transfer coefficients predicted in this article and experimental values showed an average deviation within the range of ±20% as shown in Figure 12. This deviation is largely due to the fact that the simulation parameters were not the same as the prevailing experimental conditions. However, better results would be envisaged if experimental data for PV/T existed in the same location considered for the simulation. ARD=1n∑i=1n|εexp−εthεexp|×100% (40) Figure 12. View largeDownload slide Predicted vs experimental mass transfer factors for the regenerator. Figure 12. View largeDownload slide Predicted vs experimental mass transfer factors for the regenerator. 5 CONCLUSION This article has dealt with a theoretical analysis of desiccant regeneration system powered by solar energy via a hybrid PV/T. From the air and desiccant temperature profiles at entry and exit of the regenerator, the highest inlet temperature attained was 67.22oC with a corresponding outlet temperature of 36.14oC. This difference demonstrated that PV/T module could significantly raise desiccant temperature to required levels for regeneration process. The regenerator effectiveness improved proportionally with air and desiccant mass flow rates. However, the maximum effectiveness of 69.3% was achieved at a corresponding air flow rate of 0.0847 kg/min, beyond this value, the effectiveness begun to drop due the system’s reliance on the solar energy. The optimum flow mix for effective regeneration was therefore established to be 0.847 and 0.00331 kg/min for air and desiccant solution, respectively. The liquid desiccant solution increased by 30% during peak hours when solar radiation was at its maximum. The final concentration achieved was 82%. Increase in air and desiccant mass flow rates caused an exponential increase in the overall mass transfer coefficient over the regeneration period. The comparison of heat transfer coefficients predicted in this article and experimental values showed an average deviation within the range of ±20%. This deviation is largely due to the fact that the simulation parameters were not the same as the prevailing experimental conditions. However, better results would be envisaged if experimental data for PV/T existed in the same location considered for the simulation. REFERENCES 1 Lof GOG , Lenz TG , Rao S . 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Int J Photoenergy 2016 ; 1 : 1 – 15 . Google Scholar CrossRef Search ADS 18 Zhang L , Hihara E , Matsuoka F , et al. . Experimental analysis of mass transfer in adiabatic structured packing dehumidifier/ regenerator with liquid desiccant . Int J Heat Mass Transf 2010 ; 53 : 2856 – 863 . Google Scholar CrossRef Search ADS © The Author(s) 2018. Published by Oxford University Press. This is an Open Access article distributed under the terms of the Creative Commons Attribution Non-Commercial License (http://creativecommons.org/licenses/by-nc/4.0/), which permits non-commercial re-use, distribution, and reproduction in any medium, provided the original work is properly cited. For commercial reuse, please contact journals.permissions@oup.com
TI - Interfacial heat and mass transfer analysis in solar-powered, packed-bed adiabatic liquid desiccant regeneration for air conditioning
JF - International Journal of Low-Carbon Technologies
DO - 10.1093/ijlct/cty029
DA - 2018-09-01
UR - https://www.deepdyve.com/lp/oxford-university-press/interfacial-heat-and-mass-transfer-analysis-in-solar-powered-packed-HiRxt7UnYE
SP - 277
EP - 285
VL - 13
IS - 3
DP - DeepDyve
ER -